Injection Mold Design Engineering - David O. Kazmer (Hanser, 2016)

Kazmer Injection Mold Design Engineering

David O. Kazmer

Injection Mold Design Engineering 2nd Edition

Hanser Publishers, Munich

Hanser Publications, Cincinnati

The Author: Professor David Kazmer, Department of Plastics Engineering, 1 University Ave., Lowell MA 01833 USA

Distributed in the Americas by: Hanser Publications 6915 Valley Avenue, Cincinnati, Ohio 45244-3029, USA Fax: (513) 527-8801 Phone: (513) 527-8977 www.hanserpublications.com Distributed in all other countries by: Carl Hanser Verlag Postfach 86 04 20, 81631 München, Germany Fax: +49 (89) 98 48 09 www.hanser-fachbuch.de The use of general descriptive names, trademarks, etc., in this publication, even if the former are not especially identified, is not to be taken as a sign that such names, as understood by the Trade Marks and Merchandise Marks Act, may accordingly be used freely by anyone. While the advice and information in this book are believed to be true and accurate at the date of going to press, neither the authors nor the editors nor the publisher can accept any legal responsibility for any errors or omissions that may be made. The publisher makes no warranty, express or implied, with respect to the material contained herein. The final determination of the suitability of any information for the use contemplated for a given application remains the sole responsibility of the user.

Cataloging-in-Publication Data is on file with the Library of Congress

All rights reserved. No part of this book may be reproduced or transmitted in any form or by any means, electronic or mechanical, including photocopying or by any information storage and retrieval system, without permission in writing from the publisher. © Carl Hanser Verlag, Munich 2016 Editor: Cheryl Hamilton Production Management: Jörg Strohbach Coverconcept: Marc Müller-Bremer, www.rebranding.de, München Coverdesign: Stephan Rönigk Typesetting: Kösel Media GmbH, Krugzell Printed and bound by Hubert & Co GmbH, Göttingen Printed in Germany ISBN: 978-1-56990-570-8 E-Book ISBN: 978-1-56990-571-5

To my Dad, Andrew James Kazmer, who started me in manufacturing and inspires me to this day.

Preface to the 2nd Edition

I would like to thank the many students, practitioners, and colleagues who have provided their ongoing support and input to the 2nd edition, including Carol Barry, Mark Berry, Maria Virginia Candal Pazos, Yasuo Ishiwata, Steve Johnston, Shmuel Kenig, Adam Kramschuster, Francis Lai, Robert Malloy, Roger Manse, Steve Orroth, Nick Schott, Steven Silvey, and Robert Stack. I’d also like to recognize Cheryl ­Hamilton and Mark Smith for their patience and care in this project. Since the publication of the 1st edition, three major trends have continued with respect to plastic product and mold design: ƒƒFirst, supply-chains are tightly integrated, with rapid flow of information b ­ etween the product designers, molders, and mold designers. The landscape remains highly competitive, with firms differentiated by technical capability and efficiency. ƒƒSecond, advanced manufacturing is broadly recognized as a societal strategy for improving economic growth and human well-being. Of particular note is the broad interest in rapid prototyping processes (and 3D printing in particular) for supplying mold components and even low volume production of plastic parts. ƒƒThird, the plastics industry is under increasing public pressure to minimize environmental impact. Designers of plastic products and their molds should strive to reduce, reuse, and recycle the resources that we are so fortunate to have. The second edition has been extensively revised while reflecting on these trends. The intent has remained to provide a practical yet reasoned engineering approach. I continue to hope that Injection Mold Design Engineering is accessible and useful to all who read it. I welcome your ongoing feedback and future cooperation. Best wishes, David Kazmer, P. E., Ph. D. Dandeneau Professor for Sustainable Manufacturing Department of Plastics Engineering University of Massachusetts Lowell March 2016

VIII

P  reface to the 1st Edition

„„Preface to the 1st Edition Mold design has been more of a technical trade than an engineering process. Traditionally, practitioners have shared standard practices and learned tricks of the trade to develop sophisticated molds that often exceed customer expectations. However, the lack of fundamental engineering analysis during mold design frequently results in molds that may fail and require extensive rework, produce moldings of inferior quality, or are less cost effective than may have been possible. Indeed, it has been estimated that on average 49 out of 50 molds require some modifications during the mold start-up process. Many times, mold designers and end-users may not know how much money was “left on the table.” The word “engineering” in the title of this book implies a methodical and analytical approach to mold design. The engineer who understands the causality between design decisions and mold performance has the ability to make better and more informed decisions on an application by application basis. Such decision making competence is a competitive enabler by supporting the development of custom mold designs that outperform molds developed according to standard practices. The proficient engineer also avoids the cost and time needed to delegate decision to other parties, who are not necessarily more competent. The book has been written as a teaching text, but is geared towards professionals working in a tightly integrated supply chain including product designers, mold designers, and injection molders. Compared to most handbooks, this textbook ­provides worked examples with rigorous analysis and detailed discussion of vital mold engineering concepts. It should be understood that this textbook purposefully investigates the prevalent and fundamental aspects of injection mold engineering. I hope that Injection Mold Design Engineering is accessible and useful to all who read it. I welcome your feedback and partnership for future improvements. Best wishes, David Kazmer, P. E., Ph. D. Lowell, Massachusetts June 1, 2007

Contents

Preface to the 2nd Edition . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .



V

Preface to the 1st Edition . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . VIII

Nomenclature

. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .

1 Introduction . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .

XV



1

1.1 Overview of the Injection Molding ­Process . . . . . . . . . . . . . . . . . . . . . . . . . . .

1

1.2 Mold Functions . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .

4

1.3 Mold Structures . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1.3.1 External View of Mold . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1.3.2 View of Mold during Part Ejection . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1.3.3 Mold Cross-Section and Function . . . . . . . . . . . . . . . . . . . . . . . . . . . . .



6 6 8 9

1.4 Other Common Mold Types . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1.4.1 Three-Plate, Multicavity Family Mold . . . . . . . . . . . . . . . . . . . . . . . . . . 1.4.2 Hot Runner, Multigated, Single Cavity Mold . . . . . . . . . . . . . . . . . . . 1.4.3 Comparison . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .



11 12 14 15

1.5 The Mold Development Process . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .

16

1.6 Mold Standards . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .

18

1.7 Chapter Review . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .

20

1.8 References . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .

20

2 Plastic Part Design . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .



21

2.1 The Product Development Process . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2.1.1 Product Definition . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2.1.2 Product Design . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2.1.3 Development . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2.1.4 Scale-Up and Launch . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2.1.5 Role of Mold Design . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .

21 22 23 23 24 24

X

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2.2 Design Requirements . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2.2.1 Application Engineering Information . . . . . . . . . . . . . . . . . . . . . . . . . . 2.2.2 Production Planning . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2.2.3 End-Use Requirements . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2.2.4 Design for Manufacturing and Assembly . . . . . . . . . . . . . . . . . . . . . . 2.2.5 Plastic Material Properties . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .



25 26 27 28 30 30

2.3 Design for Injection Molding . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2.3.1 Uniform Wall Thickness . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2.3.2 Rib Design . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2.3.3 Boss Design . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2.3.4 Corner Design . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2.3.5 Surface Finish and Textures . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2.3.6 Draft . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2.3.7 Undercuts . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .



31 31 33 34 34 36 38 39

2.4 Chapter Review . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .

41

2.5 References . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .

42

3 Mold Cost Estimation . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .



43

3.1 The Mold Quoting Process . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .

43

3.2 Cost Overview for Molded Parts . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 3.2.1 Mold Cost per Part . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 3.2.2 Material Cost per Part . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 3.2.3 Processing Cost per Part . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 3.2.4 Defect Cost per Part . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .



45 47 48 49 52

3.3 Mold Cost Estimation . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 3.3.1 Mold Base Cost Estimation . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 3.3.2 Cavity Cost Estimation . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 3.3.2.1 Cavity Set Cost . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 3.3.2.2 Cavity Materials Cost . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 3.3.2.3 Cavity Machining Cost . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 3.3.2.4 Cavity Discount Factor . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 3.3.2.5 Cavity Finishing Cost . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 3.3.3 Mold Customization . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .

53 54 55 56 56 58 62 63 64

3.4 Manufacturing Strategy . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 3.4.1 Breakeven Analysis . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 3.4.2 Prototyping Strategy . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .

69 69 72

3.5 Chapter Review . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .

76

3.6 References . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .

77

Contents

4 Mold Layout Design . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .



79

4.1 Parting Plane Design . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 4.1.1 Determine Mold Opening Direction . . . . . . . . . . . . . . . . . . . . . . . . . . . 4.1.2 Determine Parting Line . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 4.1.3 Parting Plane . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 4.1.4 Shut-Offs . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .



79 80 83 84 86

4.2 Cavity and Core Insert Creation . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 4.2.1 Height Dimension . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 4.2.2 Length and Width Dimensions . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 4.2.3 Adjustments . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .



87 87 88 89

4.3 Mold Base Selection . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 4.3.1 Cavity Layouts . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 4.3.2 Mold Base Sizing . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 4.3.3 Molding Machine Compatibility . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 4.3.4 Mold Base Suppliers . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .



91 91 93 95 97

4.4 Material Selection . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 98 4.4.1 Strength vs. Heat Transfer . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 100 4.4.2 Hardness vs. Machinability . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 101 4.4.3 Material Summary . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 102 4.4.4 Surface Treatments . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 103 4.5 Chapter Review . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 105 4.6 References . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 107

5 Cavity Filling Analysis and Design . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .



109

5.1 Overview . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 109 5.2 Objectives in Cavity Filling . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 5.2.1 Complete Filling of Mold Cavities . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 5.2.2 Avoid Uneven Filling or Over-Packing . . . . . . . . . . . . . . . . . . . . . . . . . 5.2.3 Control the Melt Flow . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .



110 110 111 112

5.3 Viscous Flow . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 5.3.1 Shear Stress, Shear Rate, and Viscosity . . . . . . . . . . . . . . . . . . . . . . . . 5.3.2 Pressure Drop . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 5.3.3 Rheological Behavior . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 5.3.4 Newtonian Model . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 5.3.5 Power Law Model . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .



112 112 113 115 117 119

5.4 Process Simulation . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .

121

5.5 Cavity Filling Analyses and Designs . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 124 5.5.1 Estimating the Processing Conditions . . . . . . . . . . . . . . . . . . . . . . . . . 124 5.5.2 Estimating the Filling Pressure and Minimum Wall Thickness . . 127

XI

XII

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5.5.3 Estimating Clamp Tonnage . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 130 5.5.4 Predicting Filling Patterns . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 133 5.5.5 Designing Flow Leaders . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 135 5.6 Chapter Review . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 138 5.7 References . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 139

6 Feed System Design

. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .

141

6.1 Overview . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .

141

6.2 Objectives in Feed System Design . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 6.2.1 Conveying the Polymer Melt from Machine to Cavities . . . . . . . . . 6.2.2 Impose Minimal Pressure Drop . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 6.2.3 Consume Minimal Material . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 6.2.4 Control Flow Rates . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .



141 141 142 143 145

6.3 Feed System Types . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 6.3.1 Two-Plate Mold . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 6.3.2 Three-Plate Mold . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 6.3.3 Hot Runner Molds . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .



145 146 148 153

6.4 Feed System Analysis . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 6.4.1 Determine Type of Feed System . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 6.4.2 Determine Feed System Layout . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 6.4.3 Estimate Pressure Drops . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 6.4.4 Calculate Runner Volume . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 6.4.5 Optimize Runner Diameters . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 6.4.6 Balance Flow Rates . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 6.4.7 Estimate Runner Cooling Times . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 6.4.8 Estimate Residence Time . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .



156 158 159 163 165 166 170 173 175

6.5 Practical Issues . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 6.5.1 Runner Cross-Sections . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 6.5.2 Sucker Pins . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 6.5.3 Runner Shut-Offs . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 6.5.4 Standard Runner Sizes . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 6.5.5 Steel Safe Designs . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .



176 176 180 182 183 184

6.6 Advanced Feed Systems . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 6.6.1 Insulated Runner . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 6.6.2 Stack Molds . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 6.6.3 Branched Runners . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 6.6.4 Dynamic Melt Control . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .



185 185 186 188 190

6.7 Chapter Review . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 192 6.8 References . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 194

Contents

7 Gating Design

. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .

197

7.1 Objectives of Gating Design . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 197 7.1.1 Connecting the Runner to the Mold Cavity . . . . . . . . . . . . . . . . . . . . . 197 7.1.2 Provide Automatic De-gating . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 197 7.1.3 Maintain Part Aesthetics . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 198 7.1.4 Avoid Excessive Shear or Pressure Drop . . . . . . . . . . . . . . . . . . . . . . . 198 7.1.5 Control Pack Times . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 199 7.2 Common Gate Designs . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 7.2.1 Sprue Gate . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 7.2.2 Pin-Point Gate . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 7.2.3 Edge Gate . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 7.2.4 Tab Gate . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 7.2.5 Fan Gate . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 7.2.6 Flash/Diaphragm Gate . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 7.2.7 Tunnel/Submarine Gate . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 7.2.8 Thermal Gate . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 7.2.9 Valve Gate . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 7.3 The Gating Design Process . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 7.3.1 Determine Gate Location(s) . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 7.3.2 Determine Type of Gate . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 7.3.3 Calculate Shear Rates . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 7.3.4 Calculate Pressure Drop . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 7.3.5 Calculate Gate Freeze Time . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 7.3.6 Adjust Dimensions . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .

200 200 201 202 203 204 205 206 209 212

213 213 215 217 219 221 224

7.4 Chapter Review . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 225 7.5 References . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 226

8 Venting

. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .

227

8.1 Venting Design Objectives . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 8.1.1 Release Compressed Air . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 8.1.2 Contain Plastic Melt . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 8.1.3 Minimize Maintenance . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .

227 227 228 228

8.2 Venting Analysis . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 8.2.1 Estimate Air Displacement and Rate . . . . . . . . . . . . . . . . . . . . . . . . . . 8.2.2 Identify Number and Location of Vents . . . . . . . . . . . . . . . . . . . . . . . . 8.2.3 Specify Vent Dimensions . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .



228 228 229 232

8.3 Venting Designs . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 8.3.1 Vents on Parting Plane . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 8.3.2 Vents around Ejector Pins . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 8.3.3 Vents in Dead Pockets . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .



236 236 238 239

XIII

XIV

  Contents

8.4 Chapter Review . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 241 8.5 References . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 242

9 Cooling System Design

. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .

243

9.1 Objectives in Cooling System Design . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 9.1.1 Maximize Heat Transfer Rates . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 9.1.2 Maintain Uniform Wall Temperature . . . . . . . . . . . . . . . . . . . . . . . . . . 9.1.3 Minimize Mold Cost . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 9.1.4 Minimize Volume and Complexity . . . . . . . . . . . . . . . . . . . . . . . . . . . . 9.1.5 Maximize Reliability . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 9.1.6 Facilitate Mold Usage . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .



243 243 244 244 245 245 245

9.2 The Cooling System Design Process . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 9.2.1 Calculate the Required Cooling Time . . . . . . . . . . . . . . . . . . . . . . . . . . 9.2.2 Evaluate Required Heat Transfer Rate . . . . . . . . . . . . . . . . . . . . . . . . . 9.2.3 Assess Coolant Flow Rate . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 9.2.4 Assess Cooling Line Diameter . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 9.2.5 Select Cooling Line Depth . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 9.2.6 Select Cooling Line Pitch . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 9.2.7 Cooling Line Routing . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .



246 246 252 253 254 257 260 262

9.3 Cooling System Designs . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 9.3.1 Cooling Line Networks . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 9.3.2 Cooling Inserts . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 9.3.3 Conformal Cooling . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 9.3.4 Highly Conductive Inserts . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 9.3.5 Cooling of Slender Cores . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 9.3.5.1 Cooling Insert . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 9.3.5.2 Baffles . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 9.3.5.3 Bubblers . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 9.3.5.4 Heat Pipes . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 9.3.5.5 Conductive Pin . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 9.3.5.6 Interlocking Core with Air Channel . . . . . . . . . . . . . . . . . . . . 9.3.6 One-Sided Heat Flow . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .



266 266 269 269 270 272 273 274 275 275 276 277 278

9.4 Mold Wall Temperature Control . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 9.4.1 Pulsed Cooling . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 9.4.2 Conduction Heating . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 9.4.3 Induction Heating . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 9.4.4 Managed Heat Transfer . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .

281 281 284 285 286

9.5 Chapter Review . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 288 9.6 References . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 289

Contents

10 Shrinkage and Warpage . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .



291

10.1 The Shrinkage Analysis Process . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 10.1.1 Estimate Process Conditions . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 10.1.2 Model Compressibility Behavior . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 10.1.3 Assess Volumetric Shrinkage . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 10.1.4 Evaluate Isotropic Linear Shrinkage . . . . . . . . . . . . . . . . . . . . . . . . . . . 10.1.5 Evaluate Anisotropic Shrinkage . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 10.1.6 Numerical Simulation . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 10.1.7 Shrinkage Analysis Validation . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .



293 294 294 297 300 301 303 306

10.2 Shrinkage Design Practices . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 10.2.1 “Steel Safe” Mold Design . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 10.2.2 Processing Dependence . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 10.2.3 Semicrystalline Plastics . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 10.2.4 Effect of Fillers . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 10.2.5 Shrinkage Range Estimation . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 10.2.6 Final Shrinkage Recommendations . . . . . . . . . . . . . . . . . . . . . . . . . . .



310 310 311 313 314 314 315

10.3 Warpage . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 317 10.3.1 Sources of Warpage . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 318 10.3.2 Warpage Avoidance Strategies . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 323 10.4 Chapter Review . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 324 10.5 References . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 324

11 Ejection System Design . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .



327

11.1 Objectives in Ejection System Design . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 11.1.1 Allow Mold to Open . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 11.1.2 Transmit Ejection Forces to Moldings . . . . . . . . . . . . . . . . . . . . . . . . . 11.1.3 Minimize Distortion of Moldings . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 11.1.4 Maximize Ejection Speed . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 11.1.5 Minimize Cooling Interference . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 11.1.6 Minimize Impact on Part Surfaces . . . . . . . . . . . . . . . . . . . . . . . . . . . . 11.1.7 Minimize Complexity and Cost . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .

330 330 330 331 331 332 332 333

11.2 The Ejector System Design Process . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 11.2.1 Identify Mold Parting Surfaces . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 11.2.2 Estimate Ejection Forces . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 11.2.3 Determine Ejector Push Area and Perimeter . . . . . . . . . . . . . . . . . . . 11.2.4 Specify Type, Number, and Size of Ejectors . . . . . . . . . . . . . . . . . . . . 11.2.5 Layout Ejectors . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 11.2.6 Detail Ejectors and Related Components . . . . . . . . . . . . . . . . . . . . . . .

333 334 334 340 343 345 348



11.3 Ejector System Analyses and Designs . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 350 11.3.1 Ejector Pins . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 350

XV

XVI

  Contents

11.3.2 Ejector Blades . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 11.3.3 Ejector Sleeves . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 11.3.4 Stripper Plates . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 11.3.5 Elastic Deformation around Undercuts . . . . . . . . . . . . . . . . . . . . . . . . 11.3.6 Core Pulls . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 11.3.7 Slides . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 11.3.8 Early Ejector Return Systems . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .



353 355 356 359 361 366 369

11.4 Advanced Ejection Systems . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 11.4.1 Split Cavity Molds . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 11.4.2 Collapsible Cores . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 11.4.3 Rotating Cores . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 11.4.4 Reverse Ejection . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .



371 371 373 375 377

11.5 Chapter Review . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 378 11.6 References . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 380

12 Structural System Design . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .



12.1 Objectives in Structural System Design . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 12.1.1 Minimize Stress . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 12.1.2 Minimize Mold Deflection . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 12.1.3 Minimize Mold Size . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .

382 382 387 388

12.2 Analysis and Design of Plates . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 12.2.1 Plate Compression . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 12.2.2 Plate Bending . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 12.2.3 Support Pillars . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 12.2.4 Shear Stress in Side Walls . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 12.2.5 Interlocks . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 12.2.6 Stress Concentrations . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .



388 389 392 395 402 404 407

12.3 Analysis and Design of Cores . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 12.3.1 Axial Compression . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 12.3.2 Compressive Hoop Stresses . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 12.3.3 Core Deflection . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .

410 410 412 414

12.4 Fasteners . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 12.4.1 Fits . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 12.4.2 Socket Head Cap Screws . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 12.4.3 Dowels . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .

381

417 417 422 424

12.5 Review . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 426 12.6 References . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 428

Contents

13 Mold Technologies

. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .

429

13.1 Introduction . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 429 13.2 Coinjection Molds . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 431 13.2.1 Coinjection Process . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 431 13.2.2 Coinjection Mold Design . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 433 13.3 Gas Assist/Water Assist Molding . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 434 13.4 Insert Molds . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 13.4.1 Low Pressure Compression Molding . . . . . . . . . . . . . . . . . . . . . . . . . . 13.4.2 Insert Mold with Wall Temperature Control . . . . . . . . . . . . . . . . . . . . 13.4.3 Lost Core Molding . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .



437

437 439 441

13.5 Injection Blow Molds . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 443 13.5.1 Injection Blow Molding . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 443 13.5.2 Multilayer Injection Blow Molding . . . . . . . . . . . . . . . . . . . . . . . . . . . . 445 13.6 Multishot Molds . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 13.6.1 Overmolding . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 13.6.2 Core-Back Molding . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 13.6.3 Multi-station Mold . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .

447 447 449 451

13.7 In-Mold Labeling . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 453 13.7.1 Statically Charged Film . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 454 13.7.2 Indexed Film . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 455 13.8 Review . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 456 13.9 References . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 457

14 Mold Commissioning . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .



459

14.1 Mold Commissioning Objectives . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 14.1.1 Certify Mold Acceptability . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 14.1.2 Optimize Molding Process and Quality . . . . . . . . . . . . . . . . . . . . . . . . 14.1.3 Develop Mold Operation and Maintenance Plans . . . . . . . . . . . . . . .

459 459 461 461

14.2 Commissioning Process . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 14.2.1 Mold Design Checklist . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 14.2.2 Component Verification . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 14.2.3 Mold Assembly . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 14.2.4 Mold Final Test . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 14.2.5 Preliminary Molding Recommendations . . . . . . . . . . . . . . . . . . . . . . .



14.3 Molding Trials . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 14.3.1 Filling Stage . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 14.3.2 Packing Stage . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 14.3.3 Cooling Stage . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .

470 471 473 475

462 465 465 466 466 467

XVII

XVIII

  Contents

14.4 Production Part Approval . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 14.4.1 Quality Assurance . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 14.4.2 Gauge and Process Repeatability & Reproducibility . . . . . . . . . . . . . 14.4.3 Image-Based Dimensional Metrology . . . . . . . . . . . . . . . . . . . . . . . . . . 14.4.4 Process Capability Evaluation . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .



476 476 477 479 481

14.5 Mold Maintenance . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 14.5.1 Pre-Molding Maintenance . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 14.5.2 Molding Observation and Mold Map . . . . . . . . . . . . . . . . . . . . . . . . . . 14.5.3 Post-Molding Maintenance . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 14.5.4 Scheduled Regular Maintenance . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 14.5.5 Mold Rebuilding . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .



485 487 488 489 489 490

14.6 Summary . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 491 14.7 References . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 493

Appendix . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .

495

Appendix A: Plastic Material Properties . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 497 Appendix B: Mold Material Properties . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . B.1 Nonferrous Metals . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . B.2 Common Mold Steels . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . B.3 Other Mold Steels . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .



502 502 503 504

Appendix C: Properties of Coolants . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 505 Appendix D: Statistical Labor Data . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 506 D.1 United States Occupational Labor Rates . . . . . . . . . . . . . . . . . . . . . . . 506 D.2 International Labor Rate Comparison . . . . . . . . . . . . . . . . . . . . . . . . . 506 Appendix E: Unit Conversions . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . E.1 Length Conversions . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . E.2 Mass/Force Conversions . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . E.3 Pressure Conversions . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . E.4 Flow Rate Conversions . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . E.5 Viscosity Conversions . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . E.6 Energy Conversions . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .

508

508 509 509 509 510 510

Appendix F: Estimation of Melt Velocity . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .

511

The Author . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .



515

Index . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .



517

Nomenclature

Nomenclature Mold engineering requires analysis, and so an extensive nomenclature has been developed. L, W, and H refer to the length, width, and height dimensions as shown Mold engineering requires analysis, and so an extensive nomenclature has been developed. in Figure 1. L, W, and H refer to the length, width, and height dimensions as shown in Figure 1.

Figure 1: 1 Length, Length,width, width, and and height Figure height nomenclature nomenclature

Variable names have been selected and consistently used as expected (e.g., T for temperature, have been consistently used as(with expected (e. g., T for CVariable for cost, names P for pressure, etc.).selected R refersand to rate-related constants time dependence) temperature, C for cost, P for pressure, etc.). R refers to rate-related constants (with and κ refers to monetary constants (with cost dependence). To provide clarity, subscripts are time dependence) and κmost refers (withforcost dependence). To unabbreviated throughout of to themonetary book. Theconstants nomenclature many of the variables and their units arevariable as follows. provide clarity, subscripts are unabbreviated throughout most of the book.

The nomenclature for many of the variables and their units are as follows.

Table 1: Nomenclature

Variable

Meaning

α

Thermal diffusivity [m2/s]

β

Compressibility [1/MPa]

δ

Deflection [m]

δbending

Deflection due to bending [m]

δcompression

Deflection due to compression [m]

XX

Nomenclature

Table 1 Nomenclature Variable

Meaning

α

Thermal diffusivity [m2/s]

β

Compressibility [1/MPa]

δ

Deflection [m]

δbending

Deflection due to bending [m]

δcompression

Deflection due to compression [m]

δtotal

Deflection due to bending and compression [m]

ε

Strain [m/m]

εplastic

Plastic’s strain to failure [%]

g

Shear rate [1/s]

g max

Maximum allowable shear rate for a plastic melt being molded [1/s]

η

Viscosity [Pa s]

κ

Thermal conductivity [W/m°C]

κinsert

Cost per unit volume of core and cavity insert materials [$/m3]

κmold

Cost of mold metal per kilogram [$/kg]

κplastic

Cost of plastic per kilogram [$/kg]

λ

Tolerance limit [m]

μ

Apparent viscosity for Newtonian model [Pa]

μstatic

Coefficient of static friction [–]

ρ

Density [kg/m3]

ρinsert

Density of core and cavity insert materials [kg/m3]

ρplastic

Density of plastic [kg/m3]

ϕ

Draft angle [°]

θ

Draft angle [°]

σbuckling

Stress level at which column buckles [MPa]

σcyclic

Imposed cyclic stress [MPa]

σendurance

Maximum allowable stress given cyclic loading [MPa]

σhoop

Hoop stress [MPa]

σlimit

Maximum allowable stress given cyclic loading or yielding [MPa]

σyield

Maximum allowable stress given yield failure [MPa]

τ

Shear stress [Pa]

ν

Specific volume [–]

Ωejectors

Total perimeter of all ejectors [m]

Acavity_Projected

Projected area of the mold cavity [m2]

Acompression

Area exposed to compressive stress [m2]

Aeff

Effective area under stress [m2]

Aejectors

Total area of all ejectors [m2]

Nomenclature

Variable

Meaning

Apart

Total surface area of the molded part

Ashear

Area exposed to shear stress [m2]

C

Tolerance coefficient for a standard fit [m2/3]

Cauxiliaries

Total cost of all auxiliaries [$]

Cinserts

Total cost of all cavities [$]

Cinsert_finishing

Cost of finishing one set of core and cavity inserts [$]

Cinsert_machining

Cost of machining one set of core and cavity inserts [$]

Cinsert_materials

Cost of materials for one set of core and cavity inserts [$]

Cmold

Total cost of purchasing mold [$]

Cmold/part

Cost of purchasing mold amortized across total production quantity [$]

Cmold_base

Total cost of mold base and modifications [$]

Cmold_customization

Total cost of all customizations of mold base [$]

Cmold_steel

Initial purchase cost of mold base or steel [$]

Cpart

Total cost per molded part [$]

Cplastic/part

Cost of material used in molding one part [$]

Cprocess/part

Cost of machinery and labor used to mold one part [$]

CPplastic

Plastic’s specific heat [J/kg °C]

CTE

Coefficient of thermal expansion [1/°C]

CVTE

Coefficient of volumetric thermal expansion [1/°C]

D

Diameter [m]

Dhydraulic

Hydraulic diameter of runner segment [m]

Dpin

Diameter of ejector pin [m]

E

Elastic modulus [GPa]

f

Factory of safety [–]

fcavity_complexity

Factor related to the complexity of the cavity

fcavity_discount

Discount factor related to production of multiple sets of core and cavity inserts

fcycle_efficiency

Factor associated with the efficiency of operating the molding machine

fefficiency

Factor related to the overall efficiency of all machining operations

ffeed_waste

Factor associated with material wasted in molding the feed system

f icavity_customizing

Factor associated with customization of one set of core and cavity inserts

f

i

f

i

finishing

Percentage of the molded part’s surface area to be finished in the manner i

mold_customizing

Factor associated with customizing mold base

fmachine

Factor associated with cost of operating different molding machines and auxiliaries

fmachining

Factor related to the average material removal rate of all machining ­processes relative to standard milling

fmaintenance

Mold lifetime maintenance factor

XXI

XXII

Nomenclature

Variable

Meaning

fwear

Factor associated with maintenance due to mold wear

fyield

Fraction of molded parts that are good

F

Force [N]

Fbuckling

Critical load at which column buckles [N]

Fclamp

Mold force tonnage [metric tons, t]

Feject

Ejection force [N]

Finsertion

Insertion force for interference fit [N]

Ftensile

Maximum tensile force for a socket head cap screw [N]

h

Nominal cavity wall thickness [m]

h∞

Heat transfer coefficient [W/°C]

Hcavity

Height of cavity inserts [m]

Hcore

Height of core inserts [m]

Hinserts

Combined height of core and cavity inserts [m]

HLine

Distance from cavity surface to the center of cooling line [m]

Hmold

Total stack height of mold [m]

Hpart

Maximum height of molded part [m]

I

Moment of inertia [m4]

K

Stress concentration factor [–]

k, n

Reference viscosity and power law index per the power law model [Pan, –]

kplastic

Plastic’s thermal conductivity [W/m °C]

Linserts

Length of core and cavity inserts [m]

Lmold

Length of mold [m]

Lpart

Maximum length of molded part [m]

MFI

Plastic’s melt flow index [g/min]

Mmold

Total mass of mold base [kg]

ncycles

Number of molding cycles [–]

n, t*, D1, D2, D3, A1, A3 WLF model coefficients ncavities

Number of cavities in mold [–]

ncavities_length

Number of cavities in the length direction [–]

ncavities_width

Number of cavities in the width direction [–]

ncycles

Total number of mold cycles that a mold is operated [–]

nj

Number of j-th portions of mold cavity in mold [–]

nlines

Number of cooling lines [–]

nparts

Total production quantity of parts to be molded [–]

P

Pressure [Pa]

Pinject

Pressure required to fill the cavity [Pa]

Qmolding

Total thermal energy of moldings [J] Cooling power per cooling line [W]

line

Nomenclature

Variable

Meaning

Qmolding

Cooling power [W]

rv

Relative change in the specific volume [–]

R

Radius [m]

Re

Reynold’s number [–]

Rfinishing_cost

Hourly cost of finishing [$/h]

Rifinishing

Rate of finishing the part’s surface in the manner i [m2/h]

Rmachining_cost

Hourly rate of machining [$/h]

Rmachining_area

Rate of machining per unit area [m2/h]

Rmachining_volume

Rate of machining per unit volume [m3/h]

Rmolding_cost

Hourly cost of operating molding machine and operator if required [$/h]

RW

Radius of curvature due to warpage [m]

s

Linear shrinkage rate [m/m]

s^

Shrinkage rate perpendicular to flow [m/m]

s//

Shrinkage rate parallel to flow [m/m]

save

Average shrinkage rate [m/m]

tc

Cooling time [s]

tcycle

Cycle time of molding machine [s]

tinsert_area

Time required to machine the cavity surface area for one set of core and cavity inserts [h]

tinsert_finishing

Time required to completely finish one set of core and cavity inserts [h]

tinsert_machining

Time required to perform all machining for one set of core and cavity inserts [h]

tinsert_volume

Time required to machine the cavity volume for one set of core and cavity inserts [h]

tp

Packing time before gate solidification [s]

tresidence

Residence time of the polymer melt [s]

Tc

Mold coolant temperature [°C]

Te

Plastic’s ejection temperature [°C]

Tg

Plastic’s glass transition temperature [°C]

THDT

Plastic’s heat deflection temperature [°C]

Tmelt

Melt temperature [°C]

Twall

Mold wall temperature [°C]

v

Linear melt velocity [m/s]

Vinserts

Combined volume of one set of core and cavity inserts [m3]

Vj

Volume of the j-th portion of mold cavity [m3]

Vpart

Volume of molded part [m3]

v

Volumetric flow rate [m3/s]

Wcavity

Width of core and cavity inserts [m]

Wcheek

Distance from cavity side wall to side of mold [m]

XXIII

XXIV

Nomenclature

Variable

Meaning

Wmold

Width of mold [m]

Wpart

Maximum width of molded part [m]

Wpitch

Distance between parallel cooling lines [m]

1

Introduction

Injection molding is a common method for mass production and is often preferred over other processes, given its capability to economically make complex parts to tight tolerances. Before any parts can be molded, however, a suitable injection mold must be designed, manufactured, and commissioned. The mold design directly determines the molded part quality and molding pro­ ductivity. The injection mold is itself a complex system comprised of multiple components that are subjected to many cycles of temperature and stress. There are ­often trade-offs in mold design, with lower-cost molds sometimes resulting in lower product quality or inefficient molding processes. Engineers should strive to design injection molds that are “fit for purpose”, which means that the mold should produce parts of acceptable quality with minimal life cycle cost while taking a minimum amount of time, money, and risk to develop. This book is directed to assist novice and expert designers of both products and molds. In this chapter, an overview of the injection molding process and various types of molds is provided so that the mold design engineer can understand the basic operation of injection molds. Next, the layout and components in three of the  more common mold designs are presented. The suggested methodology for mold engineering design is then presented, which provides the structure for the ­remainder of this book.

„„1.1 Overview of the Injection Molding ­Process Injection molding is sometimes referred to as a “net shape” manufacturing process because the molded parts emerge from the molding process in their final form with no or minimal post-processing required to further shape the product. An ­operating injection molding machine is depicted in Fig. 1.1. The mold is inserted and clamped between a stationary and moving platen. The mold typically is con-

2

1 Introduction

nected to and moves with the machine platens, so that the molded parts are formed within a closed mold, after which the mold is opened so that the molded parts can be removed.

Figure 1.1 Depiction of an injection molding machine and mold, adapted from [1]

The mold cavity is the “heart” of the mold where the polymer is injected and ­solidified to produce the molded part(s) with each molding cycle. While molding processes can differ substantially in design and operation, most injection molding processes generally include plastication, injection, packing, cooling, and ejection stages. During the plastication stage, a screw within the barrel rotates to convey plastic pellets and form a “shot” of polymer melt. The polymer melt is plasticized from solid granules or pellets through the combined effect of heat conduction from the heated barrel as well as the internal viscous heating caused by molecular deformation as the polymer is forced along the screw flights. Afterwards, during the filling stage, the plasticated shot of polymer melt is forced from the barrel of the molding machine through the nozzle and into the mold. The molten resin ­travels down a feed system, through one or more gates, and throughout one or more mold cavities where it forms the molded product(s). After the mold cavity is filled with the polymer melt, the packing stage provides additional material into the mold cavity as the molten plastic melt cools and contracts. The plastic’s volumetric shrinkage varies with the material properties and application requirements, but the molding machine typically forces 1 to 10 % addi-

1.1 Overview of the Injection Molding ­Process

tional melt into the mold cavity during the packing stage. After the polymer melt ceases to flow, the cooling stage provides additional time for the resin in the cavity to solidify and become sufficiently rigid for ejection. Then, the molding machine actuates the moving platen and the attached moving side of the mold to provide access to the mold cavities. The mold typically contains an ejection system with moving slides and pins that are then actuated to remove the molded part(s) prior to mold closure and the start of the next molding cycle. A chart plotting the timing of each stage of the molding process is shown in Fig. 1.2 for a molded part approximately 2 mm thick having a cycle time of 30 s. The filling time is a small part of the cycle and so is often selected to minimize the ­injection pressure and molded-in stresses. The packing time is of moderate duration, and is often minimized through a shot weight stability study to end with freeze-off of the polymer melt in the gate. In general, the cooling stage of the molding process dominates the cycle time since the rate of heat flow from the polymer melt to the cooler mold is limited by the low thermal diffusivity of the plastic melt. However, the plastication time may exceed the cooling time for very large shot ­volumes with low plastication rates. The mold reset time is also very important to minimize since it provides negligible added value to the molded product. To minimize the molding cycle time and costs, molders strive to operate fully automatic processes with minimum mold opening and ejector strokes. The operation of fully automatic molding processes requires careful mold design, making, and ­commissioning. Not only must the mold operate without any hang-ups, but the quality of the molded parts must consistently meet specification.

Figure 1.2 Injection molding process timings

Figure 1.2 also shows the possible cycle timings for a more advanced mold design using additional investment in technology. Hot runner feed systems, for example, allow the use of less plastic material while also reducing injection and pack times.

3

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1 Introduction

Conformal cooling and highly conductive mold inserts can significantly reduce cooling times. Molds and molding processes can also be optimized to minimize mold opening, part ejection, and mold closing times. The net result of additional engineering is a reduction in the cycle time from 30 to 18 s. While some cycle time improvements are often possible just through careful engineering design, many productivity improvements require additional upfront investment in mold mate­ rials, components, or processing. There are also many variants of the injection molding process (such as gas assist molding, water assist molding, insert molding, two shot molding, coinjection molding, injection compression molding, and others discussed later) that can be used to provide significant product differentiation or cost advantages. These more ­advanced processes can greatly increase the quality of the molded parts but at the same time can increase the complexity and risk of the mold design and molding processes while also limiting the number of qualified suppliers. As such, the product design and mold design should be conducted concurrently while explicitly addressing manufacturing strategy and supply chain considerations. The cost of advanced mold designs must be justified either by net cost savings or increases in the customer’s willingness to pay for advanced product designs. Cost estimation thus serves an important role in developing appropriate manufacturing strategies and mold designs.

„„1.2 Mold Functions The injection mold is a complex system that must simultaneously meet many ­demands imposed by the injection molding process. The primary function of the mold is to contain the polymer melt within the mold cavity so that the mold cavity can be completely filled to form a plastic component whose shape replicates the mold cavity. A second primary function of the mold is to efficiently transfer heat from the hot polymer melt to the coolant flowing through the mold, such that in­ jection molded products may be produced as uniformly and rapidly as possible. A third primary function of the mold is to eject the part from the mold in an efficient and consistent manner without imparting excessive stress to the moldings. These three primary functions—contain the melt, transfer the heat, and eject the molded part(s)—also place secondary requirements on the injection mold. Figure 1.3 provides a partial hierarchy of the functions of an injection mold. For example, the function of containing the melt within the mold requires that the mold: ƒƒresist displacement under the enormous forces that will tend to cause the mold to open or deflect. Excessive displacement can directly affect the dimensions of the

1.2 Mold Functions

moldings or allow the formation of flash around the parting line of the moldings. This function is typically achieved through the use of rigid plates, support pillars, and interlocking components. ƒƒguide the polymer melt from the nozzle of the molding machine to one or more cavities in the mold where the product is formed. This function is typically fulfilled through the use of a feed system and flow leaders within the cavity itself to ensure laminar and balanced flow.

Figure 1.3 Function hierarchy for injection molds

It should be understood that Fig. 1.3 does not provide a comprehensive list of all functions of an injection mold, but just some of the essential primary and secondary functions that must be considered during the engineering design of injection molds. Even so, a skilled designer might recognize that conflicting requirements are placed on the mold design by various functions. For instance, the desire for ­efficient cooling may be satisfied by the use of multiple tightly spaced cooling lines that conform to the mold cavity. However, the need for part removal may require the use of multiple ejector pins at locations that conflict with the desired cooling line placement. It is up to the mold designer to consider the relative importance of the conflicting requirements and ultimately deliver a mold design that is satis­ factory. ­ esign. There are significant compromises and potential risks associated with mold d In general, smaller and simpler molds may be preferred since they use less material and are easier to operate and maintain. Conversely, it is possible to under-­ design molds such that they may deflect under load, wear or fail prematurely, or require extended cycle times to operate. Because the potential costs of failure are

5

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1 Introduction

often greater than the added cost to ensure a robust design, there is a tendency to over-design with the use of conservative estimates and safety factors when in doubt. Excessive over-designing should be avoided since it can lead to large, costly, and inefficient molds.

„„1.3 Mold Structures An injection mold has many structures to accomplish the functions required by the injection molding process. Since there are many different types of molds, the ­structure of a simple “two-plate” mold is first discussed. It is important for the mold designer to know the names and functions of the mold components, since later chapters will assume this knowledge. The design of these components and more complex molds will be analyzed and designed in subsequent chapters.

1.3.1 External View of Mold An isometric view of a two-plate mold is provided in Fig. 1.4. From this view, it is observed that a mold is constructed of a number of plates bolted together with socket head cap screws. These plates commonly include the top clamp plate, the cavity insert retainer plate or “A” plate, the core insert retainer plate or “B” plate, a support plate, and a rear clamp plate or ejector housing. Some mold components are referred to with multiple names. For instance, the “A” plate is sometimes ­referred to as the cavity insert retainer plate, since this plate retains the cavity ­inserts. As another example, the ejector housing is also sometimes referred to as the rear clamp plate, since it clamps to the moving platen located towards the rear of the molding machine. In some mold designs, the ejector housing is replaced with a separable rear clamp plate of uniform thickness and two parallel ejector “rails” that replace the side walls of the integral “U”-shaped ejector housing. This alternative rear clamp plate design requires more components and mold-making steps, but can provide material cost savings as well as mold design flexibility. The mold depicted in Fig. 1.4 is referred to as a “two-plate mold” since it uses only two plates to contain the polymer melt. Mold designs may vary significantly while performing the same functions. For example, some mold designs integrate the “B” plate and the support plate into one extra-thick plate, while other mold designs may integrate the “A” plate and the top clamp plate. As previously mentioned, some mold designs may split up the ejector housing, which has a “U”-shaped profile to house the ejection mechanism and clamping slots, into a rear clamp plate and tall

1.3 Mold Structures

rails (also known as risers). The use of an integrated ejector housing as shown in Fig. 1.4 provides for a compact mold design, while the use of separate rear clamp plate and rails provides for greater design flexibility.

Figure 1.4 View of a closed two-plate mold

To hold the mold in the injection molding machine, toe clamps are inserted in slots adjacent to the top and rear clamp plates and subsequently bolted to the stationary and moving platens of the molding machine. A locating ring, usually found at the center of the mold, closely mates with an opening in the molding machine’s ­stationary platen to align the inlet of the mold to the molding machine’s nozzle. The opening in the molding machine’s stationary platen can be viewed in Fig. 1.1 around the molding machine’s nozzle. The use of the locating ring is necessary for at least two reasons. First, the inlet of the melt to the mold at the mold’s sprue bushing must mate with the outlet of the melt from the nozzle of the molding ­machine. Second, the ejector knockout bar(s) actuated from behind the moving platen of the molding machine must mate with the ejector system of the mold. Molding machine and mold suppliers have developed standard locating ring spe­ cifications to facilitate mold-to-machine compatibility, with the most common ­locating ring diameter being 100 mm (4 in). When the molding machine’s moving platen is actuated, all plates attached to the rear clamp plates will be similarly actuated and cause the mold to separate at the parting plane. When the mold is closed, guide pins and bushings are used to

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1 Introduction

closely locate the “A” and the “B” plates on separate sides of the parting plane, which is crucial to the primary mold function of containing the melt. Improper design or construction of the mold components may cause misalignment of the “A” and “B” plates, poor quality of the molded parts, and accelerated wear of the in­ jection mold.

1.3.2 View of Mold during Part Ejection Another isometric view of the mold is shown in Fig. 1.5, oriented horizontally for operation with a horizontal injection molding machine. In this depiction, the ­plastic melt has been injected and cooled in the mold, such that the moldings are now ready for ejection. To perform ejection, the mold is opened by at least the height of the moldings. Then, the ejector plate and associated pins are moved ­forward to push the moldings off the core. From this view, many of the mold components are observed, including the “B” or core insert retainer plate, two different core inserts, feed system, ejector pins, and guide pins and bushings.

Figure 1.5 View of molding ejected from injection mold

1.3 Mold Structures

Figure 1.5 indicates that the plastic molding consists of two different molded parts (like a cup and a lid) attached to a feed system. This mold is called a two-plate, cold-runner, or two-cavity family mold. The term “family mold” refers to a mold in which multiple components of varying shapes and/or sizes are produced at the same time, most commonly to be used in a product assembly. The term “two-­cavity” refers to the fact that the mold has two cavities to produce two moldings in each molding cycle. Such multicavity molds are used to rapidly and economically produce high quantities of molded products. Molds with eight or more cavities are common. The number of mold cavities is a critical design decision that impacts the technology, cost, size, and complexity of the mold; a cost estimation method is provided in Chapter 3 to provide design guidance. In a multicavity mold, the cavities are placed across the parting plane to provide room between the mold cavities for the feed system, cooling lines, and other ­components. It is generally desired to place the mold cavities as close together as possible without sacrificing other functions such as cooling, ejection, etc. This ­usually results in a smaller mold that is not only less expensive, but is also easier for the molder to handle while being usable in more molding machines. The number of mold cavities in a mold can be significantly increased by not only using a larger mold, but also by using different types of molds such as a hot runner mold, three-plate mold, or stack mold as later discussed with respect to mold layout design in Chapter 4.

1.3.3 Mold Cross-Section and Function Figure 1.6 shows the top view of the mold, along with the view that would result if the mold was physically cut along the section line A-A and viewed in the direction of the arrows. Various hatch patterns have been applied to different components to facilitate identification of the components. It is very important to understand each of these mold components and how they interact with each other and the molding process.

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1 Introduction

Figure 1.6 Top and cross-section views of a two-plate mold

Consider now the stages of the molding process relative to the mold components. During the filling stage, the polymer melt flows from the nozzle of the molding machine through the orifice of the sprue bushing. The melt flows down the length of the sprue bushing and into the runners located on the parting plane. The flow then traverses across the parting plane and enters the mold cavities through small

1.4 Other Common Mold Types

gates. The melt flow continues until all mold cavities are completely filled. Chapters 5, 6, and 7 provide analysis and design guidelines for flow in the mold cavity, feed system, and gates. As the polymer melt fills the cavity, the displaced air must be vented from the mold. Some analysis and design guidelines are provided in Chapter 8. After the polymer melt flows to the end of the cavity, additional material is packed into the cavity at high pressure to compensate for volumetric shrinkage of the plastic as it cools. The estimation of shrinkage and guidelines for steel-safe design are described in Chapter 9. Typically, the injection molding pressure, temperature, and timing are adjusted to achieve the desired part dimensions. The duration of the packing phase is typically controlled by the size and freeze-off of the gate ­between the runner and the cavity. During the packing and cooling stages, heat from the hot polymer melt is transferred to the coolant circulating in the cooling lines. The heat transfer properties of the mold components, together with the size and placement of the cooling lines, determines the rate of heat transfer and the cooling time required to solidify the plastic. At the same time, the mold components must be designed to resist deflection and stress when subjected to high melt pressures. Chapters 10 and 11 describe the analysis and design of the mold’s cooling and structural systems. After the part has cooled, the molding machine’s moving platen is actuated and the moving half of the mold (consisting of the “B” plate, the core inserts, the support plate, the ejector housing, and related components) moves away from the stationary half (consisting of the top clamp plate, the “A” plate, the cavity inserts, and other components). Typically, the moldings stay with the moving half since they have shrunken onto the core. This shrinkage results in tensile stresses, like a ­rubber band stretched around a cylinder or box, that will tend to keep the moldings on the core. After the mold opens, the ejector plate is pushed forward by the molding machine. The ejector pins are driven forward and push the moldings off the core. The moldings may then drop out of the mold or be picked up by an operator or robot. Afterwards, the ejector plate is retracted and the mold closes to receive the melt during the next molding cycle. The ejector system design is analyzed in Chapter 12.

„„1.4 Other Common Mold Types A simple two-plate mold has been used to introduce the basic components and functions of an injection mold. About half of all molds closely follow this design, since the mold is simple to design and economical to produce. However, the twoplate mold has many limitations, including:

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1 Introduction

ƒƒrestriction of the feed system route to the parting plane; ƒƒlimited gating options from the feed system into the mold cavity or cavities; ƒƒrestriction on the tight spacing of cavities; ƒƒadditional forces imposed on the mold by the melt flowing through (and being pressurized within) the feed system; ƒƒincreased material waste incurred by the solidification of the melt in the feed system; and ƒƒincreased cycle time related to the plastication and cooling of the melt in the feed system. For these reasons, molding applications requiring high production quantities often do not use two-plate mold designs, but instead use mold designs that are more complex, yet provide for lower-cost production of the molded parts. Such designs include three-plate molds, hot runner molds, stack molds, and others. Three-plate molds and hot runner molds are the next most common types of injection molds, and so are introduced next.

1.4.1 Three-Plate, Multicavity Family Mold The three-plate mold is so named since it provides a third plate that floats between the mold cavities and the top clamp plate. Figure 1.7 shows a section of a threeplate mold that is fully open with the moldings still on the core inserts. As shown in Fig. 1.7, the addition of the third plate provides a second parting plane between the “A” plate assembly and the top clamp plate for the provision of a feed system that traverses parallel to the parting plane. During molding, the plastic melt flows out the nozzle of the molding machine, down the sprue bushing, across the p ­ rimary runners, down the sprues, through the gates, and into the mold cavities. The feed system then freezes in place with the moldings. When the mold is opened, the molded cold runner will stay on the stripper plate due to the inclusion of sprue pullers that protrude into the primary runner. As the mold continues to open, the stripper bolt connected to the “B” plate assembly will pull the “A” plate assembly away from the top clamp plate. Another set of stripper bolts will then pull the stripper plate away from the top clamp plate, stripping the molded cold runner off the sprue pullers. The ejector plate may be designed and actuated as in a traditional two-plate mold to force the moldings off the core. The three-plate mold eliminates two significant limitations of two-plate molds. First, the three-plate mold allows for primary and secondary runners to be located in a plane above the mold cavities so that the plastic melt in the cavities can be gated at any location. Such gating flexibility is vital to improving the cost and ­quality of the moldings, especially for molds with a high number of cavities.

1.4 Other Common Mold Types

­ econd, the three-plate mold provides for the automatic separation of the feed S ­system from the mold cavities. Automatic de-gating facilitates the operation of the molding machine with a fully automatic molding cycle to reduce molding cycle times.

Figure 1.7 Section of an open three-plate mold

There are at least three significant potential issues with three-plate molds, however. First and most significantly, the cold runner is molded and ejected with each molding cycle. If the cold runner is large compared to the molded parts, then the molding of the cold runner may increase the material consumption and cycle time, thereby increasing the total molded part cost. Second, the three-plate mold r­ equires additional plates and components for the formation and ejection of the cold runner, which increases the cost of the mold. Third, a large mold-opening stroke is needed to eject the cold runner. The large mold-opening height (from the top of the top clamp plate to the back of the rear clamp) may be problematic and require a molding machine with greater “daylight” between the machine’s stationary and moving platens than would otherwise be required for a two-plate or hot runner mold.

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1 Introduction

1.4.2 Hot Runner, Multigated, Single Cavity Mold Hot runner molds provide the benefits of three-plate molds without their disadvantages, yet give rise to other issues. The term “hot runner” is used since the feed system is typically heated and so remains in a molten statue throughout the entire molding cycle. As a result, the hot runner does not consume any material or cycle time associated with conveying the melt from the molding machine to the mold cavities. A section of a multigated single-cavity mold is provided in Fig. 1.8. This mold ­contains a single cavity, which is designed to produce the front housing or “bezel” for a laptop or tablet computer. The hot runner system includes a hot sprue bushing, a hot manifold, and two hot runner nozzles as well as heaters, cabling, and other components related for heating. The hot runner system is carefully designed to minimize the heat transfer between the hot runner system and the surrounding mold through the use of air gaps and minimal contact area. Like the three-plate mold design, the primary and secondary runners are routed in the manifold above the mold cavities to achieve flexibility in gating locations. Since the polymer melt stays molten, hot runners can be designed to provide larger flow bores and excellent pressure transmission from the molding machine to the mold cavities. As such, hot runner system can facilitate the molding of thinner parts with faster ­cycle times than either two-plate or three-plate molds, while also avoiding the scrap associated with cold runners. During the molding process, the material injected from the machine nozzle into the hot sprue bushing pushes the existing material in the hot runner system into the mold cavity. When the mold cavities fill, the hot runner’s thermal gates are designed to solidify and prevent the leakage of the hot polymer melt from inside the hot runner system to the outside of the mold when the mold is opened. The melt pressure developed inside the hot runner system at the start of the next molding cycle will cause these thermal gates to rupture and allow the flow of the polymer melt into the mold cavity. There are many different hot runner and gating designs that can provide advantages that include gating flexibility, improved pressure transmission, reduced ­material consumption, and increased molding productivity. However, there are also at least two significant disadvantages. First, hot runner systems require added investment for the provision and control of the hot runner temperature. The added investment can be a significant portion of the total mold cost, and not all molders have the auxiliary equipment or expertise to operate and maintain hot runner molds. The second disadvantage of hot runner systems is extended change-over times associated with the purging of the contained polymer melt. In short run production applications having aesthetic requirements, the number of cycles required to start up or change resins or even color may be unacceptable.

1.4 Other Common Mold Types

Figure 1.8 Section of hot runner mold

1.4.3 Comparison The type of feed system is a critical decision that is made early in the development of the mold design. From a mold designer’s perspective, the choice of feed system has a critical role in the design of the mold, the procurement of materials, and the mold making, assembly, and commissioning. From the molder’s perspective, the choice of feed system largely determines the purchase cost, molding productivity, and operating cost of the mold. Table 1.1 compares the different types of molds with respect to several perform­ ance measures. In general, hot runner molds are excellent with respect to molding cycle performance but poor with respect to initial investment, start-up, and on-going maintenance. By comparison, two-plate molds have lower costs but provide limited molding cycle productivity. The evaluation of three-plate molds in T ­ able 1.1 warrants some further discussion. Specifically, three-plate molds do not provide as

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1 Introduction

high a level of molding productivity compared to hot runner molds, and at the same time have higher costs than two-plate molds. For this reason, there has been a trend away from three-plate molds with the penetration of lower cost hot runner systems. Table 1.1 Feed System Comparison Performance measure

Two-plate

Three-plate

Hot runner

Gating flexibility

Poor

Excellent

Excellent

Material consumption

Good

Poor

Excellent

Cycle times

Good

Poor

Excellent

Initial investment

Excellent

Good

Poor

Start-up times

Excellent

Good

Poor

Maintenance cost

Excellent

Good

Poor

„„1.5 The Mold Development Process Given that there is substantial interplay between the product design, mold design, and the injection molding process, an iterative mold development process is frequently used as shown in Fig. 1.9. To the extent possible, the product design should follow standard design for injection molding guidelines as described in Chapter 2. To reduce the product development time, the product design and mold design are often performed concurrently. In fact, a product designer may receive a reasonable cost estimate for a preliminary part design given only the part’s overall dimensions, thickness, material, and production quantity. Given this information, the mold designer develops a preliminary mold design and provides a preliminary quote as discussed in Chapter 3. This preliminary quote requires the molder and mold maker to not only develop a rough mold design but also to estimate important processing variables such as the required clamp tonnage, machine hourly rate, and cycle times. Once a quote is accepted, the detailed engineering design of the mold can begin in earnest as indicated by the listed steps on the right side of Fig. 1.9. First, the mold designer will lay out the mold design by specifying the type of mold, the number and position of the mold cavities, and the size and thickness of the mold. Afterwards, each of the required sub-systems of the mold is designed, which sometimes requires the redesign of previously designed subsystems. For example, the placement of ejector(s) may require a redesign of the cooling system while the design of the feed system may affect the layout of the cavities and other mold components.

1.5 The Mold Development Process

Multiple design iterations are typically conducted until a reasonable compromise is achieved between size, cost, complexity, and function.

Figure 1.9 A mold development process

To reduce the development time, the mold base, insert materials, hot runner ­system, and other components may be ordered and customized as the mold design is being fully detailed. Such concurrent engineering should not be applied to unclear aspects of the design. However, many mold makers do order the mold base and plates upon confirmation of the order. As a result of concurrent engineering practices, mold development times are now typically measured in weeks rather than months [1]. Customers have traditionally placed a premium on quick mold delivery, and mold makers have traditionally charged more for faster service. With competition, however, customers are increasingly requiring guarantees on mold delivery and quality, with penalties applied to missed delivery times or poor ­quality levels. After the mold is designed, machined, polished, and assembled, molding trials are performed to verify the basic functionality of the mold. If no significant deficiencies are present, the moldings are sampled and their quality assessed relative to specifications. Usually, the mold and molding process are sound but must be tweaked to improve the product quality and reduce the product cost. However, sometimes molds include “fatal flaws” that are not easily correctable and may ­necessitate the scrapping of the mold and a complete redesign. Some guidelines for mold commissioning and first article inspection are provided in Chapter 13.

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1 Introduction

„„1.6 Mold Standards The designs depicted in this chapter were designed from computer aided design (CAD) files of mold bases from DME Company. A “mold base” is essentially a blank mold or template design that includes all the plates, pins, bushings, and other components that may be purchased as a fully assembled system and modified for a specific molding application. Figure 1.10 is provided as the prototypical mold base. This particular design [2] was made in 1944 by Ivar Quarnstrom, the founder of Detroit Mold Engineering (DME Company). It is remarkable how similar the design of Fig. 1.10 is to that of Fig. 1.6 and other designs commonly observed today.

Figure 1.10 Mold base

There are many benefits for mold designs that rely on the use of the standard mold bases. These include: ƒƒFirst and foremost, the design of the mold base including its many detailed fits and tolerances would require extensive analysis and care in manufacturing. In other words, most mold designers and mold makers would have difficulty designing as good a mold base at as low a cost as a standard product that could be purchased off the shelf with minimal risk and lead times.

1.6 Mold Standards

ƒƒSecond, the use of standards provides for potential interoperability of mold bases and mold base components across molding applications as well as different molding facilities. For example, a mold designer may wish to provide six identical molds so that two copies of each mold are operable in Europe, the Americas, and Asia. The use of a mold base not only supports the mold design with respect to a CAD library, but also the provision of the replacement components using the mold base supply chain should mold components become damaged or worn. ƒƒThird, the use of a standard mold base supports for more rapid communication of the mold design with other industry practitioners. For example, consider the use of a mold base with a sprue bushing compared to a custom design that threads the molding machine nozzle directly into a mold cavity. The use of the sprue bushing may increase the component count, but supports ready replacement and standard interfacing to a variety of molding machines. Conversely, the ­directly threaded nozzle eliminates the sprue altogether and so provides better molding produc­tivity, but requires more skill in designing and operation. There is certainly the oppor­tunity for mold designers, mold makers, and molders to outperform mold bases using custom mold designs from scratch. These masters have developed significant experience and insight into their molding applications that motivates and supports their custom designs. For all these reasons, mold designers and makers in developed countries, where labor is relatively expensive compared to the mold materials and components, will typically use standard mold bases. There are many suppliers of mold bases who compete with different strategies including material technology, quality, lead time, cost, size, breadth of product line, unit systems, regional distribution, and others. Product designers, mold designers, mold makers, and molders should verify what system of suppliers are to be used in a given application. It should be noted that the costs of fully realized molds will vary greatly, and not solely as a function of design and quality. The author has conducted research into mold quoting, and so is aware of instances where fully designed, machined, and finished molds have been purchased for less than the cost of just the mold base in the United States. These occurrences are often the result of inferior designs, ­materials, and labor practices that require extensive rework and still perform at marginally acceptable levels. With further globalization of industry, labor rates and material costs will continue to equilibrate so product and mold designers may expect to best compete on the innovation and efficiency of their designs [3]. Adherence to standards and good engineering practices are vital to long-term competitiveness. The Society of the Plastics Industry (SPI) has provided specifications for Class 101, 102, and 103 molds intended for production of more than 1,000,000 cycles, 500,000 cycles, and 250,000 cycles, respectively. Some of the specifications are quantified. For example, Class 101 and 102 molds are required to have a Brinell

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1 Introduction

Hardness Number (BHN) of 280 while Class 103 molds only require a BHN of 165. Other specifications are not quantitatively specified. For example, Class 101 molds are to have adequate channels for temperature control. Meanwhile, other specifi­ cations (like melt flow balancing and energy efficiency) are completely omitted. The engineering design and analysis methodologies presented throughout this book will assist product and mold designers to attain the best possible molds and molded products.

„„1.7 Chapter Review After reading this chapter, you should understand: ƒƒThe basic stages of the injection molding process, ƒƒThe primary functions of an injection mold, ƒƒThe most common types of injection molds (two-plate, three-plate, hot runner, single cavity, multicavity, and multigated mold), ƒƒThe key components in an injection mold, ƒƒThe mold development process, and ƒƒThe motivation for standards in mold design and mold making. In the next chapter, the typical requirements of a molded part are described along with design for injection molding guidelines. Afterwards, the mold layout design and detailed design of the various systems of a mold are presented.

„„1.8 References [1] Catanzaro, J. C. and R. M. Sparer, Clamp force control, U.S. Patent No. 5,149,471, Sept. 22 (1992) [2] Quarnstrom, I. T., Mold Base, U.S. Patent No. 2,419,089, filed 6/9/44 (1947) [3] Kazmer, D. O., Manufacturing outsourcing, onshoring, and global equilibrium, Business Horizons (2014) 57(4): pp. 463–472

2

Plastic Part Design

„„2.1 The Product Development Process Mold design is one significant activity in a much larger product development ­process. Since product and mold design are interdependent, it is useful for both product and mold design engineers to understand the plastic part development process and the role of mold design and mold making. A typical product development process is presented in Fig. 2.1, which includes different stages for product definition, product design, business and production development, ramp-up, and launch. While there are many product development processes, most share two critical ­attributes: ƒƒa structured development plan [1] to coordinate concurrent design activities to ensure tracking and completeness of the design and manufacturing according to schedule and performance requirements, and ƒƒa gated management process [2] to mitigate risk by allocating larger budgets only after significant reviews at project milestones. The product development process shown in Fig. 2.1 is split into multiple stages separated by approval toll-gates. An overview of each stage is next provided.

22

2 Plastic Part Design

Figure 2.1 A product development process

2.1.1 Product Definition The product development process typically begins with product definition [3, 4], which includes a formal analysis of the market, benchmarking of competitors, ­definition of the product specifications, and assessment of potential profitability. If management agrees that a new product is to be developed, then an appropriate team is assembled to perform the early concept design and business development. During this first stage, the approximate size, properties, and cost of the product are estimated. Sketches, mock-ups, and prototypes may be produced to communicate and assess the viability of the concept. With respect to profitability, market studies during the early product development stage will strive to predict the potential sales at varying price points. At the same time, labor and project cost estimates will establish the budget required to develop and bring the product to market. A management review of the concept design,

2.1 The Product Development Process

sales forecast, and budget is usually performed to assess the likelihood of the ­commercial success of continued product development. At this time, the proposed product development project may be approved, declined, shelved, or modified ­accordingly.

2.1.2 Product Design If the project is approved and a budget is allocated, then the product development process continues, usually with additional resources to perform further analysis and design. During this second stage, each component in the product is designed in detail. The design of plastic components may include the consideration of aesthetic, structural, thermal, manufacturing, and other requirements. Design for manufacturing methods [5] may be used to identify issues that would inhibit the effective manufacturing of the components. Design for assembly methods [6] may be used to reduce the number of components, specify tolerances on critical dimensions, and ensure the economic assembly of the finished product. The outcome of the initial product design stage (through the second management approval in Fig. 2.1) is a detailed and validated product design. The term “detailed design” implies that every component is fully specified with respect to material, geometric form, surface finish, tolerances, supplier, and cost. If a custom plastic component is required, then quotes for these molded parts are often requested during this stage. These costs are presented to management along with the detailed design for approval. If the product design and costs are acceptable, then the required budget is allocated and the product development now focuses on manufacturing.

2.1.3 Development While mold design and mold making are a focus of those in the plastics industry, all these activities are encompassed by the single activity titled “Tooling Fabrication” in Fig. 2.1. At the same time, vital business development and production planning are being performed. Specifically, business development is required to fully define the supply chain and establish initial orders to support the product launch. Production planning is required to lay out assembly lines, define labor ­requirements, and develop the manufacturing infrastructure. When the mold tooling is completed, “alpha” parts are produced, tested, and ­assembled. This “first article inspection” includes a battery of tests to verify per­ formance levels, regulatory compliance, and user satisfaction. If the individual components or assembled alpha product are not satisfactory, then the manufacturing

23

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2 Plastic Part Design

processes, associated tooling, and detailed component designs are adjusted as ­appropriate. Typical issues discovered at this stage include [7]: ƒƒinappropriate performance with respect to stiffness, impact, thermal, color, assembly fits, or other attributes. These issues are often due to uncertain material properties, unidentified customer preferences that require changes to the design specification, or errors in analysis or simulation of the product performance. ƒƒproduction of defective product due to mold design or tooling issues. Common examples include dimensions that are outside of specification due to shrinkage and warpage, as well as poor product aesthetics due to knit-lines, poor gating, or surface finish. ƒƒexcessive production costs related to material consumption or processing time. When quality issues are encountered, it is often possible to provide remedies through processing strategies that include increasing the temperatures, pressures, or cycle times, which then increase processing cost. Similarly, it is pos­ sible that only a fraction of the sampled products are acceptable, which results in increased material and inspection costs. Mold designers will work with product designers and injection molders to optimize the molded product quality. Concurrently, the operations staff develops detailed plans governing quality control and worker training.

2.1.4 Scale-Up and Launch A management review is often used to verify that the developed product designs and production plans are satisfactory. Prior to commercial sale, a pilot production run may be implemented at each manufacturing site to produce a moderate quantity of products to verify quality and define standard operating processes [8]. These manufactured “beta” products are frequently provided to the marketing department, sales force, and key customers to ensure product acceptability. As before, the design and manufacturing of the product may be revised to address any remaining issues. When all stakeholders (marketing, sales, manufacturing, critical suppliers, and critical customers) are satisfied, the pilot production processes are ramped up to build an initial inventory of the product (referred to as “filling the channels”), after which the product is released for sale.

2.1.5 Role of Mold Design Mold quoting, mold design, and mold making support this larger product development process. Requests for mold and/or part cost quotes are usually made towards the end of the concept design stage or near the beginning of the detailed design

2.2 Design Requirements

stage. It is somewhat unusual for the molder or the mold maker to be given fully detailed designs at this time, since 1) much of the mold design could have been performed concurrently with a less developed product design, and 2) the mold ­engineering process may suggest significant changes in the design related to ­manufacturability or part performance. The mold development process (first introduced in Fig. 1.9) often begins with a preliminary design that is lacking in detail and would result in an unsatisfactory product if used directly. The critical part design information required to begin the mold concept design includes the part size, wall thickness, and expected production quantity. Given just this information, the mold designer can begin to develop initial mold layouts, cost estimates, and product design improvements [9]. To ­accelerate the product development process, mold design can be performed concurrently with the procurement and customization of the mold components. For better or for worse, mold making and commissioning occurs near the end of the product development process. For this reason, there can be significant pressure on mold suppliers and molders to provide high quality moldings as soon as possible. This task can be extremely challenging given potential mistakes made earlier in the product design process. As such, mold designers may be required to redesign and change portions of the mold and work closely with molders to qualify the mold for production.

„„2.2 Design Requirements There are many requirements of an injection molded part that need to be considered during the mold design. Figure 2.2 provides a worksheet to document the molded part requirements that subsequently guide the material selection and mold design. The design documentation discussed in this section is largely motivated for two reasons. First, detailed and available documentation will improve the design and reduce the cost of the mold engineering. Second, ISO and other regulatory agencies often require formal documentation and approval of product development. ­Accordingly, the plastic product designer and mold designer should not consider these worksheets as static pages, but rather as living documents that are linked to design decisions and decision-making processes with routing from and to the ­related coworkers for information and approval.

25

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2 Plastic Part Design

MOLD DESIGN APPLICATION - PROJECT _______________________ (#_________________) Customer: Technical contact:

Molder: Technical contact:

Project manager:

Project manager:

PRODUCTION PLANNING #Molds Planned: #Cavies/mold: Cycle me: Producon hours: Annual producon: Target producon: PART REQUIREMENTS General tolerance, %: Crical tolerance(s): Temperature range: Max. deflecon: Max. stress: Creep resistance:  Fague resistance:  Impact resistance:  Chemical resistance:  Comments: DESIGN FOR MOLDING Wall thickness: Flow length: Tolerances: No sharp corners:  Effecve rib design:  Effecve boss design:  Dra‘ applied:  Undercuts avoided:  Gang known:  DESIGN FOR ASSEMBLY Minimize # parts:  Top down assembly:  Snap fits/adhesives:  Uniform fasteners:  Part (a)symmetric:  Ejecon/handling:  Comments:

TARGET PART COSTS Material/part: Mold/part: Molding/part: 2nd ops/part: Indirect/part: Total cost/part:

SCHEDULED DATES Project iniated: Cavies required: Mold inspecon: Molding trials: First arcle approval: Producon start:

DESIGN STANDARDS Drawing: ANSI ISO DIN JIS GB ______ ANSI/ASTM: FDA: IEC: ISO: Milspec: UL: Color: DIN RAL NCS Pantone ______ Gloss level: Color match:

MOLD REQUIREMENTS Mold lifeme: SPI mold class: Target IMM: #Gates/Cavity: Family mold Precision mold 2 plate 3 plate Hot runner Slides Li‘ers Cores Interlocks In-mold sensors Other: _____________________ Other: _____________________ Surface treatment: SPI Finish: Surface Texture:

PART MATERIAL SPECS Trade name: Generic grade: Fillers: Cost: Melt temp: Mold temp: Density: DTUL: Strain to yield: Modulus: Yield stress: CTE: Shrinkage rate, %: Max. shear rate: Comments:

MOLD MATERIAL SPECS Trade name: Generic grade: Alloys: Density: Modulus: Yield stress: Fague limit stress: Hardness, Brinell: Strain to yield: Density: Specific heat: Thermal conducvity: Cužng speed: Feed per tooth: Comments:

Figure 2.2 Mold application engineering worksheet

2.2.1 Application Engineering Information The mold design engineer should understand the overall application and development schedule. As indicated at the top of Fig. 2.2, the mold designer should know the appropriate technical contact for questions related to the part design and its requirements. Ideally, this contact should understand the critical requirements of the molded part or be able to refer the mold designer to other, more knowledgeable

2.2 Design Requirements

people. Alternatively, the mold designer may contact the internal project or ­applications manager responsible for supporting the customer so as to avoid continuously contacting the customer regarding what may be considered as potentially trivial issues. The mold designer should also know the technical contact and project manager at the molder to verify molding preferences, machine specifications, and project coordination. As indicated near the top right of Fig. 2.2, there are several critical milestones that should be tracked, including the dates for project initiation, machined cavities, mold inspection, molding trials, first article inspection/approval, and volume production. These dates are frequently negotiated since they are related to technical feasibility, market success, and also payment terms.

2.2.2 Production Planning The production planning data requested near the top left of Fig. 2.2 is very im­ portant with respect to the selection of the mold layout and mold technology. In particular, the application lifetime and total production quantity is related to the determination of the type of tooling (Class 101, etc.) and with it the specification of the mold materials and treatments as well as the detailed design of the mold. The maximum monthly production quantities given depend on the available pro­duction time, the number of cavities in the mold, and the estimated cycle time. This production capability can then be compared to the target production quantity to determine the number of molds to be made, or otherwise to guide the mold design with respect to number of cavities and cycle time requirements. It should be noted that the cycle time and other mold design data in Fig. 2.2 may not be available at the start of the mold design process. In fact, these data are ­intermediate results from the mold design process. However, some customers will provide these details as specifications that the mold designer must satisfy. If these items are not specified by the customer, then the mold designer should perform a preliminary analysis and design with cost analyses to provide the customer with a reasonably efficient mold design proposal. Part cost estimation is crucial in many plastic part design applications. Product designers, mold designers, and molders are all conscious of trade-offs between the cost of the mold, material, and processing. Their documentation near the top center of Fig. 2.2 will help the mold designer to understand the intent of the part designer with respect to cost, and guide the mold design to an appropriate solution. P ­ otential variances between the target and actual costs should be discussed at design reviews as well as the end of the project for continuous improvements.

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2 Plastic Part Design

2.2.3 End-Use Requirements A typical molded part may have literally dozens, if not hundreds, of specifications. A few common end-use part requirements are provided at the left of the center of Fig. 2.2. Some of these requirements drive the geometry, material selection, and other design details about which the mold design engineer may seem to have little control. Even so, the mold designer should be aware of these requirements, as they can influence selection of the mold materials, surface finish and treatments, mold design, and performance evaluation of the finished mold. Manufacturers are generally required to use design standards to ensure that the product being designed and manufactured will perform as intended after com­ mercial sale. In many segments of the plastics industry, such as medical devices, regulatory agencies have developed extensive standards governing the design, manufacturing, and testing of plastic products. A detailed discussion of regulatory compliance is beyond the scope of this book. However, the mold designer should be aware of any regulatory compliance issues that may affect the mold engineering. A few common regulatory agencies and their compliance programs are provided at the center of Fig. 2.2. These include the American National Standards Institute (ANSI, http://www.ansi.org/), ASTM International (http://www.astm.org/), the U.S. Food and Drug Administration (FDA, http://www.fda.gov/), the International Electrotechnical Commission (IEC, www.iec.ch/), the U.S. Department of Defense Index of Specifications and Standards (MIL-SPEC, http://stinet.dtic. mil/), the International Standards Organization (ISO, http://www.iso. org/), the Underwriters Laboratories (UL, http://www.ul.com/), and many others. The mold design engineer does not usually need to know every detail of these ­specifications since they generally pertain to the use of the molded product and not specifically to the injection mold. However, the mold designer should inquire about any governing regulations that may affect the mold design. Ideally, the customer should provide a copy of any such regulations and highlight the specific requirements related to the molded product design. The specification of dimensions and tolerances is of critical importance to both the mold designer and injection molder. Dimensions in product designs are typically specified with absolute tolerances. Figure 2.3 provides an example of four different methods for specifying tolerances. The most common method is the general tolerance, typically specified in the signature block, which is applied to any dimension without an explicitly specified tolerance. In this case, the width offset of 25 mm does not have a specified tolerance and so would be governed by the general tolerance of ± 0.2 mm. The height offset of 20.0 mm has a specified bilateral tolerance of ± 0.1 mm, which is actually redundant with the general specified tolerance of ± 0.1 mm for dimensions specified with one decimal place. If the height offset was specified to a different tolerance, this explicit specification would override the ­general tolerance.

2.2 Design Requirements

Figure 2.3 Tolerance specification methods

The hole with a diameter of 18 mm is specified with limit dimensions, such that the hole diameter must be between 17.90 mm and 18.00 mm. The two decimals of precision do not reflect any additional precision on the tolerance but rather the absolute range of acceptable dimensions for the hole diameter. Finally, the shaft diameter is specified with a unilateral tolerance, meaning that the diameter of the shaft must be between 18 and 18.10 mm. The product and mold designer would both understand from these tolerance specifications that the shaft is meant to nominally fit into the hole with an interference fit such that there is no diametral clearance between the two. While product designers will usually consider tolerances in absolute terms as described with respect to Fig. 2.3, plastics molders will tend to consider tolerances in relative terms. The reason is that molded plastics will shrink as a percentage of their length, so tight absolute tolerances will become more difficult to achieve as the part’s length dimensions increase. For instance, a typical tolerance may be considered as ± 0.4 % of the nominal dimension, such that a 100 mm length would be specified as 100 ± 0.4 mm. A tight tolerance may be considered as ± 0.1 %, such that a 10 mm diameter may be specified as 10 ± 0.01 mm. The achievement of very tight tolerances requires careful mold design, process engineering, and consistent material properties. For this reason, product designers are encouraged to specify a single general tolerance governing most dimensions along with only a few tighter tolerances on specific dimensions that are critical to product. Just because a tolerance is specified does not mean that it is achievable. In fact, it is not uncommon for product designers to over-specify the tolerances on many dimensions [10]. Mold designers should discuss tight tolerance specifications with the product development team, and communicate that such specifications may require prototype molding to characterize the shrinkage behavior, nonuniform profiling of shrinkage rates in different areas of the mold, and mold modifications during mold commissioning. Product designers will often provide specifications on the aesthetics, including requirements on color, color matching across multiple components, and gloss levels. It is common for the product design to specify the mold surface finish and mold

29

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2 Plastic Part Design

surface texture, which may add significant cost to the injection mold. Also, the mold design engineer should be made aware of critical aesthetic surfaces in which aesthetic defects (such as from knit-lines, gate blemish, sink, witness marks, etc.) should be avoided.

2.2.4 Design for Manufacturing and Assembly Ideally, the mold design engineer should be involved with the product design from the early stage of concept design. Such involvement often provides for significantly improved plastic part designs that are more functional and efficient. Unfortunately, mold designers are often provided “finished” product designs that are really substandard with respect to design for injection molding and assembly. Rather than assume that the product design is finished and unchangeable, the mold designer should check that the part has been specifically designed for injection molding. Some of the most common guidelines with respect to design for manufacturing are provided at the bottom left of Fig. 2.2 [11, 12]; these guidelines can significantly improve the function and reduce the cost of the molded product, and so will be discussed in Section 2.3. With regard to design for assembly guidelines [13, 14], the mold designer should inspect the part design(s) that have been provided and check that the design for assembly is reasonable. There are two goals for this task. First, it may be possible to improve the overall design of the product by consolidating multiple components, facilitating top-down assembly, etc. Second, the mold designer can reduce the number of late design changes that can cost money and time by verifying that such design considerations have been performed. Some mold designers may be aware of product design issues, yet choose to directly implement molds for the provided part designs. For instance, a mold designer may understand that the cost of the mold could be reduced by slightly changing an ­angle on a surface to eliminate an undercut but remain silent to justify the need for a core pull and a higher priced mold. While such a strategy may result in additional work and profit for the mold designer in the short term, it is a losing long term strategy. Rather, the most successful mold designers seek to add value to their customers by providing services that improve the quality and reduce the cost of their customers’ products.

2.2.5 Plastic Material Properties The plastic material is usually specified not by the molder or mold designer, but rather by the customer, original equipment manufacturer (OEM), or original design manufacturer (ODM). However, the mold designer needs to know some specific

2.3 Design for Injection Molding

properties of the plastic, identified at the bottom center of Fig. 2.2, which will ­govern the function of the mold. Sample data for some generic grades of plastic are provided in Appendix A. Other sources from which to gather this information include resin suppliers and database suppliers (such as http://www.ulprospector. com/, http://www.matweb.com, and http://www.campusplastics.com/). The materials used in mold construction are usually specified by the mold designer. The molder will often have substantial input to the mold materials given their prior experience with molding the target plastic resin. Mold materials selection will be discussed in Chapter 4; sample data for some commonly used metals are provided in Appendix B. The mold designer should document the mold material properties listed at the bottom right of Fig. 2.2 with input from the mold material supplier, since these will later be used in analysis of the mold design.

„„2.3 Design for Injection Molding A detailed review of the plastic part design should be conducted prior to the design and manufacture of the injection mold. The design review should consider the fundamentals of plastic part design, as well as other concerns related specifically to mold design. Some of the most basic part design considerations are next discussed.

2.3.1 Uniform Wall Thickness Parts of varying wall thickness should be avoided due to reasons related to both cost and quality. The fundamental issue is that thick and thin wall sections will cool at different rates: thicker sections will take longer to cool than thinner sections. When ejected, parts with varying wall thickness will exhibit higher temperatures near the thick sections and lower temperatures near the thin sections. These temperature differences and the associated differential shrinkage can result in significant geometric distortion of the part given the high coefficient of thermal expansion for plastics. Extreme differences in wall thicknesses should generally be avoided if at all possible since internal voids may be formed internal to the part due to excessive shrinkage in the thick sections even with extended packing and cooling times. (There are exceptions, of course, such as insert molding and gas/ water assist molding that have purposefully-designed thick wall sections, some of which are later discussed in Sections 13.3 and 13.4.) Figure 2.4 provides several part designs with different thicknesses. The worst part design, shown at top left, has the melt gated into a thin section and then flowing to a thick section with a sharp transition in the thickness. This design may lead to

31

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2 Plastic Part Design

moldings with poor surface finish due to nonuniform flow of the melt as well as poor surface replication and dimensional control in the thick section related to premature solidification of the plastic molded in the thin section. The quality of the molded product would be greatly improved as shown at the top center. Just by ­gating into the thicker section, the molded product would have much better aesthetics and dimensional stability since the thicker section would allow the packing of the thinner section prior to its own solidification. The design would be further improved by gradually transitioning the thick section to the thin section. Even so, any product design with significant variations in wall thickness will exhibit extended cooling times and different shrinkage rates in the thick and thin sections.

Figure 2.4 Wall thickness design

A standard approach is to increase the nominal thickness of the molded part so as to eliminate the need for thick sections in local areas as shown at the center left in Fig. 2.4. The decision to increase the wall thickness will eliminate many issues ­related to part quality, but can lead to excessive material consumption and extended cooling times. For these reasons, the best design may be to use a thinner wall thickness together with vertical ribs in those areas requiring greater stiffness and strength as shown at the center right in Fig. 2.4. The height and/or density of the ribs may be altered to change the relative stiffness throughout the part [15]. The injection molding process is unique compared to other molding process in its ability to economically provide very complex structures. The bottom two part designs in Fig. 2.4 show alternative strategies that are increasingly common. At the bottom left is a thinner wall section with a matrix of thin, short ribs. At the bottom right is the same thicker wall section that has been dimpled on both sides to r­ educe

2.3 Design for Injection Molding

the effective wall thickness. Both strategies are useful reducing the wall thickness while still increasing the amount of material away from the part’s neutral axis in bending, thereby contributing to a significant increase in stiffness without an ­increase in the material consumption. Furthermore, both strategies provide a ­significant increase in surface area, which will result in improved mold cooling and molding productivity.

2.3.2 Rib Design Consider the ribbed part design shown at the center of Fig. 2.5 relative to a part having a greater uniform thickness shown at left. In this ribbed part design, the base thickness of the rib is 70% of the wall thickness of the part, H, and the height of the rib is four times the wall thickness of the part. The two ribs are spaced at ten times the wall thickness of the part. Analysis of this design indicates that this ­design has a stiffness equivalent to the part that is 30 % thicker but does not have ribs. However, the 30 % thicker part will consume approximately 15 % more material and have a 70 % longer cycle time than the thinner part with ribs. As such, the addition of ribs can provide significant performance and economic advantages.

Figure 2.5 Effective rib design

Product designers will tend to maximize the stiffness of the ribs by making the thickness of the rib equal to the nominal wall thickness of the part, H, while also minimizing the draft angle. Ribs thicker than 70 % of the wall thickness will tend to draw material away from the center of the opposite wall when the rib cools. The volumetric shrinkage in this region can cause internal voids or sink to appear on the side of the part opposite the rib. In nonaesthetic applications that use highly filled materials with lower shrinkage, the rib thickness can be increased. Otherwise, a rib thickness less than 70 % of the nominal thickness should be used in molding applications with unfilled materials [16]. Similarly, draft of 1° per side is often used to facilitate ejection of the molded ribs. This amount of draft may be insufficient for deep ribs or for parts molded with heavily filled resins. Conversely, parts with very short ribs (such as shown at the bottom left of Fig. 2.5) can often be made with zero draft.

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2 Plastic Part Design

2.3.3 Boss Design Bosses are typically used to secure multiple components together with the use of self-threading screws. Some different boss designs are provided in Fig. 2.6. The left-most design provides a boss near a corner with two ribs and a gusset placed at 120°. The center design shows a boss on a rib with two gussets at 90°. The rightmost design shows a free-standing boss with gusseted ribs that provide for an ­elevated assembly surface. All these boss designs utilize a boss, rib, and gusset thickness of 70 % times the nominal wall thickness. Similar to the guidelines for rib design, the wall thickness and draft angle for gussets can be modified in view of their height and the material being molded. Designed bosses must be able to withstand the torque applied during insertion of the self-threading screws as well as the potential tensile pull-out forces applied during end-use. At the same time, however, bosses should not be designed with overly thick sections that may require extended cycle times or cause aesthetic problems. In the designs of Fig. 2.6, no draft was utilized on the bosses and g­ ussets. These design features are vital to the structural integrity of the part, yet are small relative to the size of the entire part. As such, using less draft on these features can aid in increasing the stiffness and strength of the molding without significantly increasing the ejection forces. Still, ejection of bosses can be an issue in injection molding, so draft and ejector sleeves can be used to assist in ejection of tall bosses that will require large ejection forces.

Figure 2.6 Effective Boss Design

2.3.4 Corner Design Sharp corners are often specified in product design to maximize the interior ­volume of a component, to facilitate mating between components, or to improve the aesthetics. However, sharp corners in molded products should be avoided for many reasons related to product performance, mold design, and injection molding:

2.3 Design for Injection Molding

ƒƒRelative to product performance, sharp corners will result in a stress concentration that may cause many (and especially brittle) materials to fail under load. Furthermore, a box with sharp corners and tall sides may not have the torsional stiffness of a rounded box with shorter sides. ƒƒRelative to mold making, sharp corners can be very difficult to produce, requiring the use of electrical discharge machining or the use of multiple cutting passes with tools of decreasing size. ƒƒSome common guidelines for filleting and chamfering corners are provided in Fig. 2.7. As shown, the fillet radius on an external corner should be 150 % of the wall thickness. To maintain the same thickness around the corner, the fillet on the internal corner is set to 50 % of the wall thickness. In most modern solids-based CAD systems, these fillets can be readily achieved by filleting the ­outside edges prior to shelling of the part. These fillet recommendations are only guidelines. In fact, even larger fillets can be used to encourage more uniform mold cooling. In all cases, the mold designer should suggest a fillet radius that corresponds to readily available tooling geometry so that custom tools need not be custom-made.

Figure 2.7 Comparison of fillet designs

Chamfers are often used to break sharp corners with a single beveled surface ­connecting the outer surfaces, often at a 45-degree angle. As shown in Fig. 2.8, a shallow chamfer of less than one-half the wall thickness is often utilized on external corners to provide for adequate relief while avoiding potential negative issues related to melt flow and part strength. Similar to fillets, larger chamfers (such as the one shown at right in Fig. 2.8) may be applied prior to shelling to provide improved part stiffness and heat transfer near the corners.

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Figure 2.8 Comparison of chamfer designs

2.3.5 Surface Finish and Textures Surface finish and texture are commonly specified by the part designer, yet have a significant impact on the mold design and cost. Most mold-making companies are capable of providing high-quality surface finishes, though polishing can be outsourced to lower-cost companies and countries due to its high labor content. Surface texturing requires a higher level of skill and technology, with a relatively small subset of companies providing a significant portion of mold texturing surfaces. Table 2.1 SPI Surface Finishes and Roughness SPI Finish

Finishing method

Microfinish (µm)

Surface roughness (µm)

A1

#3 diamond polish

~1

~0.01

A3

#15 diamond polish

~2

~0.04

B3

#320 grit cloth

~6

~0.12

C3

#320 stone

~12

~0.3

D2

#240 oxide blast

~30

~0.8

D3

#24 oxide blast

~160

~4

Surface finishes are commonly evaluated according to standards of the Society of the Plastics Industry (http://www.plasticsindustry.org). These finishes range from the D3, which has a sand-blasted appearance, to A1, which has a mirror finish. Table 2.1 provides some common SPI finishes, the finishing method, and the measurable surface roughness. The cost of molded parts can increase dramatically with higher levels of surface finish. The reason is that the application of a given surface finishing requires the mold maker to successively apply all the lower-level surface finishing methods. For example, to obtain an SPI C3 finish, the mold would first be treated with coarse and fine bead blasts followed by polishing with a #320 stone. For this reason, higher

2.3 Design for Injection Molding

levels of surface finish cost significantly more than lower levels. Furthermore, molds with high levels of finish can produce moldings in which defects are highly visible, thus adding cost to the injection molding process and mold maintenance requirements. As an alternative to smooth surface finishes, many product designs specify a ­textured finish. One common reason is that textures may be used to impart the appearance of wood, leather, or other materials as shown in Table 2.2. As a result, textures may increase the perceived value of the plastic molding by the end-user [17]. Another reason is that textured surfaces provide an uneven depth which may be used to hide defects such as knit-lines, blemishes, or other flaws. In addition, textures may be used to improve the function of the product, for instance, by providing a surface that is easy to grip or hiding scratches during end-use. Texturing does add significantly to the cost of the mold. To apply a texture, mold surfaces must first be finished typically to SPI class B for shallow textures (in which the texture depth is on the order of a few microns) or class C for rough ­textures. Otherwise, the underlying poor surface finish may be visible after the applied texture. After surface finishing, the texture is imbued to the mold surfaces using chemical etching or laser machining processes. Since dedicated processing equipment is required, the mold development process must provide adequate time 2.3 Design for Injection Molding 33 and money for the mold texturing. 2.3 33 2.3 Design Design for for Injection Injection Molding Molding 33 Table 2.13: Texture examples Table 2.2 Texture Examples Table Table 2.13: 2.13: Texture Texture examples examples

Texture Texture Texture Texture

Sand Sand Sand Sand

Leather Leather Leather Leather

Netting Netting Netting Netting

Wood grain Wood grain Wood Wood grain grain

Image Image Image Image

Texture depth Texture depth Texture Texture depth

SPI finish required required depth SPI finish required SPI finish finishSPI required

50 µm 50 µm 50 50 µm µm

B B B B

125 µm 125 µm 125 125 µm µm

C C C C

150 µm 150 µm 150 150 µm µm

C C C C

250 µm 250 µm 250 250 µm µm

D D D D

2.3.6 Draft 2.3.6 2.3.6 Draft Draft Draft refers to the angle of incline placed between the vertical surfaces of the plastic moldings

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2.3.6 Draft Draft refers to the angle of incline placed between the vertical surfaces of the plastic moldings and the mold opening direction. Draft is normally applied to facilitate ejection of the moldings from the mold. Product designers frequently avoid the application of significant draft, since it alters the aesthetic form of the design and reduces the molding’s internal volume. Even so, draft is commonly applied to plastic moldings to avoid ejection issues and extremely complex mold designs. Draft angles on ribs must be carefully specified. In the previous rib design shown in Fig. 2.5, for instance, a 2° draft angle was applied to facilitate the ejection of the molded part from the mold. In terms of product functionality, a lesser draft angle may be desired since this allows for taller and thicker ribs with greater stiffness. Unfortunately, lower draft angles (such as ½ or 1°) may cause the part to ­excessively stick in the mold. This issue of sticking upon part ejection can be compounded when molding with mica- and/or glass-filled materials that have low shrinkage and high surface roughness. As such, the allowable draft angle is a complex function of the material behavior, processing conditions, and surface finish. Table 2.3 Draft Examples Surface finish

Resin

Roughness (µm)

Draft

Class A1

Acrylic

  0.01

0.5°

Class B3

ABS

 12

1.5°

Sand texture

20 % GF PC

 12

2°

Leather texture

Soft PVC

125

4°

Leather texture

ABS

125

7.5°

A minimum draft angle of 0.5° is normally used, with 1 to 2° commonly applied according to material supplier recommendations. Rough and textured surfaces ­typically require additional draft, with an additional 1° of draft commonly applied per 20 μm of surface roughness or texture depth. Table 2.3 provides some recommended draft angles for a few different surface finishes and materials; the recommended draft angle increases with the surface roughness. With respect to the ­material properties, the draft angle should increase for glass-filled and/or low shrinkage materials but may be decreased for highly flexible materials such as soft PVC.

2.3 Design for Injection Molding

2.3.7 Undercuts An undercut is a feature in the product design that interferes with the ejection of the molding from the mold. Four typical design features that require undercuts are shown in Fig. 2.9. These design features include, for example, a window in a side wall, an overhang above the bottom wall of the part, a horizontal boss, and a snap beam or “finger.”

Figure 2.9 Some common features with undercuts

To provide some insight into why these features are commonly avoided, consider three different but common mold design strategies for molding a snap beam as shown in Fig. 2.10 with the mold closed prior to ejection and Fig.2.11 with the moving side retracted and the ejectors extended forward. Because the snap beam is narrower at its neck than at its tip, there is an undercut in the mold that the mold designer must be aware of and make the design such that the part can be ejected after the mold opens. Please note that the provided designs are not intended to suggest the use of all three strategies in a single mold, but only provide a basis for discussion. The design at the left is the simplest of the three, in which an opening or window at the base of the snap beam allows a protrusion from the stationary side to core out the area cavity beneath the undercut. This is a reliable technique, but leaves a hole in the part that alters its function and aesthetics. At design at the right is also very common, which uses an ejector pin with a profiled or contoured surface on its side adjacent the snap beam. This contoured profile on the ejector pin provides a miniature cavity to allow the molding of the head of the pin. When the mold is opened and the ejectors are extended, the pin and part will move together until the part fully clears the mold cavity and can clear the height of the undercut. When using contoured ejectors, a dowel pin or some other design feature must be used to maintain proper orientation of the contoured surface with respect to the mold cavity. Otherwise, the cavity surfaces would not align and defective parts would occur.

39

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2 Plastic Part Design

The design at the center is also a very common design in which a sliding, angled pin or “lifter” is used to core out the volume of the mold trapped beneath the ­undercut. After the mold is opened, the forward movement of the lifter acting on the inclined surface of the mold causes the lifter to move laterally, thereby clearing the undercut upon ejection. There are three issues associated with lifters that the plastic part and mold designer should consider when adopting their use. First, there is the added design time and complexity used to implement the design. Second, there is the potential for wear, sticking, and increased maintenance associated with the sliding surfaces on the slider itself, on the mold surfaces, and within the ejector assembly. Third, the use of the lifter requires adequate clearance between features within the molded part. As indicated in Fig. 2.11, this particular design likely does not provide sufficient clearance, such that the lifter will interfere with the rightmost snap beam if the lifter is further extended.

Figure 2.10 Three mold designs for producing a snap beam, with mold closed

Figure 2.11 Three mold designs for producing a snap beam, with mold opened

2.4 Chapter Review

It may seem that such complexities in mold design would dictate the avoidance of snap beams, but in practice, such designs are not usually problematic for expe­ rienced mold designers. More challenging than such undercutting features is the horizontal boss of Fig. 2.9 that is oriented internal to the part and designed with a molded internal thread. Together, these design features would require a com­ plicated mold design with timed actuation of various cores prior to ejection of the product. Alternatively, a “lost core” could be used in which the mold core that forms the internal features of the part is melted away after the part is molded as described in Section 13.4.3. When possible, these types of product design features should be avoided since complex mold mechanisms must be designed and machined for forming and ejecting the molded part. These additional mold components can make the mold more difficult to use and even damage the mold if used improperly. For these reasons, the mold design engineer should identify problematic features, alert the customer, and work with the product design engineer to remove the undercuts. However, such undercuts should not be designed out of the product if the function provided by the feature(s) with the undercut is vital to the product or the removal of the undercut would necessitate additional post-molding operations or the redesign of a single part into multiple pieces [18].

„„2.4 Chapter Review After reading this chapter, you should understand: ƒƒThe basic stages of the molded part development process, including the role of management reviews, ƒƒWhat information is needed to begin the mold design process, ƒƒThe common specifications applied to a molded product, ƒƒWhere to find additional information relevant to product and mold design, and ƒƒBasic part design guidelines for injection molding In the next chapter, the mold cost and the part cost will be analyzed with respect to critical mold design decisions. The results of this analysis will be used to design the layout of the mold. In later chapters, the design and analysis of various mold subsystems are conducted.

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2 Plastic Part Design

„„2.5 References [1] Krishnan, V., S. D. Eppinger, and D. E. Whitney, A model-based framework to overlap product ­development activities, Manage. Sci. (1997) 43(4): pp. 437–451 [2] Cooper, R. G., Stage-gate systems: a new tool for managing new products, Business Horizons (1990) 33(3): pp. 44–54 [3] Wilson, E., Maximizing designers’ impact on market success through product definition, Des. ­Manage. J. (Former Series) (1993) 4(4): pp. 62–68 [4] Tabrizi, B. and R. Walleigh, Defining next-generation products: an inside look, Harvard Business Review (1996) 75(6): pp. 116–124 [5] Herrmann, J. W., et al., New directions in design for manufacturing, in ASME 2004 International Design Engineering Technical Conferences and Computers and Information in Engineering ­Conference, American Society of Mechanical Engineers (2004) [6] Boothroyd, G., Product design for manufacture and assembly, Comput. Aided Des. (1994) 26(7): pp. 505–520 [7] Rosato, D. V., D. V. Rosato, and M. G. Rosato, Injection molding handbook, Springer Science & Business Media (2000) [8] Haller, M., A. Peikert, and J. Thoma, Cycle time management during production ramp-up, Robotics and Computer-Integrated Manufacturing (2003) 19(1): pp. 183–188 [9] Fagade, A. and D. O. Kazmer, Early cost estimation for injection molded parts, J. Injection Molding Technol. (2000) 4(3): pp. 97–106 [10] Suri, R., D. Frey, and K. Otto, Key Inspection Characteristics, J. Mech. Des. (2001) 123: p. 479 [11] Malloy, R. A., Plastic part design for injection molding, Hanser, Munich (1994) [12] Chen, Y. M. and J. J. Liu, Cost-effective design for injection molding, Robotics and Computer-­Integrated Manufacturing (1999) 15(1): pp. 1–21 [13] Boothroyd, G., P. Dewhurst, and W. Knight, Product design for manufacture and assembly, Marcel Dekker, New York (1994) [14] Boothroyd, G., Making It Simple: Design for Assembly, Mech. Eng. (1988) pp. 28–31 [15] Kazmer, D. O., Wall Thickness Optimization In Molded Product Design, In SPE ANTEC (2003) [16] Shi, L. and M. Gupta, Approximate prediction of sink mark depth in rib-reinforced plastic parts by empirical equations. J. Injection Molding Technol. (1999) 3(1): pp. 1–10 [17] Kwahk, J. and S. H. Han, A methodology for evaluating the usability of audiovisual consumer ­electronic products. Appl. Ergonomic (2002) 33(5): pp. 419–431 [18] Roser, C., D. Kazmer, and J. Rinderle, An economic design change method, J. Mech. Des. (2003) 125(2): pp. 233–239

3

Mold Cost Estimation

„„3.1 The Mold Quoting Process The quoting process for plastic parts can be difficult for both the mold customer and supplier. Consider the view of the mold customer. The procurement specialist for the product development team sends out requests for quotes (RFQs) to several mold makers. After waiting days or weeks, the quotes come back and the customer discovers that the development time and cost of the mold may vary by a factor of three or more. In such a case, prospective mold purchasers should ask about the details of the provided quotes and check if the costs can be reduced through product redesign. To reduce uncertainty related to pricing and capability, many prospective customers maintain a list of qualified suppliers who have been found to provide satisfactory lead times, quality, and pricing across multiple projects. Long-term trust-based partnerships can provide for rapid application and mold development by avoiding the quoting process altogether and invoicing on a labor cost plus materials cost (referred to as “cost plus”) basis. Now consider the view of the mold supplier. The mold designer may need to invest significant time developing a quote that may have a relatively small chance of being accepted. Sometimes, the mold designer may have to redesign the product and perform extensive analysis to provide the quote. While the quote may seem high to the prospective customer, the design may correspond to a mold of higher-quality materials and workmanship that can provide a higher production rate and longer working life than some other, lower-cost mold. This more expensive mold may quickly recoup its added costs during production. From time to time, mold makers and molders will adjust their quote based on whether or not they want the business. If the supplier is extremely busy or idle, then the estimated number of hours and/or hourly rate may be adjusted to either discourage or encourage the potential customer from accepting the quote. Such adjustments should be avoided since the provided quote does not represent the true costs of the supplier, which would become the basis for future engagements between the mold supplier and the customer. Thus, the development of a long-term and mutually beneficial partnership will begin with justifiable project quotes.

44

3 Mold Cost Estimation

The provided mold purchase contract typically states payment and delivery terms for the mold(s) and perhaps even the molded part(s). A typical mold purchase agreement may specify that the cost of a mold is paid in three installments: ƒƒthe first third: on acceptance of the quote (after which the mold base and key materials are typically purchased); ƒƒthe second third: halfway through the mold making project (often when cavity inserts have been machined); and ƒƒthe final third: upon acceptance of the quality of the molded parts.

Figure 3.1 Schedule of mold and molding expenses

After the mold is purchased, molds are typically shipped to the specified molder or the customer’s facility where the parts are molded and marginal costs are incurred on a per-part basis. The cash outlays for a typical project are plotted in Fig. 3.1 on a monthly basis. The material and processing costs in month 3 are related to molding trials by which the mold design is validated and improved; a batch of pre-­ production parts are sampled at this time for marketing and testing purposes. Later, monthly processing and material costs are incurred during production. Maintenance costs may appear intermittently throughout production to maintain the quality of the mold and moldings. There has been a trend in the industry towards large vertically integrated molders with tightly integrated supply chains that can supply molded parts and even ­complete product assemblies. As such, the structure of the quote can vary sub­ stantially with the structure of the project and business requirements. With a vertically integrated supplier, there is typically an upfront fee for the costs associated with the development of the mold, followed by a fee for each molded part. To protect the supplier, contracts are typically developed that specify minimum production quantities with discounts and/or fees related to changes in the production schedule.

3.2 Cost Overview for Molded Parts

Some prospective mold customers may purposefully choose to disintegrate their supply chains in order to minimize the “leakage” of intellectual property. In this model, they may have one firm perform analysis or simulation of one component in the design, a second firm develop a mold for the same component, a third firm develop other designs and molds for other components in the design, yet other firms for molding different components, and then perform the assembly internally. Such a disintegrated supply chain can raise significant issues with respect to scheduling and product qualification. Since the structure and magnitude of quotes will vary substantially with the ­supply chain strategy and supplier(s), a prospective buyer of plastic parts should solicit quotes from multiple vendors and select the quote from the supplier that provides the most preferable combination of design capability, molded part quality, and payment/delivery terms.

„„3.2 Cost Overview for Molded Parts There are three main cost drivers for molded products: 1. the cost of the mold and its maintenance, 2. the materials cost, and 3. the processing cost. Figure 3.2 provides a breakdown of these primary cost drivers and their under­ lying components. It is important to note that these costs do not include indirect costs such as facilities, administrative overhead, fringe benefits, or profits. However, such indirect costs may be accounted for through the adjustment of hourly rates or application of indirect cost rates.

Figure 3.2 Cost drivers for injection molded products

45

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3 Mold Cost Estimation

Even though most molded products have the same cost drivers, the proportion of costs varies widely by application. Figure 3.3 shows the cost breakdown for a ­commodity application (such as a cable tie with a production volume of 10 million pieces) and a specialty application (such as a custom electrical connector with a production volume of 100,000 pieces). While these two products are approximately the same weight, it is observed that the magnitude and proportion of costs are vastly different. The commodity part will tend to have lower costs due to economies of scale that allow (1) amortization of the mold cost across vast production quantities, (2) optimization of the molding process for lower molding costs, and (3) lower material costs associated with bulk purchases of resin. As Fig. 3.3 suggests, the material costs represent the majority of the total molded part cost in commodity applications whereas the mold/tooling costs can dominate for custom moldings with low production quantities.

Figure 3.3 Cost comparisons for a commodity and specialty part

For analysis, the total part cost of a molded product, Cpart , can be estimated as Cpart =

Cmold/part +Cmaterial/part +Cprocess/part yield

(3.1)

where Cmold/part is the amortized cost of the mold and maintenance per part, Cmaterial/part is the material cost per part, Cprocess/part is the processing cost per part, and yield is the fraction of molded parts that are acceptable. Each of these terms will be subsequently estimated. To demonstrate the cost estimation method, each of these cost drivers is analyzed for the laptop bezel shown in Fig. 3.4. The example analysis assumes that 1,000,000 parts are to be molded of ABS from a single-cavity hot runner mold. Some relevant application data required to perform the cost ­estimation is provided in Table 3.1.

3.2 Cost Overview for Molded Parts

Figure 3.4 Isometric view of laptop bezel Table 3.1 Laptop Design Data Parameter

Laptop bezel

Material

ABS

Production quantity

1,000,000

Lpart

240 mm

Wpart

160 mm

Hpart

10 mm

Apart_surface

45,700 mm2

Vpart

27,500 mm3

Hwall

1.5 mm

3.2.1 Mold Cost per Part The cost of the mold for a given application is estimated in Section 3.3. Given the estimate or a quote for the mold cost, Ctotal_mold , the cost of the mold per part can be assessed as Cmold/part =

Ctotal_mold ntotal

´ fmaintenance (3.2)

where ntotal is the total production quantity of parts to be molded, and fmaintenance is a factor associated with maintaining the mold. Most molders perform several ­levels of maintenance, including: ƒƒpreventive maintenance after every molding run, ƒƒinspections and minor repairs on an intermittent basis, ƒƒscheduled general mold maintenance on a quarterly or semiannual basis, and ƒƒmold rebuilding as necessary.

47

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3 Mold Cost Estimation

The need for mold maintenance and repair is related to the number of molding cycles performed, the properties of the plastic and mold materials, the processing conditions, and the quality of the mold. As a general rule, annual maintenance costs can be estimated as 10 % of the mold purchase cost [1], but will vary with the design, materials, and processing conditions in application. As the resin becomes more abrasive relative to the hardness of the mold, the wear of the mold accelerates and more maintenance is required. Conversely, a well-designed, hardened mold should exhibit lower maintenance costs when used with an unfilled low-viscosity plastic. Table 3.2 provides some maintenance estimates. Table 3.2 Mold Maintenance Coefficient, fmaintenance, per Million Cycles Unfilled, low ­viscosity plastic

High viscosity or ­ articulate filled plastic p

High viscosity and fiber filled plastic

Soft mold material, such as aluminum or mild steel

4

16

64

Standard mold steel, such as P20

2

 4

16

Hardened surface or tool steel, such as H13

1

 2

 4

Example: Estimate the amortized cost of the mold base per molded laptop bezel. ABS is a moderate viscosity, unfilled material. If the mold inserts are made from D2 tool steel with a hardened surface, then a mold maintenance coeffi­cient of 2 is estimated. Given that the mold has a single cavity, one million ­cycles are ­required. The amortized cost of the mold per molded laptop bezel (including the initial ­purchase cost and maintenance costs) is then estimated as:

Cmold/part =

$75,900 × 2 = $0.152/part 1,000,000parts 

3.2.2 Material Cost per Part The cost of the material per part can be estimated as: Cmaterial/part = Vpart × rpolymer × kpolymer × fscrap (3.3)

where Vpart is the volume of the molded part, ρpolymer is the density of the molded polymer at room temperature, kpolymer is the cost of the molded polymer per unit weight, and fscrap is the total proportion of material consumed including startup, defects, and scrap associated with the feed system.

3.2 Cost Overview for Molded Parts

Table 3.3 provides estimates of the total material consumption for various types of feed systems. A cold runner is simple and low-cost but results in molded plastic that must be either discarded or recycled. Utilizing the recycled plastic as regrind reduces the waste but incurs some cost related to the labor and energy of ­recycling. As later described, hot runners have the potential to significantly reduce material costs but consume significant material during start-up and so are less effective in short runs. Table 3.3 Material Waste Coefficient Type of feed design

Feed system waste factor, ffeed_waste

Cold runner

1.25

Cold runner, fully utilizing regrind

1.08

Hot runner with short runs

1.05

Hot runner with long runs

1.02

Example: Estimate the cost of the plastic material per molded laptop bezel. Since a hot runner system is used and the production quantity is one million parts, large production runs are assumed with a feed waste factor of 1.02. Using the cost and density from Appendix A, the cost of the plastic material per molded part is estimated as

æ 0.01 m ö÷3 kg $ Cmaterial/part = 27.5 cm3 × çç ÷ × 1044 3 × 2.80 × 1.02 = $0.082/part çè cm ÷ø kg m The cost of the plastic material per part is quite low since the part has a very low thickness (1.5 mm) and low part weight (28.7 g). 

3.2.3 Processing Cost per Part The processing cost per part is a function of the number of mold cavities, the cycle time, tcycle , and the hourly rate of the machinery and labor, Rmolding : Cprocess/part =

tcycle ncavities

´

Rmolding 3600 s h

(3.4)

The cycle time is effected primarily by the thickness of the part, hwall , and, to a lesser extent, by the size of the part and the type of feed system. While the cycle time will be more accurately estimated during the cooling system design, a reasonable estimate is provided by

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3 Mold Cost Estimation

é s ù 2 ú (h é mm ûù ) ´ fcycle_efficiency (3.5) tcycle = 4 ê êë mm2 úû wall ë

where the cycle efficiency, fcycle_efficiency , is a function of the type of feed system and process that is being operated according to Table 3.4. While it is desirable to operate a fully automatic molding cell with a hot runner, many molders continue to use cold runner molds operating in semiautomatic mode. Table 3.4 Cycle efficiency coefficient Type of feed system and mold operation

Cycle efficiency factor, fcycle_efficiency, cold runner

Cycle efficiency factor, fcycle_efficiency, hot runner

Semiautomatic molding with ­operator removal of molded parts

2.25

2.0

Semiautomatic molding with ­gravity drop or high speed robotic take-out

1.5

1.25

Fully automatic molding

1.25

1.0

The hourly rate for the molding machine is primarily a function of the clamp ­tonnage, which drives the size and cost of the machine. The following model was developed relating the clamp tonnage and capability to the machine hourly rate: Rmolding = (43.3 + 0.095 × Fclamp )× fmachine (3.6)

where Fclamp is the clamp tonnage in metric tons (mTon), and fmachine is a factor relating to the capability of the machine and the associated labor. This equation was derived using published U.S. national hourly rate data [2] for twelve different sized molding machines ranging from 20 to 3500 metric tons; the described model has a coefficient of determination, R2, equal to 0.979. The hourly rate data is also a function of the geographic region, machine and molder costs, and other factors. To account for these variances, the machine capability factor, fmachine , is estimated according to Table 3.5. In general, molding machines with advanced capabilities and higher clamp tonnage cost more to purchase and operate, and so command a price premium. Machines with specialized capa­ bility (such as multiple injection units or very high injection pressures/velocities) are more expensive to purchase and so likewise command a price premium per hour of operation. The cost of all auxiliaries should be added to the appropriate machine coefficient. While advanced technology can increase the hourly rate of the molding process, it should provide a net savings by improving quality and reducing the processing and materials costs. Variances due to geographic locale may be accounted by scaling the machine factor by the labor rate data provided in Appendix D relative to the U.S. cost data.

3.2 Cost Overview for Molded Parts

Table 3.5 Molding Machine Capability Type of molding machine and labor required

Machine factor, fmachine

Old hydraulic machine (purchased before 1985) without operator or profit

0.8

Standard hydraulic machine or older electric machine (before 1998) operator or profit

1.0

Modern electric machine without operator or profit

1.1

Molder profit

Add 0.1

Take-out robot and conveyor

Add 0.05

Hot runner temperature control

Add 0.05

Gas assist control

Add 0.1

Injection-compression control

Add 0.1

Dedicated operator/assembler

Add 0.3

Foaming or induction heating unit

Add 0.3

Two-shot molding machine

Add 0.6

Three-shot molding machine

Add 0.9

The clamp tonnage required for molding will be analyzed during the filling system design. However, the clamp tonnage can be conservatively estimated assuming an average melt pressure of 80 MPa (11,600 psi) applied to the projected area, Aprojected , of the mold cavities. If the projected area is unknown, it can be estimated as the product of the part length and width. The clamp force in metric tons, t = 9800 N, is then Aprojected æ ö÷  ç ét ù ÷ ç (3.7) Fclamp = 80 ×10 ëéPaûù × ççncavities × Lpart × Wpart êé m2 úù÷÷÷× ë û ë û ÷ 9800 é Nù çç ÷ ë û è ø 6

Example: Estimate the processing cost per molded laptop bezel. The analysis assumes that a hot runner system is used with a take-out robot to fully automate the molding process. The corresponding cycle efficiency factor is 1.5. The cycle time is then estimated as

é s ù 2 ú (1.5 éë mm ùû ) ×1.5 = 13.5 s tcycle = 4 ê 2 ëê mm ûú If a modern electric machine is used with a take-out robot/conveyor, and a hot runner controller, then, allowing for molder profit, the machine techno­l­ ogy factor is

fmachine = 1.1 + 0.05 + 0.05 + 0.1 = 1.3

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3 Mold Cost Estimation

The clamp tonnage is estimated as

é mTon ù û = 294 mTon Fclamp = 75 ×106 éëPaùû × 1× 0.24 m × 0.16 m éê m2 ùú × ë ë û 9800 é Nù ë û

(

)

It should be noted that the true required clamp tonnage is likely less than 294 metric tons since the laptop bezel has a large window in it. The ana­ lysis, however, is conservative. The molding machine rate is then estimated as

Rmolding_machine = (43.3 + 0.095 × 294)×1.3 = $92.60 /hr The processing cost of the molded part can then be estimated by Eq. (3.4) as

Cprocess/part =

13.5 s cycle $92.60 hr ´ = $0.347/part 1 part cycle 3600 s hr 

3.2.4 Defect Cost per Part There are many reasons that molded parts are rejected. Some common defects include short shot, flash, contamination, improper color match, surface striations due to splay or blush, warpage and other dimensional issues, burn marks, poor gloss, and others. Since customers demand high quality levels on the molded parts they purchase, molders often internally inspect and remove any defective parts that are molded before shipment to the customer. The cost of these defects can be incorporated into the part cost by estimating the yield. Typical yields vary from 50 to 60 % at start-up for a difficult application with many quality requirements to virtually 100 % for a fully matured commodity product. Table 3.6 provides yield estimates according to the number of molding cycles and quality requirements. Table 3.6 Yield Estimates Total number of molding cycles

Low quality requirements

High quality requirements

~10,000

0.95

0.90

~100,000

0.98

0.95

~1,000,000

0.99

0.98

3.3 Mold Cost Estimation

Example: Estimate the yield and total cost for each molded laptop bezel. Since the production quantity is on the order of one million pieces and the quality requirements are assumed to be high, the yield factor is estimated as

fyield = 0.98 Substituting the prior results into Eq. 3.1, the total part cost is estimated as

CPart =

$0.152 + $0.082 + $0.347 = $0.593 0.98

The large cost of the mold relative to the material and processing costs indicates that the mold may have been over-designed. Further cost analyses are performed in Section 3.4 to analyze the effectiveness of a cold runner mold design with a lower initial mold cost but increased material and processing costs. 

„„3.3 Mold Cost Estimation Many cost estimation methods have been developed for molded plastic parts with varying degrees of causality and accuracy [3–14]. The following cost estimation method was developed to include the main effects of the part design and molding process while being relatively simple to use. To use the developed method, the practitioner can refer to the cost data provided in Appendices A, B, and D or provide more application-specific data as available. The total mold cost, Ctotal_mold , is the sum of the cost of the mold base, Cmold_base , the cost of all cavities, Ccavities , and the cost of their customization, Ccustomization Ctotal_mold = Cmold_base + Ccavities + Ccustomization (3.8) Example: Estimate the total cost of a single-cavity hot runner mold for producing the laptop bezel. This example corresponds to the mold design shown in Fig. 1.8. Subsequent analysis will show that the cost of the core and cavity inserts are $28,100, the cost of the mold base is $5,250, and the cost of the customizations including the purchase and finishing of all associated components is $42,500. As such, the estimated total cost the mold is

Ctotal_mold = Ccavities + Cmold_base + Ccustomization Ctotal_mold = $28,100 + $5,250 + $42,500 » $75,900 

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3 Mold Cost Estimation

3.3.1 Mold Base Cost Estimation A mold base is a template or blank mold that is ready to be customized. Referring to the mold design in Fig. 1.8, the mold base includes the majority of the mold with the exception of the core insert, cavity insert, hot runner, and related components such as ejector pins, support pillars, and cooling plugs. The cost of the mold base is a function of the mass of the mold and the cost of the steel per unit mass. Statistical cost analysis of mold bases was conducted and found that cost could be closely modeled as Cmold_base = $910 + M mold × kmold_material (3.9)

where Mmold is the mass of the mold base in kg, and kmold_material corresponds to the cost of the mold material per kilogram. Cost data for some commonly used mate­ rials is provided in Table 3.7. The coefficients for Eq. 3.9 were derived by statistical regression of actual mold base costs for several different mold bases (from small to large size) and four standard mold base materials [15]. The provided model has a coefficient of determination, R2, of 0.922 across 24 different mold base quotes and provides reasonable cost estimates of the mold base. Table 3.7 Mold Material Cost Coefficients Material

Similar commodity

Mold metal cost ­coefficient (US$/kg)

Hardness

Alcoa QC10

7075-T651

10.80

~160 Bhn

DME #1

SAE 1030

 3.99

~180 Bhn

DME #2

AISI 4130

 6.46

~290 Bhn

DME #3

AISI P20

 8.08

~310 Bhn

DME #7

AISI 400

13.19

~330 Bhn

Given the various mold dimensions, the mass of the mold base was estimated statistically from regression of eight differently sized mold bases as M mold = 1330

kg kg ´ Lmold ´ Wmold + 17,200 3 ´ Lmold ´ Wmold ´ H mold (3.10) 2 m m

While the mold dimensions are finalized during the mold layout design process, they can be initially estimated as Lmold = Lcavity ´ ncavities_length ´ 1.33 Wmold = Wcavity ´ ncavities_width ´ 1.33 (3.11) H mold = 0.189 + 2H cavity

3.3 Mold Cost Estimation

where ncavities_length and ncavities_width are the number of cavities across the length and width dimensions. The factor of 1.33 is included to provide for an allowance around the mold cavities. If the layout of the mold cavities across the mold is unknown, a grid layout is initially assumed as ncavities_length = ncavities_width = ceiling

(

)

ncavities

(3.12)

where the function ceiling(·) rounds any noninteger number up to the next integer. This estimate will tend to make the mold have larger size and cost than would ­actually be realized, but will provide at least a reasonable estimate. Example: Estimate the cost of the mold base for the laptop bezel. To estimate the cost of the mold base, it is first necessary to estimate the size of the mold base. Since this is a single-cavity mold, the mold base ­dimensions are estimated as

Lmold = 0.264 m × 1 × 1.33 = 0.351 m Wmold = 0.176 m × 1 × 1.33 = 0.234 m H mold = 0.189 + 2 × 0.057 m = 0.30 m The mass of the mold is then estimated as

kg kg × 0.35 m × 0.23 m + 17,200 3 × 0.35 m × 0.23 m × 0.30 m m2 m = 538 kg

M mold = 1330 M mold

If a DME#3 (AISI P20) steel is used for the mold base construction, then the cost of the mold is $8.08 per kg. The cost of the mold base is then estimated as

Cmold_base = $910 + 538 kg × 8.08

$ = $5250 kg



3.3.2 Cavity Cost Estimation The cost of the core and cavity inserts is typically the single largest driver of the total mold cost. The reasons for their expense are that they need to contain every geometric detail of the molded part, are made of very hard materials, and are ­finished to a high degree of accuracy and quality.

55

56

3 Mold Cost Estimation

The total cost of all the cavity and core inserts is driven by the cost of each set of inserts, Ccavity , multiplied by the number of cavity sets, ncavities , and a discount ­factor, fcavity_discount , which decreases the cost per cavity with the number of cavities that are made as indicated in Table 3.10. Ccavities = (Ccavity ´ ncavities )´ fcavity_discount (3.13)

Example: Estimate the total cost of all core and cavity insert sets for the laptop bezel. Since there is only one cavity and no cavity discount, the cost of all insert sets is:

Ccavities = ($28,100 ×1)×1 = $28,100 

3.3.2.1 Cavity Set Cost The cost of each cavity set is estimated as the sum of the materials costs, Ccavity_material , the insert machining costs, Ccavity_machining , and the insert finishing costs, Ccavity_finishing . Ccavity = Ccavity_material + Ccavity_machining + Ccavity_finishing (3.14)

Example: Estimate the cost of one set of core and cavity inserts for the laptop bezel. Subsequent analysis will show that the cost of the materials is $440, the cost of the cavity machining is $25,800, and the cost of the cavity finishing is $1,700. As such, the total cost for one core and cavity set is

Ccavity = $440 + $25,800 + $1900 » $28,100 

3.3.2.2 Cavity Materials Cost The cost of the cavity insert materials is the simplest and least significant term to evaluate as the product of the volume of the cavity set, Vcavity_material , the density of the insert material, ρcavity_material , and the cost of the insert material per kilogram, kcavity_material . Ccavity_material = Vcavity_material × rcavity_material × kcavity_material (3.15)

Cost data for some common metals are provided in Appendix B.

3.3 Mold Cost Estimation

The cavity insert volume is the product of the cavity length, Lcavity , the cavity width, Wcavity , and the cavity height, Hcavity . Vcavity_material = Lcavity × Wcavity × H cavity (3.16)

The size of the cavity set is finalized during the mold layout design process as is discussed in Chapter 4. From generalization of the later analysis, these dimensions can be roughly estimated as a function of the part size as follows. Lcavity = Lpart + max éê0.1× Lpart or H part ùú ë û é Wcavity = Wpart + max ê0.1× Wpart or H part ùú (3.17) ë û H cavity = max êé0.057or 2H part úù ë û

It should be noted that for the formula to work with the data provided in the ­appendices, all dimensions must be stated in meters or converted with the data to another consistent set of units. As previously suggested, the analysis should be conducted using application-specific data for the material properties, part geo­ metry, mold geometry, or manufacturing processes when such data is available. Example: Estimate the cost of the core and cavity insert materials for the laptop bezel. First, the dimensions of the core and cavity inserts are estimated. From the ­dimensions provided in Table 3.1, the preliminary dimensions of the inserts are

Lcavity = 0.24 m + max éë0.1 ´ 0.24 m or 0.01 m ùû = 0.264 m Wcavity = 0.16 m + max ëé0.1 ´ 0.16 m or 0.01 m ûù = 0.176 m H cavity = max éë0.057 or 2 ´ 0.01 m ùû = 0.057 m which provides a volume of

Vcavity_material = 0.264 m ´ 0.176 m ´ 0.057 m = 2.65 ´ 10-3 m3 To calculate the cost of the core and cavity insert materials, the type of material must be known. Since this is a tight tolerance part with a high production quantity, tool steel D2 is selected for its wear and abrasion resistance. This material has a density of 7670 kg/m3 and a cost of 21.4 $/kg, which leads to a cost for the core and cavity insert materials of

Ccavity_material = 2.65 ´ 10-3 m3 ´ 7670

kg $ ´ 21.4 = $440 3 kg m 

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3 Mold Cost Estimation

3.3.2.3 Cavity Machining Cost The cavity machining cost, Ccavity_machining , is often the single most significant driver of the total mold cost, and is a function of many variables including ƒƒthe volume and geometric complexity of the part to be molded, ƒƒthe core and cavity inserts’ material properties, ƒƒthe machining processes, ƒƒthe labor cost, and ƒƒthe quality of the inserts required. The approach used here is to estimate the cavity machining cost by multiplying the machining time, tcavity_machining , with the machining labor rate, Rmachining_rate . Ccavity_machining = tcavity_machining × Rmachining_rate (3.18)

The machining labor rate, Rmachining_rate , varies substantially with the cost of living in the location where the mold is manufactured. A mold maker in a high cost of living area (such as Germany) will tend to have a higher labor cost than a mold maker in a low cost of living area (such as China). Furthermore, the labor rate will also vary with the toolset, capability, and plant utilization of the mold maker. For example, a mold maker using a five-axis numerically controlled milling machine will tend to have more capability and charge more than a mold maker using manually operated three-axis milling machines. Some approximate cost and efficiency data for machining and labor rates is provided in Appendix D, though application-specific data with the negotiated machinist’s rate should be used if this data is available. The cavity machining time is driven by the size and complexity of the cavity details to be machined, as well as the speed of the machining processes used. In theory, the exact order and timing of the manufacturing processes can be planned to provide a precise time estimate. In practice, however, this approach is fairly ­difficult unless the entire job can be automatically processed, for instance, on a numerically controlled mill. The cavity machining time is estimated as the sum of the volume machining time, tcavity_volume , and the area machining time, tcavity_area . To take application-specific ­requirements into account, the cavity machining time is then multiplied by a ­geometric complexity factor, fcavity_complexity , and a machining factor, fmachining , then divided by an efficiency factor, fmachining_efficiency . tcavity_machining = (tcavity_volume + tcavity_area )×

fcavity_complexity × fmachining fmachining_efficiency

(3.19)

3.3 Mold Cost Estimation

The cavity volume machining time is a function of the volume of material to be removed and the material removal rate. To provide an approximate but conservative estimate, the assumption is made that the removal volume is equal to the entire volume of the core and cavity inserts. This may seem an overly conservative estimate, but in fact much of the volume must be removed around the outside of the core insert and the inside of the cavity insert. The material removal rate is a function of the processes that are used, the finish and tolerances required, as well as the properties of the mold core and cavity insert materials. To simplify the analysis, a geometric complexity factor will later be used to capture the effect of different machining processes and tolerances needed to produce the required cavity details. As such, the volume machining time captures only the time to require the material removal as follows. tcavity_volume =

Vcavity_material Rmachining

(3.20)

where Rmachining is the volumetric mold material machining rate measured in cubic meters per hour. Machining data for different materials are provided in Appendix B, though application-specific material removal rates can be substituted if the depth of cut, speed, and feed rates are known [22]. The cavity area machining time, tcavity_area , estimates the time required to machine all the cavity surfaces, and is similarly evaluated as follows. tcavity_area =

Apart_surface Rmaterial_area

(3.21)

where Apart_surface is the total surface area of the part measured in square meters and Rmaterial_area is the area mold material removal rate measured in square meters per hour. Modern 3D computer-aided design systems can provide an exact measure of the part’s surface area and volume. The cavity complexity factor, fcavity_complexity , adjusts the cavity machining time to account for the design and manufacturing of the myriad of features that will ­compose the mold cavity. Some of the activities that the complexity factor accounts for include: ƒƒDecomposition of the mold cavity into multiple machining tasks; ƒƒGeneration of machining tasks and NC programs, including electrodes for electrical discharge machining; ƒƒExecution of machining tasks, including multiple machine setups, electrical discharge machining, milling, etc.; ƒƒInspection and rework to obtain all the specified geometry.

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Previous research [5, 23] has found that the complexity of the cavities is related to the total number of dimensions and/or features specified in the design of the part to be molded. Unfortunately, these former cost estimation approaches are time-­ consuming and dependent upon the subjective opinion as to what constitutes a ­dimension or feature. As such, this cost analysis uses a complexity factor that is based on the ratio of the expected volume of the part (the surface area, Apart_surface , times the wall thickness, hwall) compared to the actual volume of the part, Vpart . fcavity_complexity =

Apart_surface × hwall Vpart

(3.22)

This complexity factor increases with the addition of features, since each added feature (such as a rib, boss, or window) increases the surface area of the part without causing a significant increase in the actual part volume. To demonstrate different levels of complexity, Table 3.8 provides the calculated complexity factors for part designs of varying complexity. Table 3.8 Complexity Factor for Various Example Part Designs Part design

Complexity factor 1.02

1.9

2.5

3.1

3.3 Mold Cost Estimation

The machining factor, fmachining , accounts for the discrepancy in the material removal rates for various type of machining. The volumetric removal rates provided in Appendix B assume a carbide, two fluted, 19.05 mm (¾ inch) diameter end mill with a depth of cut of 3.2 mm (0.125 inch); the surface area removal rate assumes a carbide, four fluted, 6.35 mm (¼ inch) diameter end mill operating at half the nominal feed rate recommended for the various materials. Since the cavity and core inserts are typically produced with a variety of machining operations, the overall machining factor for a given application is the weighted average of each of the machining factors provided in Table 3.9 in proportion to its use. Table 3.9 Machining Factor for Various Processes Machining process

Machining factor

Turning

0.5

Drilling

0.5

Milling

1

Grinding

2

EDM

2

The machining efficiency factor, fmachining_efficiency , accounts for the fraction of time that labor and machine time are spent on non-machining activities. In theory, the efficiency of a fully automated numerical control machining cell will approach 100 %. In reality, the efficiency rarely exceeds 50 %. The reason is that a significant amount of time is required to develop the sequence of machine operations, procure and check cutting tools, perform setups, verify cutting paths, create electrodes, operate EDM, and other tasks. As such, a machining efficiency rate of 25 % is ­recommended for cost estimation reasons. Example: Calculate the machining cost of the core and cavity inserts for the laptop bezel. Using the application data Table 3.1 and the removal rates for tool steel D2 from Appendix B, the machining times are estimated as

tcavity_volume =

tcavity_area =

2.65 ×10-3 m3 = 3.78 h 7.00 ×10-4 m3 h

0.0457 m2 = 2.69 h 0.0170 m2 h

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3 Mold Cost Estimation

The cavity complexity factor is evaluated as

fcavity_complexity =

45700 mm2 ´1.5 mm = 2.5 27500 mm3

Since the laptop bezel contains many narrow ribs that will be produced ­primarily with EDM, a machining factor of 4 is used. An efficiency factor of 25 % is also assumed. The estimated machining time is

tcavity_machining = (3.78 h + 2.69 h)×

2.5 × 4 = 258 h 0.25

The statistical cost data in Appendix D indicates that the average hourly wage for a tool and die maker in the United States is $23.94 per hour. This hourly wage is the direct salary to the employee, and does not include the employee’s fringe benefits (such as medical benefits, vacation, etc.), the cost of the equipment and plant (such as land, building, machinery, etc.), the cost of supplies (such as cutting tools/fluids, electricity, water, etc.), and other overhead (including management salary, profit, etc.). As such, mold-makers will charge significantly more than their employee’s direct ­wages. Assuming a billed hourly rate for the machinists of 100 $/h, the ­estimated cavity machining costs are

Ccavity_machining = 258 h ×100

$ = $25,800 h 

3.3.2.4 Cavity Discount Factor The cavity discount factor stems from the fact that there are fixed costs associated with the design and tooling of the first cavity set. Manufacturing productivity will then improve as additional sets are machined. Accordingly, a set of discount factors is provided in Table 3.10 as a function of the number of cavity sets to be made. For each doubling in the number of cavity sets, the cost is reduced by 15 %. However, after 16 cavities, it is difficult to further improve manufacturing productivity. This table is based on generic human factors research [24], so the discount factor may be replaced with application-specific data if available. Table 3.10 Discount Factor as a Function of Number of Cavity Sets Number of cavity sets

Discount factor

1

1

2

0.85

4

0.72

8

0.61

16 or more

0.52

3.3 Mold Cost Estimation

Example: Since the laptop bezel is produced in a single cavity mold, there is no quantity discount and the discount factor is set to one. 

3.3.2.5 Cavity Finishing Cost The cavity finishing cost, Ccavity_finishing , is also a significant cost driver representing 5 to 30% of the total mold cost [25]. The finishing cost is the product of the time required to finish the cavity surface area, tcavity_finishing , and the finishing labor rate, Rfinishing_rate . Ccavity_finishing = tcavity_finishing × Rfinishing_rate (3.23)

The finishing time is a function of each area of the part to be finished, Aipart_surface , divided by the rate at which the area is finished, Ricavity_finishing . tcavity_finishing = å i

i Apart_surface i Rcavity_finishing

(3.24)

Since the finishing rate depends on the surface finish and texture to be applied, the use of the summation over the index i in Eq. (3.24) indicates that the time ­required to finish each portion of the mold to various finishes must be added ­together. Some representative finishing rates are provided in Table 3.11, which were adapted from Rosato [25] to account for the various finish levels. Approximate labor rates for finishing are provided in Appendix D. Since finishing can be quite labor-intensive, the finishing of core and cavity inserts is sometimes outsourced. Table 3.11 Finishing Rates Finish

Finishing method

Surface ­roughness (µm)

Finishing rate (m2/h)

Finishing rate (in2/h)

Texture

Electrochemical engraving

50–200

0.0002

0.3

SPI A-1

#3 diamond polish

1

0.0005

1

SPI A-3

#15 diamond polish

3

0.001

2

SPI B-3

#320 grit cloth

8

0.002

4

SPI C-3

#320 stone

40

0.005

8

SPI D-2

#240 oxide blast

30

0.01

20

SPI D-3

#24 oxide blast

200

0.02

30

63

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3 Mold Cost Estimation

Example: Calculate the cavity finishing cost for the laptop bezel. Assume that the laptop bezel is to be finished to SPI B3 on all surfaces, which together have a surface area of 0.0457 m2, except for an improved surface finished of SPI A1 to be applied to the front surface of the bezel, which has an approximate area of 0.01 m2. The estimated finishing time is

tcavity_finishing =

0.0357 m2 0.01 m2 + = 38 h 2 0.002 m /h 0.0005 m2 /h

In the preceding equation, the area of the front surface of the bezel is subtracted from the total area of the bezel to avoid double-computing the time to finish the front surface. If the labor rate for finishing is 50 $/h, then the cost of cavity finishing is

Ccavity_finishing = 38 h ´ 50

$ = $1900 h 

3.3.3 Mold Customization The mold base customization includes many design, machining, and assembly steps. Some of the specific steps in the mold customization include: ƒƒCutting pockets and bolt holes in the mold plates to receive the core and cavity inserts. This cost is proportional to the number of mold cavities and the mold dimensions. ƒƒMilling a cold runner system into the mold plates or purchasing a hot runner system and modifying the mold accordingly. This cost is related to the type of feed system, the number of gates, and the mold dimensions. ƒƒDrilling, tapping, and plugging the cooling lines in the mold. This cost is related to the number and layout of the cooling lines, which is related to the number of cavities and their geometry. ƒƒDrilling and reaming holes in the core inserts and support plates to accept ejector pins, and providing appropriate counter bores in the ejector retainer plate. This cost is related to the number of ejector pins, which is related to the number of cavities and their geometry. ƒƒMilling holes in the ejector plate and the ejector retainer plate to provide support pillars, if needed. This cost is related to the number of cavities and their geo­ metry. ƒƒDesigning and machining other necessary mold components such as stripper plates, slides, core pulls, etc. These costs are related to the specific part geometry and application requirements.

3.3 Mold Cost Estimation

A detailed cost analysis of all the customizations is too lengthy to present given the necessary discussion of the assumptions and equations. However, a review of the above customizations indicates that the costs are generally related to the size of the mold base, the cost of the inserts, and the specific technologies required. ­Accordingly, a reasonably simple model is i Ccustomization = Ccavities × å fcavity_customizing i

i +Cmold_base × å fmold_customizing

(3.25)

i

where the coefficients fcavity_customizing correspond to the factors governing the costs of customizing the cavity inserts, and the coefficients, fmold_customizing , correspond to the factors governing the costs of modifying the mold base. The summation over i represents the added customization for each of the mold subsystems, with i including the [feed system, cooling system, ejector system, structural system, and other items]. It should be noted that these customization factors have been developed so as to include the procurement cost of the required components and system assemblies, such as hot runners, fittings, core pulls, etc. Feed systems are discussed in detail in Chapter 6. The cost factors associated with modifying the cavity inserts and mold base for accommodating different types of feed systems are provided in Table 3.12. A simple molding application with one to four cavities might use a two-plate cold runner system with fcavity_customizing equal to 0.05 and fmold_customizing equal to 0.1. For a molding application with high production volume and sixteen or more cavities, a thermally gated hot runner might be used with fcavity_customizing equal to 0.5 and fmold_customizing equal to 1.0. Cooling systems are discussed in Chapter 9. The cost factors for various cooling system designs are provided in Table 3.13. Many molds use straight lines with O-rings and fittings, adding 5 % to the cost of the cavity inserts and 20 % to the cost of the mold base. As the cooling system becomes more complex, the implemen­ tation cost increases. Ejector systems are discussed in Chapter 11. The cost factors for various ejection system designs are provided in Table 3.14. Most molds can be assumed to use a mix of round ejector pins, blades, and sleeves though ejection requirements will vary significantly depending on the part geometry and application requirements.

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3 Mold Cost Estimation

Table 3.12 Feed System Cost Coefficients Feed system design

Cavity cost coefficient,

Mold cost coefficient,

Two-plate cold runner system

0.05

 0.1

Three-plate cold runner system

0.1

 1.0

Hot runner system with thermal gate

0.4

 2.0

Hot runner system with valve gates

0.5

 4.0

Hot runner stack mold with thermal gates

0.5

 8.0

Hot runner stack mold with valve gates

0.9

12.0

feed_system fcavity_customizing

feed_system fmold_customizing

Table 3.13 Cooling System Cost Coefficients Cooling system design

Cavity cost coefficient,

Mold cost coefficient,

Straight lines with O-rings and fittings

0.05

0.2

Straight lines with bubblers or baffles, O-rings, and fittings

0.10

0.2

Circuitous cooling lines with O-rings, plugs, and fittings

0.15

0.4

Circuitous cooling lines with bubblers or baffles, O-rings, plugs, and fittings

0.20

0.4

Complex cooling line layout with thermally conductive inserts or contoured cooling inserts

0.25

0.8

cooling_system fcavity_customizing

cooling_system fmold_customizing

Table 3.14 Ejector System Cost Coefficients Ejector system design

Cavity cost coefficient,

Mold cost coefficient,

Round ejector pins

0.1

0.1

Mix of round ejector pins, blades, and ­sleeves

0.2

0.2

Stripper plate

0.2

0.4

External slide or lifter

0.2

0.4

Internal slide or lifter

0.4

0.4

Actuated core pull

0.4

0.5

Reverse ejection system

0.5

1.0

ejector_system fcavity_customizing

ejector_system fmold_customizing

3.3 Mold Cost Estimation

The structural design of molds is detailed in Chapter 12. The cost factors for various structural system designs are provided in Table 3.15. Most molds with high production volumes can be assumed to use support pillars and parting plane interlocks. In this cost estimation method, the sealing of the cavity by the core insert and cavity insert is considered as part of the structural system. The cost of the mold will increase with the complexity of the parting surface, the design of which will be discussed in Section 4.1. There are many other customizations that can be performed on the mold. Some of these factors are provided in Table 3.16 and are applied as necessary. For most molds, none of these customizations are required. Table 3.15 Structural System Cost Coefficients Structural system design

Cavity cost coefficient,

Mold cost coefficient,

No additional support structures and ­planar mold parting surface

0.0

0.0

Multistepped parting surface

0.2

0.0

Complex contoured parting surface

0.4

0.2

Support pillars

0.0

0.1

Support pillars and interlocks

0.1

0.2

Split cavity mold

0.5

1.0

structural_system fcavity_customizing

structural_system fmold_customizing

Table 3.16 Other Customization Cost Coefficients Required mold customization

Cavity cost coefficient,

Mold cost coefficient,

Mold temperature sensors

0.05

0.1

Mold pressure sensors

0.05

0.1

Gas assist molding

0.2

0.5

Runner shut-offs

0.0

0.1

Dynamic melt control

0.2

1.0

Insert molding

0.4

0.4

In-mold labeling

0.4

0.4

Two-shot molding

2.0

4.0

Three-shot molding

3.0

6.0

miscellaneous fcavity_customizing

miscellaneous fmold_customizing

67

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3 Mold Cost Estimation

Example: Estimate the cost of customizing the mold base and inserts for the laptop bezel. The mold will use a hot runner with thermal gates, so the appropriate ­customization factors are: feed_system fcavity_customizing = 0.4 feed_system fmold_customizing = 2.0

The mold will use a cooling system with circuitous cooling lines, O-rings, and plugs, so the appropriate customization factors are: cooling_system fcavity_customizing = 0.15 cooling_system fmold_customizing = 0.4

The mold will use an ejector system with a mix of round ejector pins, blades, and sleeves, so the appropriate customization factors are: ejection_system fcavity_customizing = 0.2 ejection_system fmold_customizing = 0.2

The mold will use a structural system with support pillars and interlocks. Also, the mold will require a stepped parting plane to form the details along the side of the molding as shown in Figure 3.4. As such, the appropriate customization factors are: structural_system fcavity_customizing = 0.1 + 0.2 = 0.3 structural_system fmold_customizing = 0.2 + 0.0 = 0.2

The mold will use a melt thermocouple at the end of flow and a melt ­pressure transducer near the gate for process control purposes, so additional customization factors are: miscellaneous fcavity_customizing = 0.05 + 0.05 = 0.1 miscellaneous = 0.1 + 0.1 = 0.2 fmold_customizing

The cost of all customizations may then be calculated as:

Ccustomizations = $28,100 × (0.4 + 0.15 + 0.2 + 0.3 + 0.1) +$5250 × (2 + 0.4 + 0.2 + 0.2 + 0.2) = $42,500 To summarize the above analysis, the total cost of the mold is estimated as:

Ctotal_mold = Ccavities + Cmold_base + Ccustomization Ctotal_mold = $28,100 + $5250 + $42,500 » $75,900 The estimate seems reasonable for a mold produced in the United States. On the other hand, this result may over-estimate the cost of the mold if made in Asia, especially if not including a hot runner system. 

3.4 Manufacturing Strategy

„„3.4 Manufacturing Strategy This chapter presented an overview of the mold quoting process, including a ­detailed mold cost and part cost estimation methodology. The methodology was developed to utilize minimal information and yet provide the causal analysis relating critical mold design decisions to the mold cost and part cost. It is recommended that multiple cost estimates be developed for different mold designs until an effective mold specification is established.

3.4.1 Breakeven Analysis In theory, the production quantities should be known beforehand and used to ­design an “optimal” mold for the specified quantity. In reality, the production schedules and quantities are not precisely known, so the molder and customer must carefully consider the possible result of using molds that are over- or ­under-designed. For this reason, alternative mold designs should be considered and analyzed with respect to breakpoints in production quantity where different mold design approaches are preferred. Minimization of the total molded part cost is not a simple task since injection molds and molding processes are optimally designed for different target production quantities. Typically, there is a tradeoff between the upfront investment in the mold and later potential savings related to the processing and material costs per part. Consider alternative mold design and molding strategies for producing the bezel part shown in Fig. 3.4 with (1) a one-cavity cold runner mold or (2) a two-­ cavity hot runner mold. The described methods have been applied to estimate the mold, material, and processing cost for these alternative mold designs with the results summarized in Table 3.17. Table 3.17 Part Cost Data for Two Different Bezel Molds Runner system

Cold runner

Hot runner

Number of mold cavities

1

2

Mold insert material

AISI4130

D2

Effective cycle time/part

27 s

9s

Molding machine hourly rate

$85.46/h

$128.88/h

Processing cost/part

$0.641

$0.16

Material cost/part

$0.101

$0.082

Mold maintenance cost/part

$0.176

$0.132

Estimated yield

95%

98%

Marginal cost/part

$0.965

$0.382

Initial mold investment

$58,600

$131,700

69

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3 Mold Cost Estimation

It is useful to consider the total costs, Ctotal , incurred to produce a given quantity that is computed as Ctotal = Cfixed + ntotal × Cmarginal (3.26)

where Cfixed is the total cost of the mold and its maintenance, ntotal is the total production quantity across the life of the mold, and Cmarginal is the total marginal cost of the resin, machine, labor, and energy on a per-part basis. For a given mold ­design, the marginal cost per piece will remain fairly constant across the life of the application (though there may be cost decreases related to elimination of defects, reductions in cycle times, etc. as well as cost increases due to material pricing or shipping costs). To provide the best possible mold design and quote, multiple mold designs should be developed for different target production quantities, and the ­total production costs estimated and compared via breakeven analysis. Example: Consider the cost data provided in Table 3.17. Calculate the production ­volume where a two-cavity hot runner mold becomes more economical than a one-cavity cold runner mold. Equation 3.26 is used to calculate the costs with the cold runner and hot runner as: cold_runner cold_runner cold_runner Ctotal = Cfixed + ntotal × Cmarginal hot_runner hot_runner hot_runner Ctotal = Cfixed + ntotal × Cmarginal

Equating these two costs and solving for the production volume provides the breakeven quantity. breakeven ntotal =

hot_runner cold_runner Cfixed - Cfixed

cold_runner hot_runner Cmarginal - Cmarginal

The analysis assumes that the marginal cost per molded part consists pri­ marily of the processing and material costs. Then, the marginal costs for the cold and hot runners are $0.55 and $0.16, respectively. Substituting these values provides breakeven ntotal =

$131,700 – $58,600 $73,100 = = 125,000 parts $0.965/part – $0.382/part $0.583/part

The costs for the cold and hot runner mold designs are provided in Fig. 3.5. While the cost function of Eq. 3.26 is linear, a log-log scale has been used in the figure to provide better resolution of the cost across a wide range of production volumes. In this example, the total production cost for the one-cavity

3.4 Manufacturing Strategy

cold runner mold and the two-cavity hot runner mold are plotted as a function of the “realized” production quantity, Q. For this example, the one-cavity cold runner mold has a lower total cost up to the 125,000 production quan­ tity, after which the two-cavity hot runner mold provides a lower total cost.

Figure 3.5 Breakeven analysis for bezel



The cost analysis will typically indicate the need for different mold designs at extremely low and extremely high production quantities. In the previous example, the upfront cost of the two-cavity hot runner system cannot be justified at low or moderate production quantities. At higher production quantities, however, a hot runner system is essential to maximizing profitability since the marginal costs of operating the hot runner mold are significantly less than those of the cold runner mold. While the breakeven analysis supports clear design decisions at very low and very high production quantities, the mold design can be less certain at intermediate production volumes. If the production quantity is on the order of 100,000 or 200,000 parts, then either mold design may be acceptable from a cost perspective. The customer can be provided both designs to select the alternative that best fits their technical requirements and business strategy. Many molders and customers require a quick return on investment and so will examine the total cost curve to accept the use of a hot runner system with higher cavitation only if a desirably short payback period can be achieved. Sometimes, however, mold design decisions are not based solely on economics but rather by other concerns such as: ƒƒThe need for a mold to permit rapid color changes, for which a hot runner feed system may not be desirable. The color change issue in hot runners will be revisited in Section 6.4.8.

71

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ƒƒThe capability and preference of the molder that will use the mold. If the molder does not have the experience or auxiliaries required to utilize a hot runner system, then a cold runner mold may best be utilized. ƒƒThe lean manufacturing strategies of the molders to reduce costs and improve quality. For instance, it is not uncommon for molders to standardize on a specific type and size of mold and molding machine to maximize production flexibility and reduce setup times. As a general practice, the mold should be designed to maximize the molder’s productivity unless the application requirements and cost constraints dictate otherwise. When an advanced molding application has special requirements, it may be critical to select a molder with an advanced capabilities that can operate the mold without undue experimentation. Chapter 13 provides a survey of mold technol­ ogies, many of which require special molding machines and auxiliary systems.

3.4.2 Prototyping Strategy The described cost estimation methods were primarily developed for class 101 injection molds [16] that are intended for mass production of millions of molded parts. In many applications, there is the need to develop mold tooling for lower production volumes or to “bridge” prototype and production. For many molding applications, the four common manufacturing strategies are: 1. Class 101 production molds [16], 2. Aluminum tooling for moderate volumes [17], 3. Prototype molds made by additive manufacturing with photopolymers [18] of direct metal laser sintering, and 4. Products directly manufactured by 3D printing via fused deposition modeling, laser sintering, or stereolithography [19]. For each strategy, the upfront cost, marginal cost, and total production time as estimated can be a function of different target production volumes as an update to prior work [20]. Cost estimation is a well-established process in which the true costs required to procure a manufactured product are fully accounted. The accuracy of a cost estimate will vary based on the level of detail taken in accounting for all the activities and expenses in manufacturing, as well as the precision of the associated times and costs [21]. Again, the standard form for the total production cost, C, is applied for each of the alternative manufacturing strategies according to Eq. (3.26). The total production time can be similarly calculated as

3.4 Manufacturing Strategy

t total = t initial + ntotal × t marginal (3.27)

where tinitial is the initial lead time to procure the first part, and tmarginal is the production time for each additional part. The average part cost may also be estimated as Cpart = Ctotal ntotal (3.28)

Two significant issues in the evaluation of mold prototyping strategies are (1) the quality of the procured parts and (2) the longevity of the prototype tooling. First, prototype molds rarely provide the surface finish and performance of parts molded from production tooling. In particular, 3D printing processes have a finite resolution or line width that can result in rough surfaces that limit the part aesthetics and performance. Directly produced parts (by FDM, selective laser sintering, or stereolithography) have a reduced portfolio of grades as well as somewhat inferior properties related to the integrity of the constituted materials. Second, prototype tooling can have a shorter lifetime due to stress-related failures associated with injection of the polymer melt and/or ejection of the molded plastic parts. For these reasons, it is important to develop reasonable cost and time estimates for each manufacturing strategy. Table 3.17 provides some firm cost estimates for ­alternative strategies for production of the molded bezels. The first two rows correspond to the two-cavity hot runner and one-cavity cold runner molds that were previously analyzed. An alternative is the use of printed inserts that were produced by a polyjet process with ultraviolet-cured resin. This type of mold insert has a lifetime on the order of 100 molding cycles, so the mold inserts must be continuously reproduced to achieve high production quantities, which leads to a relatively high marginal production cost of $33 and marginal production time of 1 h. Alternatively, the bezels may be directly produced by fused-deposition modeling or selective laser sintering with a marginal cost of $60 and a marginal production time of 4 h. Table 3.17 Part Cost Data for Two Different Bezel Molds Type of mold

Initial cost

Marginal cost

Initial time

Marginal time

Two-cavity hot runner

$131,700

$0.382

84 days

9s

One-cavity cold runner

$58,600

$0.965

28 days

27 s

Printed inserts

$2900

$33

5 days

1h

Direct 3D printing

$100

$60

2 days

4h

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The average part production costs for the four alternative strategies are plotted in Fig. 3.6 as calculated according to Eqs. 3.26 and 3.28 for the model coefficients of Table 3.17; a log-log scale is applied given that the production volume and total production costs span a large range. It is observed that the average part costs vary significantly as a function of total production volume, and that each of the four strategies provides the lowest average part cost at some production volume.

Figure 3.6 Average part production cost

The direct 3D printing has the lowest production costs up to production volumes of 100 parts, after which molding with the printed inserts is the lowest alternative up to 1000 parts. The reason for the switchover to printed inserts at a production quantity of 100 is that the cost of the printed inserts is significant, and can only become fully amortized with this production volume. The total production cost ­using the printed inserts is nearly linear with increased production quantities as additional inserts are made and consumed repeatedly. Above a production volume of 1000, however, the recurring cost of the printed inserts is so significant that the one-cavity cold runner mold is preferred. The two-cavity hot runner mold becomes preferred for production volumes at 125,000 parts. It is remarkable how investment in custom tooling can reduce the $60/part cost to less than $0.40 at high production volumes. Production time is an important consideration in many applications, such that a plastics manufacturer may trade off higher costs to shorten delivery times. Figure 3.7 plots the total production time as a function of the production volume. It is observed that direct 3D printing is preferred for up to 50 parts. However, each part requires about 12 hours to print, so even assuming a four-hour production time

3.4 Manufacturing Strategy

with parallel printing, molding with printed inserts can provide a lower production time up to about 500 parts. At higher production volumes, aluminum or simple single-cavity tooling will provide the fastest delivery for production volumes up to 200,000 cavities. For higher production volumes, however, the production mold tooling with its two cavities will deliver lower total production time at very high production times.

Figure 3.7 Total part production time

There is currently widespread interest in 3D printing, with significant investments being made by hobbyists, manufacturers, venture firms, and government. The ­described analysis can be applied to alternative manufacturing strategies with ­updated cost coefficients on a time-dependent and application basis. The results suggest that, at least for the time being, production in most commercial applications will continue to be performed by conventional injection molding, especially when considering quality issues associated with 3D printing. Specifically, the line width in 3D printing is on the order of 100–200 µm. The deposition and welding of successive filaments results in an inherent surface roughness factor, Ra, on the order of 20% of the filament diameter. Surface finishes on the order of 40 µm are aesthetically and functionally poor [22]. While surface finish can be improved by sanding and other processes, cost and consistency are significant issues for large production volumes. The trend to smaller filament diameters to provide better precision and surface finish does impose the adverse effect of slower printing speeds, since more filament passes are required to deposit the same part volume. Still, plastics product and mold designers should not ignore 3D printing for at least three reasons.

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ƒƒFirst, 3D printing is an extremely useful process for prototyping both product and mold design concepts. It can certainly be used for pilot production to proof the designs prior to moving to other manufacturing strategies that are more suitable to higher production volumes. ƒƒSecond, 3D printing of mold cavity inserts, especially by stereolithography of UV-curable resins, holds significant promise. The reason is that tolerances in this process are significantly better than fused deposition. Furthermore, the costs and benefits of secondary processing of the printed inserts are applied across the molding of many plastic parts. Current efforts in metal plating and more advanced material compositions for 3D-printed mold inserts are likely to extend their longevity, and with it their manufacturing competitiveness for higher production volumes. ƒƒThird, 3D printing does provide potential benefits that are difficult to achieve in traditionally injection molded products. Some of these benefits include the in­ corporation of hollow features (e. g., internal cavities or conduits), the deposition of multiple materials (e. g., color and/or electrical properties), functionally gra­ diated materials (e. g., soft to rigid), and others. 3D printing technology is evolving rapidly and so it is possible that parallel processing and other deposition techniques may resolve the cost, processing, and quality issues [16].

„„3.5 Chapter Review This chapter presented an overview of the mold quoting process, followed by a ­detailed mold cost and part cost estimation methodology. The methodology was developed to utilize minimal information and yet provide causal analysis relating critical mold design decisions to the mold cost and part cost. It is recommended that multiple cost estimates be developed for different mold designs until an effective mold specification is established. After reading this chapter, you should understand: ƒƒThe primary cost drivers for injection molds, ƒƒThe primary cost drivers for molded parts, ƒƒThe mold quoting process and the typical schedule of payments required to make and operate a mold, ƒƒHow to estimate the cost of molded parts, including mold, material, and processing, ƒƒHow to estimate the cost of an injection mold, and ƒƒHow to use the cost estimation methodology to develop a manufacturing strategy that seeks to minimize the total cost of the mold, material, and process per part.

3.6 References

The burden to minimize the total part cost is shared between the mold designer, molder, and end user of the molded parts. The mold designer must contemplate trade-offs between the mold costs, material costs, and processing costs. In the long run, significant inefficiencies in the mold design brought about by poor design decisions will lead to lost profitability for all parties. In the next chapter, the specification resulting from the cost analyses will be used to design the layout of the mold. Afterwards, the design and analysis of the underlying systems in injection molds is conducted.

„„3.6 References [1] Bryce, D. M., Determining Injection Molding Costs, in Plastic Injection Molding, Society of Manufacturing Engineers (1997) p. 145 [2] Naitove, M., Semi-Annual Injection Molding Machine Hourly Rate Survey, Plast. News (2014) [3] Banker, R. J., et al., Cost of product and process complexity, Measures of Manufacturing Excellence, T. Kaplan, (Ed.), Harvard Business School Press Cambridge, Massachusetts (1990) [4] Beiter, K. A. and K. Ishi, Incorporating dimensional requirements into material selection and d ­ esign of injection molded parts, in ASME Design Automation Conference (1996) [5] Fernandez, R. L., The effect of part design on tooling cost in injection molding, in Department of Mechanical Engineering Univ. of Massachusetts, Amherst (1987) [6] Raviwongse, R. and V. Allada, Artificial neural network based model for computation of injection mould complexity, Int. J. Adv. Manuf. Technol. (1997) 13(8): pp. 577–586 [7] Xu, T., Functional feature-based approximate structural analysis for injection molded LFRTP part, in ANTEC (1994) [8] Fagade, A. and D. O. Kazmer, Economic Design Of Injection Molded Parts Using DFM Guidelines – A Review Of Two Methods For Tooling Cost Estimation, in Society of Plastics Engineers Annual Technical Conference, Atlanta, GA (1998) [9] Archer, D., Early cost estimation of injection molded components, in Manufactur. Engineer., Univ. of Rhode Island, Providence (1988) p. 175 [10] Beiter, K. A., J. M. Cardinal, and K. Ishi, Design for Injection Molding: Balancing Mechanical ­Requirements, Manufacturing Costs, and Material Selection, in ASME Design Automation Con­ference (1996) [11] Butler, M. J., Mold costs and how to estimate accurately, Plast. Polym., April (1973) pp. 60–61 [12] Fagade, A. and D. Kazmer, Economic design of injection molded parts using DFM guidelines–A review of two methods for tooling cost estimation, in ANTEC, SPE, Atlanta, GA (1998) [13] Fagade, A. and D. Kazmer, Effects of complexity on tooling cost and time-to-market of plastic in­jection molded parts, in ANTEC, SPE, New York (1999) [14] Bush, J. V., et al., Computer aided part cost analysis for injection molding, SPE Reg. Tech. Conf., Oct. (1988) [15] D-M-E American Standard Mold Base Price Catalog, Madison Heights, MI, March 30 (2015) [16] Customs and Practices of the Mold Making Industry, The Society of the Plastics Industry (1996) [17] Zhong, Z., M. Leong, and X. Liu, The wear rates and performance of three mold insert materials, Mater. Design (2011) 32(2): pp. 643–648

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[18] Martinho, P., et al., Hybrid moulds: the use of combined techniques for the rapid manufacturing of injection moulds, Virtual Modelling and Rapid Manufacturing, (2005) pp. 421–427 [19] Choi, J., et al., Development of a Mobile Fused Deposition Modeling System with Enhanced Manufacturing Flexibility, J. Mater. Process. Technol. (2010) [20] Karania, R. and D. Kazmer, Low Volume Plastics Manufacturing Strategies, J. Mech. Des. (2007) 129: p. 1225 [21] Fagade, A. and D. O. Kazmer, Early cost estimation for injection molded parts, J. Inject. Mold. Technol. (2000) 4(3): pp. 97–106 [22] Ahn, D., et al., Representation of surface roughness in fused deposition modeling. J. Mater. Process. Technol. (2009) 209(15): pp. 5593–5600

4

Mold Layout Design

During the mold layout stage, the mold designer confirms the type of mold and determines the dimensions and materials for the cavity inserts, core inserts, and mold base. Mold bases are only available in discrete sizes, so iteration between the inserts’ sizing and mold base selection is normal. The goal of the mold layout ­design stage is to develop the physical dimensions of the inserts and mold so as to begin procurement of these materials. Mold material selection is also an important decision, since the material properties largely determine the mold making time and cost as well as the mold’s structural and thermal performance. The mold layout design assumes that the number of mold cavities and type of mold has been determined. To develop the mold layout, the mold opening direction and the location of the parting plane are first determined. Then, the length, width, and height of the core and cavity inserts are chosen. Afterwards, a mold base is selected and the inserts are placed in as simple and compact a layout as possible. It is important to develop a good mold layout design since later analysis and detailed design assumes the layout design, and subsequent changes to the cavity and mold dimensions can quickly become difficult and expensive.

„„4.1 Parting Plane Design The parting plane is the contact surface between the stationary and moving sides of the mold. The primary purpose of the parting plane is to tightly seal the cavity of the mold and prevent melt leakage. This seal is maintained through the application of literally tons of force (hence the term “clamp tonnage”) that are applied normal to the parting plane. While the term “parting plane” implies a flat or planar surface, the parting plane may contain out-of-plane features. Prior to determining the parting line and designing the parting plane, the mold designer must first ­determine the mold opening direction.

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4.1.1 Determine Mold Opening Direction Examination of any of the previous mold designs (e. g., Figs. 1.4 to 1.8) indicates that the mold opening direction is normal to the parting plane. In fact, the mold usually opens in a direction normal to the parting plane since the moving platen of the molding machine is guided by tie bars or rails to open in a direction normal to the platen. Accordingly, guide bushings and/or mold interlocks are almost always located on the parting plane to guide the mold opening in a direction normal to the parting plane. It may appear that there is nothing about the mold opening direction to determine since the mold opens normal to the parting plane. However, it is necessary to determine the mold opening direction relative to the mold cavity and direction of part ejection. There are two factors that govern the mold opening direction. ƒƒFirst, the mold cavity should be positioned such that it does not exert undue stress on the injection mold. The mold cavity is typically placed with its largest area parallel to the parting plane. This arrangement allows the mold plates, ­already being held in compression under the clamp tonnage, to resist the force exerted by the plastic on the surfaces of the mold cavity. ƒƒSecond, the mold cavity should be positioned such that the molded part can be ejected from the mold. A typical molded part is shaped like a five-sided open box with the side walls, ribs, bosses, and other features normal to its largest area. If so, then the part ejection requirement again supports the mold opening direction to be normal to the part’s largest projected area. Consider the cup and lid shown in Fig. 4.1. A section of the core and cavity inserts used to mold these parts was previously shown in Fig. 1.6. There are only two ­potential mold opening directions relative to the part. One mold opening direction is in the axial direction of the cup, while the second direction is in the radial direction of the cup. A section of a cavity block with an axial mold opening direction is shown in Fig. 4.2. The two bold horizontal lines indicate the location of the parting plane where the two halves of the insert are split to form the cavity insert (top) and the core insert (bottom).

4.1 Parting Plane Design

Figure 4.1 Sectioned isometric view of cup assembly

Figure 4.2 Axial mold opening direction for cup

Consider next the same cavity block but with a radial mold opening direction for a portion of the cavity insert as shown in Fig. 4.3. For this design, four bold lines separate the sides from the top and bottom. Since the metal core is located inside the molded part, there is no way to remove the core other than in the part’s axial direction. The cavity insert, however, can be separated into three pieces that move along two different axes in order to remove the molded part.

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Figure 4.3 Radial mold opening direction for cup

Of these two designs, the axial mold opening direction shown in Fig. 4.2 is the simplest design and is usually preferred. However, the second design is sometimes used in practice since it allows for a more complex part design as well as more ­options in locating the parting line. For instance, the second design might be ­required if a handle were added to the cup or if it were necessary to move the parting line to a location away from the top lip. This second design is known as a “split cavity mold” and is discussed in more detail in Section 13.9.1. As another example, consider the laptop bezel shown in Fig. 3.5. There are again two potential mold opening directions. The first opening direction is in the screen’s viewing direction, as indicated by the section view shown in Fig. 4.4. In this case, the mold section is split by two horizontal lines into a cavity insert forming the outside surface of the bezel and a core insert that forms the inner surface and ribs of the bezel. When the core and cavity inserts are separated as indicated by the arrows, the molded bezel can be readily removed.

Figure 4.4 Normal mold opening direction for bezel

Alternatively, the cavity block for the tablet bezel can be split as indicated with the three vertical lines shown in Fig. 4.5. In this case, the former cavity insert is split into two pieces, resulting again in a split cavity mold design. The two halves of the former cavity insert must now be removed in oblique directions in order to remove the molded part; the mold opening direction is inclined in order to allow the mold surfaces to separate from the molded part without excessive surface friction or shearing of features on the molded part. This movement requires several addi-

4.1 Parting Plane Design

tional mold components to control the moving cavity inserts, which add significantly to the cost of mold design, manufacture, and operation.

Figure 4.5 Complex mold opening directions for bezel

4.1.2 Determine Parting Line The term “parting line” refers to the location at which the cavity insert, the core insert, and the plastic molding meet. Since the core and cavity insert meet at this location, any significant deflection of the cavity insert away from the core insert will result in a gap into which the plastic will flow and form a thin film of plastic known as “flash.” Imperfections in the core and cavity inserts at this location, for instance due to wear or improper handling, will also create gaps into which the plastic will flow. Even with new and well-crafted molds, the location of the parting line usually results in a very slight “witness line” along its length. For this reason, the parting line should be located along a bottom edge of the part, or some other nonvisual, nonfunctional edge. Consider the previous cup shown in Fig. 4.1. Placing the parting line very close to the lip as indicated by the dashed line in the left drawing of Fig. 4.6 would result in a witness line and possible flash that might make the molded cup unusable. A better location for the parting line is at the bottom of the rim as indicated in Fig. 4.2, corresponding to the parting line shown in the right drawing of Fig. 4.6.

Figure 4.6 Two parting line locations for cup

For the laptop bezel, the parting line will be located around the bottom edge of the part as shown in Fig. 4.7. It is observed that, unlike the cup, the parting line for the

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bezel is not in a single plane. Rather, the parting line follows the profile of the ­features on the side walls. This nonplanar parting line is required to fit the core insert, which hollows out the mold cavity to form the holes required for the various connectors. As will be seen in the next section, this complex parting line shape will cause a more complex parting plane.

Figure 4.7 Parting line location for bezel

4.1.3 Parting Plane Once the parting line is identified, the parting plane is projected outwards from the part so as to separate the core insert from the cavity insert. The preferred parting plane for the cup is shown in Fig. 4.8. The cavity insert will form the outer and top surfaces of the part, while the core insert will form the rim and inner surfaces.

Figure 4.8 Parting plane for cup

4.1 Parting Plane Design

For the laptop bezel, the parting line in Fig. 4.7 can be radiated outward to form the parting surface shown in Fig. 4.9. It can be observed that all of the out-of-plane features along the parting line now become complex surfaces on the parting plane. These surfaces pose two significant issues during mold operation. First, any misalignment between the sharp features on core and cavity inserts will cause wear between the sliding surfaces if not an outright impact between the leading edge of the core and the mating cavity surface. Second, the clamp tonnage exerted on the core and cavity inserts can cause the surfaces to lock together with extreme force, causing excessive stress and potential mold deformation during mold operation.

Figure 4.9 Parting plane for bezel

To avoid excessive stress, interlocking features on the parting plane should be ­inclined at least five degrees relative to the mold opening direction. The parting surface is now typically created via three-dimensional computer-aided design (“3D CAD”) using lofted surfaces. Each lofted surface blends a curved feature along the parting line to a line of corresponding width on the parting plane. The result is a surface with the needed profile at the parting line and the necessary draft down to the parting plane. The lofted surfaces are then knit together with the parting plane to provide a parting surface, as shown for the bezel in Fig. 4.10. These ­surfaces may be directly machined using computer numerically controlled (CNC) machining or via plunge electrical discharge machining (EDM) using one or more electrodes that were CNC machined.

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Figure 4.10 Modified parting surface for bezel

4.1.4 Shut-Offs Shut-offs are contact areas between the core insert and the cavity insert that separate portions of the cavity formed between the core and cavity inserts. A shut-off will need to be defined for each window or opening in the molded part. Conversely, if a part has no windows, like the cup, then no shut-offs are defined. The edge of each shut-off is also defined by a parting line, which should be located in a nonvisual area where a witness line or slight flashing would not reduce the value of the molded part. For example, the laptop bezel has one large opening above the parting plane for the display. A shut-off is necessary across the entire area of the opening. As indicated in Fig. 4.11, there are essentially two possible locations for the shut-off’s parting line, corresponding to the top and bottom of the shelf that supports the display. Either location (or even any location in between) would likely be acceptable since the entire shelf is hidden from view. If the parting line is placed at the top of the shelf as indicated at the right of Fig. 4.11, then a shut-off surface as shown in Fig. 4.12 will result.

Figure 4.11 Shut-off surface for bezel

4.2 Cavity and Core Insert Creation

Figure 4.12 Shut-off surface for bezel

„„4.2 Cavity and Core Insert Creation With the definition of the parting plane and all necessary shut-offs, the core insert and the cavity insert have been completely separated. To create the cavity and core inserts, the length, width, and height of the inserts must be defined. Cavity and core insert sizing guidelines are described that have been developed so that the length and width of the cavity and core inserts are large enough to: ƒƒenclose the cavity where the part is formed, ƒƒwithstand the forces resulting from the melt pressure exerted upon the area of the cavity, ƒƒcontain the cooling lines for removing heat from the hot polymer melt, and ƒƒcontain other components such as retaining screws, ejector pins, and others. All of these requirements suggest making the core and cavity inserts as large as possible. For smaller molded parts, increasing the sizing the core and cavity inserts may have little added cost. However, the cost of larger core and cavity inserts can become excessive with increases in the number of cavities or molded part size.

4.2.1 Height Dimension The height dimension is often determined by two requirements. First, the core and cavity insert should have enough height above and below the molded part to safely pass a cooling line. Cooling line diameters typically range from 4.76 mm (3/16”) for smaller molds to 15.88 mm (5/8”) for large molds. Generally, large inserts with larger cooling lines will provide faster and more uniform cooling as will be analyzed in Chapter 9. While cooling line design will be later discussed, the minimum

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height dimension between the molded part and the top or bottom surface of the insert is typically three times the diameter of the cooling line to avoid excessive stress as analyzed in Chapter 12. The initial height dimensions for the core and cavity inserts are shown in Fig. 4.13.

Figure 4.13 Insert height allowance

Second, the core and cavity insert should have a height that is matched with the height of available cavity and core insert retainer plates (the “A” and “B” plates). These plates are commonly available in ½" increments for mold bases designed in English units, and in 10 mm increments in metric units. As such, the insert heights should be selected up in size such that the faces of the cavity and core inserts are flush or slightly raised with respect to the “A” and “B” plates on the parting plane. It should be noted that the height of the core insert as indicated in Fig. 4.13 is not its total height but rather the height dimension from the rear surface to the parting plane. For materials procurement and cost estimation, the total height of the core insert should also include the height of the core above the parting plane plus some safety stock to allow for machining and finishing.

4.2.2 Length and Width Dimensions The length and width dimensions are similarly determined by two requirements. First, if a cooling line is needed around the exterior of the mold cavity, then the inserts should be sized large enough to accommodate such a cooling line. As for the height allowance, the length and width allowances of three cooling line dia­ meters per side are typical. Second, the width and length dimensions of the inserts should provide side walls, also known as “cheek,” that are thick enough to withstand the lateral loading of the melt pressure exerted on the side walls of the mold cavity. This requirement will become dominating (meaning that it will exceed the

4.2 Cavity and Core Insert Creation

allowance for the cooling lines) for deep parts that need tall side walls. While the structural design will be discussed in detail in Section 12.2.4, a safe guideline is that the thickness of the side wall in the length and width dimension should equal the depth of the mold cavity. Figure 4.14 demonstrates an allowance that should be added to the length and width of the mold cavity to derive the length and width of the core and cavity inserts. It can be observed that for the laptop bezel, the requirement of fitting a ­cooling line will exceed the structural requirement. For the molded cup, however, the insert length and width dimension are driven by the structural requirement.

Figure 4.14 Insert length and width allowance

4.2.3 Adjustments The core and cavity inserts can now be created with the prescribed dimensions. However, it is sometimes desirable to adjust the cavity insert dimensions to p ­ rovide a more efficient mold design. In general, the length and width dimensions of the inserts are more critical than the height dimension, since these dimensions will ƒƒdrive the size of the mold base in multicavity applications and ƒƒcontribute more to the material and machining costs. As such, these dimensions may be decreased somewhat by effective cooling and structural designs, which will be further investigated with later engineering ­analysis. Figure 4.15 provides the core and cavity inserts for the cup. Since the molded part is round, the design of the core and cavity insert may also be round. This shape provides a benefit with respect to ease of manufacturing, since both the core and cavity inserts can be turned on a lathe. While the allowances in the axial and radial dimensions are sufficient to fit cooling lines, the allowance in the radial dimension for the cavity insert may not be sufficient to withstand the pressures exerted on the side wall by the melt. The side walls of the cavity insert will tend to deflect

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outward during molding unless it is closely fit to a pocket in the mold base that provides lateral support.

Figure 4.15 Core and cavity inserts for cup

There is no fundamental requirement on the external shape of the core and cavity inserts. While the insert design in Fig. 4.15 shows round inserts, the mold design for the cup shown previously in Fig. 1.4 used square inserts. Rectangular inserts with or without filleted corners are also quite common. The design of the insert should be dictated by the shape of the molded part, the efficiency of the mold ­design, and the ease of manufacture. The core and cavity inserts for the laptop bezel are shown in Fig. 4.16. In this case, rectangular inserts are designed. The length and width dimensions of the inserts have been designed quite aggressively. While the bezel is quite shallow and the inserts are structurally adequate, the thickness of the surrounding cheek will be barely sufficient to provide cooling around the periphery of the mold cavity while also providing space for other mold components.

Figure 4.16 Core and cavity inserts for bezel

4.3 Mold Base Selection

„„4.3 Mold Base Selection After the core and cavity inserts have been initially sized, the mold layout can be further developed and the mold base selected. It is critical to order a mold base with appropriately sized plates and materials, since any mistakes in the mold base selection can consume significant time and expense. To determine the appropriate size, the mold designer must first arrange the mold cavities and provide ­allowances for the cooling and feed systems. Afterwards, the mold designer should select a standard size from available suppliers and verify suitability with the molder’s molding machine.

4.3.1 Cavity Layouts The goal of cavity layout design is to produce a mold design that is compact, is easy to manufacture, and provides molding productivity. If a single cavity mold is being designed, then the cavity is typically located in the center of the mold, though ­gating requirements may necessitate placing the mold cavity off center. For multicavity molds, there are essentially three fundamental cavity layouts: ƒƒcavities placed along one line ƒƒcavities placed in a grid, or ƒƒcavities placed around a circle. Placing all the cavities along a line, as shown in Fig. 4.17, is a simple but poor ­design. Unless the insert geometries are long and narrow, the resulting mold layout produces a mold that has a high aspect ratio. In general, the width-to-length ratio of the bounding envelope around all cavities should be kept less than 2 : 1. Higher aspect ratios will require the use of large molds that are significantly ­underutilized while at the same time producing structural loadings across the mold for which molding machine platens may not be designed. Furthermore, the use of such a line layout can result in an unbalanced feed system with uneven ­cavity filling and poor molded part quality as discussed in Chapter 6.

Figure 4.17 Series layout of cavities

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As an alternative to a linear layout of all cavities, it is common to place cavities in a grid as shown in Fig. 4.18. This design is most common for applications requiring high production volumes when the number of cavities is a multiple of 2, for example, 4, 8, 16, 32, etc. There are two primary benefits to a grid layout. First, the grid layout will result in a compact mold with an acceptable aspect ratio. Second, the grid layout lends itself well to naturally balanced feed system layouts as discussed in Chapter 6.

Figure 4.18 Grid layout of cavities

While the grid layout is compact and very common, it can result in a feed system design with multiple branches. To reduce the feed system complexity and ensure more balanced melt filling, a circular layout is sometimes used when the molded parts are relatively small or when the number of mold cavities is relatively low, for example, eight or less. Figure 4.19 shows one such layout in which all the cavities are provided at an equal distance from the center of the mold. The primary dis­ advantage is that such a circular layout requires a larger mold surface area than the previously discussed grid layouts.

Figure 4.19 Circular layout of cavities

4.3 Mold Base Selection

While the previously discussed layouts are the most common, mold designers can utilize other layout designs, including combinations of the above layouts. For ­example, Fig. 4.20 shows a combination of a line layout plus a circular layout. The resulting layout is a very compact and balanced design for six cavities. Again, the designer should develop the layout that is best for the application’s geometry and requirements.

Figure 4.20 Hybrid layout of cavities

4.3.2 Mold Base Sizing The size of the mold base is determined primarily by the area required to accommodate all the cavity inserts per the designed cavity layout. A primary issue, however, is the potential for conflict between the placements of the cavities and other mold components (such as leader pins, guide bushings, and others). Furthermore, there is the potential for conflict between cavity support systems (such as cooling lines, ejector pins, support pillars, etc.) and other mold components (such as leader pins, guide bushings, and others). Due to these conflicts, mold bases are often sized larger than what would first be considered. The shaded area in Fig. 4.21 represents the usable area of the parting plane into which the core and cavity inserts can be placed. Ejector return pins are located to the left and right of this area, while guide pins and socket head cap screws are ­located above this area. A dimensional allowance equal to at least one-half of each component’s diameters is provided between the mold cavity and the surrounding components to avoid excessive stress during the mold’s operation.

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Figure 4.21 Usable parting plane area

Given the cavity layout and its geometric envelope, a mold base with a feasible length and width is selected. Standard mold bases are widely available from 200 mm up to 1000 mm on a side. When specifying a mold base, it is also necessary to specify the height of the “A” plate, the height of the “B” plate, the height of the support plate, “S,” and the distance of the ejector travel, “E,” as shown in Fig. 4.22. The total stack height is defined as the distance from the bottom of the rear clamp plate to the top of the top clamp plate.

Figure 4.22 Height dimensions to specify

4.3 Mold Base Selection

With respect to mold base selection, the height of the “A” and “B” plates are respectively matched to the height of the cavity and core inserts as previously discussed. The height of the support plate, “S,” is normally determined from the mold base supplier based on the height of the “A” and “B” plates, though the height of the support plate can be special ordered to varying dimensions. The travel of the ­ejector plate should be selected to eject the part from the mold. Often, the ejector travel is set to be equal to the depth of the molded part. From the ejector travel, the height of the ejector housing, dimension “C,” is assigned by the mold base ­supplier. When selecting a mold base, it is also necessary to specify an orifice diameter for the sprue, which is not shown in Fig. 4.22. This dimension is of lesser importance since the sprue bushing may be replaced or machined, or the molding machine nozzle changed, to match the sprue to the nozzle as later discussed in Chapter 6.

4.3.3 Molding Machine Compatibility When selecting a mold base, the mold designer should verify that the mold will fit in the available molding machine(s). There are many requirements that should be considered when matching a mold to a molding machine. First, the mold must physically fit in the machine. Perhaps the most common limitation is that the mold will not fit between the tie bars. The tie bar spacing is easily measurable on an available molding machine or can be checked in a machine drawing for a potential molding machine. For instance, Fig. 4.23 shows the tie bar spacing and bolt ­pattern for a Battenfeld HM320 molding machine. It can be viewed that the horizontal tie bar spacing is 800 mm, and that the vertical tie bar spacing is 630 mm. This means that the maximum mold width, including cooling plugs, hot runner connectors, etc., is 800 mm (some relatively small clearance between the mold and the tie bars to provide for mold insertion).

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Figure 4.23 Typical tie bar and bolt pattern (dimensions in mm)

A cross-section view of the same machine platens is shown in Fig. 4.24 with the same orientation as shown in Fig. 1.1. In this view, the nozzle of the molding machine enters the stationary platen on the right side of the drawing. The machine’s moving platen and the ejector unit are on the left side of the drawing. For the mold to be operable in the machine, the mold height must be greater than the indicated A dimension and smaller than the indicated B dimension, or between 350 and 800 mm for this machine. If the mold is smaller than 350 mm, then the molding machine platen cannot fully close the mold and build clamp tonnage. If the mold is larger than 800 mm, then the mold will not fit between the two platens when the moving mold platen is fully open.

Figure 4.24 Minimum and maximum daylight (dimensions in mm)

4.3 Mold Base Selection

Even if the mold fits in the molding machine, the molding machine may still not be operable with the mold. For instance, the injection unit of the molding machine must have sufficient shot volume and provide enough melt pressure to fill the mold cavity with the polymer melt. On the other hand, if the injection unit has too large a shot size, then the control of the injection velocity may be limited and the melt may degrade in the barrel of the molding machine. For example, for the Battenfeld HM320, the maximum shot volume is 490 cc. To provide melt homogeneity without degradation, this machine is ideally suited for molds requiring a shot volume between 120 cc and 250 cc. The molding machine must also provide sufficient clamp tonnage to hold the two halves of the mold together when pressurizing the polymer melt. For this machine, the clamp tonnage is 3200 kN, which is equal to 326 metric tons, 360 English tons, or 720,000 pounds. If the molding machine does not provide sufficient clamp ­tonnage, then the mold will open during operation and the melt will flow across the parting plane and shut-offs. If the molding machine provides too much tonnage to a very undersized mold, however, the mold may be damaged by the imposed compressive stresses.

4.3.4 Mold Base Suppliers The development of standardized mold base designs is considered a significant advance in the history of the plastics molding industry. A majority of mold makers in the U.S. use standard mold bases to reduce the time and expense of creating molds. Furthermore, mold maintenance is simplified through the availability of standard mold components that are replaceable at the molder’s facility. It is noted that many mold makers do not use mold bases for various reasons. Mold bases for very large parts, such as automobile body panels, may not be available as a standard product and so may require custom design and manufacture. Some mold makers believe that standard mold bases are inferior in quality and strive to provide a better mold with higher quality or lower lifetime cost through the development of custom designs with proprietary components. At the other end of the cost scale, some foreign mold makers can produce a simple but fully functional mold for less cost than just the standard mold base could be purchased in the United States. There are numerous mold base suppliers from which mold bases and mold components can be purchased. When selecting a mold base, the mold designer should consider: ƒƒThe range of mold base sizes and materials that are available. Not only should mold bases of varying plate lengths, widths, and heights be available, but these mold bases should also be available in different types of materials.

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ƒƒThe portfolio of mold components that can also be purchased. The mold base supplier should be able to provide insert materials, ejector pins, cooling acces­ sories, etc. ƒƒThe native system of units that the mold base was designed and the quality of the associated drawings. If the mold designer prefers U.S. customary or metric units, the mold base drawings should reflect the same system of units including the use of round numbers, fractions, and so on. Drawings should fully detail the design of the various mold components and, when appropriate, document their customization and operation. ƒƒThe inventory availability and delivery terms. Standard “quick ship” mold components should be in the supplier’s inventory. Customized mold bases with varying plate dimen­sions and material specifications should be custom manufactured and shipped within one week. Orders that are placed before noon should be shipped the same day and no later than the next day. ƒƒThe quality of the supplied mold bases. All mold plates should be supplied finish ground, heat treated, and ready for machining at the mold maker. Guide pins, ejector pins, and other mold components should be finished, hardened, and/or coated as appropriate to ensure low wear. ƒƒThe previous experience with the mold base supplier. If a company or mold ­designer has past favorable experiences with a supplier, then there may be risk or a significant learning curve associated with switching suppliers. ƒƒThe pricing should be competitive with commodity material prices. Clearly the mold base supplier adds value to the raw materials included in the mold base and is entitled to recover their costs and reasonable profit. Still, the mold designer must compare the strengths and weaknesses of various mold base suppliers to determine whether to solely source from one mold base supplier or choose from a few qualified suppliers.

„„4.4 Material Selection As part of the mold base design and procurement, the materials for the inserts and other components must also be selected. Just as there are many different plastics suitable for injection molding, there are many ferrous and nonferrous metals that are suitable for use in injection molds. Some of the more common materials and their properties are provided in A ­ ppendix B. Of all these materials, AISI P20 remains the most common due to its widespread familiarity and favorable combination of properties. However, P20 is sometimes improperly specified in many molding applications since other metals would pro-

4.4 Material Selection

vide better performance, lower mold-making cost, or lower injection molding costs. In the last two decades, high-strength aluminum alloys have been adopted in many mold design applications while 3D printed polymers have been used for inserts albeit with limited longevity [1–3]. Figure 4.25 provides a relatively simple flow chart showing the primary decisions. Class 101 production molds require the use of carbon/tool steels, which are also preferred for applications with moderate to high molding pressures. Once the de­ cision to use a steel has been made, the steel selection is often dominated by concerns related to wear and abrasion resistance that is needed for the mold’s use with fiber-filled resins. If abrasion resistance is not a particular concern, then ­corrosion resistance also guides the material selection to a stainless steel such as SS420. Otherwise, tool steels such as P20 or S7, which are relatively easy to machine and polish, are preferred.

Figure 4.25 Basic material selection flow chart

If a molding application does not present challenging pressure and wear requirements, then nonferrous metals are often considered, especially for short-run or intermediate production quantities. Aircraft grade aluminum 7075-T6 and specialty grades developed for molds (such as Alcoa QC10, Aleris Hokotol, and Vista Duramold) provide high machining and molding productivity with reasonable strength. Commodity 6061-T6 is a lower-cost option that provides somewhat lower strength but higher weldability and corrosion resistance. For rapid prototyping ­applications with very low production quantities and lower molding pressures, ­polymer inserts produced from a polyjet printing or fused deposition modeling process are feasible. Some of the important properties and trade-offs of the various mold materials are discussed next.

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4.4.1 Strength vs. Heat Transfer Strength is typically characterized by the ultimate stress that a material can endure prior to failure, or by the yield stress that can be applied to a material without causing permanent deformation. For injection molds, however, neither of these properties should be utilized. Instead, the fatigue strength (also known as the endurance limit stress) is the amount of stress that can be cyclically applied without causing failure. One issue in structural design is that the fatigue behavior differs by material. For most steels, the fatigue strength is approximately one-half of the yield stress. However, aluminum and many other materials do not have an endurance limit. Instead, these materials will eventually fail after continuous cycling, regardless of how little stress is applied. For this reason, the allowable stress is defined after a certain number of cycles (for example, ten million). Reasonable estimates of the allowable stress for fatigue are provided in Appendix B for different candidate materials; these values are later used for structural analysis in Chapter 12. Insomuch as mold makers and molders seek different material properties, there is no one perfect mold material. A common trade-off between the fatigue strength and the thermal conductivity is shown in Fig. 4.26 for various materials. In general, the materials that have the highest strength (A2, D2, H13, and P20) have the lowest heat transfer. Conversely, the materials with the highest heat transfer (­aluminum and copper alloys such as C-18200) have the lowest strength. No ­material exists that has a very high fatigue limit stress and a very high thermal diffusivity. P20, the most common of all mold materials, has good fatigue strength but low thermal diffusivity, suggesting that the mold’s performance may be improved by using other mold materials in some molding applications.

Figure 4.26 Structural and thermal performance of some mold materials

4.4 Material Selection

4.4.2 Hardness vs. Machinability To withstand wear and abrasion, it is desirable that the mold materials have very high hardness [4]. There are many ways to measure hardness, with one of the most common being the Brinell Hardness test. In this test, a carbide ball with diameter D equal to 10 mm is pushed into the test material with a force, F, of 29,500 N (3000 kg force). The diameter, d, of the resulting indentation is measured after which the Brinell Hardness, BHN, is calculated as BHN =

(

2F

p D D - D2 - d 2

)

(4.1)

Other hardness measures (including Rockwell and Vickers) are defined in a similar fashion but vary with respect to load and geometry for applicability to different materials. Generally, the hardness of a material is related to the moduli and compressive strength of the test material [5]. The hardness values for various mold materials are provided in Appendix B. (Because of the variance of the material properties, the Brinell hardness test may not be suitable for very soft or very hard materials. For this reason, some of the Brinell hardness values in Appendix B are derived from other hardness tests as appropriate.) As the material hardness increases, the materials will tend to become more d ­ ifficult to machine. Harder cutting tools and lower cutting speeds and feed rates become necessary. The volumetric machining rate can be computed from the recommended cutting speeds and feed rates for various mold materials assuming a carbide cutter is used [6]. The resulting machining rate is plotted as a function of Brinell hardness in Fig. 4.27.

Figure 4.27 Machining and wear performance of some mold materials

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The data in Fig. 4.27 indicate that the materials with high hardness have low machining rates, while materials with high machining rates have low hardness. For this reason, very hard materials such as A2, D2, and H13 should only be used when required for molding abrasive plastics that would quickly abrade the surfaces of softer mold materials. Due to their very high machining speeds, aluminum alloys can be used to quickly and economically produce molds but are rarely used when molding at moderate melt pressures (100 MPa or greater) or with even slightly abrasive plastics (such as carbon filled).

4.4.3 Material Summary The properties plotted in Figs. 4.26 and 4.27 suggest that there are inherent tradeoffs between the cost to make the mold and the cost to use the mold. Aluminum alloys provide very low mold-making costs and low operating costs. For this reason, the aluminum alloys should be seriously considered when a molding application does not require high strength or hardness. Copper alloys, such as C18200, provide even lower mold operating costs but have higher mold-making cost due to higher materials cost coupled with lower machining rates. Accordingly, copper ­alloys are good candidates for molding applications with high production volumes that require moderate strength and hardness. There are a few materials listed in Appendix B that merit additional discussion. Stainless steel, SS420, has a much higher cost than molds made of P20 due to SS420’s lower thermal diffusivity, but it should be used when corrosion resistance is needed. H13 has a very high mold-making cost due to its very high hardness and very low machining speed, but it is commonly used when abrasion resistance is desired. At the other end of the cost spectrum, aluminum 6061 is a common material that is used for prototype molds given its reasonable balance of properties, low procurement cost, and machinability. Direct 3D printing of inserts with “digital” ABS and Ultem (PEI) are becoming increasingly common, but as their properties plotted in Figs. 4.26 and 4.27 suggest, they are intolerable except for applications with very low production quantities. The data in Appendix B and the previous plots provide quantitative and qualitative compari­sons of common mold materials. The mold designer must consider the ­requirements placed on the mold in a given molding application and weigh the economic, structural, thermal, and other requirements. While one material such as P20 may have always worked well for a given mold designer, there is the possibility that significant improvements in mold performance and profitability could be realized by utilizing other mold materials. Table 4.1 recommends some of the commonly used mold materials according to the application requirements. P20 is a fine material and highly suitable for many

4.4 Material Selection

molding applications. However, other materials are better in diverse applications. In Table 4.1, all the recommendations pertain specifically to materials for the core and cavity inserts. Standard mold bases are not available in all these materials; mold bases are typically available in 1045, 4140, or P20 steel, while aluminum and stainless steel mold bases are available for lower pressure and corrosive appli­ cations, respectively. Plain 1045 steel is often chosen for molding applications with lower production volumes and moderate molding pressures. For higher production volumes and molding pressures, the alloyed steels 4140 and P20 are usually preferred. Table 4.1 Common Mold Materials by Application Low number of ­cycles (ncycles < 10,000)

Moderate number of cycles

High number of ­cycles (ncycles > 1,000,000)

Nonabrasive melt with low molding pressures

Al alloys

Al or Cu alloys

Cu alloys, P20, SS420

Slightly abrasive melt or moderate ­molding pressures

Al or Cu alloys or 1045

Cu alloys, P20, 4140, S7

SS420, S7, D2, A6

Highly abrasive melt

P20, S7

D2, A6, H13

H13

High molding pressures

1045, 4140, P20

P20, S7

D2, A6

Highly corrosive melt

P20, SS420

SS420

SS420

4.4.4 Surface Treatments When specifying the material for various mold components, the mold designer should also consider the uniformness of the stock material as well as its ability to be finished and treated. Most metals are cast and subsequently rolled/formed/slit to their supplied shape. The resulting grain structure and properties are a complex function of not only the constitutive alloying elements but also the thermal and structural history during processing. Mold designers, mold makers, and end-users should be aware that there are many issues such as porosity (voids), contaminants, inhomogeneity, and residual stress that may impact the quality of the machined mold. For these reasons, it is recommended that the mold components be machined from annealed or normalized steel with minimal residual stress and uniform properties. These treatments can provide for lower hardness and faster machining. After the mold has been machined and finished, the finished component may again be ­annealed to verify dimensional stability and then carburized (also known as case hardening) to improve surface hardness by increasing the carbon content at the surface. Carburizing is performed at 900–950°C, and can cause dimensional changes in machined components. There are many alternative surface treatments

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that are available including nitriding, boriding, plating, vapor deposition, anodizing, and others. Table 4.2 provides a comparison of several surface treatments that may be selected for specific application-specific purposes including increased hardness, reduced coefficient of friction (COF), hardness, improved corrosion resistance, and others [7, 8]. Table 4.2 Common Surface Treatments Treatment

Properties

Purpose & notes

Diamond chrome plating at ~60°C

50 µm thick; 85 RC; 0.15 COF

Super hard low-friction coating but requires anode

Hard chrome plating at ~70°C

Up to 0.5 mm thick; 72 RC; 0.2 COF

Hard coatings; can be thickened for repair of ­surfaces but requires anode

Nickel-boron nitriding at ~600°C

10 µm thick; 67 RC; 0.05 COF

Thin layer for low friction and excellent abrasion resistance

Electroless nickel coating at ~80°C

50 µm thick; 62 RC; 0.4 COF

Lower-cost alternative to hard chrome for ­corrosion and abrasion resistance

Nickel-PTFE coating

45 RC; 0.10 COF

Excellent release for deep ribs, cores with low draft, textured surfaces, and tacky polymers

Carburizing at ~900°C

60 RC surface

Used with low carbon steels to “case harden” ­surfaces to depth of 6 mm

Physical vapor deposition (PVD) of titanium nitride

5 µm thick; 70 RC; 0.1 COF

Very thin but hard coatings applied at ~500°C in reduced pressure nitrogen atmosphere

Thermoreactive diffusion (TRD) of silicon carbide

5 µm thick; 70 RC; 0.1 COF

Very thin but hard coatings applied by molten salts at high temperatures

Aluminum anodizing

50 µm thick; 65 RC

Provides hard, corrosion and wear resistant ­surface for aluminum

In most applications, mold makers will outsource the mold components to be treated by service providers that specialize in surface treatments. These treatments will increase the initial purchase cost of the mold but can greatly increase the longevity and reduce associated maintenance costs, especially when processing abrasive resins or long production runs. Molders also often rely on surface treatments to resolve issues such as improving lubricity of mold surfaces to ease part ejection, reduce wear between sliding components, repair scratches, or ­improve the surface of welded sections. Diffusion processes (e. g., carburizing, ­nitriding) do not grow the thickness of the mold surface as opposed to coatings, which may add substantial mass. Mold designers and mold makers should take note of the thickness and consistency of the applied coating, so that cavity wall thicknesses are designed to provide an allowance for its thickness as appropriate. Fortunately, many coatings may be stripped and reapplied if repair or alterations in surface properties are needed.

4.5 Chapter Review

„„4.5 Chapter Review The mold layout design process includes the examination of the part geometry to be molded to identify the parting line, parting plane, and shut-offs. The core and cavity inserts are then sized and located relative to each other. Afterwards, a ­suitable mold base is chosen or designed that can efficiently hold and support the core and cavity inserts. The mold layout process finishes with the selection of the materials used for the mold base as well as the core and cavity inserts. In many mold-making companies, these materials are immediately ordered concurrently with the detailed analysis and design of the mold subsystems. After reading this chapter, you should understand: ƒƒHow to identify the mold opening direction(s) and parting line(s) for a molded part, ƒƒHow to design a parting plane and shut-offs to separate the core insert from the cavity insert, ƒƒHow to size the length, width, and height dimensions for the core and cavity ­inserts, ƒƒThe advantages and disadvantages of different cavity layouts, ƒƒHow to lay out a given number of mold cavities, ƒƒHow to size a mold base for a given mold cavity layout, ƒƒHow to verify that a mold is appropriate for a molding machine, and ƒƒProperties and selection of mold materials and surface treatments. Figure 4.28 provides a mold design checklist for high performance, standard, and basic molds; the details are presented throughout the book. The next chapter ­examines the mold cavity filling process, which is required to 1) verify that the part design can be produced at available melt pressures, and 2) estimate the loading that will be placed on the mold components. Afterwards, the analysis and design of the feed system will be addressed.

105

Highest Volume/Quality

Standard Volume/Quality

Basic/Prototyping

      

      

              

Standard Design

Stage Start of Design

      

Performance Design

4 Mold Layout Design

Basic

106

                      



Inspecon Item ( means important,  means oponal) Number of cavies & type of mold carefully selected Mold specificaon/purchase agreement completed Part design reviewed: thickness, gang, dra, undercuts, etc. Shrinkage guidance provided by molder or end-user Mold design standards (DIN, JIS, inch) specified Supply chain preferences specified Target producon quanty/rate specified P20 (DME#2, DIN 1.2311) or stainless mold base H13 or similarly hard cavity & core insert materials Hot runner system without sub-runners Electrical connectors (male) at top of mold Limit switches on slides/pulls for confirming posion Mulple knock-out locaons spanning ejector plate Early return ejectors (posive return) Wear plates on four sides of mold Parng line interlocks with cavity/core inserts Cavity pressure/temperature transducers Thermal insulaon to minimize heat transfer to mold/platens Springs on slides to prevent accidental movement Greased ways with fi•ngs on inaccessible slides Parts/mold designed for robot interface Mold cycle counter

                          

Mold designed for fully automac operaon  Uniform cooling provided to all core & cavity inserts  Cooling "O" rings/seals recessed to avoid damage  Ejector plates guided by pins/bushings Leader pins/bushings engage prior to other components  Venng provided at end of flow and in blind pockets  Dra used to assist ejecon, esp. deep cores & textures  Mulple ejectors to push on sff secons of parts  Ejector & core pins through-hardened  Parng line interlocks  Dowel pins used for mang with locaon fits Horn pins less than 20 degrees to avoid wear Mold guide pins/bushings to engage prior to mold closure  All bolts, pins, plates, etc. standard stock parts when possible  Recessed water connectors at bo¡om of mold  Date/shi/logo inserts for run idenficaon  Mold-actuated lis/slides used instead of core pulls  Mold inserts for runners or other high-wear areas Mold cavity surfaces hardened & heat treated Hardened stainless used for corrosive resins Chrome or nitrided surface treatements used for abrasive resins Mulple support pillars with preload Early return ejectors (springs) AISI4130 (DME #2 , DIN1.1312) mold base P20 or similarly hard cavity & core materials Hot sprue bushing or hot runner system with sub-runners  Runner shut-offs for changing flow pa¡erns



 Sucker pin on cold runners & near gates  Steel/aluminum/no mold base (or master unit die)  Cavity & core from 6160 or other prototyping materials

Figure 4.28 Mold layout design checklist

4.6 References

„„4.6 References [1] Sachs, E., et al., Production of injection molding tooling with conformal cooling channels using the three dimensional printing process, Polym. Eng. Sci. (2000) 40(5): pp. 1232–1247 [2] Kovács, J. G., et al., Thermal simulations and measurements for rapid tool inserts in injection molding applications, Appl. Therm. Eng. (2015) 85: pp. 44–51 [3] Sachs, E., et al., Three dimensional printing: rapid tooling and prototypes directly from a CAD model, J. Manuf. Sci. Eng. (1992) 114(4): pp. 481–488 [4] Roberts-Austen, W., On the Hardening and Tempering of Steel, Sci. Am. (1889) 28: pp. 11520–11522 [5] Pavlina, E. and C. Van Tyne, Correlation of yield strength and tensile strength with hardness for steels, J. Mater. Eng. Perform. (2008) 17(6): pp. 888–893 [6] Oberg, E., et al., Machinery’s Handbook, Green, R. E., (Ed.), 24th ed. (1992) [7] Surface Treatments Applicable for Molds, Kata Gijutsu (1989) 10 [8] Coating Selector Sheet, Richter Precision, Inc. (2008) 7(8): p. 8

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5

Cavity Filling Analysis and Design

For an acceptable molded part to be produced, the polymer melt must completely fill the mold cavity. Accordingly, the wall thickness of the molded part and the gating locations must be specified such that the melt is able to traverse from the gates to the edge of the cavity. Mold filling analysis is used to ensure that the melt can not only fill the mold at achievable molding pressures, but fill the mold as intended to achieve the desired quality.

„„5.1 Overview Cavity filling analysis may be performed for a variety of purposes. On the most basic level, mold filling analysis is useful to ensure that the mold cavity can be filled with the plastic melt given the melt pressure that can be delivered by the molding machine. Typically, the melt pressure required to fill the cavity is less than 100 MPa (about 15,000 psi) even though most modern machines can supply twice this amount. This safety margin between the required and available melt pressures provides an allowance for the pressure drop in the feed system, and also ensures that the mold can be filled given possible variances in the material properties or molding process. Cavity filling analysis is also performed to ensure that the filling pressures are not too low, since very low melt pressures are indicative of a poor molded part design or improper processing conditions. Excessively thick wall sections will result in low pressures, excessive material costs, and extended cycle times. In such cases, the nominal wall thickness should be decreased and ribs or other features used to provide the necessary strength and stiffness. In some cases, very low melt pressures can indicate improper filling time, mold temperature, or melt temperature. These processing conditions should be adjusted to reduce the processing time and cost at the expense of higher melt pressures.

110

5 Cavity Filling Analysis and Design

On a more advanced level, cavity filling analysis is useful to predict the melt front advancement in the cavity and identify the location of knit-lines, end of fill, and other phenomena before the mold is manufactured and tested. These results can be used to adjust the gating location(s), type of gate, cavity thicknesses, ejector locations, vent locations, and other design parameters. While modern computer mold-filling simulations can provide detailed results for very complex cavity geometries, “lay-flat” cavity filling analysis remains extremely useful. This manual analysis provides a means by which the mold designer can understand the primary flow behavior and develop useful estimations to determine the mold design and process conditions or validate computer simulation results. To perform this analysis, the mold design engineer must understand the fundamentals of melt rheology and the governing equations for flow. Afterwards, a methodology for cavity filling analysis will be presented and validated.

„„5.2 Objectives in Cavity Filling 5.2.1 Complete Filling of Mold Cavities The part and mold design must be developed such that the mold cavity can be completely filled by the polymer melt at workable melt pressures. For this reason, filling analysis of the mold cavity should be performed to verify the part wall thickness for a given material and assist in the gate selection and processing conditions. Modern molding machines can typically deliver injection pressures of approximately 200 MPa (30,000 psi). However, a lower melt pressure should be assumed for filling the cavity to allow for: ƒƒa lower required clamp tonnage, ƒƒreasonable pressure drop in the feed system, and ƒƒa factor of safety for errors in assumptions. In practice, a melt pressure of 100 MPa is commonly specified as a maximum limit for the cavity filling pressure. The maximum cavity pressure may be specified higher if the molding machine is known to have a very high injection pressure, or if the mold’s feed system is purposefully designed to incur a small pressure drop (via a hot runner system for example). In the event that a mold is very difficult to fill, molders will generally try to compensate by increasing the mold and melt ­temperatures, enlarging the runner diameters, trying lower-viscosity plastics, and finally changing the wall thickness of the mold cavity. Conversely, if a mold is very easy to fill, molders will generally reduce the mold and melt temperatures while increasing the injection velocity to shorten the cycle time.

5.2 Objectives in Cavity Filling

5.2.2 Avoid Uneven Filling or Over-Packing During mold filling, the polymer melt will tend to flow radially throughout the ­cavity from the point where it is injected. In general, the mold should be designed such that the polymer melt reaches the edges of the mold cavity furthest from the gate at approximately the same time. Such even filling allows for more uniform and lower melt pressures throughout the mold cavity. If one portion of the mold fills substantially earlier than other portions of the mold, then the melt in the filled portions will stagnate with potentially serious con­ sequences. Figure 5.1 shows the filling contours from the analysis of the bezel in which the plastic melt is injected at two locations. Each contour represents the ­location of the melt front at different moments in time. As can be observed, the plastic melt flows radially out of the gate and is then constrained by the side walls. The polymer melt then flows along the side walls, then across the top and bottom walls of the bezel.

Figure 5.1 Melt front progression of laptop bezel

During the majority of the filling, the flow rates to the upper and lower halves of the part are equal. However, the plastic was injected at gates located slightly toward the lower portion of the part, such that the bottom portion of the part fills before the upper portion of the part. When the two melt fronts meet at the bottom center, very little additional plastic melt can be forced into the lower portion of the part. The flow to the bottom portion of the part stagnates, causing a surge in the melt flow to the upper portion of the part. The resulting defects can possibly include: ƒƒExcessive cavity filling pressures required to fill the mold, excessive clamp tonnage, and flashing; ƒƒInability to fill the mold cavity (short shot); ƒƒHigh residual stress and warpage; and ƒƒMelt fracture, jetting, hesitation, or other aesthetic defects associated with changes in melt velocity during mold filling.

111

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5 Cavity Filling Analysis and Design

To avoid or resolve these issues, the mold design should consider the type and location of gate, the layout and sizing of the feed system and the nominal thickness of the mold cavities, and use slight changes in the wall thicknesses to purposefully direct the flow in the mold cavity.

5.2.3 Control the Melt Flow Even when the mold fills uniformly, cavity filling analysis may be used to maximize the quality of the part. For instance, it is sometimes desirable to control the melt front advancement such that knit-lines are placed in areas of the part that are less important with respect to aesthetics or structural integrity. Similarly, cavity filling analysis may be performed to predict the last area to fill so that vents and/or ejector pins are provided for the displaced gas to exit the mold. For anisotropic plastics (such as glass-filled materials), cavity filling analysis and design can be performed to control the flow direction to affect the molded in orientation, strength, or shrinkage.

„„5.3 Viscous Flow 5.3.1 Shear Stress, Shear Rate, and Viscosity To analyze the polymer flow, it is necessary to understand the relationship between the shear stress, shear rate, and viscosity [1]. The shear stress, t, is a ­measure of how much force per unit area is being exerted by the fluid as it flows. The shear rate, g , is a measure of the rate at which the melt velocity changes. The shear stress is related to the shear rate through the viscosity, h, which is a measure of the fluid’s resistance to flow. t = hg (5.1)

Consider the flow between a moving plate and a stationary plate shown in Fig. 5.2. Assuming that the flow is fully developed and does not slip at the walls, then a linear velocity profile is observed across the fluid with the velocity, v, equal to zero at the stationary wall and equal to v at the moving plate.

5.3 Viscous Flow

Figure 5.2 Flow between two parallel plates

For a flow between one stationary and one moving parallel plate, the shear rate is defined as the change in the velocity through the thickness, or g =

dv v = (5.2) dz h Example: Compute the shear rate of a polymer melt being pulled at 100 mm/s by a moving plate 1.5 mm above a stationary plate. If the viscosity is 100 Pa·s, estimate the resulting shear stress. For a plate with a length of 200 mm and a width of 100 mm, compute the lateral force required to continue moving the plate at 100 mm/s. The shear rate is

g =

v 100 mm s = = 67 1 s 1.5 mm h

With a viscosity of 100 Pa·s, the shear stress in the melt is

t = hg = 100 Pa × s × 67 s-1 = 6700 Pa If the plate is 0.2 m in length by 0.1 m in width, the lateral force on the wall of the moving plate is

F = t A = 6700 Pa × (0.2 m × 0.1 m) = 135 N 

5.3.2 Pressure Drop The pressure drop caused by the flow of the polymer melt in a channel can be analyzed by considering the equation of motion. For steady flow, the sum of the forces must equal to zero.

å F = 0 (5.3)

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Consider the forces acting laterally on the flow in a rectangular channel as shown in Fig. 5.3. As the flow moves from left to right, there will be a pressure drop along the flow with P1 being greater than P2 . This pressure drop is being caused by the viscous flow in the channel that is generating shear stresses, t, against the side walls.

Figure 5.3 Pressures and shear stresses in channel

There are two forces on the polymer melt that must balance. First, there is the force due to the pressure drop, FDP , across the length of the melt flow. FDP = P2 (WH ) - P1 (WH ) = ( P2 - P1 )(WH )

(5.4)

Second, there is the force due to the shear stresses, Ft, acting on the top and bottom surfaces along the length. Ft = 2t ( L2 - L1 )W (5.5)

Equating the force due to the pressure drop and the force due to the shear stresses provides

( P2 - P1 )(WH ) = 2t ( L2 - L1 )W (5.6) Let dP/dL be the pressure drop per unit length. Simplifying then provides the following result. dP 2t (5.7) = dL H

5.3 Viscous Flow

Example: A polymer melt has a wall shear stress of 13,000 Pa in a mold cavity that is 1.5 mm thick. Estimate the pressure drop across a cavity that is 200 mm in length. The pressure drop per unit length is

dP 2t 2 ´ 13,000 Pa MPa = = = 17.3 0.0015 m m dL H For a cavity with a length of 200 mm, the pressure drop is

DP =

MPa dP × L = 17.3 × 0.2 m = 3.5 MPa dL m



To compute pressure drop as a function of the viscosity, it is necessary to define the viscosity as a function of the shear rate and temperature so that the shear stresses can be computed.

5.3.3 Rheological Behavior The term “rheology” refers to the study of deformation and flow of matter [2, 3]. The term “viscosity” refers to the resistance of a fluid as it deforms under shear stresses and is defined according to Eq. 5.1 as the shear stress divided by the shear rate. The viscosity behavior of polymer melts can be extremely complex, much more so than is often appreciated when practitioners contemplate melt flow indices (MFI). The melt flow index, defined by ASTM D1238, measures how many grams of polymer flow through a capillary of a specified length and diameter given a specified amount of pressure and time. A higher melt flow index usually corresponds to a lower viscosity and improved ease of processing. However, the MFI is a single point estimate of the viscosity and is not indicative of the material behavior across the broad range of shear rates, temperatures, and pressures when it is being molded [4]. For this reason, better viscosity models are applied to injection molding. The Cross-WLF model [5] is widely known as a capable model of the melt viscosity, h, as a function of shear rate, g , temperature, T, and pressure, P. h (g ,T , P ) =

h0 (T , P ) æ h g ö1-n 1 + çç 0* ÷÷÷ èç t ø÷

(5.8)

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In this model, h0 is the “Newtonian limit” in which the viscosity approaches a constant at very low shear rates, t* is a critical stress level at which the viscosity transitions from the Newtonian limit to the power law regime, and h is the power law index in the high shear rate regime. The form of the Cross model is readily understandable since these three parameters, h0 , t*, and h , can be estimated ­directly from a log-log plot of the viscosity as a function of shear rate as shown in Fig. 5.4.

Figure 5.4 Cross model with illustrated coefficients

In the Cross model, the zero shear viscosity, h0 , is itself a function of temperature, T, and pressure, P. The temperature dependence can take many forms, including the Arrhenius relation [6]. Another common model was first described by Williams, Landel, and Ferry (WLF) [7] that includes pressure dependence through the shifting of the glass transition temperature, T*. æ A (T - T * ) ö÷÷ h0 (T , p) = D1 exp ççç- 1 * ÷ èç A2 + (T - T ) ø÷÷

T > T * (5.9)

T *( p) = D2 + D3 p (5.10) A2 = A3 + D3 p (5.11)

h0 (T , p) = ¥

T < T * (5.12)

The model parameters (n, t*, D1, D2, D3, A1, A3) are typically determined by curve fitting experimental shear-viscosity data taken by a capillary rheometer at shear rates from 1 to 10,000 1/s. The material properties for many thousands of plastic resins have been characterized, and the Cross-WLF model coefficients for some

5.3 Viscous Flow

representative materials are provided in Appendix A. The Cross-WLF viscosity model for a medium viscosity PC is plotted as a function of shear rate for three different temperatures in Fig. 5.5. The viscosity exhibits a Newtonian plateau for shear rates up to 100 1/s, then ­transitions into a power law regime. For a melt ­temperature of 280°C, the viscosity decreases from 350 Pa · s at 100 1/s to 80 Pa · s at 10,000 1/s. Since the viscosity is strongly dependent on the shear rate, esti­ mation of the filling time, melt velocity, and shear rate are vital to the analysis predictions. The viscosity is also a strong function of temperature, with the zero shear viscosity increasing from 250 Pa · s at 290°C to 660 Pa · s at 270°C. Thus, knowledge of the processing temperature is also important to predicting the melt flow and pressure. While the Cross-WLF model is a very adept model and is commonly used in numerical simulation, it is not as useful in manual filling analysis. The issue is that it is difficult to operate and not amenable to analytical solution of the pressure as a function of the melt flow rate. For this reason, several other viscosity models are commonly used that have relatively simple analytical solutions.

Figure 5.5 Viscosity behavior for a polycarbonate (PC) resin

5.3.4 Newtonian Model Newton’s law of viscosity states that the shear stress, t, between parallel layers of flow is proportional to the shear rate, g . t = mg (5.13)

where the coefficient m is the apparent viscosity and is assumed constant for “­Newtonian” fluids.

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Figure 5.6 compares the Newtonian model against the non-Newtonian behavior provided by the Cross-WLF model for a medium viscosity PC at 280°C. As pre­ viously stated, the polymer melt is known to be non-Newtonian. For this reason, the Newtonian model provides an exact estimate for the viscosity only at a shear rate of 7000 1/s. The Newtonian model overestimates the viscosity at higher shear rates and underestimates the viscosity at lower shear rates. Even so, the New­tonian model is the simplest and can provide reasonable engineering estimates when a representative shear rate is used. Note that using the Cross model’s zero shear viscosity, h0 , as the apparent viscosity would result in a significant overestimation of molding shear stresses and pressures since it neglects the shear thinning behavior near the cavity walls where the majority of the flow conductance is generated.

Figure 5.6 Newtonian model of viscosity behavior

For a Newtonian flow, the velocity profile is a parabolic function of the thickness, z. æ æ 2 z ö2 ö÷ ç v ( z ) = vmax çç1- çç ÷÷÷ ÷÷÷ (5.14) ç èçç è H ø ø÷

where vmax is the velocity at the centerline and z varies from –1/2 to +1/2 of the thickness, H. The volumetric flow rate, V , is the integral of the velocity across the thickness times the width, W. æ æ 2 z ö2 ö÷ æ 2 ö ç vmax çç1- çç ÷÷ ÷÷÷ = çç ÷÷ vmaxWH = v WH (5.15) ççè çè H ÷ø ÷ø çè 3 ÷ø -H 2

V = W ò

H 2

5.3 Viscous Flow

where the average linear flow velocity, v , is equal to two-thirds of the melt velocity at the centerline. The apparent shear rate at the wall can be calculated from either the average linear flow velocity or the volumetric flow rate as g =

6v 6V = (5.16) H WH 2

Given this estimate of the shear rate, the apparent viscosity should be evaluated and used for estimation of the pressure drop. Equations 5.7, 5.13, and 5.16 can be combined to provide estimates of the pressure drop as a function of either the linear flow velocity or the volumetric flow rate. DP =

12m Lv 12m LV = (5.17) H2 WH 3

5.3.5 Power Law Model Newton’s law of viscosity assumed that the viscosity is not a function of shear rate. When a material does not obey this law, it is said to be non-Newtonian. One of the simplest and most common non-Newtonian models is the power law model, which states that the viscosity is an exponential function of the shear rate. h = k g n-1 (5.18)

where k is a consistency index representative of the value of viscosity evaluated at a shear rate of one reciprocal second and n is the power law index.

Figure 5.7 Power law model of viscosity behavior

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Figure 5.7 compares the power law model against the non-Newtonian behavior provided by the Cross-WLF model for a medium viscosity PC at 280°C. It is observed that the power law model provides excellent estimates of the viscosity at higher shear rates but overestimates the viscosity at lower shear rates. For this reason, it should be expected that the power law model will provide more accurate estimates than the Newtonian model yet overestimate the pressure drop compared to the Cross-WLF model, since it over-predicts the viscosity at lower shear rates. It should be noted that some resins, such as some grades of polypropylene, tran­ sition to a power law regime at very low shear rates. For these types of materials, there is no apparent Newtonian plateau and the power law model can be expected to provide very good estimates. For other materials exhibiting a significant Newtonian plateau, such as the polycarbonate just discussed, the power law model can purposefully fit to a smaller shear rate regime of interest to provide more accurate results. For flow of a power law fluid, the velocity profile through the thickness is a function of the power law index, n. 1ö æ çç æ 2 z ö1+ n ÷÷ ÷ ç ç v ( z ) = vmax ç1- ç ÷÷ ÷÷÷ (5.19) çç çè H ø ÷÷ èç ø÷

The volumetric flow rate is the integral of the velocity across the thickness times the width, W. æ 1ö 1ö æ çç æ 2 z ö1+ n ÷÷ ççç 1 + ÷÷÷ ÷÷ v WH = v WH (5.20) n ÷ V = W ò vmax çç1- çç ÷÷÷ ÷÷ = çç çç çè H ø ÷÷ çç -H 2 1 ÷÷÷ max èç ø÷ èçç 2 + ø÷÷ n H 2

It should be noted that a power law index equal to one reverts the power law model to the Newtonian model. As the power law index decreases, the viscosity exhibits increased shear thinning such that the polymer melt flows faster near the side wall. As the power law index approaches zero, a plug flow develops in which the melt velocity is almost constant through the thickness. These behaviors are ­graphically depicted in Fig. 5.8. Note that that the melt velocity at the center line decreases to maintain a constant volumetric flow rate as the power law index ­decreases.

5.4 Process Simulation

Figure 5.8 Velocity dependence on the power law index

With the power law model, the shear rate at the wall is not required to estimate the pressure drop, but it may be useful to calculate to avoid excessive shear rates or check the viscosity of the melt. It can be calculated from either the average linear flow velocity or the volumetric flow rate as æ æ 1ö 1ö 2çç2 + ÷÷ v 2çç2 + ÷÷V ÷ çè ç è nø n ø÷ (5.21) g = = H WH 2

Equations 5.7, 5.18, and 5.21 can be combined to provide estimates of the pressure drop as a function of either the linear flow velocity or the volumetric flow rate. n n æ æ æ æ 1ö ö 1ö ö ççç 2ççç2 + ÷÷÷ v ÷÷÷ ççç 2ççç2 + ÷÷÷V ÷÷÷ 2kL ç è 2kL ç è n ø ÷÷ n ø ÷÷ ç ç DP = ÷÷ = ÷ (5.22) ç ç è ø è H H H WH 2 ÷ø

„„5.4 Process Simulation Process simulation is a valuable tool for modeling the polymer melt flow and ­solidification in mold design. Compared with manual analysis, simulation can actually be easier to perform within modern CAD environments while also providing more accurate results. As of 2015, providers of mold filling simulations include

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Autodesk/Moldflow, CoreTech/Moldex3D, Dassault/Simpoe, Sigma Engineering/ Sigmasoft, and others. These simulations discretize the product’s design geometry into a mesh of finite elements within which the flow, temperature, and pressure may be calculated. This numerical approach allows the complex physics of the molding process to be modeled including (1) arbitrary geometry, (2) non-Newtonian flow, (3) viscous heating, (4) compressibility, (5) temperature dependent thermal conductivity and specific heat, including latent heat of fusion, (6) thermal contact resistance between the polymer melt and mold, (7) profiled injection velocity and pack pressures, (8) cooling lines, (9) core shift and plate deflection, (10) venting, and many other effects. The results of the simulation can include prediction of the filling pattern, pressure distribution, bulk temperatures, volumetric shrinkage, and part properties including residual stress, shrinkage, warpage, and others. Simulation is increasingly easy to perform with simplified user interfaces and ­wizards for setting up the analysis through the use of default inputs for process conditions and other parameters. Validation studies have been performed and indicate common failure modes related to mold geometry, material properties, and process conditions [8, 9]. Care should be taken when adopting simulation results, and the results should be challenged when they do not agree with experience. As an example of these simulation capabilities, Autodesk/Moldflow Plastics Insight 2014 was performed for the bezel design. Figure 5.9 provides the average flow velocity of the polymer melt during the filling stage. The discretized mesh is shown at upper right of Fig. 5.9, while the arrows represent the velocity vector (both magnitude and direction) as the polymer melt progresses in the mold cavity. For a filling time of 0.44 s, the melt velocity averages about 50 cm/s along the top of the bezel. This should be expected since the flow length from the gates to the center of the top side is 200 mm, so the average flow velocity is 200 mm divided by 0.44 s, or 450 mm/s. However, the simulation also provides many details of the melt flow that could be difficult to predetermine. These details include the radial flow out of gates, the local flow around windows and into bosses, the stagnation of the flow at the side walls near the gates, and the convergence of the melt flow at the last area to fill the mold (at the top center of the bezel). A common investigation in simulation is to change the number of gates from 2 to 1 or 3 or 4 or whatever to predict the resulting flow patterns, melt pressure, and part properties. Once the gate location(s) are determined, the runner system (either cold or hot runner) should also be modeled since the runner system will alter the pressure drop, flow rates, and temperature of the polymer melt entering the cavity through each leg and gate of the runner system as discussed in Chapter 6.

5.4 Process Simulation

Figure 5.9 Melt velocities in cavity at end of filling

Another standard use of process simulation is the estimation of the appropriate process conditions for injection molding, including injection time, barrel and coolant temperatures, pack pressure, pack time, and others. It is a good idea to perform a sensitivity study to characterize the process behavior with respect to the most important process conditions. Figure 5.10 shows the results of simulation across injection times ranging from 0.05 to 2 s. The input melt temperature for the ­simulation was uniformly set to 240°C, the midpoint of the recommended range for Cyclocac MG47 resin. As shown by the square symbols corresponding to the right axis, the simulation indicates that the bulk temperature is about 240°C for an ­injection time of 2 s, but increases to above 260°C for an injection time of 0.05 s. This increase is due to viscous heating of the polymer melt at higher injection ­velocities coupled with less conductive heat transfer given the shorter filling time. In molding practice [10, 11]YZ , practitioners will strive to set the injection time so that the bulk temperature remains uniform throughout the mold cavity filling. This would suggest that an injection time of 2 s is appropriate. However, Fig. 5.10 also provides a graph for injection pressure in which a minimum is observed near 0.2 s. Scientific molding practitioners [12, 13] would strive to set the injection time to minimize the injection pressure. The reason for the lower pressure at the lower injection time is that the polymer melt has not only a lower viscosity given the higher bulk temperature and shear rates, but also a thinner frozen layer.

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Figure 5.10 Required injection pressure and bulk melt temperature at end of fill as a function of the cavity filling time

While the objectives to maintain a uniform melt temperature and minimize molding pressure are rational, simulation often does not predict the many vagaries of the real molding trials when the mold is being commissioned. Figure 5.10 suggests that fill times between 0.1 and 2 s are likely for this application, with a more likely range being between 0.1 and 0.5 s given the desire to minimize cycle time. Additional simulations can be performed to further narrow the range and optimize the process once the feed system and cooling lines are designed.

„„5.5 Cavity Filling Analyses and Designs There are many applications for a filling analysis, including cost reduction, process optimization, and quality improvements. While the following examples provide a broad array of typical applications, the mold designer should customize or further develop these analyses according to the specific needs of the molding application.

5.5.1 Estimating the Processing Conditions Mold designers should verify that the mold can be filled given the cavity geometry and the material properties. However, the filling analyses require the processing conditions, including the melt temperature and either the linear velocity or volumetric flow rate of the melt. It is recommended that mold designers assume a melt temperature in the middle of the melt temperature range recommended by the material supplier, since this provides the molder with some freedom to adjust ­temperatures up or down to correct molding problems or reduce cycle time.

5.5 Cavity Filling Analyses and Designs

The true melt flow rate is not known until after the mold is made and commissioned. The maximum flow rate is typically bounded by the maximum ram velocity of the molding machine or molding defects caused by high flow rates, such as flash, jetting, or burn marks. The minimum flow rate is typically bounded by the pre­ mature solidification of the melt in the mold cavity, which results in a short shot. Typical linear flow velocities of the melt through the mold range from 0.01 to 1 m/s depending on the specifics of the molding application. Thin wall applications will generally have higher linear flow velocities because ƒƒthey require a faster injection to avoid premature solidification and ƒƒtheir thinness provides for faster linear velocities given the same volumetric flow rate from a molding machine. Melt flow rates may be estimated by computing the volume of the mold cavities and runners and dividing by the estimated filling time. This approach works well for those practitioners with experience but may not work well for new molding applications having very different geometries or material properties. Alternatively, ­additional analysis can lead to a recommended flow rate that balances the amount of shear heating with the heat loss from the melt to the mold. This result should provide not only a reasonable estimate of the melt flow rate, but also a more accurate analysis since it will tend to produce a uniform melt temperature as the melt fills the mold. The derivation of the melt velocity is provided in Appendix F. For a Newtonian ­material, the recommended velocity is v=

35(Tmelt - Twall )k 24m

(5.23)

where Tmelt and Twall are the melt and mold wall temperature, k is the thermal c­ onductivity of the plastic melt, and m is the Newtonian viscosity. Since the viscosity is a function of the shear rate and velocity, it is necessary to recompute the shear rate and viscosity until the velocity converges. Example: This analysis will now be applied to the laptop bezel, which has a wall thickness of 1.5 mm and is to be molded of ABS (Cycolac MG47) at a melt ­temperature of 239°C and a mold coolant temperature of 60°C. For the purpose of the analysis, we will initially assume that the filling time is 1 s, which provides an initial estimate of the linear velocity (defined as the flow length of 0.2 m divided by the filling time) of 0.2 m/s. At this velocity, the shear rate is computed as

g =

6v 6 × 0.2 m/s = = 800 s-1  H 0.0015 m

(5.24)

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At this shear rate, the power law model provides a melt viscosity of

(

)

(

m g = 2000 s-1 = k g n-1 = 1.71×104 Pasn × 800 s-1

n-1

)

= 219 Pas

This apparent viscosity value can then be used to provide a new estimate of the recommended injection velocity according to Eq. 5.23 as

v=

35(239°C - 60°C)0.19 W/m°C 24 × 219 Pas

= 0.44 m/s 

(5.25)

Additional iterations are useful to hone in on the recommended velocity. At a velocity of 0.44 m/s, the shear rate is 1750 1/s. The viscosity at this shear rate is 131.5 Pa s, which in turn suggests a linear melt velocity of 0.56 m/s. A further iteration for this velocity yields a shear rate of 2250 1/s, a viscosity of 111 Pa s, and a melt velocity of 0.61 m/s. With additional iterations, the solution will converge to a final velocity of 0.64 m/s. Since the flow length is approximately 0.2 m, the mold cavity for the laptop bezel will fill in approximately 0.33 s (not including the runner system). This result closely agrees with the simulation results of Fig. 5.10. 

As implied by the form of Eq. 5.23, the recommended velocity will vary with the melt temperature, the mold temperature, the thermal conductivity of the melt, and the melt viscosity. Higher temperature differences between the melt and wall ­temperatures, as well as higher thermal conductivity of the polymer melt, require faster melt velocities to maintain a uniform melt front temperature. Lower viscosity materials require a higher melt velocity to generate the shear heating needed to avoid excessive heat loss to the melt. While the melt velocity does not appear to vary with wall thickness, the effect of wall thickness is considered through the inclusion of the viscosity, which is a function of the shear rate. As the wall thickness decreases, the increasing shear rate reduces the viscosity, which thereby requires higher melt velocities to avoid cooling the melt. As expected, higher melt velocities are required as the wall thickness decreases. Figure 5.11 plots the recommended melt velocity for ABS as a function of melt temperature and wall thickness using the analysis. It is observed that the melt velocity can vary from about 0.4 m/s for a molding application with a wall thickness of 3 mm and a melt temperature of 218°C to about 1.6 m/s for a molding application with a wall thickness of 0.8 mm and a melt temperature of 260°C.

5.5 Cavity Filling Analyses and Designs

Figure 5.11 Recommended melt velocity for ABS as a function of wall thickness and ­temperature

While there is a significant range in the recommended melt velocity as a function of the molding application, it is important to recognize that the exact melt velocity and flow rate that will actually occur during the molding process is unknown. The objective should be to provide a reasonable estimate of the melt velocity and filling time and design the mold to operate under a wide variety of conditions. While the foregoing analysis may seem unnecessarily complex compared to simply assuming a filling time based on experience, the analysis is objective and provides a quan­ titative result that provides insights to the design and use of injection molds.

5.5.2 Estimating the Filling Pressure and Minimum Wall Thickness The material cost, processing cost, and environmental impact of molded parts are all reduced with reductions in wall thickness. However, minimizing the wall ­thickness can make the mold cavity difficult to fill and adversely increase clamp tonnage. To estimate the pressure required to fill a mold, the mold designer must know the total distance that the flow is required to travel to fill the mold. For this reason, the mold designer should select the gating location(s) to balance the flow between the different portions of the mold. Since this is a one-dimensional flow analysis, features such as ribs and bosses are neglected. These features are very likely to fill if they are relatively small compared to the primary flow path. Prediction of the pressure drop across the mold cavity, ∆P, can be made given the flow length and the linear flow velocity of the melt by application of either Eq. 5.17 or 5.22. The primary assumption in the estimation of filling pressures is that the melt velocity will be maintained at a constant value as the melt propagates from the gate to the end of the mold. In theory, such a uniform melt velocity could be

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achieved by careful ram velocity profiling. In practice, complex mold geometries preclude the realization of uniform melt velocities, and ram velocity profiling is seldom used towards this purpose anyways. As such, the melt velocity will vary substantially from the gate (where the velocity is initially very high due to the small cross-sectional area of the melt) to the point of end of fill. Even so, the ­estimation of filling pressures is vital to ensuring that the moldings can be made with the mold design and the plastic materials used. To predict the filling pressure in complex products manually by lay-flat analysis, it is necessary to deconstruct the geometry into a series of simple segments. The flow in each segment can then be separately analyzed using the Newtonian or power law models relating pressure drop to flow rates in the segment. Returning now to the laptop bezel shown in Fig. 5.1, it may be assumed that the flows on the left-hand and right-hand sides are symmetric. Accordingly, the analysis will consider just half of the geometry. To do the analysis, any turns in the bezel will first be straightened as shown in Fig. 5.12. While this step is not necessary for the ­analysis, it emphasizes that the analysis considers only the pressure drop along the length direction of the melt flow. Next, the edges are folded out to reveal additional flow that is required to fill the vertical sides of the mold cavity.

Figure 5.12 Lay flat of laptop bezel (dimensions in mm)

5.5 Cavity Filling Analyses and Designs

As shown in Fig. 5.12, the gate location has been selected near the center location. The lay-flat geometry for the laptop bezel is then split into two flow segments representing the flow to the upper and lower portions of the mold. It should be noted that it is possible to include changes in the channel width, such as narrower sections due to windows, as shown in the middle lay flat on the right side of Fig. 5.12. Sections of varying thickness should also be broken out into different flow segments. By analyzing the flow in each of these segments, it is possible to provide very good estimates of the melt front locations and melt pressures as the melt fills the mold. Alternatively, sections of similar width may be lumped together to ­simplify the computation of the flow rate and filling pressures as shown in the right-most lay flat. Example: Estimate the pressure drop for the laptop bezel, assuming a constant melt velocity of 0.64 m/s. The left-most lay flat of Fig. 5.12 is used, modeling one-quarter of the mold cavity as a rectangular strip with a length, width, and thickness of 200 mm, 20 mm, and 1.5 mm, respectively. The viscosity data are fit with the power law model. The coefficients for an ABS material at 239°C are a consistency index, k, equal to 17,070 and a power law index, n, equal to 0.348. According to Eq. 5.22, the pressure is then

æ ö0.348 çç 2ççæ2 + 1 ÷÷ö0.64 m/s ÷÷ ÷÷÷ 2 ´ 17070 Pas ´ 0.2 m çç èç 0.348 ø÷ çç DP = ÷÷ ÷ø çè 0.0015 m 0.0015 m DP = 82,720,000; Pa = 82.7; MPa = 12,000 psi This pressure is a fairly significant amount relative to the capabilities of most injection molding machines, especially when considering that the estimated filling pressure does not include the pressure drop through the feed system. The result also closely agrees with the process simulation results for injection pressure provided in Fig. 5.10. 

The product designer and mold designer may wish to consider the pressure required to fill for a variety of wall thicknesses, flow rates, and melt temperatures. Figure 5.13 provides the estimated filling pressure required to fill the cavity for a range of wall thicknesses at the material’s mid-range melt temperature. The minimum wall thickness allowable for a given injection pressure can be derived as indicated in Fig. 5.13. Specifically, a line indicating the maximum allowable pressure is placed on the graph with the minimum wall thickness occurring at

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the intersection of the pressure curve. The analysis in this instance indicates that the minimum wall thickness is 1.4 mm at a melt temperature of 240°C.

Figure 5.13 Filling pressure as a function of wall thickness

There are two important concepts that should be understood when minimizing the wall thickness. First, the minimum wall thickness is a function of the melt temperature. It is recommended that mold designers use the mid-range temperature for analysis, since this reserves the opportunity for the molder to increase melt ­temperature and thereby reduce the filling pressures if needed. Second, the minimum wall thickness is also a function of the feed system design, since the pressure deliverable to the cavity from the machine is dependent on the pressure drop through the feed system, as later discussed in Chapter 6.

5.5.3 Estimating Clamp Tonnage The clamp tonnage is defined as the amount of force, usually measured in units of English or metric tons of kiloNewtons, which is required to hold the mold closed during operation. The clamp tonnage, Fclamp , can be calculated as the integral of the melt pressure acting on the projected area of the mold cavities. Fclamp = ò P ( A) cos q ( A) dA (5.26) A

where P(A) is the melt pressure in the mold across the area of the cavity and q(A) is the angle between the direction normal to the mold cavity surface and the mold opening direction. The projected area of the cavity is used rather than the total area of the mold cavity, since the melt pressure acting on inclined (or vertical) side walls contribute little (or no) force in the direction of the mold opening.

5.5 Cavity Filling Analyses and Designs

The maximum clamp tonnage typically occurs at the end of the filling phase when the filling pressure is at its peak value, or at the start of the packing phase when the entire mold cavity becomes pressurized at the packing pressure. It can be difficult to discern during actual molding whether the maximum clamp tonnage will be driven by the pressures during the filling or packing stages. Consider the cavity pressure distributions along the lay-flat model of the laptop bezel shown in Fig. 5.14.

Figure 5.14 Cavity pressure distribution during filling and packing

The figure at left indicates that there will be a linear pressure drop along the flow in the cavity from about 90 MPa at the gate to 0 MPa the end of fill. The average pressure exerted in the cavity is 45 MPa. While the width of the lay flat was approximately 20 mm, the projected area of the lay flat (refer to Fig. 5.12) is approximately 12 mm. The clamp tonnage for this strip required at the end of filling is Fclamp = 45 MPa ´ 0.2 m ´ 0.012 m = 108 kN = 10.8 mTon (5.27)

During packing, a slightly lower pressure is applied, but the pressure in the cavity is much more uniform. Typically, the packing pressure is between 50 to 90 % of the filling pressure. As shown at right in Fig. 5.14, the average cavity pressure may be 75 MPa, which corresponds to a clamp tonnage required at the start of packing of Fclamp = 75 MPa ´ 0.2 m ´ 0.012 m = 180 kN = 18.3 mTon (5.28)

The analysis indicates that the peak clamp tonnage in this case will occur at the start of packing when the melt pressure in the cavity equilibrates. Since the packing pressure depends on the molding process and desired shrinkage, the exact value of the packing pressure is not known until the mold is made and operated.

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For this reason, a conservative approach is to assume that the filling pressure will be exerted everywhere in the cavity. The clamp tonnage can then be estimated as Fclamp = Pcavity × Acavity_projected (5.29)

where Pcavity is the assumed average pressure in the cavity and Acavity_projected is the projected area of the cavity. If the filling analysis suggests a reasonable filling pressure, then this value may be used for estimation of the clamp tonnage. The filling pressure for some molding applications, however, may be very low and give rise to excessive shrinkage. To avoid this issue, molders will generally use packing pressures of 50 MPa or greater. As such, the mold designer should verify the expected cavity pressures with the molder or assume a minimum cavity pressure of 50 MPa. Example: Estimate the maximum clamp tonnage required to mold the laptop bezel. Equations 5.28 and 5.29 indicate that the clamp tonnages during filling and packing for the lay-flat geometry were 10.8 and 18.3 mTon, respectively. Since the lay flat represents one quarter of the laptop bezel, the estimated clamp tonnage is four times the greater of the two, or:

Fclamp = 4 ´ max {10.8, 18.3 mTons} = 73 metric tons 

To validate the foregoing analyses of the previous three sections, the results are compared to the simulation previously discussed with respect to Figs. 5.9 and 5.10. A comparison of the analytical results with those of the numerical simulation is provided in Table 5.1. Table 5.1 Comparison of Analytical and Simulation Results Parameter

Analysis result

Simulation result

Injection time (s)

0.4 s

0.05 to 2 s

Filling pressure (MPa)

90.2 MPa

104 MPa at 0.4 s

Change in bulk melt temperature (°C)

0

+10

Average shear rate (s–1)

1760

1290

Clamp tonnage during filling (kN)

430

290

Clamp tonnage during packing (kN)

730

397

Results previously presented for a wall thickness of 1.4 mm was implemented in the laptop bezel. Two gates were located at the center of the left and right side walls. The simulation (Moldflow MPI 5.1) was performed for ABS (Cycolac MG47) with a melt temperature of 239°C, a mold temperature of 60°C, and a filling time of 0.25 s to correspond to a linear melt velocity of 0.8 m/s.

5.5 Cavity Filling Analyses and Designs

The simulation predicted a filling pressure of 104 MPa, which compares well to the analytical filling pressure of 90 MPa. The simulation would be expected to predict higher pressures since it models the development of a solidified layer as well as the flow in ribs, bosses, and other thin sections. The simulation predicted an ­increase in the bulk melt temperature of 10°C, which verifies that the analytically derived melt velocity is a reasonable estimate. The clamp tonnages predicted by the described analysis and the commercial simulation during filling are relatively close. The presented analysis is conservative and so predicted a high clamp tonnage. By comparison, the simulation models the ­solidification of the melt throughout the cavity and the resulting decay in the melt pressures. For this reason, the simulation predicted a much lower clamp tonnage of 397 kN occurring at 1.2 s into the packing stage. While this clamp tonnage may occur at this time, it does not represent the peak clamp tonnage that occurs at the very start of the packing stage or the clamp tonnage that may be required if the molding machine controller overshoots the velocity to packing changeover l­ ocation.

5.5.4 Predicting Filling Patterns Filling patterns can be readily predicted using the lay flat analysis technique and are useful to understand the behavior of the melt in filling the mold, locating gates, identifying knit-line locations, and assisting in other aspects of mold design. Analy­ sis will be performed for a five-sided container with a width, length, and height of 100 mm, 160 mm, and 60 mm, respectively. The container is shown in Fig. 5.15 and has a 2° draft with 10 mm fillets. Assuming that a two-plate mold will be used, the container will be gated at the edge of a side wall.

Figure 5.15 Container for prediction of fill patterns

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To predict the filling patterns, the sides of the container are “cut” at the corners and the side walls folded down to make a lay flat. The gate location is next identified. The flow will emanate from the gate, producing a circular melt front. As such, an arc may be drawn from the gate representing the position of the melt at a given point in time. Figure 5.16 provides the lay flat and some early melt front locations.

Figure 5.16 Lay flat and first melt front locations

Each arc in Fig. 5.16 corresponds to the location of the melt front at a different time step; the distance between the arcs is equal to the linear melt velocity times the time step. While the melt location at the first time step is correct, the melt will hit the adjacent side wall by the end of the second time step. As such, it is necessary to draw additional arcs on these adjacent side walls reflecting the position of the melt flow at various time steps. To correctly predict the flow behavior, the analysis must maintain the same flow resistance between the melt flowing in the various portions of the mold. This can be accomplished by creating a “phantom” gate and maintaining the same flow lengths from this “phantom” gate as from the real gate. For each time step, the length of flow is increased and an arc of corresponding ­radius is drawn. Intersecting arcs corresponding to the same time step are then trimmed. The flow is advanced with more phantom gates added as needed until the flow throughout the entire lay flat is created. Figure 5.17 demonstrates this melt front prediction process and the resulting melt front locations for the container. It is observed that the flow races around the side walls and will form a weld line and a gas trap on the side wall opposite the gate. This phenomenon, known as “race-tracking,” is quite common in molded parts and can occur when the length of flow around the perimeter of the molding is less than the length of flow across the centerline of the part.

5.5 Cavity Filling Analyses and Designs

Figure 5.17 Melt front locations for part of uniform thickness

In this case, race-tracking occurred because the 60 mm depth of the container is more than one-half the 100 mm width of the container. While the weld line is not desirable, a gas trap on a side wall such as shown in Fig. 5.17 is especially pro­ blematic since it is difficult to vent. As such, the trapped air will likely combust, causing a burn mark to appear at this location.

5.5.5 Designing Flow Leaders The gas trap in the previous example could have been avoided by moving the edge gate to the center of the 160 mm long side wall or by using a three-plate or hot runner mold to gate at the center of the mold cavity. Sometimes, however, the mold layout precludes these designs. As such, another alternative is to vary the thickness so that the melt purposefully flows faster in some portions of the mold [14, 15]. Such thicker sections used to control the flow velocity are generally known as “flow leaders.” It should be understood that thickness variations in molded parts are generally undesirable as discussed in Section 2.3.1. For the reasons discussed therein, the cavity thickness variation should be kept to a minimal amount. Newtonian flow analysis will now be used to redesign the wall thickness of the container to resolve the race-tracking issue. Equation 5.17 relates the pressure drop, velocity, and thickness. To eliminate the race-tracking, the pressure drop across the centerline should equal the pressure drop around the perimeter. DPcenterline = DPside_walls (5.30)

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This condition will ensure that the flow traverses across the centerline at the same time that the flow reaches the far corners of the adjacent side walls to eliminate the race-tracking phenomenon. The flow lengths are provided in Fig. 5.18. From the geometry of the container, the lengths of flow across the centerline and around the side walls are calculated to be 280 mm and 210 mm, respectively.

Figure 5.18 Lay flat showing flow lengths

From Eq. 5.17, the pressure drops across the centerline and around the side walls can be evaluated and equated as 12 mcenterline Lcenterline vcenterline 2 H centerline

=

12 mside_walls Lside_walls vside_walls 2 Hside_walls

(5.31)

The melt velocities in sections of different sections will not be equal. In fact, it is desired that the velocity of the perimeter be vside_walls = vcenterline

Lside_walls Lcenterline

(5.32)

This condition will cause the melt to arrive at the far corner of the side wall at the  same time it reaches the opposite side of the cavity along the centerline. ­Substituting this relation into Eq. 5.31 and solving for the thickness of the side walls, Hside_walls , as a function of the nominal thickness, H, Hside_walls = H

Lside_walls

mside_walls

Lcenterline

mcenterline

(5.33)

5.5 Cavity Filling Analyses and Designs

The analysis indicates that the wall thickness will be largely proportional to the ratio of the flow lengths with a lesser dependence on the melt viscosities. Assuming the same viscosity throughout the cavity, the thickness of the side walls can be evaluated as Hside_walls = 2 mm

210 mm = 1.5 mm (5.34) 280 mm

The lay flat analysis can also be used to predict the filling patterns for parts of ­varying wall thickness. When the wall thickness varies, it is necessary to increase the radii of the arcs to represent the distance that the melt traveled during the time step. For this case, the thickness of the side walls has been chosen such that the velocity of the melt in the side walls is: vside_walls = vcenterline

Lside_walls Lcenterline

= vcenterline

210 mm = 75% vcenterline (5.35) 280 mm

In the lay flat analysis, the radius of each arc in the thinner section should be incremented by 75 % of the arc in the thicker sections. Still using the same phantom gate, the resulting melt front progression in the redesigned container is shown in Fig. 5.19. The arrows along the edge of the side wall show the incremental position of the melt front in this section at various time steps. The analysis indicates that the melt does reach the end of the side walls before the melt reaches side of the cavity opposite the gate.

Figure 5.19 Melt front locations for part with flow leaders

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To validate the lay flat approach, numerical simulations were performed for the container having a uniform thickness of 2 mm and a second container in which the thickness of the side walls was decreased to 1.5 mm. The results are shown in Fig. 5.20.

Uniform thickness

Thinner side walls

Figure 5.20 Simulated melt front with and without flow leaders

As in the lay flat analyses, the simulation indicated that the container without the flow leader would exhibit race-tracking, a weld line, and a gas trap. Reducing the thickness of the side walls to 1.5 mm eliminated the problems. For reference, the reduction in the thickness of the side wall from 2 mm to 1.5 mm did increase the injection pressure by 10 % to fill out the thinner side walls but also decreased the part weight by a similar amount.

„„5.6 Chapter Review All mold engineering designs should consider the propagation of the viscous polymer melt throughout the mold cavity. Numerical simulations are preferred due to their ability to quickly and accurately consider non-Newtonian flows in complex geometries. However, analyses with Newtonian and power law viscosity models are not difficult to use and have been shown to provide reasonable results. The single most important purpose of filling analyses is to ensure that the mold cavity can be completely filled by the selected molten plastic. If the wall thickness of the cavity is too thin and the melt pressure required to fill the cavity exceeds the capability of the machine, then incomplete moldings (known as “short shots”) will be produced. The molder will try to remedy the problem by attempting to increase the melt temperature or injection pressure or by using another resin. If these ­attempts are unsuccessful, then the mold will require design changes including

5.7 References

the addition of more gates, increasing the diameters of the feed system, increasing the wall thickness of the mold cavity, or other changes. Such physical alterations of the mold can be expensive and time-consuming. Filling analyses can also be used to estimate the clamp tonnage, optimize the wall thickness, estimate the processing conditions, predict the advancement of the plastic melt throughout the cavity, and remedy filling problems by locating gates or designing flow leaders. While the governing equations for the Newtonian and power law provided in Eqs. 5.17 and 5.22 seem simple, careful application is required to obtain useful solutions. It is recommended that filling analyses utilize mid-range melt temperatures when evaluating the viscosity, and the dependence of the viscosity on shear rate be verified when using the Newtonian model. After reading this chapter, you should understand: ƒƒThe relationship between shear stress, shear rate, and viscosity; ƒƒThe relationship between cavity fill time, linear melt velocity, and volumetric flow rate; ƒƒThe assumptions made in development of the Newtonian and power law models, and potential issues associated with their use; ƒƒHow to estimate the length of flow in a mold cavity from a gate to the end of flow; ƒƒHow to calculate the shear rate, viscosity, filling pressure, and clamp tonnage for melt flow in a rectangular mold cavity using either the Newtonian or power law model; and ƒƒHow to estimate the minimum wall thickness in a molding application given the material properties and maximum filling pressure. The next chapter examines the design of the feed system for two-plate molds, threeplate molds, and hot runner molds. Flow analyses for the viscous melt in cylindrical and annular members is presented and used for feed system design. Afterwards, the analysis and design of gates will be presented before addressing cooling and other elements of mold design.

„„5.7 References [1] Spencer, R. and R. Dillon, The viscous flow of molten polystyrene, J. Colloid. Sci. (1948) 3(2): pp. ­163–180 [2] Malkin, A. Y. and A. I. Isayev, Rheology: Concepts, Methods, and Applications, ChemTec Publishing (2005) [3] Tanner, R. I., Engineering rheology, Oxford University Press (2000) [4] Bremner, T., A. Rudin, and D. Cook, Melt flow index values and molecular weight distributions of commercial thermoplastics, J. Appl. Polym. Sci. (1990) 41(7–8): pp. 1617–1627

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[5] Cross, M. M., Relation between viscoelasticity to shear-thinning behavior in liquids. Rheol. Acta (1979) 18: pp. 609–614 [6] O’Connell, P. A. and G. B. McKenna, Arrhenius-type temperature dependence of the segmental relaxation below Tg, J. Chem. Phys. (1999) 110(22): pp. 11054–11060 [7] Williams, M. L., R. F. Landel, and J. D. Ferry, Temperature dependence of relaxation mechanisms in amorphous polymers and other glassforming liquids, J. Appl. Phys. (1953) 24: p. 911 [8] DiScipio, W. and D. Kazmer, Validation of molded part shrinkage predictions by CAE simulation, in Injection Molding Division of the SPE Annual Technical Conference, Detroit, MI (1992) [9] Kazmer, D. O., et al., Validation of three on-line flow simulations for injection molding, Polym. Eng. Sci. (2006) 46(3): pp. 274–288 [10] Lord, H. and G. Williams, Mold–filling studies for the injection molding of thermoplastic materials, Part II: The transient flow of plastic materials in the cavities of injection–molding dies, Polym. Eng. Sci. (1975) 15(8): pp. 569–582 [11] Farshi, B., S. Gheshmi, and E. Miandoabchi, Optimization of injection molding process parameters using sequential simplex algorithm, Mater. Des. (2011) 32(1): pp. 414–423 [12] Orr, L. M. and D. J. Orr, When to Hire—or Not Hire—a Consultant: Getting Your Money’s Worth from Consulting Relationships, Apress (2013) [13] Sommier, E., et al., Characterization of the injection molding process of passive vibration isolators, J. Elastomers Plast. (2014) DOI: 1177/0095244314538439 [14] Seow, L. and Y. Lam, Optimizing flow in plastic injection molding, J. Mater. Process. Technol. (1997) 72(3): pp. 333–341 [15] Lam, Y. and L. Seow, Cavity balance for plastic injection molding, Polym. Eng. Sci. (2000) 40(6): pp. 1273–1280

6

Feed System Design

„„6.1 Overview The purpose of the feed system is to convey the polymer melt from the molding machine to the mold cavities. The design of feed systems can range from very ­simple to very complex. Increased investment in the feed system design will tend to provide for reduced cycle time and less material waste when using the mold. However, it is possible to overdesign the feed system, and the “best” feed system design is a function of the production volume, availability of molding pressure, and level of allowable investment. The design of the feed system follows a four-step process. First, the type of feed system (two-plate cold runner, three-plate cold runner, or hot runner) is selected if not already known; these three types of feed systems are the most common, though a few other feed system technologies are discussed in Section 13.6. Second, the routing of the feed system through the mold is determined. Third, the diameters of each segment of the feed system are specified to balance pressure drops, shear rates, and material utilization. Finally, the design details for the feed system are embodied in the mold design. To assist the design process, a discussion of the ­objectives in feed system design is provided next.

„„6.2 Objectives in Feed System Design 6.2.1 Conveying the Polymer Melt from Machine to Cavities The primary function of the feed system is to convey the polymer melt from the nozzle of the molding machine (where it is plasticized) to the mold cavities (where it will form a desired product). In most molding applications, the polymer melt must traverse portions of both the mold height and the mold width. The traversal

142

6 Feed System Design

of the height and width can be accomplished by two different layout designs for the feed systems as shown in Fig. 6.1. The feed system layout shown at left corres­ ponds to a two-plate cold runner mold design. The sprue is used to guide the poly­ mer melt from the nozzle of the molding machine to the parting plane. Runners in the parting plane are then used to guide the polymer melt across the parting plane to one or more mold cavities.

Figure 6.1 Two feed system layouts for melt conveyance

The second layout design on the right of Fig. 6.1 corresponds to a three-plate or hot runner mold. In this second design, the polymer melt is guided across the width and length dimensions of the mold by runners that are offset to the parting plane. Since the runners are offset from the parting plane, there is significant design freedom with respect to their routing and gating location. However, two sets of sprues are needed for the polymer melt to traverse the height of the mold. First, a sprue is needed to guide the polymer melt from the nozzle of the molding machine to the plane of the lateral runners. After the melt flows across the runners, a ­second set of sprues is needed to guide the melt down through a portion of the mold height to the mold cavities.

6.2.2 Impose Minimal Pressure Drop As the melt propagates through the feed system and cavities, the melt pressure in the injection molding machine will increase. The feed system must be designed so that there is sufficient melt pressure to drive the polymer melt throughout the mold cavities. As shown in Fig. 6.2, a feed system with a large flow resistance will incur a substantial pressure drop during the molding process. The flow rate of the  polymer melt will begin to decay when the molding machine reaches the ­maximum allowable injection pressure. If the flow rate decreases substantially before the end of the mold filling process, then a short shot or other defects are likely to occur.

6.2 Objectives in Feed System Design

Figure 6.2 Pressure and flow rate coupling

The feed system must be designed to incur an acceptable pressure drop to avoid short shots, extended cycle times, and other defects. The “acceptable” pressure drop through the feed system will depend on the specifics of the molding appli­ cation, especially the melt pressure required to fill the cavity compared to the melt pressure available from the molding machine. For example, a thin wall molding application [1] may use a molding machine with 200 MPa of available melt pressure. If 150 MPa is required to fill the cavity, then the pressure drop through the feed system should not exceed 50 MPa. However, if the same machine was used to mold a part requiring only 100 MPa of pressure, then the feed system could be designed to impose a pressure drop of 100 MPa. To accurately specify the acceptable pressure drop for the feed system design, the mold designer should contact the molder to obtain the molding machine’s maximum injection pressure. The mold designer should also obtain an estimate of the melt pressure required to fill the cavity through analysis, simulation, prototype molding, or prior experience. If this information is not known, then the mold ­designer can assume a maximum pressure drop through the feed system of 50 MPa (7200 psi). While this pressure drop is slightly higher than some industry practices, this specification will result in a steel-safe design with smaller feed system diameters and lower material utilization.

6.2.3 Consume Minimal Material To achieve the best feed system design, the mold designer should specify the dia­ meters of the feed system to jointly minimize the pressure drop and the feed system volume. These design constraints are represented in Fig. 6.3. As the diameters of the various segments of the feed system increase, the pressure drop decreases below the specified maximum. However, increasing the diameters of the feed system also results in an increase in the volume of the feed system, which can be undesirable for both cold and hot runner feed systems.

143

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6 Feed System Design

Figure 6.3 Coupling between volume and pressure drop

In cold runner designs, the large size of the feed system can result in extended cycle times as well as excessive material waste associated with the molding of the feed system. Some molding applications allow the use of regrind mixed with virgin material. A typical limit on regrind may be 30 %, which translates directly to a ­specification on the maximum volume of the feed system. For example, if a molding application had two cavities totaling 50 cc, then a 30 % regrind specification would limit the volume of the feed system to 15 cc. In hot runner designs, large feed systems reduce the turnover of the material in the hot runner. Low turnover is undesirable for two reasons. First, long residence times of the polymer melt in the hot runner can cause material degradation, which frequently causes black specks and reduced properties of the molded product. ­Second, large volumes of material in the hot runner system can impede color changes during molding, not only due to the large volume of the plastic melt that needs to be flushed, but also due to the low associated shear stresses in the polymer melt along the walls of the feed system. Low shear stresses during purging allow the material to stick to the walls of the hot runner, reducing the removal of old material during color changes. The maximum volume of polymer melt in a hot runner feed system can be difficult to specify since it is related to the type of material being molded, the need to ­perform color changes, and the desired pressure drop. Hot runners are being increasingly designed with smaller diameters such that the material turns over every molded cycle. For example, if a molding application has two cavities totaling 50 cc, then a turnover of the melt with every molding cycle would specify the volume of the feed system to be 50 cc. If a very low pressure drop is desired, then the volume of the feed system may be specified as 100 cc or even 200 cc if degradation and color change issues are not expected. It should be noted, however, that unlike a steel-safe designed cold runner system, high costs may be incurred to modify the diameters of a hot runner system.

6.3 Feed System Types

6.2.4 Control Flow Rates Since the primary function of the feed system is to convey the melt from the molding machine to the mold cavities, it is desirable for the feed system to control the amount of polymer melt to each mold cavity. The two most common applications pertain to multicavity and multigated molds. ƒƒIn a multicavity mold, as shown in Fig. 6.1, the molding application may require different pressure drops in each leg of the feed system to cause the different mold cavities to fill at the same time. In this example, if the cup required a higher pressure to fill than the lid, then the mold designer could provide a lower pressure drop in the portion of the feed system leading to the cavity for the cup. Such a mold design is known as “artificially balanced.” ƒƒIn a multigated mold, a common objective in the feed system design is to control the polymer melt flowing through the feed system to alter the melt front advancement in a multigated mold. For instance, it may be desirable to drive more material through one gate to move a knit-line to a different location. Other common uses include the altering of the mold filling to eliminate a gas trap or avoid overfilling a portion of the mold cavity. Using different diameters in the feed system can control the flow of the polymer melt, but there are limits as to what can be achieved. First, the pressure drop through each leg of the feed system is dependent on the viscosity of the polymer melt. As such, an artificially balanced feed system may not balance the mold filling for different materials and processing conditions. Second, differently sized feed systems will solidify at different rates and thereby provide different dynamics ­during the packing stage of the molding process; runner segments with smaller diameters will tend to freeze quickly and reduce the amount of packing to downstream cavities. For these reasons, the mold designer should strive to utilize mold cavities that have similar filling requirements. If family molds or other needs dictate very different flow rates through each gate, then the mold designer may wish to utilize dynamic melt control technology as discussed in Section 13.6.4.

„„6.3 Feed System Types The most common types of molds were first introduced in Chapter 1. In this section, the layouts of the different types of feed systems and their accompanying components are discussed in greater detail.

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6 Feed System Design

6.3.1 Two-Plate Mold The two-plate mold is so named because it consists of two assembled sections that sandwich the melt; each half of the mold can consist of one or more mold plates. A section of an isometric mold is provided in Fig. 6.4, in which the shading represents the progression of the polymer melt. During the molding process, the nozzle of the molding machine mates with the radius of the sprue bushing. The polymer melt first flows down the sprue bushing, thereby traversing the thickness of the top clamp plate and A plate. The material then flows across the parting plane through runners and gates into one or more mold cavities.

Figure 6.4 Isometric section of two-plate mold

After the polymer is fully injected and solidified, the mold is opened at the parting plane, which is located between the A and the B plates. Typically, the A half of the mold remains stationary while the B half of the mold is pulled away with the ­moldings and runner system remaining on the core. To facilitate ejection, a reverse taper is usually provided on a knock-out pin below the sprue to ensure that the sprue and attached runner remains with the B half. After the mold has opened sufficiently to remove the moldings, the ejector plate is pushed forward by the

6.3 Feed System Types

molding machine. The sprue knock-out pin pushes on the sprue, breaking the small undercut and ejecting the sprue from the B side of the mold. While not shown in Fig. 6.4, additional ejector pins and knock-out pins can be placed down the length of long runners to facilitate ejection of the feed system. Figure 6.5 shows the molding that would be produced from the mold design of Fig. 6.4. In the design of the feed system, the length of the sprue is determined by the combined thicknesses of the top clamp plate and the A plate. The lengths of the runners are determined by the position of the cavities and the layout of the asso­ ciated runners. Given this layout, the mold designer needs to specify the diameters of the feed system. In general, the diameters of the upstream runners are larger than the diameters of the downstream runners, since the flow of the polymer melt branches at runner junctions and there will be a lesser flow rate through each of the downstream runners.

Figure 6.5 Two-cavity molding with runners and sprue

While a two-cavity two-plate mold is used to demonstrate mold design concepts, the provided analysis can be applied to more complex feed system layouts. For ­example, Fig. 6.6 provides a feed system design for an eight-cavity family mold, including two primary runners, four secondary runners, and eight tertiary runners. If the flow rate through the sprue was 100 cc/s, then the flow rate through each of the primary, secondary, and tertiary runners would respectively be 50 cc/s, 25 cc/s, and 12.5 cc/s. If the flow resistance in the cavities varied substantially, however, then the flow rate in the tertiary runners could vary substantially during the mold filling stage.

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6 Feed System Design

Figure 6.6 Eight-cavity molding with runners and sprue

As with the feed system design shown in Fig. 6.5, the diameter of each downstream runner is smaller than the upstream runner shown in Fig. 6.6. There is one notable exception: the diameter of the molding machine’s nozzle orifice is typically smaller than the diameter of the sprue inlet. The smaller orifice in the machine nozzle provides a point for separation between the molded sprue and the solidified plug in the nozzle of the molding machine. If the nozzle orifice were larger than the sprue inlet, then a frozen section of plastic in the nozzle behind the sprue bushing could cause the sprue to stick to the A half of the mold. When such sticking occurs frequently, the molder may choose to perform a “sprue break” by retracting the injection unit of the molding machine from the machine nozzle prior to mold opening and part ejection. This action is undesirable since it adds complexity and variance to the molding cycle, so the mold designer should verify and/or recommend the nozzle orifice diameter appropriately.

6.3.2 Three-Plate Mold A sectioned isometric view of a fully open three-plate mold design is provided in Fig. 6.7; the view provided in Fig. 6.7 does not include the ejector housing and associated components, since these are not central to the operation of the three-plate

6.3 Feed System Types

mold. Three-plate molds are comprised of three mold sections that move relative to each other, with each section consisting of one or more plates. The addition of a second parting plane between the A plate and the top clamp plate allows for runners to be located above the mold cavities and to traverse across the width and length of the parting plane without interfering with the mold cavities. For this reason, the three-plate mold provides greater freedom with respect to gating ­locations and the feed system layout. An added benefit is that three-plate molds often provide automatic separation of the molded parts from the feed system as shown in Fig. 6.7.

Figure 6.7 Isometric section of three-plate mold

Figure 6.8 provides a section through a fully closed three-plate mold. In this d ­ esign, the polymer melt flows down the sprue bushing across the thickness of the top clamp plate and stripper (or “X”) plate. The polymer melt then flows along runners located in the parting plane (referred to here as the “A-X” parting plane) between the A plate and the stripper plate. Tapered sprues are then used to convey the melt through the thickness of the A plate and any cavity insert support plate to the mold cavities.

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Figure 6.8 Section of closed three-plate mold

Sprue pullers, also known as “sucker pins,” are used near the sprue locations and other portions of the runner to ensure that the feed system remains with the ­stripper plate; the mold designer should design the sucker pins such that they do not restrict flow. In the feed system design of Fig. 6.8, the pins have a small diameter and depth compared to the dimensions of the primary runner. To further reduce the flow obstruction in the design of Fig. 6.8, they could be moved further away from sprue bushing. Figure 6.9 provides a section through a partially opened three-plate mold. After molding, the B side of the mold is pulled away from the A side, forcing the mold to open at the parting plane between the A and B plates; the ejector system, rear clamp plate, and associated components have been omitted. A spring located ­between the A plate and the stripper plate may be used to cause early separation of the A-X parting plane. The B side continues to open, with the distance between the A and the B plates controlled by the length of a stripper bolt connecting the A plate to the B plate. The free length of the stripper bolt must be sufficient to allow for the ejection of the molded parts. A typical mold opening distance between the A and B plates is equal to two to three times the height of the molded parts. As shown in Fig. 6.9, this distance can be quite large for molded parts with even relatively shallow cores.

6.3 Feed System Types

Figure 6.9 Partial section of partially opened three-plate mold

Once the length of the stripper bolt is traversed, the A plate will move away from the stationary platen along with the B plate. The A plate will traverse the free length of the stripper bolt for the stripper plate. The free length of this stripper bolt determines the mold opening distance between the A plate and the stripper plate. As with the A plate stripper bolt, the length of the X plate stripper bolt must be sufficient to allow for the removal of the feed system. Once the A plate traverses past the free length of the X plate stripper bolt, the stripper plate will move away from the top clamp plate along with the A section, B section, and ejection system of the mold. Figure 6.10 provides a section through a fully opened three-plate mold without the ejector system or rear clamp plate. During mold operation, the mold opening velo­ city and position must be carefully determined and controlled to achieve an efficient and fully automatic cycle. If the mold opening dimensions are not carefully specified, then the feed system may not be reliably ejected or the mold can be damaged. To optimize the mold operation, the mold opening distances in many three-plate molds can be adjusted by changing the position of nuts on the stripper bolts or by adding washers between the plates and the ends of the stripper bolts.

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Figure 6.10 Partial section of a fully opened three-plate mold

It should be noted that this three-plate design has been made as compact as possible with respect to mold opening distances, selection of plate thicknesses, and stripper bolt lengths. As such, it is insightful to compare the design of the threeplate mold with that of the two-plate mold as done in Table 6.1. The additional plates and components in the three-plate mold have increased the stack height by 44 mm (1¾ inches) and the mass by 30 kg, relatively small increases (on the order of 20 %). However, the three-plate mold has a mold opening distance of 250 mm, much greater than the mold opening distance of 75 mm for the two-plate mold. This larger mold opening distance is undesirable, since it adds to the mold opening and closing time and may also prevent the mold from operating in some injection molding machines with limited daylight.

6.3 Feed System Types

Table 6.1 Two- and Three-Plate Feed System Comparison Feed system type

Two-plate mold

Three-plate mold

Mold stack height

264 mm

308 mm

Mold opening distance

 75 mm

250 mm

Total required daylight

339 mm

558 mm

Mold mass

151 kg

181 kg

Mold opening time

0.36 s

1.2 s

6.3.3 Hot Runner Molds Hot runner molds should be considered whenever gating flexibility, cycle efficiency, and material efficiency are important. In a hot runner system, the feed system is encased in a heated channel so that the plastic remains molten during the molding process. Since the plastic does not cool in a hot runner system, there is no need to ƒƒplasticize the melt that would be required to fill the feed system, ƒƒinject the material that would fill the feed system, ƒƒwait for the material in the feed system to cool, ƒƒopen the mold a substantial amount to remove the feed system as in a three-plate mold, ƒƒde-gate the feed system from the molded products, or ƒƒre-grind or discard the runner system. For all these reasons, it is not uncommon for hot runner molds to operate with 20 % faster cycle times and 20 % less scrap material than a conventional two-plate or three-plate cold runner mold. However, hot runner molds do require a higher ­initial investment than either two-plate or three-plate molds and also require hot runner controllers to maintain the melt temperature. While hot runners may seem to increase energy utilization [2], the associated gains in improved material utilization and molding productivity provide net reductions to total energy costs. Figure 6.11 provides a section through an isometric view of a hot runner system. This hot runner design includes a hot sprue bushing, manifold, two drops or “­nozzles,” four heater control zones, and other components. During operation, the material from the molding machine’s nozzle will travel down the hot sprue bushing to the primary runner located in the manifold. The melt then traverses down the length of one or more runners to downstream hot runner nozzles. The length of the nozzle is determined by the distance from the centerline of the manifold to the gating location of each mold cavity.

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Figure 6.11 Isometric section of hot runner system

Compared to cold runner designs, the diameters of the runners and drops in a hot runner system may be quite large, since all the material in the hot runner will eventually be forced into the mold cavities. Since the polymer melt is not wasted, hot runner systems can have large runner diameters to provide for very low flow resistance and excellent transmission of the injection pressure to the mold cavities. However, overly large diameters can permit the material to degrade in the hot runner and prohibit rapid change-overs between different plastic resins and colors. A section through a hot runner mold assembly is shown in Fig. 6.12. This mold design provides for the injection of the plastic melt into the left and right sides of the laptop bezel via a naturally balanced hot runner system with two drops. As can be observed, an air gap surrounds the majority of the hot runner system to minimize heat transfer from the heated manifold and nozzles to the colder mold steel. During molding, the melt pressure exerted on the faces of the mold cavity and hot runner system will result in forces that would tend to cause the cavity insert and the hot runner system to deflect. Thrust pads, sometimes machined from titanium, are used to transfer these forces from the hot runner system to the top clamp plate while transferring a minimal amount of heat. With hot runner molds, cooling lines and/or insulating sheets should be used with the top clamp plate to prevent the transfer of significant heat to the platens of the molding machine.

6.3 Feed System Types

Figure 6.12 Partial section of hot runner mold

The hot runner system design provided in Fig. 6.12 is a relatively simple design, which utilizes thermal gates that will be specified in the next chapter. In this ­design, the hot runner nozzles are concentric with the gate cut-out provided in the cavity insert. Since the manifold will expand with changes in the manifold tem­ perature, the manifold is allowed to expand and slide across the top surface of the nozzles. The manifold and nozzles are maintained in compression in the height direction to prevent any significant amount of molten polymer from escaping. There are many different hot runner system configurations, including drops that are threaded and otherwise fit to the manifold. Different configurations of hot runner manifolds are also common, as shown in Fig. 6.13. The straight-bar manifold (also shown in Fig. 6.11) is among the simplest. Two other common designs i­ nclude the “H” and “X” manifold designs. The “H” manifold provides multiple branches to feed the polymer melt via primary, secondary, and even tertiary runners located on the centerline of the manifold as similar to the design shown in Fig. 6.6. The “X” manifold uses a more direct design in which all primary runners emanate directly from the center of the manifold at the hot sprue bushing. This design typically provides for more efficient material utilization. If multiple drops are being fed, multiple manifolds may also be stacked as shown in Fig. 6.13.

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Figure 6.13 Typical hot runner manifold configurations

Hot runner designs have increased in complexity and capability with configurations far more complex that those shown in Fig. 6.13. Section 6.6.2 later presents stack molds in which two or more hot runner systems are stacked in the mold height direction to allow for multiplication of the mold cavities without an increase in clamp tonnage. Hot runner suppliers can also design “seven-leg specials” in which the lengths and branching of a hot runner are custom-designed to achieve special application requirements, typically for multigated parts such as automotive body panels. The mold designer should consult with multiple hot runner suppliers to understand the benefits and issues associated with available hot runner ­systems.

„„6.4 Feed System Analysis While the two-plate mold, three-plate mold, and hot runner mold designs differ significantly in form and function, the design of the feed systems should adhere to basic guidelines as previously discussed. To summarize, the feed system should: 1. Impose a minimal pressure drop, typically no greater than 50 % of the pressure required to fill the mold cavities or 50 MPa; 2. Consume a minimum amount of material, typically no greater than 30 % of the volume of the mold cavities for cold runner molds or 100 % of the volume of the mold cavities for hot runner molds; and 3. Not extend the mold cooling time.

6.4 Feed System Analysis

Historically, many feed systems have been designed with the intent to maintain the same linear velocity as the melt flows through the sprue, primary runner, etc. The melt velocity can be preserved in a branched runner system by setting the diameter of the downstream diameters, Ddownstream , equal to Dupstream

Ddownstream =

ndownstream

(6.1)

where Dupstream is the upstream runner diameter and ndownstream is the number of downstream runners branching off the upstream segment. Example: Consider the feed system layout provided in Fig. 6.5. If the diameter of the base of the sprue is 6 mm, suggest the diameter of the primary runners to maintain a uniform melt front velocity. According to Eq. 6.1, the downstream diameter should be

Ddownstream =

6 mm 2

= 4.24 mm

To validate this solution, the linear melt velocity can be computed in each branch. Assuming a flow rate of 50 cc/s, the linear velocity in the sprue is

vsprue =

50 ´ 10-6 m3 /s = 1.77 m/s æ p 0.006 m 2 ö÷ çç ( ) ÷÷ çç ÷ø è 4

Since the flow branches into two segments, the linear flow velocity in the primary runner is computed as

vrunner =

0.5 ´ 50 ´ 10-6 m3 /s = 1.77 m/s 2ö æ çç p (0.00424 m) ÷÷ ÷ ç èç ø÷ 4 

While this design guideline is simple and seems intuitive, the resulting designs are inferior with respect to the imposed pressure drops and the consumed plastic material. As such, an engineering methodology for feed system design is next presented based on the analysis of the abovementioned three objectives.

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6.4.1 Determine Type of Feed System The type of feed system is often specified as part of the mold quote by the mold designer since it is either obvious or specified by the customer. However, if the type of feed system is uncertain, then the mold designer should consider the requirements of the molding application and the capabilities of the molder. Table 6.2 provides a comparison of the properties for common feed system types. Table 6.2 Feed System Types and Properties Feed system type

Upfront ­investment

Molder ­capability

Material ­efficiency

Cycle ­efficiency

Two-plate cold runner

Lowest

Lowest

Low

Lowest

Three-plate cold runner

Low

Low

Low

Low

Insulated runner

Moderate

Moderate

Moderate

Moderate

Hot runner

High

Moderate

High

High

Stack mold

Highest

High

High

Highest

Some discussion is warranted regarding Table 6.2. First, the upfront investment refers not only to the cost of the mold design and associated components, but also to the time required to manufacture and test the finished mold. For instance, a twoplate mold with two cavities may cost $20,000 and require a few weeks to ­complete. By comparison, a 64-cavity stack mold may cost $1,000,000 and require several months to complete. For many accelerated product development projects, the added time may be as significant an issue as the added cost. Supply chain logistics can also be an issue. For example, a customer may prefer to construct twelve relatively simple molds, each having four cavities. A few molds can then be separately operated in Europe, Asia, and America. While the cycle time and efficiency is not as high as a single hot runner mold with high cavitation, this approach may reduce the initial mold development time, provide redundancy to mold failure, and allow for reduced tact time in the supply chain in response to fluctuations in consumer demand [3]. The capability of the molder is also an issue with respect to the selection of the type of feed system. While all molders are expected to operate two-plate molds, some molders may not be familiar with the proper setup, operation, and maintenance of three-plate molds, insulated runner molds, or hot runner molds. The operation of stack molds, while not significantly more complex than that of a conventional hot runner, may seem daunting to some molders and require auxiliary controllers that are not available. For these reasons, the mold designer should verify the capabilities of the molder if the type of feed system has not been specified. The material and cycle efficiency may be the primary driver to use more sophisticated feed systems. Since the economics are dependent upon the specifics of the

6.4 Feed System Analysis

molding applications, cost estimation should be performed for each feed system type to determine the most appropriate design. As previously discussed in Chapter 2, it may be useful to perform a sensitivity analysis to identify the risk of under- or over-designing the mold for a targeted production volume.

6.4.2 Determine Feed System Layout Section 4.3.1 provided some common layouts for mold cavities. The feed system must be designed to provide the needed amount of melt flow at the proper melt pressures to each of the cavities. For this reason, a number of feed system layouts have become common, including series, branching, radial, hybrid, and custom. Each of these types of feed system layouts is next discussed. A series layout of cavities can most compactly deliver the polymer melt to many in-line cavities through a single primary runner with many subsequent runners leading to individual cavities. Such a scenario is shown in Fig. 6.14. Unfortunately, since the secondary runners branch off at different locations down the length of the primary runner, the pressure drop along the length of the primary runner will cause lower flow rates to be delivered to cavities further from the sprue. This nonuniform flow can be abated somewhat by reducing the diameters of the secondary runners closer to the sprue as shown by the secondaries off the right primary runner in Fig. 6.14. However, such artificial balancing can be difficult to achieve and does not guarantee consistent part quality associated with different dynamics during the post-filling stages of the injection molding process. For these reasons, the series layout of runner systems is not frequently used in precision appli­cations.

Figure 6.14 Series layout of runner system

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6 Feed System Design

By branching the feed system multiple times, the melt flow to multiple cavities can be naturally balanced, as shown in Fig. 6.14. Compared to the series layout, the branched layout consumes significantly more material while also imposing a high pressure drop from the sprue to the cavities. Another problem with naturally ­balanced feed systems is the development of melt temperature imbalances asso­ ciated with the turning of the melt across multiple branches. This effect has been well documented [4, 5] and has led to the development of a “Melt FlipperTM” to ­assist in correcting flow imbalances in naturally balanced systems with multiple branches. For all these reasons, molding applications with a high number of cavities are increasingly utilizing hot runner feed systems to avoid excess material utilization and pressure drops. Melt flow imbalances have also been observed in “H” manifold layouts [6], though the effect is small due to the elevated temperate of the hot runners, and the effect can be avoided altogether through the use of stacked “X” manifolds. Radial layouts of feed systems, in which multiple primary runners emanate from the sprue as in an “X” manifold, are also quite common. The primary benefit of a radial feed system layout is that the flow rates and melt pressures are naturally balanced with only a moderate amount of runner volume. The number of primary runners that can emanate from the single sprue is somewhat limited due to the large size of the primary runners compared to the base of the sprue. To increase the number of primary runners, a disk cavity, or “diaphragm,” may be located at the base of the sprue. This diaphragm can be used to feed many primary runners, as shown in Fig. 6.15, to mold precision components. Compared to the branched layout of Fig. 6.16, this radial layout has a lower feed system volume and provides more balanced flow. However, longer primary runners and more waste is ne­cessary as the size of the cavities increases.

Figure 6.15 Radial layout of runner system

6.4 Feed System Analysis

Figure 6.16 Branched layout of runner system

Mold designers are free to develop the feed system layout to best fit their molding application. As previously discussed, the primary motivation is to provide balanced flow and minimal pressure drops while consuming the least amount of ­material. As such, many feed systems utilize a hybrid of branched and radial ­layouts. One such design is shown in Fig. 6.17, which consists of a branched feed system with primary and secondary runners, which then feeds four separate ­radial feed systems, each with four tertiary runners. Compared to the feed system layouts shown in Figs. 6.15 and 6.16, the hybrid layout of the feed system design utilizes less material while also providing naturally balanced flow.

Figure 6.17 Hybrid (branched-radial) layout of runner system

Many molding application requirements are best fulfilled by custom feed systems that do not comply with any of the previous feed system layouts. For example, many multigated parts require the feed system to deliver melt to different locations across the mold cavity. In such molding applications, there is no reason to adhere

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to either branched or radial or even naturally balanced layouts. Indeed, the mold designer should purposefully choose a feed system layout and specify dimensions that deliver the desired amount of polymer melt at the desired melt pressures to different portions of the mold cavities. One example of a custom layout is shown in Fig. 6.18. This feed system consists of four primary runners. The two longer primary runners feed the polymer melt via four secondary runners to a relatively large part surrounding the feed system. The two small primary runners closer to the sprue are used to feed smaller mold cavities that provide optional components for assembly with the large molding. These secondary runners may be fitted with rotating shut-offs that can be installed in the mold to change the connectivity of the feed system, and thereby produce different combinations of moldings while the mold is in the molding machine.

Figure 6.18 Custom layout of runner system

The performance of the feed system is ultimately determined by the creativity and care of the mold designer according to the requirements of the molding application. The mold designer has significant freedom in the design of the feed system. However, some general guidelines are as follows: ƒƒThe total length of the feed system should be as short as possible to minimize (a) material consumption and (b) pressure drop through the feed system; ƒƒNaturally balanced feed systems provide greater cavity-to-cavity consistency with respect to melt flow, melt pressure, and molded part quality than artificially balanced designs; ƒƒThe total number of branches in a feed system should be minimized to avoid ­excessive runner volume and potential melt temperature imbalances;

6.4 Feed System Analysis

ƒƒTo minimize pressure drop for a given feed system volume, the diameters of the feed system are generally largest with the sprue and subsequently become smaller with the primary, secondary, and other runners with decreasing flow rates; ƒƒEconomic analysis is vital to determine the correct number of mold cavities, the layout of the mold cavities, and the type of feed system; and ƒƒHot runner and three-plate molds should be considered when cavities in a twoplate mold obstruct the desired layout of the feed system.

6.4.3 Estimate Pressure Drops Once the layout and lengths of the feed system have been determined, the diameters of each portion of the feed system should be determined according to analysis. The flow of polymer melt through the feed system is in the laminar flow regime. To verify laminar flow, the Reynolds number, Re, should be less than 2300. Re =

4 × rmelt × Vmelt < 2300 (6.2) p × mmelt × D

where Vmelt is the volumetric flow rate (typically on the order of 50 · 10–6 m3/s), ρmelt is the density (typically on the order of 1000 kg/m3), mmelt is the apparent viscosity (typically on the order of 100 to 1000 Pa·s), and D is the runner diameter (typically on the order of 0.01 m). Substituting typical values for the variables in Eq. 6.2 indicates that the Reynold’s number is on the order of 0.1. As such, the flow regime is far from turbulent, inertial effects are negligible, and the pressure drop, ΔP, can be estimated using the well-known Hagen-Poiseuille equation. DP =

8 × mmelt × L × Vmelt p × R4

(6.3)

where L and R are the length and radius of a portion of the runner. To provide an accurate estimate of the pressure drop using the Newtonian model, the apparent viscosity should be evaluated for the polymer melt at an appropriate shear rate. g =

4V (6.4) p R3

For a power law fluid, the pressure drop can be estimated directly without calcu­ lation of the shear rate as

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ææ ön çççç3 + 1 ö÷÷V ÷÷ 2kL ççèç n ø÷ melt ÷÷÷ ç DP = ÷ (6.5) ø÷ R èç p R3

where k and n are the reference viscosity and power law index of the polymer melt at the melt temperature, respectively. Example: Estimate the pressure drop through the hot runner system design shown in Fig. 6.19 during the molding of the laptop bezel.

Figure 6.19 Dimensions of hot runner feed system The analysis assumes that ABS is molded with a volumetric flow rate at the inlet of 125 cc/s. To avoid calculating the shear rate in each portion of the runner, the power law model is used with k equal to 17,000 Pa·sn and n equal to 0.35. The bore of the hot sprue bushing is 90 mm in length and has a radius of 6 mm. The volumetric flow rate through the hot sprue bushing is 125 cc/s, so the pressure drop through the sprue is ææ 1 ö÷ ÷125 ´ 10–6 m3 ççççç3 + n ç 2 ´ 17,000 Pa·s ´ 0.09 m çè 0.35 ÷ø çç DPsprue = 3 çç 0.006 m p (0.006 m) è

ö0.35 s ÷÷÷ ÷÷ ÷÷ = 5.9 MPa ÷÷ ø

6.4 Feed System Analysis

After the hot sprue bushing, the melt branches into two flow streams. Since the multigated laptop bezel is nearly symmetric, the flow rate through each leg of the hot runner system is assumed to be 50 % of the inlet flow rate, or 62.5 cc/s. Each leg of the manifold is 118 mm in length with a radius of 5 mm, so the pressure drop through the manifold is

DPmanifold

æ ççççæ3 + 1 ÷÷ö62.5´10-6 m3 n ç 2 ´ 17,000 Pa × s ´ 0.118 m çèç 0.35 ø÷ çç = 3 çç 0.005 m p (0.005 m) è

ö0.35 s ÷÷÷ ÷÷÷ = 8.8 MPa ÷÷ ÷ø

To calculate the pressure drop through the nozzle, the most accurate ­estimate may be provided by analyzing each segment of the tapered bore. Given this particular nozzle bore geometry, however, a reasonable estimate may be obtained by modeling the tapered bore as a constant section with a radius of 3.5 mm and a length of 108 mm. The pressure drop is then æ ççççæ3 + 1 ÷÷ö62.5´10-6 m3 2 ´ 17,000 Pa × s ´ 0.108 m ççèç 0.35 ø÷ çç DPnozzle = 3 çç 0.0035 m p (0.0035 m) è n

ö0.35 s ÷÷÷ ÷÷÷ = 16.7 MPa ÷÷ ÷ø

The total pressure drop through the hot runner system is the sum of the pressure drops through each portion of the hot runner system.

DPtotal = DPsprue + DPmanifold + DPnozzle DPtotal = 5.9 MPa + 8.8 MPa + 16.7 MPa = 31.4 MPa » 4500 psi This pressure drop is significant but reasonable compared to typical ­injection pressures of 150 MPa. 

6.4.4 Calculate Runner Volume Given the number, lengths, and radii of the feed system, the total feed system volume, Vtotal , can be computed as m

m

j =1

j =1

(

)

Vtotal = å N j × Vj =å N j × L j p R 2j (6.6)

where m is the number of different types of segments in the feed system, j is an index referring to a specific type of runner segment, Nj is the number of times that the runner segment j occurs in the feed system, Lj is the length of segment j, and Rj

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is the radius of segment j. As previously discussed, the mold designer should minimize the total volume of the feed system to avoid the production of excess material waste or regrind in cold runner molds or long residence times in hot runner molds. Example: Calculate the volume of the hot runner system design shown in Fig. 6.19. In this hot runner design, there is one sprue with a radius of 6 mm and a length of 90 mm. The primary runner in the manifold consists of two ­segments each with a radius and length of 5 mm and 118 mm, respectively. There are two nozzles, each with a bore length of 108 mm and an approximate radius of 3.5 mm. Accordingly, the total volume is calculated as m

Vtotal = å N j × Vj = j =1

2

= 1× 9 cm × p × (0.6 cm)

2

+ 2×11.8 cm × p × (0.5 cm)

2

+ 2×10.8 cm × p × (0.35 cm) = 37 cm3

The volume of 37 cc is slightly larger than the 27.5 cc of the part and should not lead to extended residence times or degradation. Indeed, it may be ­preferable in this thin wall molding application to redesign the hot runner system with slightly larger diameters for an even lower pressure drop. While it may be unnecessary to reduce the volume of the feed system in this hot runner application, the large runner diameters in this design would likely be unacceptable if applied to a three-plate cold runner system. 

6.4.5 Optimize Runner Diameters Once the pressure drop through the feed system is analyzed, it is possible to adjust the feed system design to improve the performance. Multiple iterations of design and analysis may be conducted to obtain a design that provides a low pressure drop while consuming very little material. Multivariate optimization is a numerical technique that could be employed to simultaneously minimize the pressure drop while minimizing the runner volume [7, 8]. However, this approach requires time to implement and validate while suppressing the details of the analysis from the designer.

6.4 Feed System Analysis

The approach recommended here is to utilize constraint-based methodology [9] to directly solve the minimum runner system diameters given a specified constraint on the pressure drop. If the maximum pressure drop for a portion of the runner is specified as ΔPmax , then for a Newtonian material the radius of the runner could be directly solved as R=4

8 × mmelt × L × Vmelt (6.7) p ×DPmax

A difficulty with this approach, however, is that the apparent viscosity, mmelt , is a function of the shear rate and the runner radius. To avoid iterative estimation of the shear rate and viscosity, the power law model can be used to calculate the ­radius in a single step as 1

æ æ 1 ÷ö  ö÷3 + 1 1 ç çç + 3 ÷V ÷ n (6.8) ç çæ 2kL ö÷n èç n ÷ø melt ÷÷÷ ÷÷ R = çççççç ÷÷ p èççè DPmax ÷ø ø÷

An issue remains, however, as to what the maximum pressure drop should be in each segment of the feed system. Knowing the specification on the total pressure drop from the machine nozzle to the cavity, various schemes can be developed to allocate the pressure drop across each portion of the feed system. The simplest approach is to divide the maximum pressure drop for the entire feed system by the number of segments between the nozzle and the cavity. For instance, if the polymer melt flowed through a sprue, a primary runner, and a secondary runner, and the maximum pressure drop for the feed system was 30 MPa, then the mold ­designer could choose to allocate a maximum pressure drop of 10 MPa for each of the segments of the feed system. The problem with this approach, however, is that it does not account for the length of each portion of the feed system. A very short secondary runner, for instance, would be allocated the same pressure drop as a long primary runner. The resulting design would be suboptimal with the diameter being too small for the secondary runner and too large for the primary runner. Another simple approach is to distribute the pressure drop across the feed system in proportion to the length of each runner segment. DPi = DPmax ×

Li

Ltotal

= DPmax ×

Li

m

å Lj j =1

(6.9)

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6 Feed System Design

where ∆Pi is the maximum pressure drop allocated to runner segment i with length Li, and m is the number of runner segments between the inlet and outlet of the feed system. As such, longer runner segments will be allowed a proportionally greater portion of the pressure drop through the feed system. Example: Calculate the minimum diameters in the hot runner system design shown in Fig. 6.19 so that the pressure drop through the feed system does not ­exceed 30 MPa. Assume ABS is molded with the molding machine providing a volumetric flow rate of 125 cc/s. The total length of the feed system from the inlet to the outlet is 3

Ltotal = å L j = 90 mm + 118 mm + 108 mm = 316 mm j =1

The maximum pressure drop through the sprue is allocated as

DPsprue = 30 MPa ´

90 mm = 8.5 MPa 316 mm

Given this pressure drop for the sprue, the sprue diameter can be calculated from Eq. 6.8 as

æ 1 ÷ö 1 æ -6 3 ç ç ççççæ 2×17,000 Pa × s × 0.09 m ÷ö0.35 èçç3 + 0.35 ø÷÷125 ×10 m R = çççç ÷÷ çç p çèè 8.5 ×106 Pa ø÷

1

1 ö s ÷÷÷3 + 0.35 ÷ ÷÷ ÷÷ø

Rsprue = 0.005 m = 5 mm Similarly, the maximum pressure drop through the manifold is allocated as

DPmanifold = 30 MPa ´

118 mm = 11.2 MPa 316 mm

Given this pressure drop for the manifold, the primary runner diameter in the manifold can be calculated from Eq. 6.8 as

æ æ 1 ö÷ -6 3 ç 1 ç çççæ 2×17,000 Pa × s × 0.1188 m ö0.35 èçç3 + 0.35 ÷ø÷62.5 ×10 m ÷÷ R = çççç çèèç ø÷ p 11.2 × 106 Pa Rmanifold = 0.0044 m = 4.4 mm

1

1 ö s ÷÷÷3 + 0.35 ÷÷ ÷÷ ÷ø

6.4 Feed System Analysis

Similarly, the maximum pressure drop through the nozzle is allocated as

DPnozzle = 30 MPa ×

108 mm = 10.3 MPa 316 mm

Given this pressure drop for the nozzle, the nozzle bore diameter can be ­calculated from Eq. 6.8 as

æ æ 1 ÷ö 3 –6 ç 1 ç çççæ 2×17,000 Pa × s × 0.108 m ö0.35 èçç3 + 0.35 ø÷÷62.5 ×10 m ÷÷ R = çççç çèèç ø÷ p 10.3 ×106 Pa

1

1 ö s ÷÷÷3 + 0.35 ÷÷÷ ÷÷ø

Rnozzle = 0.0044 m = 4.4 mm It should not be surprising that the diameter of the runner in the manifold and nozzle are the same since the two runners 1) have the same melt flow rate and 2) were purposefully assigned the same pressure drop per unit length according to Eq. 6.9. The resulting hot runner system design has a volume of 35 cc and a pressure drop of 30 MPa, both of which are about 5 % less than the previous design (which had a volume of 37 cc and a pressure drop of 31.4 MPa). Furthermore, by maintaining the same runner diameter in the manifold and the nozzle, more uniform shear stresses are maintained with a lower likelihood for dead zones and material degradation. To further reduce the runner system volume, it is necessary to specify ­smaller feed system diameters. This action will result in a larger pressure drop through the feed system unless a higher melt temperature, lower ­viscosity material, or lower flow rate is assumed. If a 50 MPa pressure drop through the feed system was specified, then the above optimization methodology can be repeated to achieve the following results.

Rsprue = 4 mm Rmanifold = Rnozzle = 3.5 mm Vtotal = 21.3 cc The mold designer may repeat the analysis to evaluate the volume of the feed system for different pressure drops. Fig. 6.20 provides a plot of the ­volume of the feed system as a function of the maximum pressure drop. To achieve a low pressure drop, larger runner diameters are necessary, which results in a very high volume for the feed system. As the allowable pressure drop increases to 100 MPa, the volume of the feed system decreases ­substantially, though a runner volume of 10 cc remains necessary to convey the melt at a flow rate of 125 cc/s.

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6 Feed System Design

Figure 6.20 Feed system volume as a function of pressure drop and flow rate 

In optimizing the feed system design, the mold designer needs to assume the flow rates during the filling stage and the expected pressured drop. Figure 6.20 also ­indicates how the feed system designs will change with the volumetric flow rates during the filling stage. Lower flow rates will result in lower pressure drops, which in turn allow for a reduction in the radii and volume of the feed system. Since the actual flow rates are determined by the molder after the mold is designed and built, the molder should verify the expected fill time of the cavity with the molder and calculate the expected flow rates through the feed system. If the flow rates are un­ certain, then the mold designer can estimate the linear melt velocity in the c­ avity per Eq. 5.23 and assume that the flow rate is constant throughout the filling stage.

6.4.6 Balance Flow Rates The previous analysis and examples applied to a naturally balanced branching feed system. However, the analysis can also be applied to “artificially balanced” feed system designs for family molds and multigated parts. In these applications, ­different flow rates and pressure drops may be desired for each branch of the feed system. To properly balance the flow rates and melt pressures in a mold with complex cavity geometries, it is necessary to ensure that the polymer melt completes the filling of each portion of the mold at approximately the same time. As such, the first step

6.4 Feed System Analysis

of the analysis is to calculate the desired volumetric flow rate to each cavity, or for a multigated part, in each portion of the cavity. The filling pressure at each gate is then estimated assuming this flow rate. Once the cavity pressures are known, then the pressure drops through each portion of the runner system can be designed with the previously described analysis to yield the desired cavity pressure and flow rate. While this analysis approach is as simple as possible, it does not account for discrepancies in the filling time of the feed system itself. This error is often negligible since the feed system has a small volume compared to the mold cavities for cold runner molds and is already filled for hot runner molds. Even so, the total filling time and pressure of each branch of the feed system and the mold cavities should be evaluated to ensure a truly balanced design; multiple iterations may be needed to achieve an acceptable design. The mold designer should recognize that a truly optimal, balanced mold design is extremely difficult to achieve. Since polymer melts are non-Newtonian and the shear rates vary with runner diameter and flow rates, the imbalance across the feed system is a function of the material properties and the processing conditions. Furthermore, there is no guarantee that a feed system designed to balance the flow rates during the filling stage will also balance the packing pressures during the post-filling stage. As such, the mold designer should strive to reduce the amount of balancing required by the feed system by ensuring the uniformity of the mold ­cavity designs with the understanding that there will be limits to the performance of static feed systems. Example: Artificially balance the feed system in the two-plate cup and lid family mold. ­Assume ABS is molded at its mid-range melt temperature and the cavity filling time is 1 s. First, the pressures required to fill the cup and lid cavities are estimated. For a fill time of 1 s, the flow length from the gate of the cup cavity to the opposite side of the cavity (refer to Fig. 6.5) is approximately 175 mm (equal to the ­diameter of the base plus twice the height of the side wall). The average linear melt velocity will be 175 mm/s (equal to the flow length divided by the filling time) corresponding to a volumetric flow rate of 44 cc/s (equal to the volume of the mold cavity divided by the filling time). Using the power law model for ABS, the pressure required to fill the cup cavity according to Eq. 5.22 is

æ æ ö0.35 çç 2çç2 + 1 ö÷÷0.175 m/s ÷÷ ÷÷ 2 ×17,000 Pa × sn × 0.175 m çç çè 0.35 ÷ø ÷÷ = 18.2 MPa çç DPcup = è ø÷ 0.003 m 0.003

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6 Feed System Design

The lid is not as deep as the cup, so the flow length for the lid is approximately 110 mm. Assuming the same 1 s filling time, the linear melt velocity for the lid is 110 mm/s with a volumetric flow rate of 19 cc/s. The pressure required to fill the lid cavity is

æ æ ö0.35 çç 2çç2 + 1 ö÷÷0.11 m/s ÷÷ ÷÷ 2 ×17,000 Pa × sn × 0.11 m çç çè 0.35 ÷ø ÷÷ = 16.8 MPa çç DPlid = è ø÷ 0.002 m 0.003 These two filling pressures are quite similar. However, the flow rates are signi­ ficantly different. To achieve the different flow rates, the diameters of the primary runner must be designed to restrict the melt flow to the lid cavity. A first design can be produced by applying Eq. 6.7 using different pressure drops and flow rates for each branch of the feed system. Since the filling pressures are low, a pressure drop of 30 MPa across the primary runner to the cup cavity is assumed with a volumetric flow rate of 44 cc/s. With a length of 38 mm, the resulting radius for the primary runner to the cup cavity is

æ æ 1 ÷ö -6 3 çç 1 ç ççæ 2 ×17,000 Pas × 0.038 m ö0.35 ççè3 + 0.35 ÷÷ø 44 ×10 m ÷ ÷÷ R = çççç ÷ø p 30 ´ 106 Pa èççè

1

1 ö s ÷÷÷3 + 0.35 ÷÷÷ ÷ ø÷

Rrunner_to_cup = 0.0015 m = 1.5 mm The radius of the primary to the lid cavity is similarly computed, but with a ­different pressure drop and flow rate. For the side of the mold including the cup and its runner, the total pressure drop from the edge of the cup cavity to the bottom of the sprue bushing was 30 MPa plus 18.2 MPa or 48.2 MPa. As such, the pressure drop across the primary to the lid cavity will be designed to provide 31.4 MPa, computed as the 48.2 MPa pressure at the base of the sprue minus the 16.8 MPa to fill the lid cavity. The required volumetric flow rate is 19 cc/s. The radius of the runner to the lid cavity is then

æ 1 ÷ö 1 æ -6 3 çç ç çççæ 2 ´ 17,000 Pa × s × 0.038 m ö÷0.35 ççè3 + 0.35 ÷÷ø19 ×10 m ÷ R = çççç ÷÷ø çç p çèè 31.4 ×106 Pa

1

1 ö s ÷÷÷3 + 0.35 ÷÷ ÷÷÷ ø

Rrunner_to_lid = 0.00126 m » 1.25 mm Next, it is necessary to check the fill times through both branches of the feed system. The volumes of the primary runners to the cup and the lid cavities are on the order of 0.3 cc. Since this volume is very small, the filling times are on the order of 0.01 s. Any discrepancies in the filling time of the runners will not significantly affect the filling of the two mold cavities.

6.4 Feed System Analysis

To complete the design, Eq. 6.7 can be used to specify the diameter of the sprue. A reasonable pressure drop across the sprue may be assessed at 20 MPa. For conservation of mass, the flow rate through the sprue is required to be the sum of the flow rates through the primary runners, which total to 63 cc/s. With a length of 76 mm, the radius of the sprue can be com­ puted as

æ æ 1 ÷ö 1 ç3 + ÷63 ×10-6 m3 ççç ç ÷ ç æ ö 0.35 × × × 2 17,000 Pa s 0.076 m è ø 0.35 ÷÷ R = ççççç çèèç ø÷ p 20 ×106 Pa

1

1 ö s ÷÷÷3 + 0.35 ÷÷ ÷÷ ÷ø

Rsprue = 0.0027 m = 2.7 mm Finally, the volume of the cold runner system can be compared to the volume of the moldings to estimate the percentage of regrind, pregrind .

pregrind = pregrind =

Vsprue + Vrunner_to_cup + Vrunner_to_lid Vcup + Vlid 1.7 cc + 0.26 cc + 0.2 cc = 3.5% 44 cc + 19 cc

which is a very low percentage. It should be noted, however, that for a twoplate mold with more cavities (as pictured in Fig. 6.6), or for a three-plate mold (as pictured in Fig. 6.7), the scrap associated with a cold runner system will tend to be substantially greater. The mold designer may use a ­higher pressure drop through the feed system to reduce the feed system ­volume when necessary or recommend a hot runner system to the end-user of the mold to reduce resin costs if high production quantities are planned. 

6.4.7 Estimate Runner Cooling Times For cold runner mold designs, the mold designer should estimate the time required to solidify the cold runner as well as the time required to solidify the cavity. The solidification times can be estimated through one-dimensional heat transfer analysis as discussed in Chapter 9. Table 6.2 provides the cooling time equations for strips and cylinder geometries, where h is the wall thickness of the cavity, D is the diameter of a portion of the feed system, Teject is the specified ejection temperature (usually taken as the deflection temperature under load, or DTUL), Tcoolant is the coolant temperature, and Tmelt is the melt temperature.

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Table 6.2 Equations for Estimation of Cooling Times Geometry

Cooling times

strip

tc =

æ 4 T -T ö h2 çç melt coolant ÷÷ ln ÷ ç p 2 × a çè p Teject - Tcoolant ÷÷ø

tc =

æ T -T D2 ÷ö ç ln çç0.692 melt coolant ÷÷÷ 23.1× a çè Teject - Tcoolant ÷ø

cylinder

During the molding process, the cooling time will be dominated by whatever ­portion of the mold requires the longest time to cool. For this reason, it is not ­necessary to calculate the cooling times for every portion of the feed system and every mold cavity thickness. Instead, the mold designer can simply check the cooling time for the thickest mold cavity section and the largest feed system diameter (usually the diameter at the base of the sprue). If the cooling time of the feed system greatly exceeds the cooling time of the mold cavities, then the mold designer should redesign the feed system to avoid extending the molding cycle time. Example: Verify that the feed system design of the cup and lid family mold will not ­extend the cycle time. Assume that the material is ABS with melt, cooling, and ejection temperatures of 239°C, 60°C, and 96.7°C, respectively. The cooling time for the molding cycle will be dominated by the time to cool either the 3 mm cup or the 5.4 mm diameter sprue. These times are ­estimated as 2

t ccup =

2

(

(0.003 m) -8

p × 8.69 ×10

æ 4 239 - 60 ÷ö ln çç ÷÷ = 18.9 s ç m s è p 97.6 - 60 ø 2

)

2

t csprue =

(0.0054 m)

æ 239 - 60 ö÷ ln çç0.692 ÷ = 26.7 s 97.6 - 60 ø÷ 23.1× 8.69 ×10-8 m3 s çè

(

)

This analysis indicates that the two cooling times are relatively close, but that the cycle may be extended due the solidification of the sprue. However, the feed system does not need to be as rigid as the molded part for ejection. If the mold is opened before the sprue is sufficiently solidified, the feed system may be difficult to eject, either because the sprue has stuck to the A side of the mold or the feed system is overly flexible. To avoid this problem, the diameter of the sprue can be reduced, albeit with a higher pressure drop. 

6.4 Feed System Analysis

6.4.8 Estimate Residence Time For hot runner mold designs, the mold designer should check the residence time of the polymer melt in the hot runner to ensure that the plastic will not degrade. The residence time is directly related to the number of turns required to turn over the polymer melt in the hot runner system, defined as nturns =

Vhot_runner Vcavities

(6.10)

If the volume of the hot runner is large compared to the volume of the mold cavities, then many molding cycles may be required to force new material through the feed system. The number of turns does not represent the actual number of molding cycles required to purge the hot runner of the old resin, but rather the minimum number of molding cycles before a substantial amount of the new resin is delivered to the mold cavities. A high number of turns is undesirable for molding applications in which the color of the resin is frequently changed. To facilitate frequent color changes, mold designers should optimize the feed system diameters and keep the number of turns to a minimum. If the number of turns is less than or close to one, then the use of the hot runner system is unlikely to impede color changes relative to the color change issues associated with purging the injection unit of the molding machine. If the number of turns is large, on the order of 10 or more, then purging the hot runner may become a very significant issue with hundreds (or thousands) of molding cycles required to completely purge low-viscosity resins. A high number of turns is also undesirable for molding applications with resins that have short allowable residence times. The residence time of material in the hot runner system is approximately t residence = (1 + nturns )× tcycle (6.11)

This residence time is approximate, since material flows through the hot runner system at various rates; the polymer melt near the walls and in dead spots of the hot runner can have much longer residence times than those predicted by Eq. 6.11. Furthermore, the mold designer should remember that the material flowing into the hot runner system has already resided in the barrel of the molding machine for a significant amount of time. Accordingly, the mold designer should strive to minimize the number of turns to reduce the residence time and potential degradation of the polymer melt.

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Example: Compute the number of turns and residence time of the hot runner designed in Section 6.4.5. The hot runner design resulting from an allowed 50 MPa pressure drop had a volume of 21.3 cc. Since the volume of the bezel cavity is 27.5 cc, the number of turns is

nturns =

21.3 cc = 0.77 27.5 cc

which is very low. New material will flow through the hot runner system and into the mold cavity with every cycle. In Section 3.4.3, the cycle time was estimated as 13.5 s. The residence time in the hot runner system is estimated as

t residence = (1 + 0.77 cycles)×13.5

s = 24 s cycle

This residence time is very low compared to the allowable residence time of most polymers, which is typically on the order of 15 min [10], though many bioplastics may possess much less thermal stability [11]. 

„„6.5 Practical Issues While this chapter so far has discussed the purpose, types, and analyses of feed systems, there are some practical issues that the mold designer should consider before completing the feed system design.

6.5.1 Runner Cross-Sections The provided analysis applies to “full round” circular runners, since these are extremely common in mold designs and provide for simple analysis. However, other runner cross-sections are also fairly common in practice, since they are easier to machine. In particular, the trapezoidal, round-bottom trapezoid, and half-round runners are often machined into only the moving side of the mold as shown in Fig. 6.21. This mold design strategy not only reduces the amount of machining but also reduces the design time and potential for machining or misalignment mistakes associated with matching the two sides of a full round runner.

6.5 Practical Issues

Figure 6.21 Common runner cross-sections

The primary drawback associated with these noncircular runners is that they give rise to nonuniform shear rates and shear stresses across their cross-section. For example, the trapezoidal runner is easy to machine, but the sections near the four corners conduct very little flow down the length of the runner. The performance of the trapezoidal runner can be improved by rounding the bottom surface to eliminate two of the corners. However, all these noncircular types of runners will need to be slightly larger and consume additional material to provide the same pressure drop as a full round runner. The previously described analysis can be adapted for use with noncircular runner sections. While the results will not be as precise as for a full-round runner, the hydraulic diameter, Dh , for each runner type can be calculated as Dh =

4 × Asection (6.12) psection

where Asection is the cross-sectional area of the runner and psection is the perimeter of the cross-section of the runner. For reference, Table 6.3 provides equations relating the specified dimensions of the different sections in Fig. 6.20 to the hydraulic diameter. It should be noted that the equations in Table 6.3 have been derived assuming a 5 degree taper angle to assist with the ejection of the runner from the mold. This assumption allows for a reduction in the number of design variables.

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6 Feed System Design

Table 6.3 Hydraulic Diameter for Different Runner Sections Runner section (% efficiency)

Equation

Full round

Dh =

(100 %) Trapezoidal

4 × p D2 4 =D pD

4 × W × H + 0.09 × H 2 2 × W + (2.01× H ) If W = H , then Dh » H Dh =

(78.5 %)

Round-bottom trapezoid (87.9 %)

Dh =

(

2

2

4 × 1.57( R1) + 2 × ( R1)× ( H1) + 0.09 × ( H1)

)

5.14 ( R1) + 2.087( H1)

If R1 = H1, then Dh » 2( R1) Half-round

2

(61.2 %)

Dh =

0.5 × p ( R2)

(2 + p)( R2)

= 0.306 × ( R2)

The efficiencies listed in Table 6.3 are defined as

Efficiency =

æ D2 ççp h 4 èç Asection

ö÷ ÷÷ ø÷

(6.13)

The results indicate that the full round runner is the most efficient section design, followed by the round bottom trapezoid, the trapezoid, and the half-round. Example: The primary runner in the three-plate mold of Fig. 6.7 has a trapezoidal ­section. Calculate the pressure drop through a 120 mm length of primary runner with a width of 6 mm, a depth of 8 mm, and a 5 degree taper angle. Assume the use of ABS with a flow rate of 44 cc/s. First, the hydraulic diameter is calculated as 2

Dh =

4 × 6 mm × 8 mm + 0.09 × (8 mm) 2 × 6 mm + 2.01× 8 mm

= 7.04 mm

Then, the pressure drop is calculated using the power law model using the hydraulic diameter as if the trapezoidal runner were circular.

6.5 Practical Issues

DPrunner

ææ çççç3 + 1 ÷÷ö 44 ×10-6 m3 n ç 2 ×17,000 Pas × 0.12 m ççè 0.35 ÷ø çç = 3 çç 0.00704 m p (0.00704 m) çç è

ö0.35 s ÷÷÷ ÷÷ ÷÷÷ = 3.9 MPa ÷÷ ø÷

The dimensions of this trapezoidal design are too large, providing a low pressure drop but consuming excess material and cycle time. The depth and width of the runner should be reduced. 

There is one other runner section that is quite common in hot runner systems: the annulus. Specifically, many hot runner systems incorporate valve pins down the length of the nozzles to physically shut off the gate as subsequently discussed in Section 7.2.9. In this design, the polymer melt flows between a cylindrical drop and the cylindrical valve pin, forming an annulus as shown in Fig. 6.22.

Figure 6.22 Annular section in valve gated hot-runner

The polymer melt flow through an annular section may be closely approximated by adapting the equation for viscous flow in a strip. Specifically, the width of the strip can be replaced by the circumference of the mean diameter of the melt annulus, while the thickness of the strip is replaced by the distance between the valve pin and the nozzle bore. Making these replacements in Eq. 5.17 results in the following relation between pressure drop and flow rate in an annular section for a Newtonian fluid. DP =

12m LV

(

3

)

0.5p ( Dpin + Dbore ) 0.5( Dbore - Dpin )

(6.14)

179

180

6 Feed System Design

where Dpin is the diameter of the valve pin and Dbore is the diameter of the bore through the nozzle. The power law model for an annulus can be similarly derived as æ ö÷n æ ö çç çç2 + 1 ÷÷V ÷ 2 çç ÷÷÷ 2kL èç n ÷ø çç DP = ÷÷ (6.15) 0.5( Dbore - Dpin ) çç 0.5p ( D + D ) 0.5( D - D ) 2 ÷÷ ÷ bore pin bore pin ç èç ø÷

(

)

Example: Calculate the pressure drop through a valve-gated nozzle having a length of 150 mm, a bore diameter of 10 mm, and a valve pin diameter of 5 mm. ­Assume a material with a viscosity of 100 Pa·s flowing at a rate of 50 cc/s. Substituting these values into Eq. 6.14, the estimated pressure drop is

DP =

12×100 Pas × 0.15 m × 50 ×10-6 m3 s 3

0.5p (0.005 m + 0.010 m)(0.5(0.010 m - 0.005 m))

= 24.5 MPa 

6.5.2 Sucker Pins Three-plate mold designs, as shown in Fig. 6.8, often use sprue pullers, or “sucker pins,” to adhere the cold runner system to the stripper plate upon the opening of the mold. In this instance, the use of sucker pins is needed to provide sufficient tensile force along the sprue such that excessive tensile stresses break the gate between the sprue and molding. Without the sucker pins, the cold runner system would travel with the cavity plates and be difficult to remove, since ƒƒthe gates would still be attached and ƒƒthere is no mechanism provided on the A plate to eject the runner system. Similarly, mold designers should consider the necessity of sucker pins during the design of two-plate molds. The primary concern is that the cold runner system may adhere to the A half of the mold due to either vacuum suction to the A plate surface or to the solidification of the plastic melt to the machine nozzle at the top of the sprue. If the cold runner system stays with the stationary side of the mold and all the ejection mechanisms are on the moving side of the mold, then the runner ­system cannot be automatically ejected. The molding machine operator will likely need to delay the molding machine to manually remove the runner system. Furthermore, if the machine is operating on an automatic cycle, then the molding machine may try to close the mold with the runner system still in the mold.

6.5 Practical Issues

To avoid these issues and improve the reliability of the molding operation, sucker pins may be placed at multiple locations along the feed system. Perhaps the most important sucker pin is the sprue puller, located at the bottom of the sprue as shown in Figs. 6.4 and 6.5, which most effectively serves to detach the sprue from the machine nozzle and retain the sprue with the moving side of the mold. In this design, the reverse taper at the bottom of the sprue causes an undercut that retains the sprue. This undercut is later sheared off with the forward actuation of the sprue knock-out pin. Other sucker pins may be placed at various locations along the cold runner system and, if necessary, in the mold cavities. As shown in Fig. 6.23, the design is quite similar to that of the sprue puller. With respect to the design, it is recommended that the diameter of the sucker be slightly less than the diameter of the associated runner to avoid increased cooling times. The height and taper angle of the sucker pin should be sufficient to pull the runner off the stationary side of the mold without excessive material utilization or causing buckling of the associated ejector pins upon forward actuation. Typical heights and taper angles are one half the runner diameter and 5 degrees, respectively.

Figure 6.23 Two sucker pin designs for a cold runner

The implementation shown at the left of Fig. 6.23 merits a brief discussion. In this implementation, an ejector pin has been placed below the runner and slotted with a reverse taper to retain the runner until ejection. Compared with the implementa-

181

182

6 Feed System Design

tion provided at right, the use of the slotted ejector is much simpler to machine and easier to maintain. There are two common issues, however. First, the pin as shown protrudes slightly into the runner section. While this protrusion will not significantly alter the flow rates or pressure drop through the runner, there is a slight chance that it may inadvertently cause an undesired disruption or instability in the flow front. For this reason, it is preferred to align the top of the ejector pin with the bottom of the runner. Second, if multiple slotted ejector pins are used to retain and eject the runner system, then the mold designer should consider the relative alignment of the undercutting slots. If the alignment of the slots are not controlled and provided at random angles, then the runner system may inadvertently bind to the sucker pins at ejection in a random fashion, hampering the adoption of a fully automatic molding cycle.

6.5.3 Runner Shut-Offs Mold designers should consider the use of cold runner shut-offs to provide molders with manufacturing flexibility. Some of the common uses of runner shut-offs include: ƒƒto temporarily shut-off the flow to damaged cavities until mold repair can be ­performed, ƒƒto select different combinations of mold cavities to run in a family mold pursuant to production requirements, and ƒƒto alter the gating and flow in a multigated part. As such, runner shut-offs can be used by the molder to avoid molding defective or undesired parts as well as improve the quality of multigated parts without retooling. The use of runner shut-offs to temporarily seal damaged cavities is somewhat controversial since it 1) requires changes to the molding machine process conditions (especially shot size and injection velocity), and 2) can unbalance or otherwise alter the flow and heat transfer between cavities. Unless the molder re-qua­ lifies the molding process for the new cavity configuration, cavity shut-offs should not be used in commercial production for high precision molding applications. An isometric, exploded assembly view of a runner shut-off design is shown in Fig. 6.24. In this design, a rotating cylindrical insert, item 50, is held between an outer casing, 40, and a retainer, 60, that abuts the back plane of a cylindrical pocket cut into the mold plate. The runner insert may be readily rotated with the mold installed in the molding machine. A spring-loaded ball is used to engage grooves, 54, on the back of the insert to ensure that the insert does not inadvertently rotate during subsequent molding cycles. While the design of Fig. 6.24 provides a runner with a “T” branch, the shut-off assembly comes unfinished and can be provided with various runner configurations, the most common being “T,” “L,” and straight.

6.5 Practical Issues

Commercially available shut-offs are available for approximately $150 for use with runner diameters ranging from 2 mm to 9.5 mm.

Figure 6.24 Design of runner shut-off per U.S. Patent 5,208,053

6.5.4 Standard Runner Sizes When designing the feed system for cold runner molds, the mold designer should specify runner diameters that are machined with readily available cutting tools. The most commonly available sizes for tapered, square, and ball end mills are 1/32", 1/16", 3/32", 1/8", 3/16", 1/4", 5/16", 3/8", 7/16", 1/2", 2 mm, 3 mm, 4 mm, 4.5 mm, 5 mm, 6 mm, 8 mm, 10 mm, and 12 mm. To achieve the most e­ asily produced cold runner design, it may be necessary to round the diameters to ­standard sizes and verify the performance of the design with analysis. However, if ­nonstandard runner sizes provide for less material utilization and more balanced melt flow, then nonstandard runner diameters can and should be specified. Hot runner systems are similarly available with a range of standard bore sizes, typically stepped in 1 or 2 mm increments. The standard diameters and designs

183

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6 Feed System Design

will vary by hot runner supplier. Some suppliers may provide bore diameters of 5, 7, and 9 mm while other suppliers may provide diameters of 4, 6, 8, and 10 mm. Most competitive suppliers will perform flow analysis of the feed system and provide recommendations as to the hot runner technology and sizing. However, the mold designer should verify the appropriateness of the recommendations. As with the specification of custom diameters for a cold runner, many hot runner suppliers will provide custom sizing at an added cost.

6.5.5 Steel Safe Designs The design of the feed system is critical to injection molding but is often uncertain. One common issue is the capability of an available molding machine to fill the mold with a material whose flow characteristics are unknown. Alternatively, there may be uncertainty as to the exact melt flow rates and pressures that are required to properly balance a family mold or complex multigated part. In uncertain situations, the mold designer should specify feed system dimensions that are “steel safe,” which means that the design should call for the removal of less mold steel than may ultimately be required. As such, the mold designer may wish to round the feed system dimensions down one or two standard sizes. By doing so, the mold designer will impose a greater pressure drop and use less material than predicted by the analysis. There is a reasonable chance that the undersized feed system may function properly. Furthermore, if the feed system requires one or more changes, then the “steel safe” design may be easily machined with larger runner sizes to improve the mold performance. Example: Suggest a “steel safe” runner design if the analysis indicated an optimal ­diameter of 4.6 mm for a cold runner. If the feed system analysis resulted in a runner diameter of 4.6 mm, then the mold designer may specify a diameter of 4.5 mm or even 4 mm for a “steel safe” design. By comparison, if the mold designer had rounded up to 5 mm, the design would have provided a lower pressure drop but consumed unnecessary material throughout the mold’s entire lifetime. Furthermore, if the molder desired to reduce the 5 mm diameter, then the mold would require more extensive rework including pocket milling of the old feed system, the manufacture and fitting of an appropriately sized insert, welding and/or the addition of fasteners, and finally the provision of the new, smaller feed system. While this example focused on steel safe design of cold runners, the steel safe concept is also applicable to hot runner designs. 

6.6 Advanced Feed Systems

„„6.6 Advanced Feed Systems 6.6.1 Insulated Runner As previously described, the most common types of feed systems are cold runners and hot runners. Both types of feed systems have disadvantages. With cold runners, there is considerable material waste associated with the formation of the feed system as well as the potential for extended cycle times. With hot runners, there is the additional cost and complexity associated with the temperature control systems, as well as the potential for temperature variations and color change issues. As an alternative to both cold runner and hot runner designs, the insulated runner was designed in an attempt to eliminate these disadvantages. An insulated runner design is shown in Fig. 6.25 [12]. The design layout is very similar to a three-plate mold with a runner section 15, a cavity section 16, and core sections 17. The runner layout is also similar with a sprue 19 conveying the melt through the plate thickness to primary and secondary runners 18 that convey the melt across the parting plane to a second set of sprues, 22 and 23, which convey the melt down to the mold cavities. Compared to a traditional three-plate mold, however, all segments of the feed system are purposefully designed to have large diameters. In addition, the runner section 15 is secured to the cavity section 16 and does not open at all during normal molding.

Figure 6.25 Insulated runner design

During the molding process, the melt is injected from the nozzle of the molding machine and completely fills the feed system. A skin, 18a and 18b, immediately forms on the surface of the runners. However, the solidified skin does not fully propagate throughout the runner, since ƒƒthe thermal conductivity of plastic is very low and ƒƒeach molding cycle conveys heated polymer melt from the molding machine throughout the feed system.

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As a result, the diameter of the molten core remains nearly consistent during the molding cycle. In this manner, the insulated runner can be operated as a hot runner, albeit without any heaters, thermocouples, or temperature controllers. The color change issue is resolved by removing the fully solidified feed system with the release of runner section 15 from the cavity section 16. The design of Fig. 6.25 was specifically intended for the molding of semicrystalline polymers such as polyethylene and polystyrene. Experiments were conducted with runner diameters of approximately 25 mm and cycle times in the vicinity of 60 s; the thickness of the skin was approximately 6 mm. Of course, the optimal specification of runner diameters will depend on the material properties, the melt and mold temperatures, and the flow rates and cycle times. The use of internal heaters and insulating layers (such as the air gaps, 40a and 40b, around the sprue inserts, 39a and 39b, as shown in Fig. 6.25) can provide greater process robustness, albeit with increased design complexity. Perhaps because of these processing uncertainties, the use of insulated runner systems has decreased with the commoditization of hot runners. Even so, insulated runners can provide good performance at low cost; the author suggests that mold designers consider their application in future prototype molding applications.

6.6.2 Stack Molds When the plastic melt is injected into the mold cavity at high pressure, significant clamp force is required to keep the mold closed so that the melt does not escape the mold cavity. Because the clamp force is proportional to the projected area of the mold cavities, the clamp force increases proportionally with the number of mold cavities across the parting plane. However, if the cavities are “stacked” one on top of another, then the clamp force used to close one set of cavities can also be used to close any sets of cavities that are in the stack. One such stack mold design is shown in Fig. 6.26, which was designed to mold two vinyl records with the clamp force and cycle time normally used to mold one vinyl record [13]. In this design, two sets of stampers are mounted between an inner plate, 12, and two outer plates, 14 and 16; the inner plate, 12, is guided by bearings, 20. The melt flows from the nozzle, 54, of the molding machine through extended sprue, 40, to two sets of cavities where the records are formed. After the plastic solidifies, the melt shut-off rod, 65, is actuated to seal the sprue inlet, 51, with the shut-off, 66. This action also connects the sprue, 40, to the chute, 64, such that the sprue may be stripped from the moldings with actuation of the sprue knock-out rod, 75. The molded records are then ejected after retracting the sprue knock-out rod and opening the mold.

6.6 Advanced Feed Systems

Figure 6.26 Early stack mold design

There are two deficiencies in the mold design shown in Fig. 6.26. First, the stack mold requires the formation of a sprue, which is scrap. Second, the melt flow to the two cavities is not balanced, due to the additional length of the sprue to the left cavity. Both these deficiencies are resolved in modern stack mold designs that ­utilize hot runner systems; one such stack mold design is shown in Fig. 6.27 [14]. In this design, a central moving plate, 56, houses two sets of cavities, 60, on opposing parting planes, 62 and 64. A hot manifold, 65, delivers the polymer melt to the cavities through the runner, 70, and subsequent drops. The design uses two single axis valve gates to deliver the melt from the molding machine nozzle, 17, across the parting plane, 62, and to the manifold, 65. During filling and packing stages of the molding process, the actuators, 50 and 54, retract the valve pins, 24, to deliver the melt from the nozzle to the manifold. Otherwise, the valve pins seal the feed system during the plastication, cooling, and mold reset stages. While the stack mold design increases the stack height and complexity of the mold, it enables the molding of two sets of cavities with the same cycle time and clamp force as a single set of cavities. Furthermore, the flow to both sets of cavities is completely balanced and there is no material waste associated with the hot runner feed system. Given the significant part cost reductions afforded by this type of stack mold design, stack molds are now quite common with two, three, and four levels of cavities. Clearly, the stack mold design requires the careful balancing of potential processing cost savings with issues related to investment, maintenance, color change, stack height, and injection volume.

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Figure 6.27 Hot runner stack mold design

6.6.3 Branched Runners A potential issue in “naturally balanced” branched runners, such as those shown in Fig. 6.28 [15], is flow imbalances due to thermal variations caused by the flow and related shearing of the melt [16]. Despite the geometrical balance of the feed system, it has been observed that parts formed in cavities may be larger and heavier depending on their location in the branched feed system. The flow imbalance is created by a nonsymmetrical shear distribution within the laminar plastic melt as it flows through the runner system. Specifically, in the feed system there is a ­distribution of shear rates and temperatures across the radius of the runner: a hot polymer melt at the center of the runner is surrounded by a layer of more highly sheared, hotter, and lower-viscosity plastic melt. When the laminar melt flow reaches a branch in the runner system, the lower-viscosity melt remains in its

6.6 Advanced Feed Systems

outer position, while the more viscous melt at the core is split and flows to the ­opposite side of the branch, 14. This lateral variation in viscosity will cause a nonuniform flow distribution at the next downstream branch, 16 and 22.

Figure 6.28 Branched runner system

To resolve the flow imbalance, it is necessary to eliminate the lateral viscosity ­variation in the polymer melt. One approach is the “melt flipper” design shown in Fig. 6.29 [15] that imposes a level change just prior to the branch. Specifically, the upstream section, 100, of Fig. 6.29 corresponds to the primary runner, 12, of Fig. 6.28 while the downstream section, 104, corresponds to the secondary runner, 14. Prior to the branch, a flow diverter, 106, forces the melt upwards into the runner extension, 102. When the melt subsequently flows down into the runner, 104, the more viscous inner core is directed to the side of the runner that is opposite the level change. Since the viscosity variation is now distributed vertically through the runner, the melt flow is balanced when the downstream runners branch laterally.

Figure 6.29 Level change mold inserts, also known as Melt Flipper

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Figure 6.29 provides a design for a set of inserts to accomplish the level change. The cavity insert, 150, and the core insert, 156, are placed at any necessary junction between the upstream and downstream runners. An indented cavity, 164, and a protruding core, 162, accomplish the level change. Because the viscosity variation is only reoriented and not eliminated, the use of multiple level-changing ­inserts at consecutive runner branches will re-establish the flow imbalances. More recently, research has shown that the flow imbalance and the ability to control the melt flow is related to the melt rheology and the processing conditions [4]. For this reason, additional designs have been developed to adjust the viscosity distribution in the feed system [17].

6.6.4 Dynamic Melt Control Injection molding with static mold components prevents the direct control of melt pressure and temperature across the mold. There are many possible concepts for adding degrees of freedom [18], but one approach is to provide a means for instantaneously modifying the flow resistance in each branch of a runner system. As shown in Fig. 6.30 [19], one design uses a set of strategically located variable-impedance melt valves, each individually controlled with a rapid-response hydraulic actuator. The valves are designed with an adjustable annular clearance, 81, between a tapered valve stem, 45, and a tapered surface, 47, of a bore, 19. Since the resistance to flow is determined by the annular gap between the valve stem and the mold wall, axial displacement of the valve stem can be used to selectively vary the flow rate and pressure drop through each valve. When used in a closed-loop control system, this method can provide simultaneous control of multiple cavity pressures. This system implementation introduces three new characteristics into the molding process. First, the independent control of each valve allows the pressure and flow in multiple regions of the cavity to be decoupled. Previously, changes aimed at improving one area of the part could result in detrimental effects elsewhere in the cavity, since process changes could not be controlled independently. With this process, the flow through each valve can be controlled independently, bringing extra degrees of freedom to the molding process [20]. Second, the capabilities of this system can be leveraged by dynamic re-positioning of the valve within the molding cycle. For instance, this strategy can be used to specify one set of valve positions to profile flow rates in the filling stage, followed by a completely different set of valve positions to profile pack pressures so as to balance filling times and then maximize weld-line strength [21]. Third, the dynamic capabilities of this process allow the valves to be quickly controlled in response to feedback from process sensors in the mold cavity, thus providing closed-loop control of the cavity state variables, which directly determine the product quality [22]. Variation in molding machine input parameters, machine behavior, or material properties can be dynamically compen-

6.6 Advanced Feed Systems

sated to produce consistent parts. Moreover, the control of cavity variables directly enables the use of pressure measurements as a process-control technique for automated detection of quality problems. This could eliminate the need for manual inspection of part quality in many circumstances. Since the dynamics of the molding machine are decoupled from the cavity, details of molding machine performance are made less significant.

Figure 6.30 Dynamic feed control

The size, complexity, and cost of a closed-loop melt control system can raise significant barriers to implementation in many molding applications. To reduce the cost and complexity of the system, a self-regulating valve design [23] was developed, as shown in Fig. 6.31 [24], to work with an open-loop control system design and not require any melt pressure transducers. The melt entering the flow channel, 14, will flow into the aperture, 20b, and around the head of the valve pin to apply a dynamic force, 24, against the projected area of the valve pin, which will tend to shut-off the melt flow and reduce the melt pressure. At the same time, an opposing control force, 26, is applied to the stem, 16, of the valve pin, which will tend to ­increase the melt flow and the melt pressure. As a result, the pin will move until equilibrium is established between the dynamic force, 24, and the control force, 26. In other words, any difference between the control force, 26, and the dynamic force, 24, will cause movement of the control member 16 until the control force and the dynamic force equilibrate, thereby regulating the melt pressure.

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Figure 6.31 Self-regulating valve design

The outlet melt pressure will be approximately equal in magnitude to the control force divided by the projected area of the valve. Previous research [25] has shown that shear stresses and pressure drops along the length of the valve pin, 16, may contribute to errors in the output melt pressure of a few percent. If the control force is provided via a hydraulic or pneumatic cylinder, then the output melt pressure is equal to the pressure supplied to the actuator times an intensification factor, typically on the order of 100 : 1, as determined by the ratio of the push area of the actuator to the area of the head of the valve pin. By controlling the actuation pressure to each valve cylinder during the molding process, the molding process can be made more consistent and more flexible compared to conventional injection molding while not requiring cavity pressure transducers or a closed loop control system.

„„6.7 Chapter Review The selection of the type of feed system is one of the most critical decisions in a mold’s design, since it determines the type of mold and largely impacts the mold’s purchase and operating costs. Two-plate cold runner molds are the simplest design, are readily produced, and can be quite effective for a small number of cavities. Three-plate and hot runner mold designs provide for increased flexibility in the feed system design and are more suitable for a greater number of cavities and/

6.7 Chapter Review

or gates. Of all the designs, the hot runner mold provides the least pressure drop, least material utilization, and fastest cycle times. However, the hot runner system requires a significant up-front investment and greater molder capability, and can impede production of small batches of moldings. All feed systems should minimize the feed system length to reduce both material utilization and pressure drops. The optimization of the diameters along the feed system requires a trade-off between the pressure drop and volume of the feed system. Smaller diameters provide for less material consumption but higher pressure drops. If the pressure drop through the runner is too high, then the molding ­machine may not be able to complete the filling of all the mold cavities with the available injection pressure. For this reason, the mold designer should perform analysis appropriate to the molding application and provide a “steel safe” feed system design that may be readily altered if needed. When possible, feed system designs should be naturally balanced by using radial, branching, or hybrid layouts. The artificial balancing of melt flow rates, in a family mold or complex multigated part, for example, can be accomplished by using different diameters to purposefully impose different pressure drops and flow rates through each branch of the feed system. Depending on the molding application, shut-offs may be placed at multiple junctions in the runner system to direct the flow to different combinations of runners and mold cavities. After reading this chapter, you should understand: ƒƒThe objectives to be considered in feed system design, including melt conveyance, minimizing pressure drop, minimizing material consumption, and balancing melt flow rates and/or pressures; ƒƒThe form, function, advantages, and disadvantages of two-plate, three-plate, and hot runner mold designs; ƒƒThe different layouts of feed system designs, including series, branching, radial, hybrid, and custom designs; ƒƒHow to analyze pressure drop in a feed system using the Newtonian and power law models; ƒƒHow to optimize the feed system diameters to reduce material consumption without imposing excessive pressure drops; ƒƒHow to artificially balance the melt flow rates in a multigated or multicavity mold; ƒƒHow to estimate the cooling time of a cold runner; ƒƒHow to estimate the residence time in a hot runner; ƒƒHow to select the runner cross-section, calculate the hydraulic diameter, and ­estimate the pressure drop in a feed system with a noncircular section; ƒƒHow to use sucker pins in two-plate and three-plate mold designs;

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ƒƒWhen and how to use runner shut-offs; ƒƒHow to adjust analysis results to provide standard and “steel safe” feed system designs; and ƒƒThe potential for insulated runners, melt flippers, and dynamic melt control systems to improve quality or competitiveness. Chapter 5 focused on the filling analysis and design of the mold cavity. This chapter focused on the filling and design of the feed system. The next chapter connects the feed system to the mold cavity through the design and analysis of gates. Afterwards, the book moves away from the mold filling system to other mold sub­systems including venting, cooling, ejection, and others.

„„6.8 References [1] Liao, S., et al., Optimal process conditions of shrinkage and warpage of thin-wall parts, Polym. Eng. Sci. (2004) 44(5): pp. 917–928 [2] Mattis, J., et al., A framework for analyzing energy efficient injection-molding die design. in Electronics and the Environment, ISEE-1996, Proceedings of the 1996 IEEE International Symposium, IEEE (1996) [3] Azevedo, S. G., H. Carvalho, and V. Cruz-Machado, A proposal of LARG supply chain management practices and a performance measurement system, International Journal of e-Education, e-Business, e-Management and e-Learning (2011) 1(1): pp. 7–14 [4] Takarada, K. R., et al., The Effect of Primary Runner Length on Fill-Imbalance in a Geometrically ­Balanced Eight Cavity Polymer Injection Mold, in SPE ANTEC (2006) [5] Tsai, K.-M. and Y.-W. Lin, Quality improvement of optical lenses using innovative runner design, Int. J. Manuf. Res. (2013) 8(4): pp. 337–356 [6] Yang, S. Y., et al., Study on Flow Imbalance during Filling a Multi-Cavity Mold Using a H-type Runners, Key Eng. Mater. (2008) 364: pp. 1306–1311 [7] Irani, R. K., S. Kodiyalam, and D. O. Kazmer, Runner system balancing for injection molds using ­approximation concepts and numerical optimization, in Proc. 18th Annual ASME Des. Autom. Conf., Montreal, Sept. (1992) [8] Li, C. and Y. Shen, Optimum design of runner system balancing in injection molding, International Communications in Heat and Mass Transfer (1995) 22(2): pp. 179–188 [9] Finger, S. and J. R. Dixon, A review of research in mechanical engineering design. Part I: Descriptive, prescriptive, and computer-based models of design processes, Res. Eng. Des. (1989) 1(1): pp. 51–67 [10] Capone, C., et al., Thermal and mechanical degradation during polymer extrusion processing, Polym. Eng. Sci. (2007) 47(11): pp. 1813–1819 [11] Leroy, E., et al., Rheological characterization of a thermally unstable bioplastic in injection molding conditions, Polym. Degrad. Stab. (2012) 97(10): pp. 1915–1921 [12] Peters, D. L., Injection molding of plastic materials, in U.S. Patent No. 3,093,865 (1963) [13] Salzman, S., Automoatic stockmold and curing press, in U.S. Patent No. 2,992,455 (1961) [14] Gellert, J. U., Stack injection molding melt transfer system, in U.S. Patent No. 4,212,626 (1980) [15] Beaumont, J. P., Method and apparatus for balancing the filling of injection molds, in U.S. Patent No. 6,077,470 (2000)

6.8 References

[16] Beaumont, J. P., J. H. Young, and M. J. Jaworski, Mold Filling Imbalances in Geometrically Balanced Runner Systems, J. Reinf. Plast. Compos. (1999) 18(6): p. 572 [17] Beaumont, J. P., Adjustable Melt Rotation Positioning Device and Method, WO Patent WO/2006/124,940 (2006) [18] Kazmer, D. O., Axiomatic design of the injection molding process, Proceedings of the ICAD2000, Cambridge (2000) pp. 123–129 [19] Kazmer, D. O. and M. D. Moss, Manifold system having flow control, in U.S. Patent No. 6,361,300 (2002) [20] Kazmer, D. O. and P. Barkan, Multi-Cavity Pressure Control in the Filling and Packing Stages of the Injection Molding Process, Polym. Eng. Sci. (1997) 37(11): pp. 1865–1879 [21] Kazmer, D. O. and D. Roe, Exploiting melt compressibility to achieve improved weld line strengths, Plast., Rubber Compos. Process. Appl. (1998) 27: pp. 272–278 [22] Kazmer, D. and P. Barkan, The process capability of multi-cavity pressure control for the injection molding process, Polym. Eng. Sci. (1997) 37(11): pp. 1880–1895 [23] Kazmer, D., et al., Design and performance analysis of a self-regulating melt pressure valve, Polym. Eng. Sci. (2006) 46(4): pp. 549–557 [24] Kazmer, D. O., Methods and devices for melt pressure regulation, in U.S. Patent Application No. 2005/0255187 (2005) [25] Kazmer, D., V. Kudchadkar, and R. Nageri, Validation of moulding productivity with two self-regulating melt pressure valves, Plast., Rubber Compos. (2004) 33(9–10): pp. 446–451

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Gating Design

Gates provide the important function of connecting the runner to the mold cavity, and initiating the flow of the melt into the cavity. There are many different types of gates, with the most common types of gates being the edge and pin-point gates. Referring back to the cost estimation of Section 3.3, gating represents a small portion of the mold cost but has a significant impact on the operation of the mold. Knowing when to use what type of gate, and how to properly dimension the gate(s), can reduce the likelihood of mold re-work.

„„7.1 Objectives of Gating Design 7.1.1 Connecting the Runner to the Mold Cavity The primary function of the gate is to connect the runner to the mold cavity, so that the polymer melt can flow throughout the cavity to form the molding. While this is a simple function, the design of the gate provides a means by which the flow of the melt can be fine-tuned through the adjustment of its location or dimensions.

7.1.2 Provide Automatic De-gating For the economical use of an injection mold, the gate and runners should be ­automatically disconnected from the molding at the time of ejection. Otherwise, an operator may need to handle the moldings to remove the gate and runner from the molded part with a gate cutter. Such use of an operator clearly imposes a higher labor cost for the molder. Furthermore, the handling and de-gating of moldings by the operator can also limit the cycle time and induce defects into the moldings. There are three common approaches to providing automatic de-gating. First, it is possible to use the opening action of the mold to separate the moldings from the

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feed system. Such use is common in two-plate molds with tunnel gates or threeplate molds with pin-point gates. Second, it is possible to use a hot runner with ­either thermal or valve gates to completely eliminate the need for post-mold de-gating. Third, it is possible to use robots equipped with cutters to de-gate as an added step in the removal and placement of the moldings. If this third approach is to be used, then the mold designer should discuss alternative gate types and locations with the molder to provide access for pick-up of the molding and de-gating of the feed system.

7.1.3 Maintain Part Aesthetics Since gates are physically attached to the moldings, their removal will leave a witness mark on the surface of the molding. One common approach is to use a very small gate (such as a pin-point gate) in combination with a coarse texture such that the gate vestige is less apparent. Another common approach to resolve this issue is to locate gates on nonvisible surfaces such as underneath a side wall instead of into the side wall. Figure 7.1 demonstrates the relocation of a gate to a nonaesthe­ tic surface. It should be noted, however, that careful gate removal may be required since any significant gate vestige may interfere with mating surfaces in the product assembly.

Gating on side wall

Gating below side wall

Figure 7.1 Relocating gates for improved aesthetics

7.1.4 Avoid Excessive Shear or Pressure Drop Both aesthetics and de-gating suggest the need for gates with small dimensions. From a flow perspective, however, small gates can provide excessive shear rates and pressure drops [1]. Some of the resulting defects may include

7.1 Objectives of Gating Design

ƒƒmaterial degradation, ƒƒnonlaminar flow and jetting of the melt into the mold cavity, ƒƒsplay and other visual defects, ƒƒextended mold filling times, and ƒƒshort shots. For these reason, the shear rate should be calculated and verified that it is below the maximum permissible value. While Appendix A provides maximum shear rates for some materials, the mold designer should consult with the material supplier for application-specific data. If the shear rate is permissible, then the pressure drop is usually acceptable as well. However, the mold designer should calculate the pressure drop to ensure that it is not excessive. A typical pressure drop through a gate is on the order of 2 MPa (300 psi), with 6 MPa (900 psi) potentially excessive dependent on the availability of melt pressure to fill the mold cavity.

7.1.5 Control Pack Times Another important function of the gate is to control the post-filling time (generally known as the packing time) of the melt into mold cavity [2]. After the mold is filled with the polymer melt, the molding machine maintains a high melt pressure to force additional melt into the mold cavity to compensate for volumetric shrinkage as the melt in the cavity cools. It is really the gate, and not the molding machine, that determines the necessary packing time for the polymer melt in the cavity. Consider gate designs that are either too small or too large. If the gate is too small, then the melt in the gate will prematurely solidify and prevent the conveyance of additional melt into the mold. As a result, the polymer in the cavity may experience excessive volumetric shrinkage resulting in poor dimensional and aesthetic properties. Conversely, if the gate is too large, then the gate will not solidify in a timely manner. In this case, the molding machine is required to maintain a very long pack time. If a shorter packing time is used, then the melt in the cavity will flow out of the cavity and back into the runner system and the molding machine. As a result, the melt in the cavity may again experience excessive volumetric shrinkage resulting in poor dimensional and aesthetic properties. For these reasons, the theoretical minimum packing time of the gate should be calculated and checked against the expected process parameters. If the packing time is unexpectedly short or long, then the dimensions should be adjusted even if the shear rates and pressure drops were found acceptable.

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„„7.2 Common Gate Designs The most common types of gate designs are next discussed. Many additional kinds of gates exist; the design of the gate types next discussed should be customized to best meet each molding application’s requirements.

7.2.1 Sprue Gate The sprue gate provides the flow of melt from a sprue directly into the mold cavity as shown in Fig. 7.2. The sprue gate is most commonly used in single cavity molds in which the mold’s sprue bushing directly abuts the surface of the mold cavity. The sprue gate itself is the interface between the bottom of the sprue and the top of the cavity. Since it has no length, there is no pressure drop associated with the sprue gate. For the verification of the shear rate, the smallest diameter of the sprue should be used. Given relatively large dimensions of most sprue designs, the pressure drops and shear rates are relatively low such that high flow rates into the cavity can be achieved.

Figure 7.2 Sprue gate design

A significant disadvantage of the sprue gate is the difficulty of de-gating due to its large diameter. While the operator may manually remove the sprue in many applications with a gate cutter, powered cutters are necessary for many applications with large sprue diameters or tough engineering materials. Furthermore, the removal of the sprue gate can leave a large vestige that can interfere with the product usage. In the design provided in Fig. 7.2, a small rim is provided around the perimeter of the base so the cup may sit flat after sprue removal. If such a rim is not desired, then a recess around the sprue gate may be designed as in Fig. 7.3 to provide clearance for the gate vestige.

7.2 Common Gate Designs

Figure 7.3 Recessed gate well around sprue gate

7.2.2 Pin-Point Gate The pin-point gate is a common type of gate used to connect a sprue or runner to the mold cavity via a small cylindrical opening as shown in Fig. 7.4. The pin-point gate is frequently used due to its small size which provides for ease of de-gating and minimal gate vestige. Pin-point gates are often used with three-plate molds having sprues with a reverse taper. Due to the pin-point gate’s small size, the de-gating is readily accomplished upon the opening of the mold as discussed in Section 6.3.2. Pin-point gates are also often used in two-plate molds to connect the runner to the side walls of the mold cavity. Compared to other types of gates, however, the flow of the melt through such a small orifice will incur high pressure drops and shear rates.

Figure 7.4 Pin-point gate designed with inverted sprue

The diameter of the pin-point gate should be specified so as to be large enough to avoid excessive shear rates yet small enough to provide the desired de-gating and aesthetics. The length of the pin-point gate is typically on the order of its diameter,

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and need only be long enough to provide for the manufacturability of the mold. A properly designed pin-point gate will have a reverse taper between the cavity surface and the gate breakpoint as shown in Fig. 7.4. A recess may be provided in a gate well as shown in Fig. 7.3 to provide a clearance for any gate vestige after de-gating. A smooth transition should also be designed between the gate and the sprue or runner.

7.2.3 Edge Gate The edge gate is a very common type of gate used to connect a cold runner to the edge of a mold cavity. The design and redesign of an edge gate for the cup has been previously discussed with reference to Fig. 7.1. Another edge gate design is shown in Fig. 7.5. In this design, the edge gate connects to the inner periphery of the ­bezel’s supporting frame. Since this gate location is internal to the screen assembly, any vestige remaining after the gate removal will not be observed by the enduser of the molding. Therefore, the edge gate can and should utilize the full ­thickness of the adjacent wall section, and need not be gated underneath the lower surface of the frame.

Figure 7.5 Edge gate design

Compared to the pin-point gate, the edge gate has greatly reduced shear rates and pressure drops. The mold designer can select the thickness, length, and width values according to the needs of the application. In general, the thickness of the edge gate, H shown in Fig. 7.5, should be less than the wall thickness of the molding,

7.2 Common Gate Designs

but may approach the thickness of the molding if shear rates are a concern. The width of the gate, W, should be less than the diameter of the runner but wide enough to avoid excessive shear rates. The length of the edge gate should be kept to a minimum, but long enough to provide the molding machine operator access for de-gating with gate cutters.

7.2.4 Tab Gate The tab gate can be considered a variant of the edge gate, in which a tab is per­ manently added to the molding for the purpose of improved gating. For example, the edge gate design in Fig. 7.5 could be problematic since the melt flows from the runner into the thin inner frame of the bezel, which can cause premature freeze-off of the flow and excessive volumetric shrinkage in the surrounding thicker sections. To improve the flow, a tab, rib, or other feature is added to the mold cavity for the sole purpose of gating as shown in Fig. 7.6. In this design, a rib with a thickness, H, shown in Fig. 7.6, equal to the thickness of the nominal thickness of the part, has been provided that connects the runner to the thicker portion of the molding outside the thin inner frame. Since the thickness of the tab gate is greater than the thickness of the thin frame, sink will likely develop on the top surface. However, this issue is not significant in this application since this area is hidden by the screen assembly.

Figure 7.6 Tab gate design

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Tab gates can be extremely effective with respect to cost and molding performance. The key to their effectiveness is to establish potential gating areas where their remnants will not affect the aesthetics or functionality of the resulting moldings. Once such gating areas are established, the mold designer should select whatever tab geometry and dimensions are appropriate for the application. Typically the thickness and width are on the order of the nominal thickness of the part while the length is minimized to connect to the closest available runner.

7.2.5 Fan Gate The fan gate can be considered as another variant of the edge gate, in which the width of the fan gate at the molding exceeds the diameter of the runner. One fan gate design is shown in Fig. 7.7. In this design, the width of the fan gate has been selected to avoid excessive shear rates when the melt flows into the cavity at a high volumetric flow rate. Given the large width of most fan gates, the feed system is typically removed by a powered gate cutter, reciprocating saw, or router.

Figure 7.7 Fan gate design

One common use of fan gates is to provide a linear melt flow from the gate instead of the radial flow that will result with the previous gate designs. A simple molded plaque application with a fan gate is shown in Fig. 7.8. For this design to be effective, two criteria must be met. First, the fan gate must span the width of the molding across which linear flow is desired. Second, the flow resistance across the width of the fan gate must be negligible. Various fan gate geometries have been developed to adjust the flow rates into the cavity across the width of the fan gate. The design shown in Fig. 7.8 is typical, and consists of a simple loft between the circular section of the runner and the rectangular section of the mold cavity.

7.2 Common Gate Designs

Figure 7.8 Fan gate designed for linear flow

7.2.6 Flash/Diaphragm Gate While fan gates are effective, another alternative is the flash gate or film gate. The word “flash” implies a melt flow through a very thin section. Accordingly, the flash gate consists of a thick circular section that is adjacent to a thin rectangular section as shown in Fig. 7.9. During molding, the melt will proceed from the runner into the thick circular section. The thin adjacent section will cause the melt flow to slow, cool, and potentially freeze while the melt fills the thick section. Once the melt hits the end of the thick section, the melt pressure will then increase signi­ ficantly and force the frozen material in the thin section to flow. Since the flow ­resistance along the thick section is small compared to the flow resistance across the thin section, the flash gate provides a nearly linear melt flow to the cavity across its width.

Figure 7.9 Flash gate design

The concept of the flash gate can be extended to a cylindrical geometry to provide a linear melt flow without knit-lines as shown in Fig. 7.10. In this design, a solid thick “diaphragm” is used to convey the melt from the sprue to the inner periphery of the mold cavity. A thinner gate section is then used to ensure a uniform cavity filling and also assist in the removal of the diaphragm from the molding. Even though the geometry of the diaphragm gate is cylindrical, the analysis of the shear

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7 Gating Design

rate and pressure drop are correctly performed according to a strip geometry with thickness, H, and a width equal to the circumference of the diaphragm.

Figure 7.10 Diaphragm gate design

A flash gate can typically be removed by an operator without the need for power assisted cutters by cyclically flexing the molding using the flash gate as a hinge. Due to the geometry of the diaphragm gate, however, power tools or a punch press are typically required for de-gating. Both the flash and the diaphragm gates will leave a witness line, so it is desired to minimize the thickness of the gate; typical thicknesses are on the order of 0.2 to 0.6 mm, or approximately one-third the thickness of the molded part. The geometry and thinness of these gates may seem to impose excessive shear rates and pressure drops upon the melt. However, these gates’ large widths will result in relatively low linear melt velocities even at high volumetric flow rates. As a result, these gates can be effectively designed to provide moderate shear rates and pressure drops.

7.2.7 Tunnel/Submarine Gate With the exception of the pin-point gate used with a three-plate mold, all the preceding gate designs require the removal of the feed system from the molding by some post-molding system (usually the operator). The tunnel gate is a common type of gate that can be considered a variant of the pin-point gate. Its primary advantage is that the tunnel gate provides for automatic de-gating with the actuation of a simple two-plate mold. The design of a tunnel gate for the lid molding is shown

7.2 Common Gate Designs

in Fig. 7.11. Compared to the pin-point gate, the changes appear to be cosmetic with the addition of some turns and tapers. These differences are negligible with respect to the flow of the plastic melt, so the dimensions of the tunnel gate should be determined as previously discussed to provide for reasonable shear rates and pressure drops.

Figure 7.11 Tunnel gate design

At first glance, the tunnel gate seems very similar to the pin-point gate. However, they differ significantly in structure and function. The function of the tunnel gate can be understood by examining the mold design. A cross section through the closed mold with the tunnel gate is shown in Fig. 7.12. The key to the function of the tunnel gate is that the tunnel gate “tunnels” through the cavity insert. As shown in Fig. 7.13, the molding will move away from the cavity insert and stay on the core with the tunnel gate when the mold opens. At the same time, the opening of the mold forces the tunnel portion of the runner to temporarily remain with the cavity insert. The motion of the core insert away from the cavity insert causes the tunnel gate to break at its junction with the molding. The molding and the feed system can then be ejected as in a conventional injection molding process.

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7 Gating Design

Figure 7.12 Section of closed mold with tunnel gate

Figure 7.13 Section of slightly opened mold with tunnel gate

The diameter of the tunnel gate at the cavity should be designed to avoid excessive shear rates and pressure drops. For the tunnel gate to operate reliably, there are two very important angles that must be specified. First, a nominal 45 degree angle should be maintained between the centerline of the tunnel gate and the parting plane to allow for the transmission of shearing stresses to the gate. Second, the tunnel gate should have an included taper angle of 20 degrees to ensure that the tunnel gate does not stick in the mold and that the tunnel gate breaks at the ­junction with the molding. To ensure adequate structural integrity of the cavity undercut, the tunnel gate should be located at least three tunnel diameters off the parting plane.

7.2 Common Gate Designs

The tunnel gate is a clever design since it provides for automatic de-gating without significant investment. The primary risk in application is that the tunnel gate may be improperly designed or wear such that the runner system does not reliably degate from the molding. To assist the de-gating of the tunnel gate from the molded part, the runners should be designed with nearby sucker pins, as previously shown in Fig. 6.1, to retain the runner system on the core side. If the tunnel gates and the runner system remain on the cavity side, then they cannot be removed through actuation of the ejection system. There are several variations of the tunnel gate. Just as the tunnel gate burrows up into the cavity insert, the term “submarine” gate refers to a design variant in which the tunnel gate descends into the core inserts. The actuation of the ejector system and the ejection of the molding of the core then act to break the gate and strip the feed system from the molding. With both tunnel and submarine gates, it is also possible to design extended gates that curve around vertical side walls to gate onto the interior surfaces of the part. Such an extended submarine gate design, also known as a “banana” or “cashew” gate, is shown in Fig. 7.14, though such designs pose additional risk with respect to reliable de-gating.

Figure 7.14 Section of mold with extended submarine gate

7.2.8 Thermal Gate As discussed previously in Chapter 6, the use of a hot runner feed system eliminates the need for the molding and cooling of a cold runner. The design of gates for hot runners varies substantially from those for cold runners. The primary objectives are generally the same regarding the shear rate, pressure drop, and aesthetic requirements. However, thermal gates in hot runners must also provide a solidified plug that prevents the liquefied plastic melt in the hot runner from flowing out of the gate when the mold opens and the solidified plastic near the gate is removed with the molding. One of the most common types of gates used in hot runners is the pin-point t­ hermal gate formed with an internal “torpedo.” This design is shown in Fig. 7.15. In this

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210

7 Gating Design

design, a highly conductive torpedo is inserted into the nozzle near the gate. The purpose of the torpedo is to transmit heat from the nozzle towards the gate and keep the plastic molten internally. Typically, two or more orifices in the torpedo are used to convey the plastic melt in the feed system into the cavity. A thin layer of residual plastic melt is used to insulate the hot torpedo from the cold mold walls. During the filling stage, the melt pressure from the molding machine increases until the pressure within the torpedo forces any solidified plastic between the ­torpedo orifices and the gate into the mold cavity [1]. The melt can then flow from the hot runner nozzle, through the orifices, and into the mold cavity much like a conventional cold runner feed system. When the flow ceases, the heat transfer to the mold will cause the insulating plastic to partially solidify, with the plastic around the tip of the torpedo becoming solidified. When the mold opens, a small annulus of the solidified material will be broken around the torpedo tip. However, a thin solidified layer will remain that prevents the leakage of the melt from the hot runner to the environment.

Figure 7.15 Section of mold with thermal pin-point gate

The thermal pin-point gate is a clever design with respect to its dual use of the plastic to reduce heat transfer and form a solid seal. However, it does have three significant disadvantages. First, pin-point gates typically have a small gate diameter. Just as with conventional pin-point gates for cold runners, the diameter of the thermal gate and its associated orifices must be designed to provide reasonable pressure drops and shear rates. Because of the small orifices, this gate design may not be suitable for shear sensitive or heavily filled materials. The second disadvantage is related to the residence of the insulating plastic. Over time, any stagnant material will degrade with the potential to be pulled into the flow stream and ­contaminate the plastic melt, most typically as black specks in the molded parts. The residence of the insulating plastic can also cause significant issues when the

7.2 Common Gate Designs

molder performs a color change, since even small amounts of residual material may cause color streaking on subsequently molded parts. A third disadvantage of the thermal gate may also arise. Specifically, the solidified layer must be forced from the gate by increased melt pressure at the start of the molding cycle. The magnitude and timing of the melt pressure may vary slightly from gate to gate ­depending on gate tolerances, gate assembly, and gate temperature distribution. While not an issue in most molding applications, these variances may be problematic in precision molding applications. Hot runner suppliers have worked to resolve these issues, but with limited success. For molding applications involving frequent color changes or the use of shear sensitive or heavily filled materials, it is desirable to streamline the flow through the gate and reduce the shear rates. Accordingly, the thermal sprue gate design has been developed. This design is shown in Fig. 7.16. In this design, a nozzle tip is utilized that has a long contact length with the surrounding mold. This allows the gate area at the cavity to cool significantly, such that no insulating layer of plastic is required. An open flow bore within the nozzle and nozzle tip can then guide the plastic melt directly to the cavity. After the melt fills the mold, the entire sprue below the thermal gate solidifies. A set of converging-diverging tapers in the nozzle tip dictates the break point of the sprue, leaving a thin layer of solidified plastic behind to seal the plastic melt.

Figure 7.16 Section of mold with thermal sprue gate

Compared to the thermal pin-point gate, the thermal sprue gate provides an open flow bore with reduced shear rates and pressure drop. Since the shear rates are reduced, the thermal sprue gate is better suited for use with shear sensitive and heavily filled materials. Due to its open flow bore, moreover, the thermal sprue gate typically requires fewer molding cycles when colors or materials are changed. One less apparent but important advantage is that the length of the sprue can be

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7 Gating Design

designed to allow clearance for ribs or other cavity details that emanate towards the feed system and may prevent direct gating with a thermal pin-point gate. There is one significant disadvantage of the thermal sprue gate, however. Since a sprue is molded with the part, it must remain with the part as vestige or otherwise be later detached by the operator or another post-molding process.

7.2.9 Valve Gate Thermal gates are economical and generally suitable for a wide range of molding applications and materials. However, both the pin-point and sprue thermal gates have two potential limitations. First, they rely on a solidified layer of plastic to prevent leakage, and this solidified layer may not be sufficient in a variety of circumstances. Second, thermal gates provide a significant gate vestige that may not be acceptable in many applications. To resolve the limitations of the thermal gate, mechanically actuated valve gates have been developed. One such design is shown in Fig. 7.17. During operation, the valve pin is retracted to provide access to the mold cavity. After the cavity is filled and packed, the valve pin is advanced to seal the gate.

Figure 7.17 Section of mold with valve gate

Valves gates have at least three significant advantages over thermal gates [3]. First, valve gates provide a mechanical seal (steel on steel) and so are more robust with respect to preventing melt leakage. Second, the face of the valve pin presents a mold shut-off surface to the mold cavity when closed and thereby significantly ­reduces the gate vestige. Third, opening and closing of valve pins may be timed to sequentially fill multigated molds so as to eliminate knit-lines [4] or strategically control packing pressures to improve part quality [5].

7.3 The Gating Design Process

Unfortunately, valve gates increase the cost and complexity of the mold. The cost is increased due to the addition of the valve pins, actuators, and much larger top clamp plate to house the actuators, hoses, fittings, and the control system. As such, the cost of a hot runner system with valve gates may be twice the cost of a hot ­runner system with thermal gates. Complexity in operation is also increased, as the operator must correctly connect the hoses and specify timings to coincide with the process settings on the molding machine.

„„7.3 The Gating Design Process 7.3.1 Determine Gate Location(s) The product and mold designer will typically work together to determine the gating location(s) based on many factors including the type of runner system, aesthetic requirements, intended knit-line locations, strength and orientation considerations, and dimensional control. Generally, increasing the number of gates will improve molding productivity and dimensional stability albeit with increased mold complexity and greater number of knit-lines. As an example, consider the screen bezel with the number of gates equal to 1, 2, 3, or 4. For the sake of variation, assume that the designs with 1 and 3 gates use cold runners with edge gates into the internal frame of the bezel while the designs with 2 and 4 gates use hot runners with thermal sprue gates into the back face of the surrounding frame. The resulting filling patterns for the various gating options are depicted in Fig. 7.18. It is observed that for each of the gate designs, the polymer melt flows radially from the gate and then transitions to a linear flow pattern. The knit-line locations are depicted in Fig. 7.18 as thick lines; the number of knit-lines increases proportionally with the number of gates. The flow length is the distance from the gate to the knit-line. In each of the gating designs, the gating locations have been selected so that each of the flow length are approximately equal. It should be understood from the cavity filling analysis of Eqs. 5.17 and 5.22 in Chapter 5 that longer flow lengths will require higher melt pressures. One benefit of using multiple gates is the ability to inject the polymer melt more quickly into the cavity while maintaining the same polymer flow rate and shear rate in the cavity. Accordingly, the simulation depicted in Fig. 7.18 used a 1 s injection time for the single gate design, 0.5 s for the two gate design, 0.33 s for the three gate design, and 0.25 s for the four gate design. The resulting melt pressures for each of the gating designs are shown in Fig. 7.19. It is observed that each of the melt

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7 Gating Design

­ ressure histories are closely aligned throughout much of the cavity filling prop cess, but the designs with fewer gates must have longer filling times and filling pressures in order to drive the polymer melt to the flow length required. The simulation results of Fig. 7.19 closely follow the results obtained by manual analysis with Eqs. 5.17 and 5.22. The single gate design is likely an unacceptable option due to the high injection pressures of 160 MPa. Furthermore, the location of the knit-line so far from the gate will likely result in a very weak knit-line because the packing stage of the molding process will be unlikely to compensate for volumetric shrinkage of the polymer melt as it cools. As a result, the part is likely to have excessive nonuniform shrinkage, warpage, and a very weak knit-line. As indicated in Fig. 7.19, any of the other gating designs will result in much lower melt pressures and improved part quality.

Figure 7.18 Gating options and cavity filling patterns

7.3 The Gating Design Process

Figure 7.19 Melt pressure at gate for different gating options

The selection of the number and location of gates is critical with respect to the mold design and molding productivity, and so should be made with the approval of the product designer. It is often possible to improve the product quality with increased investment in technology or material cost. With respect to the bezel, for example, the appearance of the knit-lines could be improved through the induction heating of the mold surface as discussed in Section 9.4.3. Alternatively, all knitlines could be eliminated by removing the internal window in the bezel where the screen would reside. This area could then be center-gated with a single sprue, and then the internal window recreated by machining around the perimeter of the ­internal frame. While it would incur excessive labor and material costs, such a mold design strategy is technically very simple.

7.3.2 Determine Type of Gate Concurrent with determining the gating location(s), the mold designer must also determine the type of gate(s) to be designed. Often, the selection of a type of gate is obvious once the gating locations are specified. The primary factors that should be considered include the type of runner system, the desired method of de-gating, the allowable level of shear rates through the gate, and the resulting flow that is desired. To facilitate gate selection, Table 7.1 provides a summary of the types and properties of common gates.

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In Table 7.1, the de-gating method refers to the use of the mold action to de-gate the parts, and does not consider automated de-gating via robotics. The stated shear rate regimes are approximate since they are a function of the gate dimensions and process conditions; the levels of low, moderate, and high roughly correspond to shear rates on the order of 10, 40, and a 100,000 reciprocal seconds. Finally, most gates result in a radial flow pattern out of the gate with only the fan, flash, and ­diaphragm gates purposefully designed to provide linear melt flow into the mold cavity. Table 7.1 Gate Types and Properties Gate type

Runner type

De-gating method

Shear rates

Resulting flow

Sprue

Cold

Manual

Moderate

Radial

Pin-point

Cold

Automatic

High

Radial

Edge

Cold

Manual

Moderate

Radial

Tab

Cold

Manual

Moderate

Radial

Flash/diaphragm

Cold

Manual

Moderate

Linear

Fan

Cold

Manual

Low

Linear

Tunnel/submarine

Cold

Automatic

High

Radial

Thermal pin-point

Hot

Automatic

High

Radial

Thermal sprue

Hot

Automatic

Moderate

Radial

Valve

Hot

Automatic

Moderate

Radial

In selecting gate locations and types, the use of mixed hot and cold runners should also be considered. Figure 7.20 demonstrates the use of a hot drop with a thermal gate that feeds four sub-runners and parts arranged in a radial layout. Such a configuration could be used with a hot sprue bushing to save material in a conventional two-plate mold, but is quite common in the hot runner molds with an “X” manifold configuration. Such an application with the design of Fig. 7.20 would ­result in the production of 16 cavities with excellent material turnover in the hot runner, low pressure drop from the molding machine, and efficient material utilization in the cold runner.

7.3 The Gating Design Process

Figure 7.20 Use of hot drop with thermal gate to cold sub-runners

7.3.3 Calculate Shear Rates The shear rates are calculated according to the previously provided equations for flow in strips and cylinders. For reference, the formulae for Newtonian and ­power-law flows are provided in Table 7.2. These formulas are based on the volumetric flow rate rather than the linear melt velocity. The reason is that molding processes are setup with the volumetric flow rate being specified, and that the ­linear melt velocity of the polymer melt through the gate is not readily observed. As such, shear rates in gates and runners are typically calculated as a function of volumetric flow rate. Table 7.2 Shear Rate Equations Geometry

Newtonian

Strip

g =

6V Wh2

g =

4V p R3

Cylinder

Power-law

æ 1ö 2çç2 + ÷÷÷V çè nø g = Wh2

g =

æ ö çç3 + 1 ÷÷V çè n ÷ø p R3

Appendix A provides the material properties and recommended maximum shear rates for some materials. The maximum shear rates should be considered approximate at best, since in most cases these values are taken from general guidelines for various materials [6]. In reality, the maximum shear rates are dependent not just

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on the maximum shear rate, but also the entire thermal and mechanical history of the polymer melt. In many if not most cases, much higher shear rates may be possible than the maximum shear rates listed in Appendix A. Given this dilemma and the ease of increasing the size of gates, it may be desirable for the mold designer to be “steel safe” and specify a smaller gate with the intent that the mold will be tested and the gate sizes increased as necessary to avoid excessive shear rates and fine tune the flow. To calculate the shear rates, the mold designer must specify some initial gate ­dimensions. For thicker gates having low and moderate shear rates (including the sprue, edge, tab, fan, and valve gates), the initial thickness may be set equal to the wall thickness of the molding at the location of the gate. For thinner gates having moderate and high shear rates (including the pin-point, flash, diaphragm, tunnel, submarine, and thermal gates), the initial thickness may be set equal to one-half the wall thickness of the molding. Strip-type gates also require the specification of the width. For flash and diaphragm gates, the width should be set to the edge length along which linear flow is desired. For other strip-type gates, the initial width may be set to twice the gate thickness; this width can be increased or decreased to adjust the shear rates as necessary. Example: Calculate the shear rate through the two edge gates into the cavity of the ­bezel mold as shown in Fig. 7.5 assuming a volumetric flow rate for ABS at the nozzle of 125 cc/s. Since two edge gates are specified, the volumetric flow rate through each gate will be 62.5 cc/s. Assigning the thickness and width of the edge gate to be 0.75 mm and 6.0 mm, respectively, the shear rate is evaluated as:

g =

6 × 12 ×125 ×10-6 m3 s

2

0.006 m × (0.00075 m)

= 111,000 s-1

This shear rate is significantly above the maximum shear rate of 50,000 1/s. An increase in the gate width to 14 mm would bring the shear rate within the specified maximum, but require a change in the gate type to a fan gate. ­Alternatively, the flow rate can be reduced from 125 cc/s at the nozzle to 60 cc/s, which would require a doubling of the filling time. 

7.3 The Gating Design Process

Example: Calculate the shear rate through the pin-point gate in the molding of the cup as shown in Fig. 7.4. Assuming a 1 s fill time and a 44 cc cavity volume, a volumetric flow rate of 44 cc/s is used for analysis. The initial diameter of the pin-point gate is 1.5 mm (one half the wall thickness of the cup). The shear rate is then:

g =

4 × 44 ×10-6 m3 s 3

p (0.0015 m)

= 132,000 s-1

As in the previous example, the shear rates are again excessive. To achieve a specific maximum shear rate at the gate, the gate radius could be solved directly:

R=3

4 × 44 ×10-6 m3 s 4V =4 = 1.03 mm pg max p × 50,000 s-1

which corresponds to a diameter of approximately 2 mm. This larger dia­ meter would leave a larger gate vestige and require greater forces for ­de-gating. It may be reasonable to initially specify the lesser diameter of 1.5 mm, and then increase the diameter if issues with excessive shear rates are encountered. 

7.3.4 Calculate Pressure Drop If the shear rates are within the allowable limits, then the pressure drops are likely acceptable as well. However, the pressure drop through the gate should be calculated to ensure that adequate injection pressure is available to fill the mold cavity; gate dimensions can also be adjusted to tweak the location of weld-lines in the mold cavity [7]. The pressure drops are calculated according to the previously ­provided equations for viscous flow in strips and cylinders. For reference, the formulae for Newtonian and power-law flows are provided in Table 7.3.

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7 Gating Design

Table 7.3 Pressure Drop Equations Geometry Strip

Cylinder

Newtonian

Power-law

DP =

12m LV Wh3

æ æ ön çç 2çç2 + 1 ö÷÷V ÷÷ 2kL çç èç n ÷ø ÷÷÷ çç DP = ÷ h çç Wh2 ÷÷÷ ÷÷ çç è ø

DP =

8m LV pR4

n ææ 1ö ö çççççç3 + ÷÷÷V ÷÷÷ 2kL çè n ø ÷÷ ç DP = ÷ R ççç p R 3 ÷÷÷ ÷ çç ÷ø è

In estimating the pressure drop through gates, it is important to calculate the viscosity at the appropriate shear rate when using the Newtonian model, or alter­ natively use the equations for the power-law model. The pressure drop through the gates may vary from almost zero to several MPa. Pressure drops above 10 MPa are usually indicative of improperly designed gates that are either too thin or too long. Example: Calculate the pressure drop across the fan gates in the bezel mold shown in Fig. 7.7, assuming ABS is molded at midrange processing temperatures with a volumetric flow rate at the sprue of 125 cc/s. The fan gate has an initial round section with diameter of 6.35 mm and ends at the cavity as a rect­ angular section with a width of 14 mm and a height of 0.75 mm. The analysis of fan gates is complicated since the size and shape of its cross-section varies down its length. The fan gate could be broken into a number of small segments each with a different section to accurately cal­ culate the pressure drop. For approximation, this analysis will assume a ­rectangular profile with a width of 10 mm (half way between the starting diameter of 6.35 mm and the ending width of 14 mm), a thickness of 3.5 mm (half way between the starting diameter of 6.35 mm and ending thickness of 0.75 mm), and a length of 10 mm. Using the power-law model for ABS, the pressure drop through the fan gate can be calculated as:

æ æ çç 2çç2 + 1 ö÷÷62.5 ×10-6 m3 n 2 ×17,000 Pas ´ 0.01 m çç çè 0.35 ÷ø ç DP = ç çç 0.01 m × (0.0035 m)2 0.0035 m ç èç DP = 1.9 MPa = 280 psi

ö0.35 s ÷÷÷ ÷÷ ÷÷ ÷÷÷ ø÷

This pressure drop is negligible and requires no change to the gate design. 

7.3 The Gating Design Process

Example: Calculate the pressure drop through the pin-point gate in the molding of the cup as shown in Fig. 7.4 assuming ABS is molded at midrange processing temperatures. The previous analysis for the 1.5 mm diameter pin-point gate indicated a shear rate of 132,000 1/s for a volumetric flow rate of 44 cc/s. The visco­ sity at this shear rate can be calculated from the Cross-WLF model as 5.4 Pa s. The pressure drop is then:

DP =

8 × 5.4 Pa × s × 0.001 m × 44 ×10-6 m3 s 4

p (0.00075 m)

= 1.9 MPa

Even though the shear rate through the pin-point gate was extremely high, the shear thinning of the melt produced a low melt viscosity and an acceptable pressure drop. For the 2 mm diameter pin-point gate, the shear rate of 50,000 1/s yields a viscosity of 11.2 Pa × s. The pressure drop through the larger gate is approximately:

DP =

8 ×11.2 Pa × s × 0.001 m × 44 ×10-6 m3 s 4

p (0.001 m)

= 1.3 MPa

which again is normally acceptable. 

7.3.5 Calculate Gate Freeze Time After the mold cavity is filled with the polymer melt, additional material must be forced into the cavity to compensate for volumetric shrinkage as the melt cools. As the melt in the cavity cools, the melt in the gate will also tend to cool. The frozen skin will propagate from the mold wall to the centerline of the gate. Since no additional melt flow can be supplied to the cavity once the gate freezes, the molder should set up the molding machine to end the packing stage at gate freeze-off and begin the plastication stage. The gate freeze time must be sufficient to allow for packing of the polymer in the mold cavity [8]. The cooling of the melt in the cavity and the use of the associated heat equation will be discussed in the Chapter 9. Using the provided analysis, the bulk melt temperature is plotted as a function of time in Fig. 7.21 for a 2 mm diameter cylindrical gate for ABS at midrange melt and coolant temperatures. The bulk temperature of the polymer in the gate will initially be close to the set melt temperature of 240°C, and then decrease in the post-filling stage as the heat transfers to the colder mold walls. It would eventually reach the mold coolant temperature of 80°C.

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7 Gating Design

°

222

Figure 7.21 Gate temperature and viscosity history

Using the Cross-WLF viscosity model, the apparent viscosity of the polymer melt in the gate at a shear rate of 10 1/s is also plotted in Fig. 7.18. It is observed that the viscosity of the plastic melt is initially low, and then begins to increase exponentially as the temperature decreases. Eventually, the viscosity will increase such that no flow is effectively transmitted through the gate and the packing stage should end. In this case, a viscosity of 100,000 Pa s has been arbitrarily selected as a “no-flow” condition, corresponding to a gate freeze time of about 1.5 s. Appendix A provides the “no-flow temperature” for various materials estimated in this ­manner. For reference, the equations to calculate the gate freeze time are provided in Table 7.4 for rectilinear and strip geometries. These equations will provide the minimum pack times since they assume perfect heat conduction between the melt and the mold wall. In practice, gate freeze times will be longer since these equations do not consider the melt flow through the gate into the cavity and the associated con­ vection of heat that will tend to prevent the freezing of the gate. For these reasons, the duration of the packing stage should be expected to be somewhat longer than those predicted with the equations of Table 7.4. Even so, the equations are useful to provide an estimate of the order of magnitude of the polymer solidification time in the gate.

7.3 The Gating Design Process

Table 7.4 Gate Freeze Time Equations Geometry

Packing time

Strip

ts =

æ 4 T -T ö÷ h2 ln ççç × melt coolant ÷÷ 2 p × a èç p Tno_flow - Tcoolant ø÷÷

ts =

æ ö÷ T -T D2 ln ççç0.692 melt coolant ÷÷ 23.1× a èç Tno_flow - Tcoolant ø÷÷

Cylinder

Example: Estimate the solidification time of the fan gates in the bezel mold shown in Fig. 7.7 assuming ABS is molded at midrange processing temperatures. The fan gate has an initial round section with diameter of 6.35 mm and ends at the cavity as a rectangular section with a width of 14 mm and a height of 0.75 mm. The 0.75 mm thick rectangular section at the end of the fan gate will be first to solidify and determine the solidification time, so: 2

ts =

(0.00075 m) 2

-8

p × 8.69 ×10

æ 4 240 - 60 ö÷ ln çç × ÷ = 1.5 s m s èç p 132 - 60 ø÷ 2

Since the thickness of the molding is the same as the gate thickness, increas­ ing the gate thickness will have no effect on the packing of the material in the mold cavity away from the gate. It is noted that this edge gate design does gate into a thinner section of the mold cavity, which is not recommended. For this reason, a three-plate mold or hot runner mold should be ­con­sidered to provide gating into the thicker 1.5 mm section with a longer ­packing time. 

Example: Calculate the solidification time for 2 mm diameter pin-point gate in the molding of the cup as shown in Fig. 7.4. The solidification time can be estimated as: 2

ts =

(0.002 m)

-8

23.1× 8.69 ×10

æ 240 - 60 ÷ö ln çç0.692 ÷ = 1.1 s 132 - 60 ÷ø m s çè 2

This gate freeze time may be compared to the solidification time of the cup with a nominal wall thickness of 3 mm:

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2

ts =

(0.003 m) 2

-8

p × 8.69 ×10

æ 4 240 - 60 ö÷ ln çç × ÷ = 24 s m s èç p 132 - 60 ø÷ 2

It is likely that the gate will freeze prematurely and the cup may not be ­adequately packed. 

7.3.6 Adjust Dimensions After the shear rate, pressure drop, and gate freeze time have been calculated for an initial design, the type and dimensions of the gate can be modified to improve the design. For aesthetic and de-gating purposes, smaller gate sizes are desired. As the previous examples have shown, however, excessive shear rates may dictate the use of larger gate sizes. These shear rate calculations are dependent upon an assumed flow rate, and this flow rate will not be known until after the molder has optimized the process with the implemented mold. For this reason, the mold designer should assume a reasonable flow rate for analysis, select a type of gate that can be enlarged, and specify dimensions that are “steel safe.” In this way, the size of the gates can be readily increased to reduce shear rates as needed. Gate dimensions are may also be adjusted to tweak the pressure drops and flow rates at different gates. Such fine-tuning may help to balance the melt flow between multiple cavities, or to adjust the flow rates and knit-line locations within a multigated cavity [9]. The extent of the balancing that can be achieved through gate design is extremely limited due to the small size of the gate. To bring about large changes in flow, the gate dimensions must vary by such significant amounts that the shear rates and gate freeze times will vary substantially between gates, causing unintended consequences. For this reason, it may be preferred to change the dimensions of the runners or to use a dynamic flow control technology for melt control as later discussed in Section 16.6.4. Gate dimensions are often adjusted to improve the dimensional control of moldings. When the gate solidification time is significantly less than the packing time required by the melt in the cavity, then excessive volumetric shrinkage may occur. There are four approaches that are used to reduce the volumetric shrinkage. The most common approach used by the molder is to impose a very high packing ­pressure before the gate freezes, such that the residual melt pressure in the cavity will be relieved as the melt shrinks. Unfortunately, this approach can lead to excessive flashing and/or residual stresses. For this reason, a second common approach is to increase the diameter or thickness of the gate to increase the solidification time and provide packing at more moderate melt pressures. A third approach is to

7.4 Chapter Review

use a different (often particle filled) material with lower volumetric shrinkage to achieve the desired part dimensions. A fourth, and most often avoided, approach is to rework the mold to reduce the nominal thickness of the molding.

„„7.4 Chapter Review The gate design process includes the selection of the type of gate and the careful specification of the gate dimensions to balance aesthetics, shear rates, pressure drops, and packing times. The selection of the gate type will primarily be determined by the previous specification of the mold type (two-plate, three-plate, or hot runner), the need to bring about a certain type of flow in the gate and/or cavity, and finally by the desire to provide a robust and fully automatic molding cycle. The optimization of gate dimensions is driven by the tradeoff between small gate sizes (that provide for improved aesthetics and ease of de-gating) and large gate sizes (that provide for lower shear rates and pressure drops). If the specification of the gate dimensions is uncertain, then the mold designer should utilize smaller gate dimensions since they can be more readily increased in size if required after molding trials. After reading this chapter, you should understand: ƒƒThe various types and functions of gates, including sprue gates, pin-point gates, edge gates, tab gates, fan gates, flash gates, diaphragm gates, tunnel gates, submarine gates, thermal pin-point gates, thermal sprue gates, and valve gates, ƒƒThe various requirements and trade-offs associated with gate designs, ƒƒHow to select a type of gate for a given molding application, ƒƒHow to calculate the shear rate for a given gate design, ƒƒHow to calculate the pressure drop for a given gate design, ƒƒHow to calculate the minimum gate freeze time for a given gate design, and ƒƒHow to adjust the gate dimensions to optimize the gate performance. Now that the gates, feed system, and cavity have been analyzed and designed, the next chapter discusses the need for venting the air displaced by the advancing melt during the filling of the mold cavity. Afterward, the mold cooling system is developed.

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„„7.5 References [1] Gava, A. and G. Lucchetta, Numerical Simulation of a PA66 Flow Behaviour in a Hot Runner Gate, in Macromolecular Symposia (2008) Wiley Online Library [2] Ferreira, I., et al., Multidisciplinary optimization of injection molding systems, Structural and Multidisciplinary Optimization (2010) 41(4): pp. 621–635 [3] Spina, R., Injection moulding of automotive components: comparison between hot runner systems for a case study, J. Mater. Process. Technol. (2004) 155: pp. 1497–1504 [4] Javierre, C., et al., Criteria on feeding system design: Conventional and sequential injection moulding, J. Mater. Process. Technol. (2006) 171(3): pp. 373–384 [5] Kazmer, D. O. and D. Roe, Exploiting melt compressibility to achieve improved weld line strengths, Plast., Rubber Compos. Process. Appl. (1998) 27: pp. 272–278 [6] Boronat, T., et al., Influence of temperature and shear rate on the rheology and processability of reprocessed ABS in injection molding process, J. Mater. Process. Technol. (2009) 209(5): pp. 2735–2745 [7] Zhai, M., Y. Lam, and C. K. Au, Runner sizing and weld line positioning for plastics injection moulding with multiple gates, Eng. Comput. (2006) 21(3): pp. 218–224 [8] Leo, V. and C. Cuvelliez, The effect of the packing parameters, gate geometry, and mold elasticity on the final dimensions of a molded part, Polym. Eng. Sci. (1996) 36(15): pp. 1961–1971 [9] Shen, Y.-K., et al., Analysis for optimal gate design of thin-walled injection molding. Int. Commun. Heat Mass Transfer (2008) 35(6): pp. 728–734

8

Venting

Venting is normally a minor aspect of mold design, and frequently neglected until molding trials indicate mold inadequacies related to venting. An understanding of the purpose and function of vents can assist the mold designer to design vents where clearly needed and ensure that the mold may accommodate additional vents when required.

„„8.1 Venting Design Objectives 8.1.1 Release Compressed Air The primary function of the vent is to release the air in the mold that is being ­displaced by the highly pressurized plastic melt. If all the air in the cavity is not removed during the filling stage, then several defects can result. ƒƒFirst, the trapped air can form a highly pressurized pocket in the mold cavity through which the melt cannot flow, forming a short shot in the molded product. ƒƒSecond, the highly compressed, high temperature gas can combust in the presence of the plastic melt, causing a phenomenon known as “dieseling” and a defect known as “burn marks.” If the burn marks appear on an aesthetic surface, the molder should reject the molded part. ƒƒThird, the presence of gas between two converging melt fronts can reduce the part strength due to interference of the air with the two welding melts while also forming v-notches on the surface of the molded part that act as a stress concentration during the part’s end use.

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8.1.2 Contain Plastic Melt Since a lack of venting is associated with several significant defects, many wide and thick vents are desirable at different locations to facilitate the air flow out of the cavity. However, if a vent is too thick, then the polymer melt can seep out of the vent, causing a thin but sharp line of plastic flash to form at the vent locations. In many molding applications, this flash needs to be trimmed by the machine operator using a deburring tool. Such deflashing is undesirable since the operator incurs labor cost yet does not provide 100% consistency. Furthermore, if a molder continues operation with excessive flashing, then the mold’s parting plane can wear and require resurfacing. For these reasons, fewer and smaller vents are preferred.

8.1.3 Minimize Maintenance The use of vents also provides more features on the mold that can require maintenance. Many polymers will off-gas in the molten state, releasing particles that can build up and clog the venting system. Such clogged vents can occur especially quickly with the use of mold release. As a result, the molding process may intermittently develop defects related to a lack of venting. Many molders resolve this issue by incorporating vent cleaning as part of a preventive maintenance program. In any case, the mold designer should strive to design venting systems that require minimal maintenance, and are easy to maintain when required.

„„8.2 Venting Analysis A three step analysis process is recommended for the analysis and design of vents. First, the air displacement rate should be estimated relative to the melt flow rate. Second, the number, type, and location of vents must be assessed. Third, the width, length, and thickness of each vent must be specified. With respect to the thickness selection, the thickness must be greater than some minimum value to ensure adequate venting while also smaller than some maximum amount to avoid excessive flashing.

8.2.1 Estimate Air Displacement and Rate The amount of air displaced will be approximately equal to the volume of the ­injected plastic. The term “approximately” is used here to imply that the air will expand somewhat when contacted by the hot plastic melt. However, the heated air

8.2 Venting Analysis

will also cool somewhat as it flows past the surface of the mold. For these reasons, the analysis here will assume that the volumetric flow rate of the air will equal the volumetric flow rate of the melt. Example: The volumetric flow rate of the melt for the bezel was 125 cc/s. This will be assumed for the air flow rate. 

8.2.2 Identify Number and Location of Vents Next, it is necessary to identify the locations where the venting is needed. These locations may seem obvious, but on closer consideration these locations may not be so trivial to identify. There are generally three different types of locations where venting is necessary, as shown in Fig. 8.1. The first type of vent is required where the melt converges at an edge of the mold’s parting plane or other shut-off surface. The second type of vent is required where two melts converge to form a knit or weld line. The third type of vent is required where the melt converges at a dead pocket in the mold. Each of these scenarios will next be briefly discussed.

Figure 8.1 Vent locations by type

Figure 8.1 suggests many potential locations of gas traps and corresponding vent locations around the bezel’s parting plane and shut-off surfaces. Some of these vents, including the four locations near the gates and the four locations at the

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c­ orners may not be necessary since the melt flow is predominantly radial. Since the flow is radial, the melt should reach the edges of the mold without trapping any air, and continue displacing the air further into the unfilled cavity. Thus, there is no need for a vent at those locations. However, the exact melt front behavior may change slightly and it is not uncommon for the melt to trap gas at locations along the parting line as indicated as Ê in Fig. 8.2. While the vents near the gate and at the corner may be considered as optional, the mold designer may choose to specify vent locations at these locations to avoid potential mold changes later. The vent locations at the end of flow indicated at bottom left of Fig. 8.2 should be included since a significant fraction of the displaced air from the cavity will likely exit here.

Figure 8.2 Vent locations on shut-off surface

The second type of vent is required where two melt fronts converge as Ë in Fig. 8.3. In this case, two concave melt fronts can come together and form an entrapment from which the air cannot escape. As indicated in Fig. 8.3, a vent is therefore required on an internal surface of the mold cavity. Usually, ejector pins are designed to provide such venting functions on the surface of the mold cavity.

8.2 Venting Analysis

Figure 8.3 Vent locations on part interior

The third type of vent occurs at dead pockets in the mold. The exact locations are not always obvious, so three examples are provided as Ì in Fig. 8.4. In the left detail, the melt flows from the cavity surface along the length of the boss, and eventually traps the air at the end of the boss. In the center detail, two melt fronts come together at a rib, pushing the air to the top dead center of the rib. In the right detail, the melt front flows diagonally across a rib. Due to a cutout in the rib, the air can be trapped in this corner of the mold cavity. There are approximately twenty such dead pockets in the bezel design that may require venting.

Figure 8.4 Vent locations in dead pockets

The above discussion and further inspection indicate that there are about three dozen vent locations that the mold designer may wish to consider. It is unlikely that all of these vent locations are necessary. Furthermore, the addition of vents is

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usually a relatively simple operation that can be accomplished after the mold is built and tested. For this reason, it is fairly common for the mold designer to initially specify vents at only the most critical vent locations. Example: For the bezel mold, the mold design will initially specify seven vent locations as indicated in Fig. 8.5.

Figure 8.5 Initial vent locations 

8.2.3 Specify Vent Dimensions Once the number of vents is specified, the rate of air flow through each vent can be calculated. It may seem reasonable to estimate the air flow through each vent as the total volumetric air flow divided by the number of vents. However, this approach would not be conservative. The reason is that the exact location of the end of fill is not known. As such, it is possible that much of the air flow can dispro­ portionately favor any one of the four locations on each side of the part. A more conservative approach is to assume that all the local air flow exits through each available vent. Example: The assumed flow rate of the air for the bezel was 125 cc/s. This total flow rate will certainly be split into two air flows, each with a flow rate of 62.5 cc/s, towards the top and bottom sets of vents. Since the exact flow rate to each vent is unknown, the analysis will assume that each vent be ­designed for a volumetric flow rate of 62.5 cc/s. 

8.2 Venting Analysis

In general, the length and width of the vent are determined by the application geometry. The minimum vent thickness is related to the pressure drop across the vent necessary to release the displaced air. The minimum thickness can be derived using analysis of the air as a laminar, viscous flow according to the Newtonian model previously presented. While air flowing through vents may be better modeled as a compressible, turbulent flow, the following analysis is simpler while also extremely conservative [1]. The pressure drop of a Newtonian fluid in a rectangular channel is: DPair =

12mairVair L 3 Whvent

(8.1)

The minimum thickness of a vent can then be evaluated from the width and length of the vent as: min hvent =3

12mairVair L (8.2) DPairW

where mair is the apparent viscosity of the air equal to 1.8  ×  10−5 Pa  s, V is the ­volumetric flow rate of air through the vent, ∆Pair is the specified pressure drop of air across the vent, and the other variables are the vent dimensions. Example: Evaluate the minimum thickness of a typical vent required to vent the displaced air at low air pressures. A conservative analysis assumes air flow at 100 cc/s through a single vent with a width of 10 mm and length equal to 5 mm. To avoid compressing the gas and increasing increased pressure on the plastic melt, the allowable pressure drop across the vent is specified as one atmosphere (14.7 psi or 0.1 MPa). The viscosity of air at room temperature is 1.8 × 10–5 Pa s. Then, the minimum thickness is: min hvent =3

12×1.8 ×10-5 Pas ×100 ×10-6 m3 s × 0.05 m 0.1×106 Pa × 0.1 m

= 4.8 ×10-5 m = 0.05 mm = 0.001 in

The analysis indicates that a vent thickness of 0.05 mm is sufficient for this case, and could be further decreased if the vent were wider or shorter, or if there was less air flow or a higher pressure drop was tolerable. This result, derived by analysis, is a common recommendation in industry [2–5]. 

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It is again emphasized that the previous analysis and example are conservative since ƒƒthe analysis assumed laminar flow for the air and so suggests higher pressure drops and the need for thicker vents than a turbulent flow, ƒƒthe geometry and process conditions apply to a single, small vent with relatively high air flow, and ƒƒthe assumed viscosity of air at room temperature is higher than would occur if the air were heated by the polymer melt or compression. For these reasons, the minimum thickness of the vent will not generally be a limiting design constraint. The maximum size of the vent is related to the maximum amount of flashing that is tolerable at the vent locations. The formation of flashing in extremely thin channels such as vents is an advanced research topic, requiring transient simulation coupling the heat transfer and fluid flow as well as a highly refined mesh with very small time steps. No simple analytical solution exists. However, for the purpose of discussion only, consider the application of laminar viscous flow. The average volumetric flow rate of the melt during the flashing is: WL h Vflash = flash vent t flashing

Substituting this relation into the Newtonian flow model of Eq. (8.1) and solving for the thickness provides the upper constraint on the thickness of the vent: max hvent =

12m (8.3) L Pmeltt flashing flash

where Pmelt is the melt pressure at the vent inlet. When the melt first reaches the vent, the melt pressure exerted on the edge of the vent is zero. For the purpose of analysis, the melt pressure can be conservatively assumed as the melt pressure ramp rate times the time for the flashing to solidify: Pmelt =

dPmelt × t flashing (8.4) dt

For most injection molding processes, the melt pressure ramp rate is less than 100 MPa/s. The flashing time is related to the solidification time of the polymer melt in the vent.

8.2 Venting Analysis

Example: Evaluate the maximum thickness of a typical vent using Eq. (8.3). Assuming a vent thickness on the order of 0.06 mm, the gate freeze time equations provided in Table 7.4 can be used to estimate that the approximate time for the melt to solidify while flashing is 0.003 s. Given this solidification time, the flashing should solidify by the time the melt pressure reaches:

Pmelt = 100

MPa × 0.003 s = 300,000 Pa s

Since the vent is thin, there will be significant shear thinning so a low visco­ sity of 10 Pa s is assumed. Substituting these values, the maximum thickness of the vent is: max hvent =

12×10 Pas L = 0.4 × Lflash 300,000 Pa × 0.003 s flash

For example, if a flash length of 0.2 mm is allowed, then the maximum thickness of the vent is 0.08 mm. For comparison, the minimum thickness for the vent required to provide adequate air flow is 0.06 mm. If less flashing was desired, then more and wider vents could be used to reduce the required air flow, after which the vent thickness could be reduced to reduce flashing while providing adequate air flow. 

Since the above analysis may be difficult to apply given the requisite assumptions, several recommendations for vent thickness are listed in Table 8.1 from various handbooks. The differences in the recommendations are interesting and explainable in part. The majority of the variance likely stems from the fact that there has been a long term trend in the plastics industry to move to thinner walls, faster ­injection rates, and higher injection pressures; the maximum thickness of the vent decreases with increasing melt pressure. At the same time, material manufacturers have sought to reduce the viscosity of plastic resins while improving structural properties. Accordingly, it should not be surprising that the technical standards for vents change, with thinner vents being recently recommended. Table 8.1  Recommended Vent Thicknesses (mm) Plastic

Glanvill (1965, [2])

Rosato (1986, [3])

Menges (2001, [4])

Low viscosity materials:

0.08

0.1

0.015

0.2

0.3

0.03

PP, PA, POM, PE Medium viscosity materials: PS, ABS, PC, PMMA

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„„8.3 Venting Designs 8.3.1 Vents on Parting Plane The first type of vent to be considered is the vent on the parting plane. These vents are commonly provided as very thin channels directly at the end of flow. Many molds are produced with vents on the parting plane that emanate from the edge of the parting line outwards to a thicker vent “relief” or vent “channel.” Figure 8.6 provides a venting system design for the bezel, in which two vents have been provided on the inside and outside surfaces of the cavity insert. The width of the vent, W, has been made purposefully high to provide for uncertainty in the last area of the melt to fill the cavity. The thickness of the vent, hvent , has been specified as 0.05 mm. The length of the vent, L, is 2 mm, after which the air flows through a 2 mm thick channel to a 3 mm diameter outlet located at the center and top of the insert.

Figure 8.6 Vent design on parting plane

8.3 Venting Designs

While vents should be provided on the parting plane at the end of fill, it is not uncommon for vents to be placed periodically around the periphery of the cavity. For the molding of center-gated cylindrical parts, vents can be placed around the ­periphery of the entire mold cavity as shown in Fig. 8.7. In this design, the cavity for a lid is center gated as in a three-plate or hot runner mold. A vent is placed around the entire periphery of the mold cavity. Given the ample vent width, the vent is specified with a thickness of 0.015 mm and a length of 1 mm. A vent channel connects the vent ring to the side of the insert and subsequent outlets.

Figure 8.7 Vent design around cylindrical part

While the above designs are certainly effective with respect to venting the displaced air, it should be mentioned that they are susceptible to flashing with bending of the mold plates. As will later be discussed in the structural design of Chapter 12, the melt pressure exerts significant forces on the mold cavity and core. Any significant deflection will tend to increase the thickness of the vents and thereby increase the likelihood and amount of flashing. Indeed, the design of Fig. 8.7 may be especially problematic since the outside, bottom surface of the lid is an area

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observed and handled by the end user. The use of an internal vent around the periphery of a stripper plate will resolve this issue as later designed in Section 11.3.4. To avoid excessive flashing and associated maintenance, it is recommended that vents on the parting plane be used sparingly with a thickness on the order of 0.02 mm. If venting is subsequently found to be inadequate, then additional vents can be added or the thickness of existing vents increased.

8.3.2 Vents around Ejector Pins A very common practice is to use the clearance around ejector pins for venting purposes. There are many advantages to this vent design. First, the actuation of the ejector serves to at least partially clear the venting channel between the pin and the core. Second, ejector pins are commonly used and well understood. Since a clearance needs to be specified around the pin to provide a sliding fit anyways, it is economical to specify a clearance suitable for venting. Holes for ejector pins are normally drilled and subsequently reamed. In mold manufacturing, the diametral clearance between the ejector pin and ejector hole is typically 0.13 mm (0.005 in), which leaves 0.065 mm (0.0025 in) thickness on each side for venting. While this is somewhat larger than previously suggested vent thicknesses, this thicker clearance around the ejector pins is recommended for several reasons. First, the clearance is useful to avoid increased sliding friction and ejector pin buckling. Second, ejector pins are usually machined through solid steel, so increased flashing due to parting plane deflection are unlikely. Third, any witness lines associated with flashing at the ejector pins are usually located on non-aesthetic surfaces. Figure 8.8 provides some typical venting design details using ejector pins. Detail B shows an ejector blade and an ejector pin that have been assigned clearance, H, for venting. For both these ejectors, a venting channel has been provided up to 3 mm away from the mold cavity surface, after which the channel tapers down to the nominal bore of the ejector hole. Both of these elements should be present in a good ejector pin design. The larger channel serves to reduce the flow resistance of the air while also assisting in the assembly of the ejector pins to the mold. The ­taper is useful to guide the head of the pin during mold assembly.

8.3 Venting Designs

Figure 8.8 Vent design around ejector pin and blade

The vent length, L, of 3 mm has been chosen for illustrative purposes and is certainly not mandatory. The previous air flow analysis with Eq. 8.2 implies that the standard clearance of 0.05 mm between an ejector pin and its hole will provide significant air flow for venting. For this reason, it is possible to extend the length of the vent to a location that is convenient. For instance, it may be desirable to avoid a large vent channel near cooling lines. As another example, a mold may be more economically produced with the same vent section through the majority of the core insert, tapering to a larger size only where the core insert faces the support plate.

8.3.3 Vents in Dead Pockets For venting gas traps in dead pockets, one approach is to use a mold insert for the purpose of venting. Figure 8.9 shows a design in which a rectangular pocket has been machined in the core insert, into which a vented insert has been placed. As shown in Detail A, the vent only spans the width of the rib where the trapped air is expected. As shown in Details B and C, the vent has thickness, H, of 0.05 mm and

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a length, L, of 2 mm. Afterwards, a wider vent channel has been placed behind the vent. Since there is no ejector pin, there is no need for a smooth transition between the vent and the vent channel.

Figure 8.9 Vent design in core insert

It should be noted that the venting function of the insert provided in Fig. 8.9 could have also been provided by using an ejector blade at the same location. The ejector blade likely could have been provided at lower cost while also facilitating the ejection of the part. As such, venting inserts are not especially common. Another design alternative is the incorporation of a sintered vent, pictured in Fig. 8.10, which is a type of mold insert that can be used for releasing gas in dead pockets [6]. These devices are relatively small in size, typically ranging from 2 mm to 12 mm in diameter, and contain many small vent holes in sizes ranging from 0.03 to 0.1 mm in diameter. Given their small size and non-machinable top surface, sintered vents are best placed with their venting surface flush with flat mold cavity surfaces. Furthermore, sintered vents can require intermittent replacement or maintenance as the micro-channels may clog without any easy method for in-mold cleaning.

8.4 Chapter Review

Figure 8.10 Vent design in core insert

„„8.4 Chapter Review Venting design and analysis is often neglected during mold design, with venting channels placed after the mold is trialed and issues are identified. This approach has some merit since all of the required venting locations may not be known until the mold filling patterns are verified. However, a complete lack of analysis and foresight regarding venting can lead to significant mold defects, time consuming mold changes, and costly product development delays. After reading this chapter, you should understand: ƒƒThe three different types of venting required: 1) around the periphery of the part on the parting plane, 2) internal to the cavity where two or more melt fronts can form a gas trap, and 3) in dead spots where the air cannot escape. ƒƒThe different types of vents that can be designed including those on the parting plane, around ejector pins, and alongside core inserts. ƒƒHow to calculate the thickness of a vent given the required air flow without causing flash. The next chapter examines the mold cooling system, whose purpose is to provide maximum and uniform heat transfer from the hot polymer melt to the mold coolant. Afterwards, the mold’s ejector and structural systems will be designed and analyzed.

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„„8.5 References [1] Munson, B. R., T. H. Okiishi, W. W. Huebsch, and A. P. Rothmayer, Fundamentals of Fluid Mechanics, Wiley Global Education (2012) [2] Glanvill, A. B. and E. N. Denton, Injection-Mould Design Fundamentals, Industrial Press (1965) [3] Rosato, D. V. and D. V. Rosato, (Eds.), Injection Molding Handbook, Van Nostrand Reinhold Com­ pany, New York (1985) [4] Menges, G., W. Michaeli, and P. Mohren, How to Make Injection Molds, 3rd ed., Hanser (2001) p. 409 [5] Tang, S., Y. Kong, S. Sapuan, R. Samin, and S. Sulaiman, Design and thermal analysis of plastic ­injection mould, J. Mater. Process. Technol. (2006) 171: pp. 259–267 [6] Wieder, K. A., Mold vent and method, U.S. Patent 6,827,569, issued Dec. 7 (2004)

9

Cooling System Design

The cooling system is extremely important to the economics and operation of the designed mold, and yet remains one of the most underengineered systems in injection molds. Perhaps the reason for the lack of engineering is that the temperature distribution is not obvious when molding compared to defects related to flow. Improperly designed cooling systems often result in at least two undesirable outcomes. First, cooling and cycle times are much longer than what could have been achieved. Second, significant temperature gradients arise across the mold, causing differential shrinkage and warpage of the moldings. To operate effectively, cooling systems must be carefully designed to manage the heat flow throughout the mold without incurring undue cost or complexity.

„„9.1 Objectives in Cooling System Design 9.1.1 Maximize Heat Transfer Rates 

In steady state heat conduction, the heat transfer rate, Q conduction , is proportional to the thermal conductivity, k, and temperature gradient of the mold, dT/dz: dT Qconduction = k (9.1) dz

There are at least two implications of this equation. First, heat transfer rates are proportional to the thermal conductivity. A highly conductive material like copper or aluminum has a thermal conductivity several times higher than all of the steels, and should be able to reduce the cooling time in injection molds. The second implication is that a temperature gradient is required to transfer heat, which also means that heat transfer rates can be increased by moving the cooling lines closer to the surface of the mold cavity.

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9.1.2 Maintain Uniform Wall Temperature The temperature of the molded parts at the time of ejection is a complex function of the molded part design, cooling line design, material properties, and processing conditions. While high heat transfer rates are desired, an overly aggressive cooling system design can actually cause quality problems. As the cooling lines approach the mold cavity surface, the heat transfer path between the surface and the cooling line becomes more direct. As a result, there can be a great variation in the tem­ perature across the cavity surface unless the cooling lines are also placed very close together. When differential shrinkage and warpage occurs in the molded parts due to variations in the temperature of the moldings, the molder often has no choice but to run longer cycle times and use the mold as a cooling fixture [1, 2]. By running longer cycle times, often with higher mold coolant temperatures, the molder is reducing the heat transfer rate (and its variation), and then allowing the temperature of the molded plastic to fully equilibrate across the mold. The result is a cycle much longer than might have been planned, and which could have been prevented with a better cooling system design without much additional cost.

9.1.3 Minimize Mold Cost Equation 9.1 implies that infinitely high heat transfer rates can be achieved by moving to higher conductivity materials and using many cooling lines very close to the surface. However, there is a point at which further investment in the cooling system reaps no rewards. The reason is that the heat transfer rate becomes limited not by the heat conduction through the mold, but rather by the heat conduction through the plastic as well as the heat convection to the mold coolant. For these reasons, molds made out of highly conductive materials may have a significant ­reduction in the cycle time by improving heat conduction through the mold [3], but not anything near the eight-fold improvement that might be anticipated from these material’s high thermal conductivity values. The key to designing a cost effective mold is to know where to invest. Highly ­conductive materials are extremely effective in some applications and are usually easier to machine, but are not universally best. Similarly, cooling line layouts can range from the very simple to the very complex. Complex cooling line designs ­often require substantial machining, plugging, sealing, fitting, and maintenance. It is important for the mold designer to know when the added cost of a complex cooling system design is justified.

9.1 Objectives in Cooling System Design

9.1.4 Minimize Volume and Complexity A very significant issue in the design of cooling systems is that they often conflict with the placement of other components. While placing many, tightly spaced cooling lines provides for fast and uniform cooling, this design will also result in very little space in the mold to place the ejector systems, runners, fasteners, and other mold components. For this reason, the mold designer should strive to route cooling lines that parallel the geometry of the mold cavities. A smaller cooling line dia­ meter, while more difficult to machine and transferring less heat, may have a lesser impact on nearby components and allow for the use of multiple lines to achieve more uniform cooling.

9.1.5 Maximize Reliability The melt will exert significant pressures on the surfaces of the mold cavities. These forces translate directly into significant stresses within the mold plates and inserts. The mold’s structural integrity is weakened by every cooling line, each of which requires the removal of supporting mold material and also provides a stress concentration. While the impact of the cooling lines on the structure of the mold can be especially acute in molding applications with high melt pressures, the potential impact ­remains significant in most applications since the cyclic loading and unloading of the melt pressure gives rise to failure due to fatigue. Worse, corrosion of the metal by the circulating coolant tends to exacerbate the stress concentrations. Cracks can form, corrode, and propagate through the mold to the cavity and subsequently require repair.

9.1.6 Facilitate Mold Usage Molding machine operators should be able to operate the injection mold with ­minimal information. The number of external connections should be kept to a ­minimum, with preferably two connections (one inlet and one outlet) per mold half. Each of the connections should be labeled “in” and “out” to help the operator avoid forming a dead cooling circuit. Many molders color code supply lines (blue) and return lines (red), so mold setup is facilitated by similarly anodizing or providing colored labels for the mold inlets and outlets. To avoid damage to the cooling system, all external components should be recessed to avoid direct contact with tie bars, work tables, or other items.

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9 Cooling System Design

„„9.2 The Cooling System Design Process Given that there are multiple objectives in the design of the cooling system, it is unlikely that every objective will be simultaneously optimized. The goal is for the mold designer to reach a good compromise, such that fast and uniform cooling is achieved in a cost effective manner. The following seven step analysis is provided to support the mold design decisions that have to be made. Afterwards, some common cooling issues and designs are developed.

9.2.1 Calculate the Required Cooling Time The cooling time is defined as the amount of time required after the mold is filled for the plastic to become sufficiently rigid to eject. The following theory is provided to support the estimation of the theoretical minimum cooling time. In practice, cooling times can be substantially longer than those predicted by analysis for two reasons. First, the following analysis assumes perfect heat conduction between the plastic and the mold, while there is known to be a substantial thermal contact ­resistance between two dissimilar materials. The cooling time can be further increased by thin gaps, which open up between the shrunken molding and the mold walls. Second, the cooling time is often not driven by the rigidity of the part, but rather by quality requirements that may necessitate the molder to extend the cooling time to achieve the specifications. Since there is very little volumetric melt flow of the melt after the mold filling stage, there is very little heat convection. For this reason, the transfer of heat ­between the plastic and the mold is governed by the transient heat conduction equation: ¶T ¶2T = a 2 (9.2) ¶t ¶z

where a is the thermal diffusivity defined as: a=

k (9.3) r × CP

Here, T is the temperature, t is the time, z is the dimension in the thickness direction, k is the thermal conductivity, ρ is the density, and CP is the specific heat. The thermal diffusivity is essentially a measure of a material’s ability to transmit heat relative to its ability to store heat. For more rapid heat transfer, mold materials with higher thermal diffusivity are desired, though these materials (e. g., alumi-

9.2 The Cooling System Design Process

num) tend to have lower structural properties than steel as previously plotted in Fig. 4.26. The transient heat conduction Eq. 9.2 is a second-order, parabolic, partial differential equation. Analytical solutions for transient heat transfer have been developed for simple geometries such as plates and rods. The cooling plastic can be modeled as a finite slab [4] with the temperature of the melt at the centerline of the mold cavity evaluated from a series expansion: ¥ m 4 (-1) Tz =0 (t ) = Tcoolant + (Tmelt - Tcoolant ) å 2m+1 e p m=0

2

p2(2m+1) a 2

h

t

(9.4)

Taking the first six terms, m Î [0,5], in the series, a plot of the plastic’s temperature at the centerline is shown as a function of the cooling time in Fig. 9.1. The temperature of the plastic at the center of the molding is equal to the initial melt temperature at the start of the cooling process. After a brief delay, the melt at the center of the molding begins to cool. Eventually, the plastic will approach the temperature of the mold coolant.

Figure 9.1 History of melt temperature at centerline

To determine the cooling time, it is necessary to provide some criterion that indicates when the molding is rigid enough to be ejected from the mold. One reasonable approach is to consider the modulus of the material, which is a measure of the material to resist deflection [5, 6]. The effective modulus of the material as it cools is shown on the right-hand axis of Fig. 9.1. It can be observed that as the plastic

247

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9 Cooling System Design

melt approaches the mold coolant temperature, the modulus also approaches a steady state value. The temperature at which the material has significant rigidity is related to the heat distortion temperature (HDT) or the deflection temperature ­under load (DTUL) as characterized by standard tests such as ASTM D648. Equation 9.4 can be solved to provide the cooling time as a function of the melt, coolant, and ejection temperatures. For plates, the theoretical minimum cooling time is: tc =

æ 4 T -T ö÷ h2 ç ln çç melt coolant ÷÷÷ (9.5) 2 p × a çè p Teject - Tcoolant ÷ø

where h is the wall thickness, Teject is the specified ejection temperature (usually taken as the DTUL), Tcoolant is the coolant temperature, and Tmelt is the melt temperature. For rods such as cold runners, the theoretical minimum cooling time is: tc =

æ T -T D2 ÷ö ç ln çç1.60 melt coolant ÷÷÷ (9.6) 23.1× a çè Teject - Tcoolant ÷ø

When computing the cooling time, the mold designer should consider the thickest section that is likely to require the longest time to solidify. Example: Estimate the cooling time for the two cavity family mold used to produce the cup and the lid. The nominal thickness of the lid is 2 mm, the nominal thickness of the cup is 3 mm, and the diameter of the primary runner is 6.25 mm. Assuming that the material is ABS with melt, cooling, and ejection temperatures of 239, 60, and 96.7°C, then the cooling times for each of the various portions of the mold are: 2

tclid

=

(

2

(0.002m) -8

p × 8.69 ×10

æ 4 239 - 60 ÷ö ln çç ÷÷ = 8.4 s m s èç p 97.6 - 60 ø

)

2

2

tccup

=

2

(

(0.003m) -8

p × 8.69 ×10

æ 4 239 - 60 ÷ö ln çç ÷÷ = 18.9 s m s çè p 97.6 - 60 ø

)

2

2

tcrunner

=

(0.00476m)

(

-8

23.1× 8.69 ×10

æ 239 - 60 ÷ö ln çç1.6 ÷÷ = 22.9 s ç m s è 97.6 - 60 ø 2

)

These results provide some important insights. First, since the cup and the lid are different thickness, the family mold will be forced to operate at the

9.2 The Cooling System Design Process

much longer cycle time of the cup. If high production quantities are desired with parts of different wall thicknesses, then it may be more economical to use two different molds operating at different cycle times for separately ­producing the different designs. However, such a mold design strategy gives up the benefits of color matching and at-press assembly, which are very significant motivations for using family molds. Second, the cooling time of the runner is greater than that for the cup. In practice, the runner need not be as rigid as the part being de-molded so the required cooling time of the runner may be less than the 22.9 s calculated above. However, the results do indicate that the cycle time can be dominated by the cooling of the sprue and cold runners, so it is important to minimize the runner diameters not just for material savings but also to maintain a productive molding process. 

To validate the above cooling analysis, the transient heat conduction is numerically simulated for a molding with a wall thickness of 3 mm. The simulation ­assumes that the plastic is initially at the melt temperature, and that the tempe­ rature of the plastic:steel interface at the mold wall is always at the coolant temperature. As before, the simulation assumes that the material is ABS with melt, cooling, and ejection temperatures of 239, 60, and 96.7°C, respectively. The temperature of the plastic through the wall thickness of the molding is shown in Fig. 9.2 for various time steps. At the start of cooling, the temperature of the plastic is at the melt temperature. According to this simulation, the temperature of the plastic at the mold wall immediately drops to the mold coolant temperature. As heat is transferred from the plastic to the mold, the temperatures at the outside layers decrease until finally the core approaches the mold coolant temperature.

Figure 9.2 Cooling of plastic, isothermal boundary

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9 Cooling System Design

During cooling, the plastic molding must become sufficiently rigid to withstand ejection forces. As such, the cooling time can be estimated from Fig. 9.2 as the time at which the temperature at the centerline drops below the specified deflection temperature.1 The simulation results shown in Fig. 9.2 confirm the previous analytical results that the cooling time will be approximately 19 s for a centerline ejection temperature of 97°C. These results are for an isothermal boundary condition at the mold wall, meaning that plastic at the mold wall immediately drops to the mold coolant temperature. In reality, the mold steel can’t withdraw heat so quickly. As a result, the adjacent mold steel will increase in temperature. This behavior can be modeled using a convective boundary condition: ¶T hc éëTcoolant - Tmelt ùû = k (9.7) ¶z

where hc is the representative heat transfer coefficient from the melt at the polymer:mold interface to the coolant. Previous research [7] has indicated that the convective heat transfer coefficient in molding is on the order of 1000 W/°C, though the default in some molding simulations is 5000 W/°C [8]. The simulated temperature history with a convective boundary condition is shown in Fig. 9.3, and is quite similar (albeit slower) to the behavior shown in Fig. 9.2. With a convective boundary condition, the plastic at the mold wall requires additional time to approach the mold coolant temperature. This slower rate of heat transfer also limits the cooling at the center of the molding. If the core must reach a specified temperature of 97°C, then the cooling time should be closer to 24 s rather than the 19 s predicted with an assumed mold wall temperature. On a side note, there is a common rule in the plastics industry that the cooling time can be estimated as: é s ù 2 ú (h é mm ùû ) (9.8) tc = 2 ê êë mm2 úû ë

where the wall thickness, h, is evaluated in units of mm.

1

There are different forms of the cooling time equation. The two most frequently used assume either the centerline temperature or the average temperature through the thickness reach the ejection temperature. This book recommends analysis using the centerline criterion for two reasons. First, it is conservative and will yield mold designs that should perform more robustly. Second, this approach is supported by bending theory. Specifically, consider a part that is rigid at the walls but semi-molten at the centerline. Since the plastic at the centerline is not able to transmit the shear stresses from one wall to the opposing wall under ejection loads, the deformation of the molded part will be much higher than if the plastic at the centerline were solidified and able to transmit stress.

9.2 The Cooling System Design Process

Figure 9.3 Cooling of plastic, convective boundary

Example: Use Eq. 9.8 to estimate the cooling time for a molding that is 3 mm in thickness.

é s ù 2 ú (3 mm) = 18 s tc = 2 ê êë mm2 úû This result compares very well to the previous analysis results of 19.2 s ­(assuming perfect conduction) and 24 s (assuming convection of 1000 W/m°C). 

The reason that Eq. 9.8 provides a good approximation is that most thermoplastics have a thermal diffusivity on the order of 9 · 10–5 m2/s, and processing temperatures such that is around 5. Substituting these values into Eq. 9.5 provides: tc =

æ4 ö é s ù 2 h2 ú h é mm ùû ) (9.9) ln çç 5÷÷ = 2.08 ê 2 ú( ë 2 2 ù ÷ ç é ê p × 0.09 ê mm sú è p ø ë mm û ë û

The rule of thumb provided in Eq. 9.8 closely matches the typical heat conduction analysis provided in Eq. 9.9. While Eq. 9.8 is an excellent guideline, it is a good idea to use Eqs. 9.5 and 9.6 to evaluate the cooling time for the specific application’s design, material properties, and processing conditions. Also, it should be noted that Eq. 9.9 provides an estimate of the cooling time, which is roughly half of the cycle time previously estimated by Eq. 3.5.

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9 Cooling System Design

9.2.2 Evaluate Required Heat Transfer Rate Once the cooling time is known, the heat transfer rate or “cooling power” required of the feed system can be calculated. The total amount of heat to be removed by the cooling system, Qmoldings , is: Qmoldings = mmoldings × CP × (Tmelt - Teject ) (9.10)

where mmoldings is the combined mass of the molded parts and any associated cold runners, which can be estimated as the volume of these moldings times their density at room temperature. The cooling power is the defined as the amount of energy that the must be removed per second of cooling time: Qmoldings Qcooling = (9.11) tc

Typically, an injection mold has multiple cooling lines. Assuming that the mold is well designed and each cooling line removes the same amount of heat, then the heat transfer rate per cooling line may be evaluated as: Qcooling Qline = (9.12) nlines

where nlines is the total number of cooling lines in the mold. At this point, the mold designer should recognize that multiple design iterations may be necessary to perform the cooling analysis for different cooling line layouts with varying number of cooling lines. Example: Analyze the power required to cool the cup and lid family mold. The mass of the two moldings totals 62.6 g. If an ABS material is processed at mid-range temperature, then the heat to be removed is:

é J ù ú × (239°C - 96.7°C) = 20,900 J Qmoldings = 0.0626 éë kg ùû × 2340 ê êë kg ×°C úû For the cup and lid family mold designed for a cooling time of 20 s, the total cooling power is:

20,900 J Qcooling = = 1050W 20s

9.2 The Cooling System Design Process

Assuming for now that the cup and lid mold has 4 cooling lines (2 lines per side), then:

1050 W Qline = = 260 W 4 So, each side of the mold with two cooling lines will require an average ­cooling power of 500 W. 

9.2.3 Assess Coolant Flow Rate Any heat removed from the polymer melt in the mold cavity must be carried away by the coolant. As such, the coolant will increase in temperature as it travels through the mold. This temperature increase is not desirable, since the coolant will provide less cooling to the last portion of the mold through which it flows. If the coolant temperature increase is too great, then thermal gradients will arise across the molded part, which may lead to differential shrinkage and warpage. , the increase in the coolant Given a volumetric flow rate of the coolant, V ­temperature along one cooling line is:

coolant

Qline (9.13) DTcoolant =  Vcoolant × rcoolant × CP,coolant

The thermal properties of some common coolants are provided in Appendix C. The required coolant flow rate can be calculated given the allowable increase in the coolant temperature. A typical allowable increase in the coolant temperature is 1°C. For a precision molding application, the allowable increase in the coolant temperature may be 0.1°C. Much tighter control of the coolant temperature requires much higher flow rates, and yet provides little added benefit given that the mold cavity surface temperatures will tend to vary more significantly between the cooling lines as later discussed. Table 9.1 Specifications of Two Mold Temperature Controllers Model

Vactherm coolant controller

IMSelect oil controller

Minimum temperature (°C)

10

32

Maximum temperature (°C)

99

304

Heating capacity (kW)

9

16

Cooling capacity (kW)

14.6

16

Coolant flow rate [m3/s (GPM)]

1 · 10–3 (15)

3 · 10–3 (45)

Coolant pressure [kPa (psi)]

200 (29)

30 (4.3)

253

254

9 Cooling System Design

Example: Calculate the required volumetric flow rate for the mold coolant, assuming an allowable increase in the coolant temperature for the cup/lid family mold of 1°C. If water is used as the coolant, then the required volumetric flow rate is:

Vcoolant =

260 W m3 = 6.2 ×10-5 3 s 1°C ×1000 kg m × 4200 J kg V

which is equal to about 1 gallon per minute per line. It should be noted that if two cooling lines were connected in series, then the cooling power would also be doubled such that twice the flow rate would be needed to maintain the same temperature distribution. 

After estimating the required coolant flow rate, the feasibility of this value should be checked against the capabilities of commercial mold temperature controllers. The specification for a typical coolant temperature controller and an oil temperature controller are listed in Table 9.1. Example: Assess the capability of the VacTherm controller specified in Table 9.1 for the two cavity, cup/lid family mold. For the cup/lid family mold, the required volumetric flow rate of 1 gallon per minute is well within the capabilities of commercial controllers. In fact, a single mold temperature controller will be able to supply the needed flow to all four cooling lines: 3

3

m m total Vcoolant = 4 × 6.2 ×10-5 = 2.5 ×10-4 s s

Note, however, that multiple mold temperature controllers would be needed if the allowable temperature increase were set to 0.1°C, or if the number of cavities in the mold was increased from 2 to 8. 

9.2.4 Assess Cooling Line Diameter The allowable range of cooling line diameters can now be determined based on the heat transfer and fluid flow constraints. To ensure adequate heat transfer from the mold steel to the coolant, turbulent flow in the coolant is desired. If the cooling line

9.2 The Cooling System Design Process

diameter is too large, then the linear velocity of the water may not be sufficient to ensure turbulent flow. To ensure turbulent flow, the Reynolds number, Re, should be greater than 4000: Re =

4 × rcoolant × Vcoolant > 4000 (9.14) p × mcoolant × D

where D is the cooling line diameter and mcoolant is the viscosity of the coolant. This turbulence requirement implies a maximum diameter, Dmax , for the cooling line of: Dmax =

4 × rcoolant × Vcoolant (9.15) p × mcoolant × 4000

Example: For the cup/lid mold, determine the upper limit on the diameter of the ­cooling line to ensure turbulent flow. Given a flow rate of 1 GPM and a viscosity of 0.001 Pa s, then Eq. 9.15 ­provides a maximum diameter of:

Dmax =

4 ×1000 kg m3 × 6.2 ×10-5 m3 s = 20 mm p × 0.001Pas × 4000

As this example indicates, the requirement of turbulent flow is not very constraining since any diameter less than 20 mm would ensure turbulent flow in the cup/lid family mold. Most molding applications require a high rate of heat transfer and an associated high volumetric flow rate such that turbulent flow is almost given. 

A more binding constraint governs the minimum cooling line diameter, which is related to the pressure drop required to force the coolant through the cooling lines at the required volumetric flow rate. The pressure drop for water through a cooling line can be estimated from pipe flow [9] as: DPline =

2 rcoolant × Lline × Vcoolant

10 p × D5

(9.16)

where the cooling line has length, Lline . This pressure drop requirement implies a minimum diameter of: Dmin = 5

2 rcoolant × Lline × Vcoolant (9.17) 10 p ×DPline

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256

9 Cooling System Design

To calculate the minimum cooling line diameter, the line length and allowable pressure drop across the cooling line must be known. This information can a­ ctually be a bit uncertain, since it depends not only upon the configuration of the cooling lines in the mold, but also whether the cooling lines are piped in series or parallel. Table 9.2 Specifications of Typical Cooling Plugs DME plug

Normal pipe thread

Cooling line diameter

JP-250

1/16

4.76 mm (3/16”)

JP-251

1/8

6.35 mm (1/4”)

JP-352

1/4

9.53 mm (3/8”)

JP-553

3/8

11.1 mm (7/16”)

JP-554

1/2

15.9 mm (5/8”)

Example: For the cup/lid mold, determine the lower limit on the diameter of the ­cooling line to avoid an excessive pressure drop in the coolant temperature controller. The analysis will assume that the cooling lines traverse the width of the mold, and each has a length of 302 mm. The analysis further assumes that the two cooling lines on each side of the mold will be connected in series. The allowable pressure drop is set to 100 kPa, which is ½ of the maximum supply pressure from the VacTherm controller. This last assumption is made to ensure that some supply pressure is reserved for flow through the cooling hoses from the controller to the mold, as well as for pressure drops asso­ ciated with turns, plugs, and so on. The minimum cooling line diameter may then be estimated as:

Dmin =

5

(

10 p ×100 ×103 Pa

2

)

1000 kg m3 × 0.6 m × 6.2×10-5 m3 s

= 3.7 mm

Combining the turbulence and pressure drop requirement, the allowable range of cooling line diameters for the cup/lid mold is:

3.7 mm < D < 20 mm While this is quite a broad range, the allowable range may be much smaller depending on the molding application and manufacturing requirements. 

In selecting the final cooling line diameter, the mold designer should consider the manufacturability of the cooling lines and the molder’s standards regarding cool-

9.2 The Cooling System Design Process

ing plugs, connectors, and hoses. Table 9.2 provides some specifications for typical cooling plugs provided by a mold component supplier (DME). As observed in the table, the commercial plugs range from 4.76 to 15.9 mm. The mold designer should select a cooling line diameter that satisfies the above analysis and is a standard size. Example: Specify an appropriate cooling line diameter for the cup/lid mold. The analysis of the previous example indicates that any standard diameter between 3.7 mm and 20 mm is feasible, which means that any of the ­cooling plugs listed in Table 9.2 would be fine. For this reason, the mold ­designer should choose a cooling line diameter that is readily machinable and also compatible with the cooling plug standards at the molder. A reasonable cooling line diameter is 6.35 mm. 

It should be noted that the above analysis is most appropriate for water as the coolant. Ethylene glycol and oil are not as common in practice due to environmental and cost concerns. These non-water coolants are also substantially more viscous than water, such that turbulent flow is not likely to be achieved. For laminar, viscous flow, the pressure drop can be estimated using the previously developed Hagen-Poiseuille law with the coolant properties of Appendix C: DPline =

128 × mcoolant × Lline × Vcoolant p × D4

(9.18)

The mold designer should then select an appropriate cooling diameter to ensure that the maximum pressure drop across the cooling lines does not exceed the capability of the coolant temperature controller.

9.2.5 Select Cooling Line Depth After the cooling line diameter has been determined, the depth of the cooling lines must be selected. From a structural point of view, it is desirable to place the cooling lines far from the surface of the mold cavity. Deep placement avoids the stress concentrations associated with the removal of material close to the surface. For reference, Fig. 9.4 plots the stress contours for mold designs with cooling line depths, Hline , equal to one cooling line diameter at left and four cooling line diameters at right. It is observed that there are stress concentrations around the cooling line, and the magnitude of the stress, s, increases significantly as the cooling line ­approaches the mold wall.

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258

9 Cooling System Design

Hline = 1 D σ = 3.3 ∙ Pmelt

Hline = 4 D σ = 2.6 ∙ Pmelt

Figure 9.4 Stress distributions around cooling line

Stress concentration factors have been well analyzed for a variety of materials, ­geometries, and load conditions [10]. For a uniformly loaded plate with a hole, the minimum stress concentration factor is typically between 2 and 3 [11]. Even when the cooling lines are placed at a distance of four diameters, a stress concentration factor of 2.6 is observed. If a mold insert is made of P20, which has an endurance stress (to avoid fatigue) of 456 MPa. Then the mold can only be designed for a ­maximum melt pressure of: max Pmelt =

sendurance = 175 MPa (9.19) 2.6

Fortunately, this melt pressure is about equal to the maximum injection pressure for most molding machines, and is unlikely to be fully transmitted to the mold cavity. The stress concentration associated with cooling lines is very significant, and this constraint requires the cooling line to be placed far away from the mold surface in molding applications with high melt pressures. Even when the cooling lines are placed far from the cavity surface, the stress concentrations still potentially limit the melt pressures with which the mold may be operated. As another example, consider the design of an aluminum mold with a fatigue limit stress equal to 166 MPa. If the cooling line depth is equal to one diameter then the stress concentration factor is 3.3, which would allow a maximum melt pressure of just 50 MPa. This analysis does not prevent the molder from operating at higher melt pressures, but simply indicates that the mold will likely not operate for a long life without developing cracks emanating from the cooling lines. Stress concentrations in molds are discussed in more detail in Section 12.2.6.

9.2 The Cooling System Design Process

While the structural considerations suggest that cooling lines should be placed far from the mold surface, the rate of heat transfer is maximized by placing the cooling lines as close to the surface as possible. The heat conduction equation states that the thermal resistance is linear with the distance between the cooling line and the mold surface. The effective heat conduction coefficient is: hconduction =

kmold (9.20) H line

As previously discussed with the heat transfer analysis of Eq. 9.7, a typical convective heat transfer rate in molding is 1000 W/°C. To ensure that the cooling line depth does not add unnecessarily to the cooling time, the maximum cooling line depth may be estimated as: H line <

kmold (9.21) 1000W °C

A commonly used steel, P20, has a thermal conductivity of 32 W/m°C, which suggests a maximum cooling line depth of 32 mm for effective cooling. Combining the structural and heat transfer requirements for a typical 6.35 mm diameter cooling line, the recommended range for the cooling line depth is: 2 D < H line < 5 D (9.22)

which is a commonly used range in mold design. While a mold designer may choose an arbitrary cooling line depth from this range, the provided analysis should be used for special applications with diverse structural or heat transfer ­requirements. Example: Specify the cooling line depth and maximum melt pressure for the cup/lid mold if P20 is to be used as the mold material. The cooling line diameter is 6.35 mm. Let us assume that the depth will be set to four cooling line diameters, so the cooling line depth is set to 25.4 mm. This depth still imposes a stress concentration of 2.6. If the endurance limit for P20 is 456 MPa, then the maximum melt pressure for infinite life is: max Pmelt =

456 MPa = 175 MPa 2.6

which is close to the maximum injection pressures available from most molding machines. 

259

260

9 Cooling System Design

9.2.6 Select Cooling Line Pitch Once the cooling line depth is selected, the distance between the cooling lines (known as the “pitch”) is assigned. A tighter pitch, Wline , between cooling lines provides for faster and more uniform cooling. However, a tighter pitch also means more cooling lines and the likelihood of conflicts arising between the cooling lines and other mold components. The mold designer should select a cooling line pitch that is appropriate for the specific molding application using analysis. The temperature prediction of the melt during cooling involves the solution of a system of parabolic differential equations. While this is readily solved using the finite element method as above, no suitable analytical treatment has yet been developed. Menges [12] provides an estimate of the percentage variation in the heat flux, DQ , across the mold between cooling lines: ççæ Wline ÷÷ö

æ W ö2.8lnççè H line ÷÷÷ø DQ éë %ùû ¥ççç line ÷÷÷ (9.23) çè H ø÷ line

This function is plotted in Fig. 9.5 for steel and an aluminum mold materials. The analysis indicates that the variation in the heat flux is less than 5 % up to a cooling line pitch equal to twice the cooling line depth. Afterwards, the variation in heat flux increases dramatically and is indicative of slower rates of mold cooling and high temperature gradients within the molded part.

Figure 9.5 Effect of pitch on variation in heat flux

9.2 The Cooling System Design Process

To avoid a significant temperature gradient between cooling lines, it is recommended that mold designers use a cooling line pitch in the range of: H line < Wline < 2H line (9.24)

depending on the requirements of the application. A commodity product with loose tolerances would likely be fine with a cooling line pitch equal to two or three times the cooling line depth. For tighter tolerance applications or for applications requiring faster cycle times or more uniform cooling, a closer spacing equal to the cooling line depth is desirable. Figure 9.5 indicates that the use of highly conductive materials (such as aluminum or copper) actually increases the variation in heat flux by improving the heat conduction between the cooling line and the cavity surface. As such, the use of highly conductive materials does not directly allow for a wider pitch and a reduced number of cooling lines. If fewer cooling lines are desired, then this may best be accomplished by selecting a large cooling line depth and still setting the pitch to twice this amount. Highly conductive mold materials can then be utilized to accomplish high rates of heat transfer with uniform cooling.

Wline = Hline

Wline = 4 Hline

Figure 9.6 Heat flow from cavity centerline to cooling line

Example: Transient thermal simulation was performed for the cup/lid family mold for two mold designs having different pitch to cooling line depth ratios. Figure 9.6 plots the heat flow from the centerline of the molding in the cavity to the cooling. In the figure, the lengths of the arrows represent the relative amount of heat flowing out of the mold cavity at that location. As the

261

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9 Cooling System Design

c­ ooling lines are moved further apart two adverse conditions arise. First, the effective heat transfer rate at the mold wall is reduced given the finite ­capacity of the cooling lines to remove heat. Second, a significant variation in the heat transfer rate arises across the cavity surface. This variation in the heat transfer rate across the cavity surface will give rise to a gradient in the temperature of the moldings at the time of ejection as plotted in Fig. 9.7. With a tight cooling line pitch, the moldings are ejected not only with a lesser temperature variation across the molding, but also at a significantly lower temperature. With a wide pitch, the moldings exhibit a much higher temperature gradient and a much higher temperature. ­Interestingly, extending the cycle time for the mold with the wider pitch does not reduce the temperature gradients without significantly increasing the ­cooling time to allow the entire molded part to approach the coolant ­temperature.

Wline = Hline

Wline = 4 Hline

Figure 9.7 Temperature distribution in plastic and mold 

9.2.7 Cooling Line Routing Once the cooling line diameter, depth, and pitch have been determined, the cooling lines can be routed through the mold. This routing is of critical importance since it not only impacts the cost and quality of the moldings, but also limits the placement of other mold components such as ejector pins and bolts. In general, the mold design should provide at least half a cooling diameter between the surface of the cooling line and the surface of any other mold component. This design constraint is provided to maintain the structural integrity of the mold while also minimizing cooling leaks during mold operation due to corrosion. The shaded area in Fig. 9.8 represents the possible locations in the mold where cooling lines may be placed.

9.2 The Cooling System Design Process

Figure 9.8 Potential mold areas for locating cooling lines

While the shaded area of Fig. 9.8 represents a large portion of the mold, the placement of cooling lines is further constrained by potential interference with the mold cavity, cavity inserts, core inserts, ejector return pins, guide pins, sprue bushing, and other mold components. The previous analysis for the cup/lid mold suggested that the cooling system design use: ƒƒa cooling line diameter of 6.35 mm, ƒƒa cooling line depth of 12.7 mm, and ƒƒa cooling line pitch of 25.4 mm. The design implemented exactly to these recommendations is shown in Fig. 9.9. This initial design is infeasible for many reasons. Perhaps the most significant shortcoming in the design is that many of the cooling lines intersect critical mold features such as the sprue bushing or the interface between the cavity inserts and the mold plates. There are two different strategies to resolve this issue. One approach is to enlarge the cavity insert, core insert, and associated mold plates to fit all the cooling lines within the envelope of the core and cavity inserts. This option is costly since it requires redesign of the mold, procurement of a larger mold base, and more machining. However, such a design may be economically justified given the more rapid and uniform cooling.

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Figure 9.9 Infeasible initial cooling line layout

An alternative approach is to move the cooling lines further from the mold cavity while maintaining the same pitch to depth ratio for the cooling lines. The resulting design is shown in Fig. 9.10. This design requires fewer cooling lines, all of which avoid the intersection with other mold components. While this design provides poor cooling performance, it is quite common. A primary advantage is that all of the cooling lines are not only straight, but each cooling line also passes through a single mold plate as well. As such, the cooling lines can be machined in a single setup without any need for seals or gaskets. Unfortunately, the placement of the cooling lines far from the mold cavity will reduce the rate of heat transfer and ­necessitate longer cycle times. The rate of heat transfer is further reduced since the cooling lines are placed in another plate, and there will be a thermal contact resistance associated with the heat flow between the two plates [13].

Figure 9.10 Feasible but poor cooling line layout

9.2 The Cooling System Design Process

There is a second significant shortcoming in cooling line layout of Fig. 9.10, which stems from the use of a straight cooling line with a core of significant height. The source of cooling is the support plate behind the core insert, far from the heat source originating from the plastic melt in the cavity. As such, significant temperature variations will develop throughout the mold and molded part during cooling. The predicted temperature distributions at the end of the molding cycle for the cup are provided in Fig. 9.11, in which each contour line represents a 2°C change in temperature. Due to the relatively deep core, a gradient of 6°C exists from the base of the core to the top of the core. The temperature gradient in Fig. 9.11 will drive differential shrinkage along the axis of the cup as well as differential shrinkage through the wall thickness of the molding. The reason is that the temperature at the top of the core is not only 6°C hotter than the temperature at the base of the core, but is also roughly 6°C hotter than the temperature at the opposing surface on the cavity insert. Three options for rectifying the situation include using a highly conductive core insert, implementing a baffle or bubbler, or designing a cooling insert. These different cooling designs are next discussed.

Figure 9.11 Temperature gradient from poor design

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„„9.3 Cooling System Designs There are many different cooling system designs that are used in practice. While many molds use straight lines, such designs are often not optimal. Instead, the mold designer should strive to achieve uniform and high rates of cooling across the entire cavity surface. Creativity is often required to provide effective cooling ­system designs at reasonable costs. Next, some of the most common designs and components are analyzed.

9.3.1 Cooling Line Networks When molds have more than one cooling line, an issue will arise as to how the molding machine operator will connect the mold to the mold temperature controller. Consider the computer bezel, overlaid with cooling lines as shown in Fig. 9.12. The mold has been provided with eight cooling lines traversing across the width of the mold cavity. Faced with eight cooling line connections, the machine operator may use short hoses to loop the cooling lines as shown. Such a setup has two ­compounding issues. First, the flow resistance through the combined length of all the cooling lines can be extremely high, reducing the coolant flow rates per Eq. 9.16. Second, the mold coolant temperature can increase along the length of the cooling circuit at reduced coolant flow rates. As such, a significant temperature differential can arise from where the coolant enters the mold to where the coolant exits the mold.

Figure 9.12 Bezel cooling line layout in series

Aware of this problem, many if not most molders have coolant manifolds installed on the molding machine between the mold temperature controller and the mold. The operator can then use longer cooling hoses to individually connect two sets of

9.3 Cooling System Designs

cooling lines with a short return loop on the opposite side of the mold. Such a parallel setup is shown in Fig. 9.13. This configuration is extremely common since it is simple and provides effective cooling. However, the installation and removal of the mold from the machine is complicated by the number of lines that must be connected and disconnected. The high number of hoses and operator steps also increases the likelihood that the cooling system may be setup incorrectly or fail, for instance, due to a loosely connected hose.

Figure 9.13 Bezel cooling line layout with four parallel cooling circuits

There is currently a great deal of interest in the plastics industry in lean manu­ facturing, which places significant emphasis on reduced process complexity and setup times [14–16]. By investing slightly more in the injection mold, it is possible to reduce the mold setup time, reduce potential failure modes, and improve the mold performance. Figure 9.14 shows the addition of two vertical cooling lines connecting all eight horizontal lines within the injection mold; twenty pressure plugs have been installed to block the coolant flow at selected locations. The result is that a cooling manifold has been designed internal to the mold, such that only two connections are required. At the same time, the cooling uniformity is increased. This internal manifold design has very little added cost while delivering both increased performance and ease of use.

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Figure 9.14 Bezel with internal, parallel cooling line layout

Once plugging is considered an option in the routing of cooling lines, many more complex cooling line layout become available. In the cooling system designs for the bezel shown in Figs. 9.12–9.14, the portion of the cooling lines located inside the screen area of the bezel is not removing any significant heat. If a two-plate mold with a cold runner is used as shown in Fig. 7.7, then these cooling lines would cool the sprue and runners. For a three-plate or hot runner mold, however, there is no heat being generated in the central area of the mold cavity. Given that there is no need for cooling in the center of the mold, it is possible to route the cooling lines around the periphery of the part to improve the cooling of the mold cavity while reducing the mold making cost and providing a mold that is even easier to operate. Figure 9.15 shows such a cooling system design using blind drilled holes and plugs that can be produced for a cost similar to that shown in Fig. 9.14. This design ­provides ease of use, moderate flow resistance, and uniform cooling about the ­entire molding.

Figure 9.15 Bezel with drilled peripheral cooling line layout

9.3 Cooling System Designs

9.3.2 Cooling Inserts As an alternative to drilling cooling lines, cooling lines that conform to the shape of the mold cavity can be milled into the rear faces of the cavity or core inserts as shown in Fig. 9.16. In this case, a ball end mill is routed around the bottom of the core insert, after which connecting lines are drilled to one side of the mold. The cooling lines can thereby closely follow the contours of the molded part, even for curved surfaces. The location of the coolant entrance and exits has been selected to balance the pressure drop between the internal and external circuits.

Figure 9.16 Bezel core insert with milled cooling

Even though the cooling insert design shown in Fig. 9.16 provides exceptional cooling, it presents potential leakage issues. In this design, a groove has been ­provided and fitted with a gasket. When fastened tightly to the support plate, the gasket will prevent leakage outside the mold. However, leakage should be expected at any ejector pins located internal to the area surrounded by gasket. In this application, a stripper plate could be successfully used as discussed in Section 11.3.4.

9.3.3 Conformal Cooling Manufacturing technology is continuing to advance, and relatively new mold making technologies include polyjet [17], stereolithography [18], fused deposition modeling (FDM, [19]), and selective laser sintering (SLS, [20]) in the form of direct metal laser sintering (DMLS, [21]). These 3D printing technologies provide the ability to directly place cooling lines at any location as the core and cavity inserts are being printed. Returning now to the core insert for the cup, a conformal cooling

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line design is provided in Fig. 9.17 in which the coolant is routed to the center of the insert and then branches to a series of arteries like an internal water fountain, thereby eliminating the temperature gradients shown in Fig. 9.7.

Figure 9.17 Core insert for cup with conformal cooling lines

Since nearly any geometry can be made with these 3D printing technologies, internal cooling lines of arbitrary design can be made to conform to the cavity surfaces to improve heat transfer rates and uniformity. Significant ongoing issues include strength, porosity, and surface finish of the printed inserts. Specifically, polymer mold inserts made by polyjet, stereolithography, FDM, and SLS provide a low cost approach to prototype molds but have very low thermal conductivity and strength. The design of Fig. 9.17 would still require an extended cycle time compared to metal inserts, on the order of 100 s, while being unlikely to survive more than a few dozen molding cycles. Placing the cooling lines further from the mold surface would reduce the strength issues, but negate the benefit of having conformal ­cooling lines. Conversely, DMLS inserts would provide sufficient strength as suggested by the review of material properties described by Kruth et al. [20]. However, such DMLS inserts would be quite expensive while not providing sufficient surface appearance.

9.3.4 Highly Conductive Inserts Another approach to reducing temperature gradients is to utilize highly conductive insert materials, such as Cu 940 or Aluminum QC10, for portions or the entire core insert. Since these materials have much higher thermal conductivity than steel, their selective use in molds having “hot spots” will tend to reduce the temperature variation across the cavity. For example, the predicted temperature distributions at the end of the molding cycle for the cup core insert using a Cu 940 material are

9.3 Cooling System Designs

provided in Fig. 9.18. As with the results provided for a steel core in Fig. 9.11, each contour line represents a 2°C change in temperature. The results indicate that the temperature gradient has been reduced by approximately 60 % compared to the temperature gradients shown in Fig. 9.11.

Figure 9.18 Temperatures in deep conductive core

Conductive inserts can also provide improved cooling in the internal corners of moldings. Because of the heat transfer in three dimensions and limitations regarding the proximity of the cooling line to the mold wall, the cavity insert for will tend to conduct much more heat away from the molded plastic compared to the core insert. The temperature distribution for a typical design using a single material for both the core and the cavity is shown in Fig. 9.19(a). When the core and cavity inserts both consist of P20, there is a 5°C temperature gradient across the wall thickness of the molding. However, only a 1°C differential across the wall t­ hickness of the molding occurs when the core insert is specified with Cu 940 as shown in Fig. 9.19(b).

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a) P20 cavity and core

b) P20 cavity and Cu 940 core

Figure 9.19 Temperature distribution in corner

The primary advantage of highly conductive core inserts is the ability to strategically control the heat flow. While these materials increase the rate of heat transfer, their properties are not appropriate for use throughout the mold. There are two primary reasons. First, it should be noted that the improved temperature distributions achieved in Fig. 9.19(b) were the result of using different materials for the core and cavity inserts. These temperature distributions would not have been as uniform if both the core and cavity inserts were made from Cu 940. Second, these highly conductive materials tend to have lower hardness and are more susceptible to wear. As such, highly conductive inserts may be best when used in applications with high production volumes, low to moderate injection pressures, and non-­ abrasive materials.

9.3.5 Cooling of Slender Cores Mold cores with a high length to diameter ratio prevent effective heat transfer along the length of the core, even with the use of highly conductive materials. For this reason, it is desirable to provide a cooling channel along the axis of the core to conduct heat from the surface of the core and then convect the heat down the center of the core. Larger cooling channels in the center of the core generally allow for higher coolant flow rates and higher rates of heat transfer. Larger cooling channels, however, require the removal of more volume inside the core and a lessening of the core’s structural integrity. To balance these two issues, different cooling components have been developed for use in different ranges of core diameters. ­Table 9.3 lists some of the various options, which will next be discussed.

9.3 Cooling System Designs

Table 9.3 Slender Core Cooling Options Core diameter

Hole diameter

Cooling rate

Cooling insert

> 50 mm

> 25 mm

Very high

Baffle

12–75 mm

6–25 mm

Very high

Bubbler

6–30 mm

3–12 mm

High

Heat pipe

5–20 mm

3–12 mm

Medium

Conductive pin

× ( edge center ) èç W ø÷

where h is the wall thickness of the molding, and scenter and sedge are the shrinkage rates at the center and edge of the part, and W is the distance from the center to the edge of the molding. This buckling analysis models the molding as an isotropic circular plate under uniform radial edge compression with a buckling stress threshold of:

10.3 Warpage

sbuckling = K

2 E æç H ö÷ (10.20) × ÷ ç 1- n 2 èç R ø÷

where K is a constant dependent upon the Poisson’s ratio, v, E is the modulus of the material, H is the thickness of the plate, and R is the plate radius. Equation 10.19 assumes a material with a Poisson ratio of 0.4 (valid for most plastics). If a molded part buckles, then the out of plane warpage can be conservatively estimated as:

( (

2

))

d = W 2 - W 1- (scenter - sedge )

(10.21)

As is well-known in molded part design [16], the addition of ribs will increase the moment of inertia, and thereby decrease the likelihood of a molding to buckle and severely warp [17]. Example: Analyze warpage assuming that the cup lid is center gated and molded with ABS at a packing pressure of 66 MPa at the center of the part and 0 MPa at the outer rim. To evaluate the warpage, it is first necessary to calculate the shrinkage rates and check the buckling criterion. Given the 66 MPa packing pressure and a temperature of 132°C, the linear shrinkage at the center will be 0.31 % at the center. At the edge, the pressure of 0 MPa and temperature of 132°C provides a linear shrinkage of 1.66 %. Given that the cup is 2 mm in thickness and 81 mm in diameter, the buckling criterion is stated as: 2 2 mm ÷ö ÷÷ è 0.5 × 81 mm ø

æ

?

(1.66% - 0.31%)>0.44 × ççç

0.0135> 0.0011 This criterion indicates that the central portion of the lid will buckle. The estimated warpage is:

d=

2

(40.5 mm)

(

2

)

- 40.5 mm (1- (1.66% - 0.31%)) = 6.6 mm

In the actual molding of the lid, it is somewhat unlikely that the lid would warp and very unlikely that the lid would warp to this extent. The reason for the warpage in the analysis is that the analysis assumed that the pressure at the edge of the lid was 0 MPa and did not pack out at all. As such, the material around the edge was predicted to shrink at a rate much higher than would be encountered in practice. 

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As the previous warpage analyses have shown, warpage is caused by nonuniform shrinkage due to temperature gradients through the wall thickness of the molded part, pressure gradients across the area of the molded part, or temperature gra­ dients across the area of the molded part. These are the most common causes of warpage, and have been treated with the simplest possible analysis. However, there are other causes of nonuniform shrinkage including orientation and residual stress. For further information, the interested mold designer is referred to the research literature [18–25]. Reasonably accurate warpage predictions may also be obtained with computer simulation as previously discussed. Figure 10.20 provides the warpage prediction from computer simulation for the simulation described for Fig. 10.8. The results suggest that the middle of the part will bow down while the left and right sides of the bezel will bow upwards. The reason for this behavior is the upper surface of the bezel contains more material, and this upper surface is shrinking less than the lower portion of the ribs.

Figure 10.20 Simulation of predicted shrinkage and warpage

Given that warpage can be significant compared to shrinkage, mold designers may choose to compensate for warpage by reverse biasing the mold surfaces based on predicted or observed warpage. This bias is sometimes referred to as “Kentucky windage,” a shooting term that refers to the adjustments a distance shooter must make to account for the wind [26]. Thus, Kentucky windage in mold design refers to the contouring of the mold cavity surfaces such that upon warping the molded part flattens to the desired shape. The use of such bias is not without controversy, since it incurs expense in the contouring of mold cavity surfaces while also inherently changing the warpage and structural behavior of the molded parts. For this reason, it remains a somewhat rare practice.

10.3 Warpage

10.3.2 Warpage Avoidance Strategies There are several common strategies that should be used to avoid and address warpage issues. By far, the most important strategy is to design a mold that will provide uniform melt temperatures and pressures throughout the cavity, so that the shrinkage of the molded part(s) will be highly uniform. To maximize the shrinkage uniformity in tight tolerance molding applications, the mold designer should: ƒƒAvoid high flow length to wall thickness ratios by utilizing multiple gates; ƒƒMaintain uniform cavity pressures by designing a balanced feed system with low flow resistance; ƒƒMaximize the mold surface temperature uniformity with a tight cooling line pitch and highly conductive mold inserts where needed; and ƒƒFacilitate melt pressure and temperature uniformity in the molding by requiring uniform part thickness and generous fillets. If the mold is well designed, then warpage is less likely to occur. In the event that warpage is encountered, a molder may try to reduce or eliminate the warpage by: ƒƒFilling the mold cavity as fast as possible to reduce cooling in the solidified skin; ƒƒIncreasing the pack time until the part weight no longer increases; ƒƒIncreasing the packing pressure to reduce the amount of material shrinkage; ƒƒUtilizing a pack pressure profiling as discussed with respect to Fig. 10.12 to increase melt pressure and shrinkage uniformity across the part; ƒƒUtilizing different coolant temperatures on different sides of the mold or in different portions of the mold to purposefully control the temperature and shrinkage distribution; and ƒƒTrying different types of materials and filler systems with varying shrinkage behaviors to find satisfactory performance; a high flow grade of polystyrene (such as HIPS) can indicate if the warpage problem is due to a temperature or pressure gradient. Even with the all these mold design and molding actions, warpage issues may require mold design changes. There are several mold design changes that are commonly used to reduce the magnitude of the warpage. The most common approach might be the addition of one or more gates to improve the uniformity of the ­shrinkage across the cavity. Another common approach to reduce the likelihood of buckling is to increase the stiffness of the molding through the addition of shallow ribs. The use of Kentucky windage, although of increasing interest, is less common since it places a significant burden on the mold designer and mold maker, while requiring a very high level of predictive capability and very fine surface machining. Since the dimensional shifts of the part due to warpage may exceed steel safe limits, errors in this approach can incur very high costs.

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„„10.4 Chapter Review In this chapter, shrinkage and warpage analyses were provided to predict changes in the molded part dimensions based on the pressure-volume-temperature (PvT) behavior of the polymer together with the melt pressures and temperatures. These analyses provide insight into the shrinkage and warpage phenomenon, but are highly dependent upon the assumed pressures and temperatures. For this reason, many mold designers use a mid-range shrinkage value recommended by a material supplier or other source. In tight tolerance applications, prototype molding and/or steel safe mold design strategies are frequently used to converge to the ­optimal dimensions of the mold cavity to deliver the desired part dimensions. After reading this chapter, you should understand: ƒƒThe relationship between shrinkage, mold dimensions, and part dimensions; ƒƒThe PvT behavior of amorphous and semicrystalline polymers; ƒƒThe qualitative relationship between melt pressure, melt temperature, and shrinkage; ƒƒHow to calculate volumetric shrinkage from the PvT model; ƒƒHow to calculate linear shrinkage from volumetric shrinkage; ƒƒThe causes of differential shrinkage; ƒƒHow to calculate warpage from differential shrinkage; ƒƒThe effect of processing conditions and fillers on shrinkage and warpage; and ƒƒMold design strategies for managing shrinkage and warpage. The analysis of shrinkage is useful for specifying the mold dimensions. However, the shrinkage of the plastic onto the mold core(s) also determines the forces required to eject the molded part. The estimation of these ejection forces will guide the design of the ejection system so as to avoid deforming the molded part(s) upon ejection. After the ejection system is designed, the mold’s structural systems are analyzed and designed.

„„10.5 References [1] Fagade, A. and D. O. Kazmer, Early cost estimation for injection molded parts, J. Injection Molding Technol. (2000) 4(3): pp. 97–106 [2] Bushko, W. C. and V. K. Stokes, Solidification of thermoviscoelastic melts, Part I: Formulation of ­model problem, Polym. Eng. Sci. (1995) 35(4): pp. 351–364 [3] Greener, J. and R. Wimberger-Friedl, Precision injection molding, Hanser, Munich (2006) [4] Berins, M., Standards for Molding Tolerances, in SPI Plastics Engineering Handbook, The Society of the Plastics Industry, Inc., 5th ed., Kluwer Academic Publishers (1991) pp. 821–844

10.5 References

[5] Ginell, R., Derivation of the Tait equation and its relation to the structure of liquids, J. Chem. Phys. (1961) 34(4): pp. 1249–1252 [6] Zoller, P. and D. J. Walsh, Standard Pressure-Volume-Temperature Data for Polymers, CRC Press (1995) [7] Olabisi, O. and R. Simha, Pressure-volume-temperature studies of amorphous and crystallizable polymers, I. Experimental, Macromol. (1975) 8(2): pp. 206–210 [8] Speight, R., et al., Best practice for benchmarking injection moulding simulation, Plast., Rubber Compos. (2008) 37(2-4): pp. 124–130 [9] Kazmer, D. O., et al. Prediction of part dimensions using sensed melt pressure and melt temperature and estimated specific volume, in Injection Molding Division of the Society of Plastics Engineers’ ­Annual Technical Conference, Orlando, FL (2015) [10] Hoffa, D. W. and C. M. Laux, Gauge R & R: an effective methodology for determining the adequacy of a new measurement system for micron-level metrology (2007) [11] Bushko, W. C. and V. K. Stokes, Solidification of thermoviscoelastic melts, Part IV: effects of boundary conditions on shrinkage and residual stresses, Polym. Eng. Sci. (1996) 36(5): pp. 658–675 [12] Panchal, R. R. and D. O. Kazmer, In-situ shrinkage sensor for injection molding. J. Manuf. Sci. Eng. (2010) 132(6): p. 064503 [13] Zheng, R., et al., Modeling of flow-induced crystallization of colored polypropylene in injection molding, Korea-Australia Rheol. J. (2010) 22(3): pp. 151–162 [14] Fornes, T. and D. Paul, Modeling properties of nylon 6/clay nanocomposites using composite theories, Polymer (2003) 44(17): pp. 4993–5013 [15] Jacques, M. S., An analysis of thermal warpage in injection molded flat parts due to unbalanced cooling, Polym. Eng. Sci. (1982) 22(4): pp. 241–247 [16] Malloy, R. A., Plastic part design for injection molding, Hanser, Munich (1994) [17] Walsh, S., Shrinkage and warpage prediction for injection molded components. Journal of reinforced plastics and composites (1993) 12(7): pp. 769–777 [18] Zoetelief, W. F., L. F.A. Douven, and A. J.I. Housz, Residual thermal stresses in injection molded ­products, Polym. Eng. Sci. (1996) 36(14): pp. 1886–1896 [19] Chen, S.-C., et al., Integrated simulations of structural performance, molding process, and warpage for gas-assisted injection-molded parts, I. Analysis of part structural performance, J. Appl. Polym. Sci. (1998) 68(3): pp. 417–428 [20] Chang, R. Y. and B. D. Tsaur, Experimental and theoretical studies of shrinkage, warpage, and sink marks of crystalline polymer injection molded parts, Polym. Eng. Sci. (1995) 35(15): p. 1222 [21] Heberlein, D. E., Jr., Warpage scale to evaluate injection molded plastics, in Annual Technical Conference – ANTEC, Conference Proceedings, San Francisco, CA (1994) [22] Shrock, J. E., Study of corner cooling as related to warpage analysis, in Annual Technical Conference – ANTEC, Conference Proceedings, San Francisco, CA (1994) [23] Walsh, S. F., Shrinkage and warpage prediction for injection molded components. J. Reinf. Plast. Compos. (1993) 12(7): pp. 769–777 [24] Bushko, W. C. and V. K. Stokes, Dimensional stability of thermoplastic parts: Modeling issues, in ­Annual Technical Conference – ANTEC, Conference Proceedings, Indianapolis, IN (1996) [25] Fan, B., et al., Warpage Prediction in Optical Media, J. Polym. Sci., Part B: Polym. Phys. (2003) 41: pp. 859–872 [26] Lankisch, T., The Kentucky Windage Solution in Simulation Reduces Warpage, Mold Making Technology (2014)

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Ejection System Design

The ejection system is responsible for removing the molded part(s) from the mold after the mold opens. While this may seem a simple function, the complexity of the ejection system can vary widely depending on the requirements of the molding application. Many issues must be considered including the need for multiple axes of actuation, the magnitude and distribution of the ejection forces, and others. Before beginning the analysis and design of the ejection system, an overview of its function is first provided. Figure 11.1 provides a side view of a mold opening for the subsequent ejection of the laptop bezel. The ejector assembly (consisting of the ejector plate, ejector retainer plate, return pins, ejector pins, stop pins, and other components) is housed between the rear clamp plate, support plate, and rails.

Figure 11.1 Side view of opening mold

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11 Ejection System Design

At this time in the molding cycle, the molded part has shrunk onto the core side of the mold and has been pulled from the mold cavity as the moving side of the mold is retracted from the stationary side of the mold. In a few moments, the molding machine will push the ejector knock-out rod against the ejector plate to actuate the ejector assembly and strip the molded parts off the core. At this time, however, a clearance exists between the ejector knock-out rod and the ejector plate. Figure 11.2 provides a side view of the mold during the actuation of the ejection system. Prior to ejection, the opening of the molding machine platens separated the two mold halves to allow clearance for the ejection of the part. The machine then drives the ejector knock-out rod forward to make contact with the rear surface of the ejector plate. Since the machine can provide a force to the knock-out rod much greater than the force with which the moldings have shrunk onto the core, the entire ejector assembly is forced forward. The ejector pins come into contact with the molded part(s) and push the molding(s) off the core.

Figure 11.2 Side view of mold with actuated ejectors

After the moldings are ejected, the molding machine then retracts the ejector knock-out rod as shown in Fig. 11.3. A clearance is then made between the front of the knock-out rod and the back of the ejector plate, which allows the ejector assembly to be reset to its original position for the next molding cycle.

11 Ejection System Design

Figure 11.3 Side view of mold with reset knock-out rod

There are several ways of resetting the ejector system, which will be discussed later. However, one common method for returning the ejector assembly is to simply close the mold as shown in Fig. 11.4. The front surface of the return pins will then contact the opposing face of the A plate. The back surface of the return pins will then drive the ejector plate, ejector retainer plate, and all the ejector pins backwards as the mold closes.

Figure 11.4 Side view of closing mold

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11 Ejection System Design

„„11.1 Objectives in Ejection System Design The plastic moldings will tend to shrink during cooling and will usually remain on the mold cores upon the opening of the mold. As such, mechanisms are required to push the parts off the mold during the ejection stage. While this primary function is easily understood, there are several related design objectives that should be ­satisfied in the design of the ejection system.

11.1.1 Allow Mold to Open The first step in the ejection of the moldings from the mold is to open the mold at one or more parting planes. The mold designer should work with the product designer and molder to ensure that the mold design is suitable and robust. In general, the number of moving cores should be minimized by simplifying the product design and developing a suitable mold design. When moving cores are used, they should be designed, when possible, to work with the opening action of the mold rather than relying on additional actuators and control systems. Sometimes the molded part necessitates a moving core design that cannot be ­actuated by the mold opening movement. Most modern molding machines support such “core pull” sequences through the use of digital signals. After the cooling and plastication stages, when the mold is ready to open, the molding machine can be programmed to provide one or more core pull signals to the required actuators (typically pneumatic valves, hydraulic valves, electric solenoids, or electric motors). The actuators can then retract the connected mold components, which should be designed to contact a limit switch when fully retracted to provide a positive feedback signal that the moving cores are retracted and the mold is safe to open. The molding machine will typically be programmed to delay the mold opening ­until all limit switches from all core pull circuits are energized.

11.1.2 Transmit Ejection Forces to Moldings To remove the moldings from the mold, ejection forces must be applied to strip the moldings off the core surfaces. These ejection forces can be applied by many different mold components including ejector pins, sleeves, blades, lifters, air poppets, stripper plates, and other devices. The number, location, and design of these components must be developed to reliably transmit the forces from the molding machine’s knock-out rod(s) through the ejection system to the plastic moldings. With every ejection cycle, significant shear and compressive forces are applied to the ejection system components. If the components are poorly designed, these ejection

11.1 Objectives in Ejection System Design

forces may result in excessive shear stress, compressive stress, deflection, fatigue, buckling, and mold failure. For example, the use of too few, small pins will cause such high shear stresses to literally punch holes through the molded part, a defect known as push-pin.

11.1.3 Minimize Distortion of Moldings Just as the ejection forces can cause stress and deflection in the ejection system components, the ejection forces can also cause stress and deflection in the plastic moldings. To avoid permanent distortion of the plastic moldings, the number, location, and design of the ejector components must be developed to apply a low and uniform state of stress across the moldings. If the ejector force is uniformly distributed across many points in the mold cavity, then the molding will be uniformly ejected from the mold without any permanent distortion.

11.1.4 Maximize Ejection Speed The ejection stage consumes precious seconds of the molding process, without providing much value to the moldings. As such, the ejection system should be designed to remove the moldings as quickly and reliably as possible, and then reset so that the mold may be closed and the next cycle initiated. To increase the speed of the ejection system, some molders may specify the use of air poppets and/or air jets to increase part ejection velocities and reduce the cycle time. To increase the reliability of the ejection system, the mold designer should develop the mold to tightly interface with the molder’s preferred part removal system. While many molds rely on gravity drop of the moldings and the feed system to a moving conveyor, molders are increasing using sprue pickers and gantry robots for part removal. In general, these systems do not greatly reduce the molding cycle time but rather provide increased control of molding’s removal and subsequent placement while protecting the aesthetic areas. If sprue pickers or robots will be used, then the mold designer must appropriately customize the ejection system. Typically, the ejectors are used to strip the moldings off the core but then retain the moldings at a controlled position. Furthermore, mold designers should confirm and design interface geometry in the cavity and/or feed system that is easily identified and highly repeatable for interfacing with the part removal system.

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11.1.5 Minimize Cooling Interference There can be many components in an ejection system and, unfortunately, most of these components are not actively cooled. As such, the ejection system components can significantly interfere with the heat transfer path from the molding to the coolant. There are two issues that commonly arise. First, the ejection system components can be made of a hardened steel that is less thermally conductive than the core inserts. If the ejection system components are large, then the mold’s cooling effectiveness will be greatly reduced. Second, the ejection system components are assembled into the mold and provided with sliding fits. The result is that there is a thermal contact resistance across every boundary between the ejection system components and the adjacent mold. This thermal contact resistance results in lower rates of heat transfer through and around components in the ejection ­system. The effect of cooling interference by the ejection system can be very significant. Consider, for example, an ejector pin with a diameter greater than the nominal wall thickness of the molding. In this case, the ejector pin will not transfer signi­ ficant heat from the adjoining surface of the molding since ƒƒthe ejector pin has a thermal contact resistance between it and the mold, and ƒƒthe ejector pin is relatively large. As a result, the plastic in the mold cavity above the ejector pin will have to cool via heat transfer to the mold steel around the periphery of the ejector pin as well as heat transfer to the opposite side of the mold. While the local cooling of this exact area of the molding may not be the significant constraint on the cycle time, the result is that this large ejector pin will cause a hot spot in the mold and less consistent properties upon ejection. For this reason, the use of overly large ejector pins should be avoided in favor of multiple, smaller ejector pins placed so as to not interfere with the mold cooling. Sometimes, large ejection system components including stripper plates, lifters, core pulls, and others are required. Such large components should be fitted with cooling channels and actively cooled to provide consistent ejection temperatures.

11.1.6 Minimize Impact on Part Surfaces The ejection system is usually located on the moving side of the mold along with the mold cores. Since ejector pins and other components contact the molding, they leave witness marks on the adjacent surfaces, which can ƒƒreduce the visual quality of the molding’s surface, ƒƒinterfere with mating assembly surfaces, and ƒƒreduce strength in structural applications.

11.2 The Ejector System Design Process

As such, ejector pins and other components should be located and designed to have a minimal impact on the molding’s surfaces. The most common approach is to locate ejector pins on nonvisible surfaces and in low stress areas of the molding. Alternatively, larger components such as sleeves, slides, lifters, and stripper plates may be strategically used such that their witness lines coincide with features of the molding. These carefully designed components can leave no apparent witness line while providing very effective ejection across large areas of the part surface. Some applications require one side of the molding to be completely free of all witness marks. In these applications, one strategy in mold design is to locate the entire system on the stationary side of the mold along with the feed system. This “reverse ejection” design allows the surface of the moldings facing the moving side of the mold to be completely free of witness marks due to both the feed system and the ejection system. This design will be discussed in more detail in Section 11.4.4.

11.1.7 Minimize Complexity and Cost The cost of the ejection system can be either a negligible or a significant portion of the total mold cost. The simplest molds use an interchangeable set of ejector pins, each with the same diameter and length; such a design follows design for manu­ facturing and assembly guidelines [1] for part interchangeability to minimize the mold assembly time and maintenance requirements. However, most molds use a number of different ejector pins with varying diameter, section, and length. While the cost of additional cost of the pins is small compared to the cost of the mold, the mold designer and mold maker should be sure to key and label each ejector pin so that they can be readily maintained by the molder. The cost of the ejector system can increase dramatically with the use of slides, lifters, and other ejection sub-assemblies. Again, the goal of the mold designer is to provide a simple, cost-effective, and reliable design that satisfies the previously described objectives. The mold designer should not just consider the initial design and tooling costs, but the operational, maintenance, and failure costs as well.

„„11.2 The Ejector System Design Process The ejector system design is determined first by the required layout of the mold’s parting surfaces, and subsequently by the detailed design of the various components required to eject the molding(s). The following design process assumes that

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the molded part has been properly designed with a minimum number of undercuts, etc. Otherwise, the mold designer should revisit the part design to simplify the mold’s ejection system design.

11.2.1 Identify Mold Parting Surfaces As described in Section 4.1, the product geometry and orientation in the mold ­determines the number and location of the mold’s parting surfaces. If the mold has no undercuts or special requirements, then only one parting surface may be necessary. However, if the mold has internal or external undercuts, then additional parting surfaces may be necessary along with the associated ejection components to actuate the sliding cavity and/or core inserts to release the trapped areas of the moldings so that they may be ejected. Such “split cavity molds” are discussed in Section 11.4.1.

11.2.2 Estimate Ejection Forces The ejection force, Feject , required to remove a molding from a mold core is a function of the normal force between the surface of the molding and the surface of the mold, Fnormal , together with the associated draft angle, f, and the coefficient of static friction, ms , between the molded part and the core insert. To estimate the ejection force, the friction force, Ffriction , is first computed as: Ffriction = ms × Fnormal (11.1)

The ejection force is then calculated as the component of the friction force that is normal to the parting surface: Feject = cos(f)× Ffriction = ms × cos(f)× Fnormal (11.2)

The relationships between these forces are represented in Fig. 11.5. As the draft angle decreases from zero in Eq. 11.2, the ejection forces decrease with the cosine of the draft angle. The normal force acting between the molded part and the core is driven by the internal tensile stresses in the plastic, which will cause the plastic molding to hug the core like an elastic band. The normal force is estimated as the integral of the residual stresses, s, in the molded part across the area of the molded part.

11.2 The Ejector System Design Process

Figure 11.5 Ejection force vectors

Approximate values for the coefficient of friction vary from 0.3 for highly polished surfaces (with low surface roughness) to more than 1.0 for rough and/or textured surfaces [2]. Table 11.1 provides some coefficient of friction data generated according to ASTM D 1894, Standard Test Method for Static and Kinetic Coefficients of Friction of Plastic Film and Sheeting. Notably, abrasive materials such as filled PA6 have a higher coefficient of friction. Surface finish is very important. In Table 11.1, LaserForm ST-100 refers to a powdered, polymer-coated stainless steel material that is shaped into a green part with a laser and subsequently sintered and infiltrated with bronze to form a dense, strong part with a 0.2 mm surface roughness [3]. SL5170 is a liquid resin material formed into a three-dimensional mold insert using a stereolithography or polyjet process with a surface roughness of 3.6 mm. The apparent coefficient of friction increases with surface roughness. The very high coefficient of friction equal to 5.47 between HDPE and SL5170 is believed to be caused by molecular adhesion [3]. Table 11.1 Coefficient of Friction for Different Polymers and Mold Materials Polymer

P-20 Steel [4]

LaserForm ST-100 [3]

SL5170 Resin [3]

ABS

0.40

Not available

Not available

HIPS

0.23

0.25

0.38

HDPE

0.35

0.54

5.47

PA6

0.54

Not available

Not available

PC

0.31

Not available

Not available

PP

0.36

Not available

Not available

The estimation of the residual tensile stresses is a complex function of the processing conditions, mold geometry, and material properties. A detailed treatment is well beyond the scope of this book; Burke [4] provides a good thesis on the topic, and modern computer simulations can also provide estimates of ejection forces

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[5, 6]. For the purpose of mold design, conservative simplifying assumptions are ­applied to provide an estimate of the ejection force. The primary assumption is that the tensile stresses in the molding are the result of the thermal contraction of the solidifying polymer within mold. This assumption will cause the analysis to over predict the ejection forces since in practice the polymer (1) may be in a compressive state before the application of thermal shrinkage, and (2) may tend to relax. Since the polymer melt cannot support tensile stress in a fluid state, the thermal strain, e, is estimated for the solidified plastic as the coefficient of thermal ex­ pansion of the plastic material, CTE, multiplied by the difference between the solidification temperature, Tsolidification , and the ejection temperature, Tejection : e = CTE × (Tsolidification - Tejection ) (11.3)

While there will be stress relaxation as the polymer melt becomes rigid, conservative assumptions are that the strain develops with the material at its room temperature modulus, E, and that the molded part is ejected at room temperature (around 20°C). The resulting tensile stress internal to the part can then be computed as a constant throughout the entire molding as: s = E e = E × CTE × (Tsolidification - Tejection ) (11.4)

To estimate the normal and ejection forces, the cross-sectional area upon which the stress effectively acts must be calculated. This effective area, Aeff , is not the projected area of the molding, but rather the cross-sectional area of the molding upon which the residual stresses act. Figure 11.6 demonstrates the governing concept by sectioning the molding into two halves. As previously suggested, the molding is similar to an elastic band wrapped around the mold core. When the molding is sectioned, the normal forces between the two halves are relieved. As such, the normal force can be well estimated as the tensile stress multiplied by the cross-­ sectional area: Fnormal = s × Aeff (11.5)

11.2 The Ejector System Design Process

Figure 11.6 Tensile stresses pulling across effective area

Combining all the previous terms provides the following estimate of the ejection force: Feject = ms × cos(f)× E × CTE × (Tsolidification - Tejection )× Aeff (11.6)

The determination of ejection force is closely related to the concept of “hoop stress” in statics [7]. Generalizing from the prior two examples, the effective area of a ­complex molding with ribs may be estimated as: Aeff = nwall × h × H part + nrib × hrib × H rib (11.7)

where h is the wall thickness of the molding, nwall is the number of side walls, Hpart is the average height of the side walls, nrib is the number of ribs, hrib is the average thickness of the ribs, and Hrib is the average height of the ribs. Rosato [8] provides an alternative estimation of the effective area as: Aeff =

L × Acontact (11.8) P ( P 2h - l P 4h)

where h is the wall thickness of the molding, P is the perimeter of the molding, L is the length or diameter of the molded plastic, l is the Poisson’s ratio of the plastic, and Acontact is the contact area of the molding in contact with the mold core in the line of draw (in other words, the vertical surfaces).

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Example: Estimate the ejection force required to strip a cup molded from ABS off the mold core according to Eqs. 11.7 and 11.8. For Eq. 11.7, the area of the hatched cross-section of the cup in Fig. 11.6 can be assessed from CAD as 348 mm2. A smooth core surface is used with a coefficient of friction of 0.5, and a draft angle of 1°. The modulus, coeffi­ cient of thermal expansion, solidification temperature, and ejection temperature are taken from Appendix A. The ejection force is then estimated as:

( )

Feject = 0.5 × cos 10 × 2.28 GPa × Feject » 3,900 N » 870 lb

8.83 ×10-5 × (132 °C - 20 °C)× 348 ×10-6 m2 °C

For Eq. 11.8, the area of the shaded section on the core in Fig. 11.6 can be assessed from CAD as 9640 mm2, with an average perimeter of 188 mm and an average diameter of 60 mm. Given a Poisson’s ratio of 0.35 for ABS, the effective area can then be assessed as

Aeff =

60mm × 9640 mm2 = 119 mm2 188 mm (188 mm (2× 3 mm) - 0.35 ×188 mm (4 × 3 mm))

Applying this effective area to Eq. 11.6, Rosato’s analysis suggests that the ejection force for the cup would be 1320 N (300 lb). 

Example: Estimate the ejection force required to strip the laptop bezel off the mold core according to Eqs. 11.7 and 11.8 assuming it was molded from ABS. The laptop bezel is more geometrically complex than the molded cup, and so involves greater effort to estimate the effective area for the calculation of the ejection force. Some different cross-sections of the laptop bezel are shown in Fig. 11.7. At first, the mold designer may first consider using the area of only section A-A or section B-B as the effective area. However, if the molding was cut along only one of these sections, then the resulting halves of the moldings would still remain on the core due to the shrinkage along the other sections. As such, the mold designer might consider adding the areas of section A-A to that of section B-B to estimate the effective area. However, this area would still be insufficient. If the molding were cut along these two sections, then the resulting pieces would still remain on the core due to the tensile forces between the ribs. For example, the normal force between the indicated ribs in section B-B is driven the tensile stresses across the area of section C-C, which is dominated by the cross-sectional area of the top surface if the molding.

11.2 The Ejector System Design Process

Figure 11.7 Different cross-sections of laptop bezel According to Eq. 11.7, the effective area for the laptop bezel is:

Aeff = 4 × 0.0015 m × 0.01 m + 7 × 0.001 m × 0.01 m = 1.3 ×10-4 m2 This effective area can be substituted into Eq. 11.6 along with a 1 degree draft angle to estimate an ejection force of:

( )

Feject = 0.5 × cos 1 × 2.28 GPa × Feject » 5800 N » 1300 lb

8.83 ×10-5 × (132 °C - 20 °C)×1.3 ×10-4 m2 °C

For Eq. 11.8, the area of the bezel on the core can be assessed from CAD as 18,400 mm2, with a perimeter of 750 mm and a characteristic length of 182 mm (evaluated as the geometric mean of the length and width equal to (145 mm × 228 mm)^0.5). Given a Poisson’s ratio of 0.35 for ABS, the ­effective area can then be assessed as Aeff =

182 mm ×18,400 mm2 = 21.6 ×10-6 m2 750 mm (750 mm (2×1.5 mm) - 0.35 × 750 mm (4 ×1.5 mm))

Applying this effective area to Eq. 11.6, Rosato’s analysis suggests that the ejection force for the bezel would be 360 N (81 lb). 

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Reflecting on the foregoing analyses and results, the use of Eq. 11.7 for the effective area is recommended since it explicitly considers the number of vertical walls and ribs that exert drag forces on the core during molding. The analysis is conservative in that assumptions have been made regarding the solidification temperatures and material properties to provide estimates of ejection force that are higher than should be encountered. As such, the use of this analysis for the ejection force should result in effective ejection system designs without the use of safety factors. To validate the analysis, it is useful to compare the predicted ejection forces with the typical ejection forces provided by commercially available molding machines. A survey of several different sized machines available from different machine suppliers indicates that the ejection force provided by the machine is typically 2 % of the clamp tonnage of the machine. This percentage means that if a molding machine provides 1000 kN of clamp force, then the machine may provide 20 kN of ejection force. For comparison, the molded cup had an expected clamp tonnage of 400 kN and an expected ejection force of 4 kN while the molded bezel had an expected clamp force of 1400 kN and an expected ejection force of 5.8 kN. In both examples, the analysis predicted an ejection force less than 1 % of the clamping force. Since molding machines would be expected to be designed to provide a ­ejection force for worst case scenarios, the analysis results are believed to be on the right order of magnitude and appropriate for ejection system design.

11.2.3 Determine Ejector Push Area and Perimeter Once the ejection forces on the molding have been estimated, the next step is to determine the total “push area” of the ejectors onto the molded part. Specifically, there is a minimum push area that is required to avoid excessive compressive stress on the ejection system components as well as excessive shear stress on the plastic moldings. These two phenomena are illustrated in Fig. 11.8 for a single pin ejecting a portion of the laptop bezel.

Ashear = π Dpin h

Fpin

Figure 11.8 Compressive and shear stresses at ejection pin

11.2 The Ejector System Design Process

When the pin is actuated with the ejection system, a reaction force, Fpin , will develop between the pin and the molded part before the part is ejected. The magnitude of this force is related to the total ejection force required to eject the part as well as the number, location, and geometry of the ejectors. The compressive stress on the pin, spin , is the force on the pin divided by the area of the pin, or: spin =

Fpin Acompression

=

4 Fpin 2 p Dpin

(11.9)

To avoid fatigue and/or buckling of the ejection system components, compressive stress levels must be maintained below a critical threshold. This critical stress, sfatigue_limit , is dependent upon the material and treatment of the ejectors. Most ­ejector pins and sleeves are made of hardened materials, with fatigue limit stresses on the order of 800 MPa. A conservative mold design, however, may assume a lower fatigue limit stress of 450 MPa for P20. In either case, the total push area of all ejectors, Aejectors , to avoid excessive compressive stresses must meet the requirement: Aejectors >

Feject sfatigue_limit

(11.10)

Example: Calculate the combined push area of all ejectors required for the bezel mold to avoid excessive compressive stresses in the ejector pins. Also, calculate the required diameter if 20 ejector pins of the same diameter are to be used. Assuming a fatigue limit stress of 450 MPa, the required push area is ­calculated as:

Aejectors >

5800 N =12.8 ×10–6 m2 =12.8 mm2 450 MPa

If 20 pins are to be used, then each pin should have a cross-sectional area of at least 0.5 mm2. The minimum diameter is then: min Dpin >

4 ×12.8 mm2 20 pins = 0.90 mm p



The required push area to avoid excess compressive stresses in the ejection system is very small in most molding applications given the relatively high strength of steel. The compressive strength of the pins is thus not constraining the design.

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However, the ejector system must also have enough push area to avoid developing excessive shear stresses in the molded parts upon ejection. For the example of Fig. 11.8, the shear stress exerted on the molded part is the force on the pin divided by the area of the molded part directly above the circumference of the pin, or: t part =

Fpin

=

Ashear

Fpin p Dpin h

=

Fpin Wpin h

<

splastic_material 2

(11.11)

where Ωpin is the perimeter of the pin. If the shear stress in the molded part is too high, then the part can permanently distort near the pin (an effect known as “push pin”), permanently warp, or even fracture. To avoid these defects, the mold should be designed such that the perimeter around all the ejectors provides a shear stress less than one-half the yield stress of the material, splastic_yield . This requirement leads to the following relationship for the total perimeter of the ejector system, ­Ωejectors : Wejectors >

2Feject splastic_yield h

(11.12)

where h is the wall thickness of the molding above the ejector pins and splastic_yield is the yield stress of the plastic. Example: Calculate the combined perimeter of all ejectors for the bezel mold. Also, calculate the minimum required diameter to avoid excessive shear stresses in the ABS molding if 20 ejector pins of the same diameter are to be used. Assuming a yield stress of 44 MPa for ABS, the required combined peri­ meter of all ejectors is:

Wejectors >

2 × 5800 N = 0.175 m 44 ×106 Pa × 0.0015 m

If 20 pins are to be used, then the minimum pin diameter is then: min Dpin >

Wejectors npins p

=

0.175 m 20 pins = 2.8 mm per pin p 

The analysis and examples indicate that for most molding applications, the design of the ejector system is driven more by the yield stresses exerted on the plastic molding rather than by the compressive stresses on the pin. However, compressive stress can cause buckling in long, slender members such as ejector pins. For this

11.2 The Ejector System Design Process

reason, further analysis of the compressive stresses is important, and will be subsequently used to avoid pin buckling.

11.2.4 Specify Type, Number, and Size of Ejectors Once the required push area and perimeter of the ejectors are known, different ejector systems designs can be developed. The mold designer should consider different designs with a varying number and sizes of ejectors. There are advantages and disadvantages to ejector system design strategies having a large quantity of small ejector pins compared to having fewer but larger ejector pins. With respect to tooling and operation costs, a smaller number of large ejector pins are preferred. There are two primary reasons. First, a smaller number of ejectors requires a lower number of mold components and features to be machined. For this reason, the mold is less expensive to manufacture and maintain. Meanwhile, the larger size of the ejectors will tend to have very low compressive stresses and thus be less ­susceptible to buckling. With respect to design flexibility and mold operation, however, a larger number of small ejector pins is preferred. There are several reasons. First, the greater number of ejector pins allows for more frequent placement of the ejectors across the cavity. This higher density of ejectors will tend to provide for more uniform venting and ejection. At the same time, smaller sized ejectors allow greater design flexibility with respect to the placement of the ejectors. As previously discussed, molds ­contain many tightly spaced and complex features so small ejector sizes allows pins to be effectively placed between cooling lines, down narrow cores, on side walls or ribs, etc. The mold designer should remember that the above analysis only provides a lower limit for the number and size of the ejectors. The mold designer can always add ejectors or increase the ejector size to improve the uniformity of ejection or reduce stress in the molded part. The mold designer must also determine the type of ­ejector to be used at various locations. Typical components include ejector pins, ejector blades, ejector sleeves, stripper plates, slides, lifters, angle pins, core pulls, collapsible cores, expandable cavities, split cavity molds, and others. The selection of the most appropriate components is heavily dependent on the requirements and geometry of the application. For this reason, the use of each of these components will be subsequently discussed.

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Example: Analyze and discuss the design of the ejector system for the laptop bezel consisting of 10 and 40 ejector pins of the same diameter. The minimum pin diameters are calculated according to the previous example for the various number of ejector pins. Both designs provide the same total perimeter around the ejectors and so also provide the same shear stress on the molded part. If only 10 pins are used, then the minimum pin diameter would be approximately 5.6 mm. Assuming uniformly distributed ejection forces, the compressive stresses in each of the 10 pins would be 24 MPa. By comparison, if 40 pins are used, then the minimum diameter would be ­approximately 1.4 mm. The compressive stress in each of the 40 pins would be approximately 95 MPa.

Figure 11.9 Candidate ejector pin layout for laptop bezel The design for 10 evenly spaced, 5.6 mm ejector pins is shown in Fig. 11.9. Since the gates are located on the left and right side walls, the ejector pins located at the center of the top and bottom walls would provide needed ­venting at the end of flow. This design, however, may be unsuitable for two reasons. First, there may not be enough ejectors at locations near where the molding will stick in the mold. In particular, the ribs and bosses will tend to shrink onto the core and so ­require nearby ejector pins. Second, the ejector pin diameter is slightly large given the close proximity of the nearby ribs. In this design, only 1 mm of steel separates the ejector hole from the surface of the mold cavity. With high melt pressures, stresses will develop in the steel, deforming the ejector holes to be nonround, causing the ejector pins to bind. Eventually, cracks will propa­ gate between the ejector hole and the mold cavity. For these reasons, the ejector pins should be made smaller and more strategically located. 

11.2 The Ejector System Design Process

11.2.5 Layout Ejectors The previous example implied that the effectiveness of an ejector is not simply a function of its size but also its location. In general, ejectors will be more effective when placed near the locations where the ejection forces are generated. Furthermore, the ejectors will be more effective when pushing on rigid areas of the molded part. A common but ineffective placement arises when ejector pins are uniformly distributed across the mold cavity. Such an approach can give rise to the layout design shown in Fig. 11.10 with an ejector pin located relatively far from the ribs and side walls of the molding. Since the molding has shrunk onto the core, the ejection force is being generated by the friction between the molding and the mold core at the rib and side wall. By placing the ejector pin far from these two sticking points, a significant moment and deflection will be applied before the molding is stripped off the core.

Figure 11.10 Ejector pin poorly located away from sides of core

The design can be improved by adding ejector pins closer to the rib and side wall as shown in Fig. 11.11. In this case, three additional pins are added to provide ejection forces close to the molding. To avoid excessive stress in the core insert due to the provision of the ejector hole, an allowance of at least one ejector pin diameter should be specified between the surface of the mold cavity and the surface of the ejector hole. However, this ejector pin layout may lead to a potential cooling issue since there may not be enough clearance to provide a cooling line in the core insert between the rib and the side wall. As such, the diameter of the ejector pins may be reduced slightly to allow the addition of a cooling line if desired.

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11 Ejection System Design

Figure 11.11 Ejector pins better located near core side walls

Another alternative layout is to provide an ejector pin underneath the rib or side wall as shown in Fig. 11.12. This design has the direct benefit that the friction force and the ejection force are in-line, such that very little deformation of the molding will occur. One common problem arises due to the thinness of the rib and side wall compared to the larger ejector pin diameter. To avoid very small ejectors that may buckle during operation, a solid boss or “ejector pad” may be provided on the rib. When the ejector pin is actuated forward, the force is transmitted from this pad down the length of the rib and to the surrounding areas of the part. Since the ejector pin pushes directly on the ejector pad, no draft angle is required so the ejector pad diameter can be maximized.

Figure 11.12 Ejector pin located under rib with ejector pad

One issue with such a use of the ejector pad, however, is the high volumetric shrinkage that can lead to sink on the aesthetic surface of the part. For this reason, a cored out boss ejected with an ejector sleeve (subsequently discussed) can pro-

11.2 The Ejector System Design Process

vide for higher quality ejection albeit with a higher mold manufacturing cost. The need for ejector pads can also be eliminated through the use of contoured ejector pins as shown in Fig. 11.13. In this case, the ejector pin is aligned with one side of the rib or wall, and then contoured to push on the top surface of the feature. The pin is then contoured and extended down along the side of the feature so as to also push on the parting plane of the molding. Compared to the previous designs, this layout allows for effective transmission of the ejection forces and compact ejector pin spacing without any changes to the molded part design.

Figure 11.13 Contoured ejector pins located on side walls

The use of a contoured ejector pin design requires careful ejector pin design as well as careful alignment of the ejector pin to the part features. Furthermore, there is a possible problem that can arise with the use of contoured ejectors extending outside the parting line of the mold cavity as indicated in Fig. 11.13. Specifically, if the ejector pin is too short, then a gap will form between the top of the ejector pin and the opposite surface of the cavity insert. If this gap is larger than the thickness of a vent, then flash is likely to occur. Meanwhile, if the ejector pin is too long, then the pin will interfere with the opposing plate and be compressed upon mold ­closure. With repeated ejection cycles, the pin can fatigue and buckle. Given that the required length of the ejector pin is difficult to precisely determine due to the stack-up in tolerances across the mold assembly, the mold designer may wish to use a “steel-safe” approach with multiple length adjustments. Alternatively, the mold designer may choose to place the ejector pin within the mold cavity and ­contour the pin as for the rib in Fig. 11.13. In these cases, slight errors in the ­contour of the pin will be on nonaesthetic surfaces and so be less significant with respect to the quality of the molding.

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11 Ejection System Design

11.2.6 Detail Ejectors and Related Components After the number, layout, and geometry of the ejectors have been determined, the detailing of the design should be completed to ensure robust mold assembly and operation. There are several very specific issues that need to be addressed. First, the mold designer should recognize that the mold assembly is complicated by the large number of ejector system components that must be simultaneously mated to the core inserts. This issue is compounded by tolerance stack-up across multiple plates in the mold assembly. Taken together, the mold assembly can consume a fair amount of time and result in damage of valuable mold components. To facilitate the mold assembly, careful detailing is needed wherever the ejector system components interface with other components in the mold. Figure 11.14 provides a top and section view of a round ejector pin (left) and a contoured ejector pin (right). Detail B of Fig. 11.14 indicates that a clearance can be provided between the pin and the bore of the ejector hole for the purpose of venting displaced air during the molding process. The analysis of the vent’s clearance was provided in Chapter 8, indicating that typically a clearance of 0.02 mm (0.001 in) is provided for a sliding bearing length of the order of two to three diameters of the ejector pin. Beyond this bearing length, the ejector hole should step to a larger size so as to not restrict the sliding of the pin. The size of the clearance is not critical but rather only limited by the interference with other nearby components. A chamfer should be provided from the larger diameter to the bearing/venting diameter. Otherwise, the ejector pin would tend to hang up on the sharp corner during mold assembly, which can hamper the mold assembly when trying to locate a multitude of ejector pins.

Figure 11.14 Clearances around ejector pin

11.2 The Ejector System Design Process

The larger clearance between the ejector pin and the ejector through-hole not only serves to eliminate the sliding friction between the pin and the plate, but also provides needed slop to allow for misalignment between the axes of the ejector holes in various plates. The specified clearance should exceed the total stack-up of the holes’ positional tolerances across the mold plates. Since typical drilling tolerances are on the order of 0.25 mm, a diametral clearance of 0.5 mm should be sufficient in most mold making applications. Furthermore, a generous chamfer should be provided at the interface between the core insert and the support plate. As indicated in Detail C of Fig. 11.14, this chamfer assists the guiding of the ejector pin from the support plate into the core insert during mold assembly. The detailed design of the ejector retainer plate is shown in Fig. 11.15. As shown in Detail D, a counterbore is provided in the ejector retainer plate to pull the head of the ejector pin(s) away from the parting plane of the mold when the ejector system is being retracted. To provide clearance for misalignment of the positions of the ejector holes, the counterbore is provided a generous tolerance so that the centerline of each ejector pin is governed by the mating of the pin with the reamed e­ jector hole in the core inserts. If a contoured pin is used, the head of the pin is typically provided with a flat as shown in Detail E. A parallel slot and locating dowel are provided in the ejector retainer plate to maintain the correct orientation of the contoured ejector pin.

Figure 11.15 Retention and alignment of contoured ejector pin

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11 Ejection System Design

Whenever possible, the mold designer should specify the same length and diameter of ejector pins to facilitate mold assembly and maintenance. If different ejector pins are used in the mold design, the mold designer and mold maker should be sure to key and label each ejector pin and matching location on the ejector retainer plate so that the mold can be readily maintained by the molder. The mold designer should always avoid designing ejector pins that vary only slightly in their design, since similar pins may accidentally be considered interchangeable by the molder. The incorrect assembly of ejector pins may cause damage to the pins as well as the opposing mold cavity surfaces.

„„11.3 Ejector System Analyses and Designs There are many different components that can be used for ejection system design. The most common components and their usage are next discussed.

11.3.1 Ejector Pins Ejector pins are typically hot forged and cylindrically ground from hard steels (such as H13). Subsequently, the pins are nitrided and polished to provide a very hard and smooth surface for low wear and friction. Ejector pins are available from several suppliers in standard diameters (ranging from 1 mm to 25 mm) and lengths (from 150 mm to 500 mm). Typically, mold makers cut and grind standard ejector pins to the finished length and contour specified in the mold design. However, ­ejector pins may be custom ordered with varying options including different materials or surface treatments, precise diameters or lengths, threads for mating with the ejector plate, flats, grooves, etc. While ejector pins are available in a range of diameters and lengths, especially long pins with small diameters should be avoided. The reason is that such slender pins tend to buckle under load. As shown in Fig. 11.16, the loading of an ejector pin corresponds to a column with the top end supported by the bore of the ejector hole, and the bottom end pinned by the ejector retainer plate. If the compressive load become too large, then the pin may bow or buckle in an unknown direction.

11.3 Ejector System Analyses and Designs

Figure 11.16 Buckling model of ejector pin

For this load case, Euler theory [9] indicates that the critical load, Fbuckling , is:

Fbuckling =

2

p EI 2

(0.7 L)

=

I   p E p R2 4 2

2

(0.7 L)

(11.13)

where E is the modulus of the material, I is the moment of inertia, and L is the length of the ejector pin. For a circular ejector pin of radius R, the moment of inertia is pR4/4. To avoid buckling, the ejector pin must be designed so that the critical buckling force is higher than the forces applied to the pin during part ejection: Fbuckling =

(

)>

p2 E p R2 4 2

(0.7 L)

Fpin Acompression

=

Fpin p R2

(11.14)

which when solved for the ejector pin radius R provides the following result: 1

æ 2.8 × F L2 ö÷4 ç pin ÷÷ (11.15) R > çç çèç p 4 E ø÷÷

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Example: Calculate the minimum diameter of the ejector pins for the bezel molded of ABS and ejected with 20 ejector pins. An ejection force of 5800 N was estimated in a previous example, so the force on each pin is approximately 235 N. The modulus for steel is 200 GPa. The approximate length of the ejector pin is 0.2 m. Then, the minimum ­ejector pin radius to avoid buckling is: 1

æ 2.8 × 290 N × 0.2 m 2 ö÷4 ç ( ) ÷÷ R > ççç ÷ = 1.1 mm 4 9 çè p × 200 ×10 Pa ÷÷ø Given this radius, ejector pins with a diameter of 2 mm would theoretically be sufficient if all pins contributed equally to the ejection force. 

The result from this example indicated that the minimum diameter required to avoid buckling is on the same order of magnitude as the minimum diameter required to avoid excessive compressive stress in the pin or excessive shear stress in the molding as calculated in previous examples. The results from the above buckling analysis and example are strongly dependent upon the length of the ejector pin; if the pin length was longer, then the buckling constraint would be dominant. The mold designer should perform analysis for their molding application to confirm the driving constraint and ensure an adequate ejector design. In practice, there will be a variation in the forces applied to each pin. The causes of the variation may come from differences in the ejector pin lengths, surface finish of the cavity near the pins, etc. For a robust design [10], it is suggested that a ­diameter 50% greater than the minimum required for buckling be adopted. Smaller ejector pin diameters might be desired in some molding applications for aesthetic or pin positioning requirements. If the minimum pin diameter required to avoid buckling is greater than the desired pin diameter, then a stepped pin with a larger diameter shoulder can be investigated. Stepped pins typically have a shoulder approximately 1 mm larger in diameter than the head of the ejector pin, and a typical shoulder length of 50 mm. When necessary, the mold designer can custom order ejector pins with multiple steps and tapers for a given application. If a stepped ejector pin is used, however, the mold designer should ensure that a suitable hole and clearance is specified in the support plate and core insert.

11.3 Ejector System Analyses and Designs

11.3.2 Ejector Blades Ejector blades are typically larger diameter ejector pins that are contoured to ­present a rectangular cross-section to the core insert. As shown in Fig. 11.17, the ejector blade’s large width and small thickness allow for the blade to be positioned directly below ribs. This position is very effective since the blade applies the ­ejection force at the location where the friction forces between the molding and the mold core are generated. Furthermore, the rib is stiff and so will effectively eject nearby portions of the rib and part. Finally, the rib is not an aesthetic surface and so should not be adversely affected by the witness mark left by the ejector blade, though this is a potential area of stress concentration during the molding’s e­ nd-use.

Figure 11.17 Ejector blade design

The detailing of the ejector blade, shown in Fig. 11.17, is very similar to that pre­ viously discussed for ejector pins. Clearances should be provided in the support and core inserts to allow for free actuation of the ejector blade, with the mating being provided between the rectangular section of the ejector blade and the tightly mating surfaces in the core insert. To provide the rectangular hole in the core ­insert, wire or plunge EDM is necessary. The amount of EDM can be minimized by specifying the clearance hole close to the surface of the mold cavity, with a typical land length equal to twice the width of the ejector blade. The mold designer should also ensure that the length of travel between the ejector blade’s tapered shoulder and the narrowed hole in the mold insert exceed the maximum stroke of the ­ejector

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system. Otherwise, the molder may inadvertently seize and damage the ejector blades. Given that this detailing is not trivial, some mold component suppliers not sell guides for ejector blades such that the core insert is drilled with a hole and then fitted with the guide from back to provide the bearing surface for the ejector blade. The compressive stress in ejector blades and the imposed shear stress on the molding by the ejector blades do not usually constrain the design of the ejector blades. Due to their small thickness, however, buckling can be a concern. For this reason, the thickness of the ejector blade should equal the full thickness of the rib. The buckling is still governed by Eq. 11.13, but with the moment of inertia defined for a rectangular section as: I=

1 W H 3 (11.16) 12

where W and H are the width and thickness of the ejector blade (the thickness being the smaller dimension across which buckling is more likely). The governing relationship between the stress in the blade and the buckling stress is: Fblade < Fbuckling =

1 p2 E × W H 3 (11.17) 12 (0.7 L)2

The thickness of the ejector blade is usually set to the thickness of the opposing rib or wall. The maximum length of the blade section can then be verified as: 1

æ1.7 × E × W H 3 ÷ö2 ÷÷ (11.18) L < ççç ÷ø çè Fblade Example: Calculate the maximum length of the ejector blade for the bezel molded of ABS and ejected with 20 ejector blades. An ejection force of 5800 N was estimated previously, so the force on each blade is approximately 290 N. The modulus for steel is 200 GPa. The ­thickness and width of the candidate ejector blade are preliminarily designed as 1 mm and 6 mm, respectively. Then, the maximum length of the ejector blade to avoid buckling is: 1

3 ö2 æ ç1.7 × 200 ×109 Pa × 0.006 m × (0.001 m) ÷÷ ÷÷ = 84 mm L < ççç 290 N çè ÷÷ø

11.3 Ejector System Analyses and Designs

In this molding application, the maximum blade length is computed as 84 mm. In the mold design of Fig. 11.17, the actual blade length (from the cavity ­surface to the taper) is 93.8 mm. As such, this blade design is marginal. The mold designer could choose to add additional blades to reduce the ejection force per blade, use a wider or thicker blade if available, use a push pad with a constant rectangular section across the rib to allow for the use of a thicker ejector blade, or use a stepped ejector blade. 

11.3.3 Ejector Sleeves The function of an ejector sleeve is similar to that of an ejector blade, in that both are typically used to push on a vertical section of the molded part. The design of the ejector sleeve varies significantly, however, since it is a hollow cylinder that slides along a fixed core pin to provide an ejection force at the bottom surface of a molded boss. Ejector sleeves are very effective components for part ejection, since they push on a stiff bosses in the part at a location where friction forces between the molding and the core occur. A typical ejector sleeve assembly design is provided in Fig. 11.18. In this design, a highly conductive core pin is seated in the rear clamp plate and secured with a socket head set screw. The core pin passes through the rear clamp plate, ejector plate, ejector retainer plate, support plate, and core insert to hollow out the desired portion of the mold cavity. The ejector sleeve is held by the ejector retainer plate and passes through the support plate and core insert. In the design of detail A, the end of the ejector sleeve is coplanar with the top of the boss and the adjacent rib. During the ejection cycle, the ejector plate is moved forward, causing the ejector sleeve to slide along the core pin and push the boss off the core pin. Just like the ejector pins and blades, the retraction of the ejector retainer plate causes the ­ejector sleeve to retract into the core insert for the next molding cycle.

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11 Ejection System Design

Figure 11.18 Ejector sleeve design

Given the annular geometry and large moment of inertia for ejector sleeves, there are typically no issues related to stress or buckling. However, the detailed design of the ejector sleeve is especially critical since it slides along between a stationary core pin and the stationary core insert. The axial location of the ejector sleeve is governed by the concentric mating of the ejector sleeve with the ejector hole in the core insert. Since the core pin is internal to the sleeve, the wall thickness and ­concentricity of the molding around the core pin is governed by the tolerance stack-up of the ejector hole, ejector sleeve, and core pin. To reduce dimensional variations in the molded part, clearances for venting should be minimized. Details B to F Fig. 11.18 provide examples of clearances in the various mold plates. The mold designer should ensure that the core pin has a suitable clearance through the ejector plate and ejector retainer plate, otherwise a slight lack of concentricity ­between the ejector sleeve and core pin may cause the sleeve to bind.

11.3.4 Stripper Plates The function of a stripper plate is similar to that of an ejector sleeve, in that both are typically used to push on a periphery of the molded part. The design of the stripper plate varies significantly, however, since it normally pushes on most or all of the entire periphery of the molded part(s). For this reason, the stripper plate has

11.3 Ejector System Analyses and Designs

a significantly larger area than a single ejector sleeve and a completely different construction. The design of a mold with a stripper plate is shown in Fig. 11.19. In this design, the stripper plate replaces the B plate and is made to float between the A plate and the support plate. To locate the core inserts, a locating dowel has been placed to mate the center of the core inserts with the support plate. Socket head cap screws (not shown) are used to securely fasten the core inserts to the support plate. Portions of the stripper plate are designed to extend beneath the bottom surface of the molding, but not to interfere with the outer surfaces of the core inserts.

Figure 11.19 Stripper mold design

As shown in Fig. 11.20, the moldings are ejected by the opening of the mold when the stripper bolt engages the stripper plate and pulls the moldings off the cores. Since the stripper plate fully engages the bottom of the part, the ejection forces are uniformly distributed across the moldings resulting in low imposed stress, little deformation, and reliable ejection. One interesting aspect of this stripper plate ­design is that the ejector retainer plate, ejector plate, and leader pins serve no ­purpose and can be eliminated from the mold, such that the support plate may be used as the rear clamp plate. More conventional designs, however, use the forward actuation of the ejector plate to engage the stripper plate to eject the molded parts.

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11 Ejection System Design

Figure 11.20 Stripper plate actuation

There are some important items to note with regard to design details A and B, which are identified in Fig. 11.20 and magnified in Fig. 11.21. One significant ­issue with respect to this specific molding application is the location of the parting line along the top of the cup. From the viewpoint of mold design, the center of the rounded top would be the best location to mate the stripper plate with the core ­insert since it would provide a reliable sliding surface. However, this mating ­location would result in an undesirable and possibly sharp witness line. As such, the mating location has been moved towards the interior of the core insert. While this provides an improved witness line location and a significant push area for the stripper plate to push on the molded cup, it also results in a sharp edge at the parting line of the stripper plate. This sharp edge can damage the vertical surface of the core insert, and will likely quickly wear. For this reason, the mold designer may wish to avoid the use of a stripper plate or request the redesign of this section of the cup to provide a flat push area to mate with the stripper plate.

Figure 11.21 Potential stripper plate detailed design issues

11.3 Ejector System Analyses and Designs

11.3.5 Elastic Deformation around Undercuts The molding of the lid shown in detail B of Fig. 11.21 presents an entirely separate problem related to the ejection of the molded undercut from the core insert. As shown in this detail, the side wall of the molded lid must bend so that the lip of the lid can escape the undercut on the core insert. If the amount of strain caused by ejection is within the elastic limit of the material, then such undercuts can be reliably molded and ejected from the mold without special concerns. In fact, stripper plates are ideal for such ejection since they provide very uniform ejection forces that are nearly in-line with the friction force between the molding and the core. The strain, e, caused by an undercut during ejection can be readily estimated as the amount of deflection, d, that the part is required to undergo divided by the distance, L, across which the deflection is applied, or: e=

d (11.19) L

Appendix A provides some material properties for various plastics, including the strain to yield. It is observed that most plastics have a strain to yield above 2 %, which is a reasonable mold design guideline. The exception is heavily filled materials, which have a lower elastic limit and tend to fail in a brittle manner. The ejection force for a part that is elastically deformed during ejection can also be estimated. First, the stress in the deformed section of the part can be calculated as the imposed strain multiplied by the modulus of the material, E: s = E e (11.20)

This stress acts as a hoop stress around the perimeter of the part, similar to the previous analysis. The normal force is estimated as: Fnormal = s Acontact (11.21)

The ejection force may be estimated as a function of the normal force, the coefficient of friction, ms , and the draft angle, f. Combining the above terms provides: d Feject = ms × cos(f)× E × Acontact (11.22) L

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Example: Verify that the lid may be elastically ejected off the core with the undercut. Estimate the ejection force and stress exerted on the part during ejection. The relevant dimensions of the part geometry are provided in Fig. 11.22. The strain in the part is:

e=

1 mm d = = 1.3% L 77 mm

which is significant but not excessive. The height of the undercut section of the lid is approximately 3 mm. The contact area under stress is:

Acontact = p Dh = p × 77 mm × 3 mm = 726 mm2 Assuming a coefficient of friction of 0.5 and a draft angle of 0, then the ejection force is estimated as:

Feject = 0.5 × cos(0)× 2.3 ×109 Pa ×

1 mm × 726 ×10-6 m2 = 10,800 N 77 mm

This force is greater than the ejection force required to eject the cup on the opposite side of the mold. Unbalanced ejection forces can lead to uneven wear in the mold due to bending moments applied across the ejector and/ or stripper plates that cause nonuniform stress on guide pins and bushings. To minimize this wear, the thickness of the stripper plate and the size of the guide bushings can be increased. Additionally, the stripper plate should be actuated at two locations that are in-line with the axis of the cavities instead of one central point as indicated in Fig. 11.20.

Figure 11.22 Elastic ejection of undercut

11.3 Ejector System Analyses and Designs

The shear stress on the undercut portion will be approximately:

t=

Feject Acontact

=

10,800 N = 14.9 MPa 726 ×10-6 m2

which is significantly less than the 44 MPa yield stress of the ABS. Accordingly, the ejection of this lid with this undercut by a stripper plate should be ­acceptable. 

11.3.6 Core Pulls The term “core pull” or “side action” generally refers to a device that retracts a core in a direction that is not parallel to the opening direction of the mold. Core pulls allow the molding of parts with relatively large, complex undercuts that might otherwise not be economically feasible to produce. For example, the product design of the bezel may require the molding of lateral bosses and a window as shown in Fig. 11.23. To mold these features, an extra mold insert is required to provide the steel around these features. This mold insert must be removed prior to the opening of the mold and the actuation of the ejection ­system.

Figure 11.23 Undercut features in bezel

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11 Ejection System Design

To design the core pull, it is necessary to first layout the moving mold insert. One design for the moving core for the bezel is shown in Fig. 11.24. In this design, ­cavity and core features are provided to form the bosses, ribs, and window. A key is profiled along the bottom of the core insert to provide a sliding fit with a matched keyway in the B plate and core insert. This keyway will vertically retain the moving core in the mold and also guide the moving core during actuation. A chamfer is provided on the leading edges of the moving core to avoid damaging the mating surfaces of the mold.

Figure 11.24 Layout of moving core

The preliminary mold assembly with the moving core is shown in Fig. 11.25. The A plate, cavity insert, B plate, and core insert all required modifications to accommodate the moving core. Any significant vertical displacement would cause flashing along top of the molding or the bottom of the rib. For this reason, an interlock has been provided between the front of the moving core and the core insert to prevent the moving core from shifting due to the pressures imposed by the melt on the core. The sides of the moving core will prevent lateral displacement and flashing along the sides of the core. Also, a clearance has been provided between the front surface of the moving core and the core and cavity inserts. This clearance ensures that entire clamping force of the actuation cylinder is applied to the window core to prevent flashing of the window.

11.3 Ejector System Analyses and Designs

Figure 11.25 Moving core mold design layout

To complete the design of the core pull, the actuation mechanism must be designed. The first step is to estimate the required actuation force, Fcore_pull , which is directly related to the melt pressure, Pmelt , exerted on the projected area of the moving core, Acore_projected . Fcore_pull = Pmelt × Acore_projected (11.23)

Example: Calculate the maximum force required to maintain the core pull in the forward position during the molding process. The core pull is approximately 22 mm wide by 10 mm in height. While portions of the front face of the core pull are not subjected to the melt pressure, the

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analysis will make conservatively assume that the projected area is 220 mm2. Also, the analysis will conservatively assume a melt pressure of 200 MPa. Using these assumptions, the maximum expected force is:

Fcore_pull = 200 ×106 × Pa × 220 ×10-6 m2 = 44,000N which is approximately 4 tons of force! This result may be initially surprising, but the mold designer should remember that the moving core must provide a closing force equivalent to the clamp force required for the production of a similarly sized molding. Larger moving cores will require proportionally ­larger actuation forces. 

Once the actuation force is determined, the mold designer must select the type of actuator. There are three common types: hydraulic, electric, or pneumatic. While an exhaustive discussion is beyond the scope of this text, the reason for the ­predominance of hydraulic actuators will be briefly discussed. First, hydraulic ­actuators have a power density an order of magnitude above that of pneumatic or electric actuators. This increased power density means that hydraulic actuators are much more compact and common than the other types. As a result, hydraulic cylinders are widely available at low cost with an extremely broad range of bore diameters and travel lengths. Furthermore, hydraulic actuators are easily integrated with the hydraulic and electric systems on many molding machines. ­Accordingly, molders can often utilize molds with hydraulically actuated moving cores without any need for auxiliary equipment. The mold designer must select an actuator that provides the appropriate actuation force and travel. The travel must be sufficient for the moving core to clear the ­envelope of the features of the molded part. If a hydraulic actuator is used, then the diameter of the bore can be calculated as: Dhydraulic_bore =

4Fcore_pull p Phydraulic_fluid

(11.24)

where Phydraulic_fluid is the available pressure of the hydraulic fluid or compressed air.

11.3 Ejector System Analyses and Designs

Example: Calculate the travel and bore diameter of the hydraulic cylinder to actuate the moving core for the bezel. To calculate the bore diameter, it is necessary to know the available hydraulic pressure. While most hydraulic systems are designed for a pressure of 20.7 MPa (3000 psi), many molding machines and auxiliary systems are ­operated at 10 MPa. Assuming the actuation force of 44 kN, the bore dia­ meter can be calculated as:

Dhydraulic_bore =

4 × 44,000 N = 75 mm p ×10 ×106 Pa

Figure 11.26 Mold design with actuated ejectors Inspection of the layout provided in Fig. 11.25 indicates that the required ­travel is 15 mm. A standard cylinder with a bore of 76.2 mm (3 in) and a ­stroke of 25.4 mm (1 in) is selected. The finished mold design with the retracted core is shown in Fig. 11.26. The design requires risers to be inserted between the hydraulic cylinder and the mold to allow for the piston rod to be actuated with the moving core; the ­risers must be of sufficient strength to avoid flexure under load when the ­cylinder pushes the moving insert against the mold core. The height of the risers could have been reduced or eliminated by reducing the length of the moving core and allowing the piston rod to travel within the keyed profile of the mold. This height reduction is desirable since it reduces interference with the actuator, molding machine, and operator. It should also be noted that the cylinder attachments must be located on only one side of the mold. 

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To provide as safe and efficient a mold as possible, the mold designer should specify the use of limit switches to confirm that the moving core is in its forward or retracted position. These position signals can be used by the molding machine to ensure that the moving cores are properly positioned so as to not damage the molded parts or the injection mold during mold opening or part ejection. Furthermore, the mold designer should strive to design the moving core such that the mold opening or part ejection does not damage the mold if the moving core is ­improperly positioned. Consider, for example, the design shown in Fig. 11.25. If the mold is opened and the part ejected with the moving core in its forward ­position, it is most likely that the plastic part will be sheared off at the primary rib by the actuation of the nearby ejector sleeve and ejector blade. Obviously, this event is undesired and should not occur for a properly set molding process. However, such events do occur and molders greatly appreciate a robust mold design that can withstand intermittent abuse without reworking the ejector pins, blades, sleeves, or moving cores.

11.3.7 Slides Core pulls are quite common since they allow moving inserts to be actuated in different directions, strokes, and times. However, core pulls require actuators, ­auxiliary control, and significant space. For this reason, mold designers often prefer to use sliding cores that are actuated by inclined angle pins. One such mold design is shown in Fig. 11.27. In this design, a bronze gib is located in the B plate to provide a lubricated sliding surface for the moving insert. The core insert is provided with an inclined can surface that mates with the angle pin. As the mold opens and closes, the angle pin engages the sliding core, causing the core to move in and out. A retainer plate secured to the B plate prevents the sliding core from falling out of the gib. The detailed design is shown in Fig. 11.28 and warrants further discussion. The angle pin is located through the use of an angle pin insert, which has a flat surface to orient the angle pin in the direction of the sliding action. While there are many ways to design a slide for side action, the angle pin insert is retained between the core insert and a heel block by socket head cap screws in this design. On the ­moving side of the mold, the bronze gib is located with dowels and fastened with cap screws to a pocket cut in the B plate and core insert. The gib provides a keyway into which guides the sliding core. The sliding core itself is very similar to that previously shown in Fig. 11.24.

11.3 Ejector System Analyses and Designs

Figure 11.27 Moving slide layout view

Figure 11.28 Moving slide detail view

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In operation, the clamping of the mold causes two forces to be imposed on the sliding core. First, the angled surface on the heel block contacts the angled surface on the slide to force the slide laterally against the core insert; this lateral force withstands the melt pressure and prevents flashing of the ribs, bosses, and window. Second, the cavity insert contacts the top surface of the slide, which provides a downward clamping force to prevent deflection and flashing around the parting line of the molding. It should be noted that the angle pin does not provide the ­lateral force and is not subjected to significant stress in this design. Well designed clearances, tolerances, and fits are crucial to the function and longevity of the ­sliding core. The application of core slides for side action is limited with respect to the slide direction and stroke. To avoid excessive friction, the bronze gib may be drilled and filled with graphite lubricant. However, any friction can cause sticking of the core slide during actuation so the inclined angle, fangle_pin , between the axis of the angle pin and the mold opening direction is limited to about 20 degrees. The stroke of the slide, Sslide , is then a function of the contact length of the angle pin, Langle_pin , as: Sslide = Langle_pin × sin (fangle_pin ) (11.25)

Example: Calculate the required length of the angle pin for the bezel mold. The mold design uses an angle pin insert with a 20 degree incline. The ­required travel is 12 mm. Then, the contact length of the angle pin is:

Langle_pin =

Sslide

sin (fangle_pin )

=

12 mm

( )

sin 20

= 35 mm

An additional 25 mm of length is required to mate the angle pin with the angle pin insert. The length of the angle pin will be approximately 60 mm, which will be cut to length and finished during mold assembly. This length was used in the design of Fig. 11.24. 

Just as with actuation of core pulls, improper actuation of slides is a significant ­issue. For example, a curious operator or visitor may be intrigued with a mold in a molding machine, and naively move the sliding core. If the mold closes with the core not in its outwards position, then the angle pin will improperly contact the top surface of the slide rather than the inclined bore, and cause the angle pin to bend under even a relatively low mold closing force. To prevent this issue, the mold designer may place a compression spring between the front of the sliding core and the core insert to maintain the core slide in its outward position when the mold is

11.3 Ejector System Analyses and Designs

not closed. In addition, the retainer plate can be designed with a limit switch to provide a signal that the core is in its outward position. If multiple sliding cores are used, then the multiple switches can be wired in series to indicate that all cores are in the proper position and the mold is ready to be closed.

11.3.8 Early Ejector Return Systems This chapter began with an overview of a basic ejector system that used the mold’s return pins to retract the ejector assembly upon mold closure. While this design is simple and reliable, some molders prefer the ejector assembly to be returned prior to the closing of the mold. There are many ways to provide for early ejector return, but the two most common means are ƒƒpositive return with a threaded ejector knock-out rod, and ƒƒthe use of compression springs. Each of these will be briefly discussed. It should be noted that there are many other ways to provide early ejector return, including pneumatic cylinders, hydraulic cylinders, electric motors or solenoids, mechanical cams, and others. However, these systems are less common and so are not detailed. The term “positive return” refers to the confirmed resetting of the ejector system. As shown in Fig. 11.29, one design threads the molding machine’s ejector knockout rod(s) into the ejector plate. After the molding machine pushes the knock-out rods forward to actuate the ejector assembly and eject the moldings, the molding machine can then pull the knock-out rods back. Since the knock-out rods are threaded into the ejector plate, the entire ejector assembly is returned prior to the closure of the mold. As an added benefit, the molding machine’s ejector knock-out system are typically instrumented with position transducers, so positive return provides feedback as to the actual position of the ejection system prior to mold closure. To properly interface the ejection system with the molding machine, the mold designer should confirm the location(s), diameter, and thread of the knockout rods with the molder.

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Figure 11.29 Positive ejector return with threaded knock-out rods

While threaded knock-out rods are relatively simple to design and operate, some mold makers and molders use compression springs to return the ejector assembly prior to mold closure. One design is shown in Fig. 11.30, which uses several compression springs located between the support plate and the ejector retainer plate. When the knock-out rod actuates the ejector assembly, the springs are placed in  compression. When the molding machine retracts the knock-out rod(s), the ­compression springs will tend to reset the ejector assembly. A few notes on the design of compression springs are warranted. First, a support pin should be used in the center of the compression spring to avoid spring buckling when the free length of the spring exceeds four times the diameter of the spring; the support pin should be threaded into the support plate or rear clamp plate to locate the spring. Second, the range of spring compression should be limited to about 40 % of the free length of the spring. The diameter and gauge of the spring should be selected to provide a return force that is a fraction (for example, one-fourth) of the required ejection force.

Figure 11.30 Early ejector return with compression springs

11.4 Advanced Ejection Systems

Both these early return systems are very common, but the positive return with threaded knock-out rods provides several advantages. First, positive return provides feedback to the molding machine about the position of the ejector system. Second, the positive return system requires fewer changes to the mold design. Third, the compression springs limit the range of ejector travel and can be damaged or cause damage if the molding machine forces the ejector assembly beyond the compression spring’s range of free travel. Fourth, compression springs and ejector systems tend to wear such that molds with compression springs frequently fail to completely return the ejector system after an indefinite number of molding cycles. In either case, if early return of the ejectors needs to be guaranteed prior to mold closure, then the mold designer should include a limit switch that is active when the ejector system is fully reset.

„„11.4 Advanced Ejection Systems There are many types of ejection components as the analyzed and designed in the prior sections. There many specialized ejection system designs that have been developed to provide molded parts with very complex exterior details, very complex interior details, an aesthetic surface completely free of defects, and other purposes. Some of the relatively common ejection systems are next discussed.

11.4.1 Split Cavity Molds As discussed in Sections 11.3.6 and 11.3.7, core pulls and sliding inserts are ­commonly used when there is one or more external undercuts. If the section of the cavity with undercuts is very large, or if the exterior of the molded part necessitates a parting plane that is transverse to the mold opening direction, then a split cavity mold is often designed. As the term “split cavity” implies, a split cavity mold is a mold design in which the cavity insert is split into two or more pieces, such that the walls of the cavity can be moved away from the molded part during the ejection stage of the molding cycle. One split cavity mold design is shown in Fig. 11.31 for the molding of bowling pins [11]. The mold includes a top clamp plate 14, a cavity retainer plate 16, and a support plate 12, among others. The split cavity is formed by two moving cavity inserts 23 and 24 that mate with the conical bore 21 in the cavity retainer plate 16 when the mold is closed. Four elongated angle pins 30 are fastened in the top clamp plate and extend through the cavity inserts 23 and 24. Each cavity insert is fastened to two gibs 28 that can traverse in slideways 26.

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When the mold opens, the support plate 12 is moved away from the cavity retainer plate 16. Since the angle pins 30 are stationary and inclined relative to the mold opening direction, the cavity inserts 23 and 24 are forced to move away from each other through a cam action. Ultimately, a sufficient clearance is produced between the molded part 66 and the cavity side walls 44 so that the molded product may be removed.

Figure 11.31 Split cavity mold

There are a few interesting items to note regarding this particular split cavity mold design. First, there is a significant amount of mold cheek provided in the cavity retainer plate 16. The thickness of the cheek is required to avoid excessive shear stress and deflection of the cavity side walls 44. It is observed that the thickness of the cheek is approximately the same as the depth of the mold cavity, as suggested by the analysis of Section 12.2.4. Second, wear can be an issue in this mold design due to the large mass of the inserts, the length of travel, and the high number of molding cycles. For this reason, the gibs should be specified to include lubricity and be easily replaced when necessary. In addition, wear plates should be in­ corporated between the support plate 12 and the cavity inserts 23 and 24. Third,

11.4 Advanced Ejection Systems

internal cooling of the core is provided through the use of a large bubbler 75 with coolant inlet 74 and outlet 76. Split cavity molds have been designed for quite some time, and this design was not selected solely due to its incorporation of a split cavity design. Another interesting feature of the design is the forward actuation of the core pin 50 during the filling and packing stages to provide injection compression molding. This actuation is needed to compensate for the very high volumetric shrinkage during the solidification of the thick side walls 54 of the molded part. As such, the mold design includes a bearing 46 that supports the shoulder 65 of the core pin. Since the molded part 66 will tend to shrink onto the core pin 50, the core pin must be retracted after the mold is opened as shown in Fig. 11.31 to release the molded part.

11.4.2 Collapsible Cores Split cavity molds are often used when the part design includes complex and undercutting external surfaces. Collapsible cores are often used when the part design includes complex and undercutting surfaces on the interior of the part. The design of a mold which includes a collapsible core is shown in Fig. 11.32, which was ­developed to mold the head of a doll with a nearly uniform wall thickness [12]. The mold cavity (14 and 15 together) is formed by two cavity inserts 12 and 13, which are hollowed out by a collapsible core 17. In this design, the collapsible core is comprised of eight segments: 18, 19, 20, 21, 22, 23, 24, and 25. Four of the segments 18, 19, 20, and 21 are mostly triangular in section and fitted at the corners with a contoured outer surface in the desired form of the core. The other four segments 22, 23, 24, and 25 are mostly planar in section and fitted between the corner segments with a contoured outer surface to complete the desired form of the core. A core rod 37 is located at the center of the core, and prevents the radial displacement of the eight segments when the collapsible core is assembled. To prevent the axial displacement of the collapsible core, all eight segments have a stem 35 with external threads 35a that engage the internal threads 39 in a sleeve 38. The operation of the collapsible core relies upon the threads 37b of the core rod 37, and their engagement with the threaded passageway 41 of the sleeve 38. Speci­ fically, prior to molding the core rod is rotated within the sleeve so that it fully ­extends until its distal (far) end is flush with the ends of the eight segments to form a rigid core 17. The sleeve with the rigid core is then placed in the mold ­cavity and the part is molded according to conventional practice. Once the part is solidified, the mold is opened and the molded part is removed along with the core and sleeve. The core rod 37 is then unscrewed from sleeve 38 and removed from the inside of the core 17. Without any support, the eight segments can collapse and be removed from the inside of the molded part. The segments, core rod, and sleeve are then reassembled for the next molding cycle.

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Figure 11.32 Mold design with collapsible core

The collapsible core design of Fig. 11.32 allows very complex and undercutting features to be formed internal to the molded part. Because of its design, however, a significant amount of time is required to assemble and disassemble the moving core. To facilitate the design and manufacture of molds with collapsible cores, standard collapsible core designs have been developed and are available from a number of mold base and component suppliers. In typical designs, the actuation of the ejector plate slides the segments along a retaining sleeve, which provides a cam action to collapse the core segments during the ejection of the molded part. The diameter of commercially available collapsible cores ranges from 13 to 90 mm, with a collapse of approximately 6 % of the core diameter. While their collapse is not nearly as much as the design of Fig. 11.32, these standard components support fully automatic molding of small features such as internal threads for molded ­closures.

11.4 Advanced Ejection Systems

11.4.3 Rotating Cores Collapsible cores are relatively simple to incorporate into mold designs when ­using a purchased assembly, and can be used for the forming of threads, dimples, windows, and other internal features. However, one issue with collapsible cores is the formation of witness lines on the interior of the molded part where the core segments interface. Depending on the application requirements, these witness lines may prohibit the use of the collapsible cores. As such, many different mold designs have been developed with rotating cores for the formation and demolding of internal threads. One design is shown in Fig. 11.33 for a 64 cavity mold for the production of threaded caps [13]. The mold design includes cavities 16 that are formed by matching sets of cavity inserts 10 and core inserts 15. The back of each core includes an integral support 17, which is mounted upon a shaft 18 that extends from a coarsely threaded helix 19. The helix is axially located between the rear clamp plate 21 and the support plate 23, and radially supported by bearings 20 and 26.

Figure 11.33 Mold design with coarsely threaded rotating core

Follower pins 30 have been fitted to the actuated plate 29 that engage the threads of the helix 19. Since these pins cannot rotate, the actuation of the plate 29 will cause the rotation of the helix 19, and the subsequent rotation of the threaded

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cores 15. Regarding the design, a coarsely threaded helix is necessary since the torque and wear will increase substantially as the pitch decreases. As such, the required length of the helix is related to the friction between the helix and the ­follower, as well as the number of rotations in the molding application.

Figure 11.34 Mold design with planetary gearing of rotating cores

Another mold design for rotating cores is shown in the plan view of Fig. 11.34 [14]. In this mold design, a central sun gear 84 simultaneously actuates multiple planetary gears 86, which in turn drive shafts 88. The core inserts are not shown in Fig. 11.34, but are keyed to the drive shafts through slots 66. There are many possible designs to rotate the sub gear 84. In this design, a pinion 42 is attached to a shaft 74 which ends at a bevel gear 76. This bevel gear 76 meshes with another bevel gear 78 that is locked to the central sun gear 84. In operation, the mold opening stroke causes ƒƒthe rack 44 to engage the pinion 42, ƒƒthe pinion 42 to rotate the shaft 74, ƒƒthe shaft 74 to rotate the bevel gear 76, ƒƒthe bevel gear 76 to rotate the bevel gear 78, ƒƒthe bevel gear 78 to rotate the sun gear 84, ƒƒthe sun gear to rotate the planetary gears 86,

11.4 Advanced Ejection Systems

ƒƒthe planetary gears 86 to rotate the shaft 88, and ƒƒthe shaft 88 to rotate the cores keyed to slot 66. There are advantages and disadvantages of the mold design of Fig. 11.34 compared to the previous design of Fig. 11.33. The primary advantage is the use of multiple gearing stages to decouple the actuation of the rack and pinion from the rotation of the cores. As such, it is possible to delay and otherwise program the rotation of the cores during the mold opening while avoiding the very large stack height associated with the coarse helix of the previous design. The primary disadvantage is the large number, complex layout, and large volume of the gearing stages. In addition, the planetary layout suggests a radial layout of cavities and so may require very large molds for a high number of cavities. Accordingly, the planetary gear design may be preferable in a mold with a relatively low number of cavities requiring high actuation torques. With either design strategy, the mold designer should ensure that the part geometry is designed to prevent the rotation of the molded part with the rotating core. In some cases, the runner and gate may provide sufficient strength to prevent the molded part’s rotation. In other cases, however, this approach is inadequate since the ejection forces will tend to vary with the material properties, processing con­ ditions, and surface finish as analyzed earlier in Chapter 11. For this reason, the mold design may use some small undercuts or other nonasymmetric features to prevent the part rotation.

11.4.4 Reverse Ejection The cavity inserts in most molds are located within the stationary side of the mold and the core inserts are located on the moving side of the mold. Since the molded part shrinks onto the cores as the plastic cools, the molded parts will tend to ­remain with the cores on the moving side of the mold when the mold is opened. Accordingly, molds are usually designed with an ejector housing and ejector plate on the moving side of the mold such that ejector pins can remove the part from the core. However, this conventional design is problematic in that it does not provide for a purely aesthetic surface, completely free of defects, on either side of the molded part. Witness marks will typically be left on the core side of the molded part from the ejector pins while witness marks will typically be left on the cavity side of the molded part from the feed system. To provide a completely aesthetic surface, molds can be designed with “reverse ejection.” One such design is shown in Fig. 11.35 [15], which includes a mold cavity plate 68 on the moving side of the mold and a mold core plate 38 on the stationary side of the mold 36. The sprue 76 conveys the plastic melt from the machine nozzle 14 through the mold core plate 38 to the mold cavity 40 and 74. Because the

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molded part will tend to remain on the mold core, the stationary side of the mold 36 also includes ejector pins 48 and other components that operate with an ejector plate 30 located between rails 18. Since the molding machine’s ejector rod is located on the moving side of the molding machine and is useless with this mold design, the mold design also includes hydraulic cylinders 32 for actuation of the ejector plate. As a result of this mold design, the entire surface of the molded part opposite the core is free of cosmetic defects.

Figure 11.35 Mold design with reverse ejection

„„11.5 Chapter Review The most typical ejector system designs consist of the ejector plate, leader pins, ejector pins, and the ejector retainer plate. In operation, the molding machine’s ejector knock-out rod forces the ejector plate forward, which in turn drives the ejector pins forward to strip the molded parts off the core. The knock-out rod is then retracted so that the closing motion of the mold can force the leader pins to retract with the entirety of the ejector assembly. While this ejector system design is used in a predominance of molds, molding applications place many diverse requirements on the ejection system. As a result, mold designers can be expected to utilize other components including contoured ejector pins, ejector blades, ejector sleeves, stripper plates, core pulls, angle pins, slides, and springs. When neces-

11.5 Chapter Review

sary, more advanced ejection systems may be provided by third party suppliers or custom designed by the mold designer. The design of the ejection system is governed first by the parting surfaces of the mold and the required ejection directions. If more than one ejection direction is needed, then core pulls, slides, or other components must be planned prior to the detailed design. Next, the ejection forces should be estimated as a function of the geometry and thickness of the part, the draft angle, the coefficient of friction, the material properties, and the processing conditions. Given the ejection force, the mold designer determines the number, size, and location of the ejectors to prevent excessive stress and buckling of the ejection system components as well as excessive shear stress exerted on the molding(s). In general, mold designs using many small ejectors are more expensive to make and maintain but provide greater flexibility and more uniform ejection than a mold design using fewer, larger ejectors. After reading this chapter, you should: ƒƒUnderstand the design and function of basic ejection systems; ƒƒUnderstand the different objectives that must be satisfied in good ejection system designs; ƒƒBe able to identify the mold parting surfaces and ejection directions for a given molded part geometry; ƒƒKnow how to estimate the ejection forces for a given molding application; ƒƒKnow how to estimate the required push area and perimeter of all ejectors to avoid excessive compressive stresses in the ejectors and excessive shear stresses in the molded parts; ƒƒKnow how to specify the type, number, size, and location of different ejectors such as straight pins, contoured pins, stepped pins, blades, and sleeves; ƒƒKnow how to detail the design of various ejection system components to avoid interferences and facilitate mold assembly; ƒƒKnow how to design ejector pins and blades to avoid buckling; ƒƒUnderstand the function of stripper plates, core pulls, slides, and other more ­advanced ejection system components. With the ejector system design completed, the mold design is nearly completed. However, structural analysis and design is required to ensure that the mold can withstand the melt pressures that will be applied for many cycles.

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„„11.6 References [1] Boothroyd, G., P. Dewhurst, and W. Knight, Product design for manufacture and assembly, Marcel Dekker Inc., New York (1994) [2] Menges, G., W. Michaeli, and P. Mohren, How to Make Injection Molds, Hanser, Munich (2001) p. 409 [3] Kinsella, M. E., et al., Experimental determination of friction coefficients between thermoplastics and rapid tooled injection mold materials, Rapid Prototyping J. (2005) 11(3): pp. 167–173 [4] Burke, C. T., Ejection force measurement for the injection molding process, Univ. of Massachusetts, Lowell, MA (1990) p. 116 [5] Wang, H., K. Kabanemi, and G. Salloum, Numerical and experimental studies on the ejection of ­injection-molded plastic products, Polym. Eng. Sci. (2000) 40(3): pp. 826–840 [6] Pontes, A., et al., Ejection force of tubular injection moldings, Part II: A prediction model, Polym. Eng. Sci. (2005) 45(3): pp. 325–332 [7] Bickley, W. G., The distribution of stress round a circular hole in a plate, Philos. Trans. R. Soc., A (1928): pp. 383–415 [8] Rosato, D. V., D. V. Rosato, and M. G. Rosato, Ejection of Molded Products, in Injection Molding Handbook (2012) Kluwer Academic Pub., pp. 332–333 [9] Megson, T. H.G., Euler Theory for Slender Columns, in Structural and Stress Analysis, Elsevier (1996) pp. 608–615 [10] Manickarajah, D., Y. Xie, and G. Steven, Optimisation of columns and frames against buckling, ­Comput. Struct. (2000) 75(1): pp. 45–54 [11] Davis, C. C., Process for making plastic bowling pins, in U.S. Patent No. 4,012,386 (1977) [12] Levine, L., Methods for making hollow articles of plastic material, in U.S. Patent No. 1,076,681 (1913) [13] Chabotte, A. M., Molding apparatus, in U.S. Patent No. 2,984,862 (1957) [14] Sayre, G. B., Unscewing device for the molding of threaded articles, in U.S. Patent No. 2,984,862 (1944) [15] Jesse, E. L., Ejector apparatus for molding machine, in U.S. Patent No. 3,899,282 (1975)

12

Structural System Design

Injection molds are subjected to high levels of pressure from the heated polymer melt. When this pressure is integrated across the surfaces of the mold cavities, forces result that typically range from tens to thousands of tons. The structural design of the mold must be robust enough to not only withstand these forces, but also to do so while producing high quality molded products. To develop a robust structural design, the mold designer should understand the relationships between the pressures, forces, and stresses in an injection mold. Figure 12.1 shows the typical flow of stresses through the mold, platens, and tie-bars. During molding, the melt pressure is exerted against all surfaces of the mold cavities. This pressure results in both compressive and shear stresses in the cavity inserts, core inserts, and support plates. This mixed state of stress is indicated by the diagonal arrows in Fig. 12.1. The molding machine’s platens are also under a significant state of bending to transfer the forces to the tie bars, which are in tension.

Figure 12.1 Stress paths in molding machine and mold during molding

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„„12.1 Objectives in Structural System Design In general, the mold designer wishes to provide a structural design for the mold that will neither break due to fatigue across many molding cycles, excessively deflect under load, nor will not be overly bulky or expensive. These objectives can be explicitly stated as: ƒƒMinimize the stress, ƒƒMinimize deflection, and ƒƒMinimize cost. Each of these objectives is next briefly discussed after which the detailed analysis and design of molds are presented.

12.1.1 Minimize Stress The state of stress varies significantly between the moving and the stationary sides of the mold. For most molds, the cavity inserts are directly supported by the top clamp plate and the stationary platen. As such, the cavity inserts are generally in a state of pure compression so very little out of plane bending occurs. On the moving side, however, the pocket required to house the ejector assembly provides no support for the core inserts. As a result, the core inserts and support plates must transmit the load via both compressive and shear stresses, which will tend to result in significant plate bending. Figure 12.2 plots the predicted von Mises stress in a hot runner mold for the laptop bezel when the surfaces of the mold cavity are subjected to 150 MPa of melt pressure. The von Mises stress, sMises , is a commonly used criterion to predict failure, which is defined as: sMises = s12 - s1s2 + s22 < slimit (12.1)

where s1 and s2 are the first and second principle stresses. To avoid failure, the von Mises stress should be less than some specified stress limit, slimit .

12.1 Objectives in Structural System Design

Figure 12.2 Von Mises stresses during molding

The limit stress is usually determined by two primary concerns. The first concern is that the stress should not be so high so as to plastically deform the mold. When a material is subjected to stress, it will deform or strain. For most materials, the amount of deformation is proportional to the applied stress. The ratio of stress to strain is related to the elastic modulus, E, as: e=

s (12.2) E

where e is the strain that results from an applied stress, s. A material with a higher elastic modulus will tend to deform less given an applied stress. The stress-strain behavior of two common mold metals, P20 and aluminum QC7, are plotted in Fig. 12.3. P20 has a much higher modulus than QC7, meaning that it will exhibit less strain and deformation given an applied load or stress. The yield stress is the point at which the material departs from its linear behavior. The yield stress is also the stress at which the material plastically deforms, meaning that the components with a higher stress will not return to their original shape after the load is removed. Once the material exceeds the yield stress, it can continue to carry load up to the ultimate stress, after which it fails completely.

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1000 900 P20

800

QC7

700 Stress (MPa)

384

600 500 400 300 200 100 0

0

0.1

0.2

0.3

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1.0

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Strain (%)

Figure 12.3 Stress-strain behavior of P20 and Al QC7

All mold designs should be engineered to operate at stresses lower than the yield stress. With respect to specifying a safety factor to relate the yield stress to the limit stress, there are two common approaches. One approach is to simply set the limit stress equal to the yield stress, but then assume a worst case scenario with respect to the load condition. For example, a mold designer working with P20 could assume a limit stress equal to the yield stress of 830 MPa. To ensure the mold does not yield under load, the mold designer should then perform analysis assuming the highest melt pressures that the mold would ever be expected to see, perhaps 200 MPa. Another approach is to set the limit stress equal to the yield stress divided by a factor of safety: slimit =

syield f

(12.3)

where f is the factor of safety, whose value is related to the level of uncertainty and the cost of a potential failure. Typical values range from 1.5 for non-critical mold components to 6.0 for hoist rings [1]. When a factor of safety is used, the mold ­designer should apply the expected melt pressure for the mold, perhaps 100 MPa. To avoid over-designing the mold, the mold designer should not jointly apply a factor of safety with the worst case scenario. While the mold designer might expect that a design based on the yield stress with a conservative factor of safety would be robust, this approach may result in molds that fail after a significant number of molding cycles. The reason is that the contin-

12.1 Objectives in Structural System Design

ued cycling of clamping loads and melt pressure in the mold cavity causes cyclic stresses as shown in Fig. 12.4. Each stress cycle allows small cracks in the mold to open and close, causing a few more crystals at the crack tip to fail. Over the course of thousands of molding cycles, these cracks will grow and propagate through the mold like a wedge driven by a hammer. Once the crack reaches a critical size, the stress concentrations around the crack will cause the mold to fail even when the mold was properly designed with a limit stress specified well below the yield stress. This failure mode is generally referred to as fatigue.

Figure 12.4 Cyclic stresses in molds

Fatigue is a well understood mechanism, and the behavior of various materials has been characterized through cyclic stress testing to a million cycles or more. In ­general, the number of cycles that a mold can withstand will decrease with the applied stress. Figure 12.5 plots the expected number of cycles before failure due to fatigue as a function of the imposed stress for P20 steel [2] and aluminum 6061T6 [3]. This data is generally referred to as “sn” curves where the “s” implies stress and the “n” implies number of cycles. The “endurance stress” is defined as the stress at which a theoretically infinite number of stress cycles can be applied without failure. For most steels, the endurance stress is approximately one-half the yield stress. For P20, the endurance stress is approximately 450 MPa.

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Figure 12.5 Stress-failure curves for QC7 and P20

The data in Fig. 12.5 indicate that 6061-T6 has a much lower endurance stress than P20. There are two very important differences in the behaviors of aluminum and steel. First, the s-n curve for aluminum has a greater slope than that for steel. Second, aluminum does not exhibit an endurance stress limit. In other words, the continued cycling of any stress on aluminum will eventually cause failure due to fatigue. For this reason, the mold designer working with aluminum should carefully select the limit stress according to the desired number of molding cycles. If a mold insert made of aluminum 6061-T6 is to be used for less than 1000 molding cycles, then the mold designer may select a limit stress equal to the yield stress of 483 MPa. If approximately 10,000 molding cycles are expected, then the allowable limit stress drops to 240 MPa. If the mold is to be operated for up to a million molding cycles, then the limit stress should be set to 110 MPa. To summarize, the limit stress is specified according to whether issues related to yielding or fatigue will dominate during the mold’s operation: æs ö slimit = min ççç yield , sendurance ÷÷÷ (12.4) ÷ø çè f

If the mold is to be designed for a low number of molding cycles, then the limit stress can be set to the yield stress and designed using a safety factor or a worst case scenario. If the mold is to be operated for a large number of molding cycles, then the endurance stress should be used as the limit stress. These data are provided for some common mold materials in Appendix B.

12.1 Objectives in Structural System Design

12.1.2 Minimize Mold Deflection While excessive stresses in the mold components can cause damage to the mold, excessive mold deflection is an even greater concern in many molding applications. The primary reason is that excessive mold deflection can cause flashing at the parting lines between the core and cavity inserts. In tight tolerance applications, mold deflection can also cause part dimensions to be out of specification. As such, the mold design may be driven more by the need to minimize deflection rather than minimizing stress. Figure 12.6 plots the deflection of the mold and platens for the stress distribution shown in Fig. 12.2. It is observed that the maximum mold deflection occurs at the center of the mold cavity, with the core surface deflecting 0.24 mm to the left and the cavity surface deflecting 0.12 mm to the right. As a result, the melt pressure in the mold cavity causes the surfaces to separate by a total of 0.36 mm (0.014 in). This deflection will cause any nearby parting lines to open a similar amount. Since this amount is much greater than the vent thickness (typically on the order of 0.02 mm), a significant amount of flashing is expected. The mold design must be improved to reduce this deflection.

Figure 12.6 Deflection during molding

The plotted contours of Fig. 12.6 indicate that there is platen deflection as well. Mold designers and molders usually assume platens to be flat and infinitely rigid. While outside the scope of this book, platen deflection can be a significant issue. In Fig. 12.6, the deflection of the stationary platen is approximately 0.04 mm, roughly

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twice that of the moving platen. The reason is that the mold’s ejector housing tends to transfer forces closer to the sides of the moving platen so there is less applied load and deflection in the center of this platen compared to the deflection of the stationary platen.

12.1.3 Minimize Mold Size The mold designer can investigate different mold materials and designs to reduce the mold stress and deflection. A review of the material properties in Appendix B will reveal that harder materials (such as H13) can withstand much higher stresses than softer materials (such as 1045 or QC7). However, the mold designer should be aware that all steels have nearly the same elastic modulus, around 200 GPa. As a result, the mold designer cannot change the deflection by steel selection, but rather must resort to changing the geometry of the mold. The simplest method to reduce deflection is to increase the thickness of plates. As later analysis will show, this approach is effective since the stiffness of the mold plates is related to the cube of the plate thickness. Even so, the repeated use of very large and thick plates can result in an overly heavy and expensive mold with a stack height that limits the availability of molding machines. For this reason, the mold designer should seek to minimize the size of the mold through appropriate analysis and careful specification of plate thicknesses and support structures such as support pillars and interlocks.

„„12.2 Analysis and Design of Plates In mold design, the term “plate” refers to a prismatic or rectangular structural member with a length and width typically greater than the thickness. The “face” of the plate generally refers to the top and bottom surfaces of the plate that span the width and length directions. The “sides” of the plate refer to the four outer surfaces of the plate that traverse the thickness direction. Mold plates are widely available from a number of suppliers in a variety of sizes and materials. Plates can be provided oversized with a slight (1 mm) stock allowance, or finish ground to tolerances on the order of ±0.02 mm. The majority of the mold consists of plates, including the top clamp plate, A plate, cavity inserts, core inserts, B plate, support plate, ejector plate, ejector retainer plate, and the rear clamp plate. A notable exception is the design of molds with deep cores, in which the core insert is not constructed from a plate but rather from

12.2 Analysis and Design of Plates

a rod; this type of mold design has a separate set of issues that are subsequently discussed in Section 12.3. Each of the mold plates is typically subjected to a load on one face of the plate. While the sides of the plate may be constrained by surrounding plates, the majority of the applied load is carried by compressive and shear stresses and thus ­transmitted through the thickness and across the plate. Plate compression and bending are next separately analyzed.

12.2.1 Plate Compression If the plate is fully supported by underlying mold plates and the mold platen (as typical on the stationary side of the mold), then all plates are in compression and there is negligible plate bending. It should be noted that compressive forces due to mold clamping will tend to cause uniform compressive stresses through the mold plates. The compressive stress, s, is defined as the force, F, per unit of compressed area, Acompression : s=

F (12.5) Acompression

The strain, e, that develops is the stress divided by the elastic modulus, E: e=

s (12.6) E

The amount of deflection, dcompression , is equal to the strain multiplied by the length across which the strain exists: dcompression = e L (12.7)

Deflection due to compression is not usually an issue since 1) it is relatively small and 2) it is uniform across the mold. As such, it does not normally cause flashing or significant dimensional change in the molded parts. As the following example will show, however, the mold designer should slightly increase the depth of the mold cavity to compensate for plate compression if a tight tolerance is specified on the thickness of a part with a deep cavity.

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Example: Estimate the change in the stack height of the bezel mold when clamped with 200 metric tons of force. To provide an accurate estimate of the mold deflection, the compressive stress and strain in each plate and rail could be calculated. Each component’s deflection in the height direction would then be computed and summed to provide the total mold compression. However, a faster but ­approximate estimate can be readily provided by assuming the mold is a monolithic block in a uniform state of compression. Figure 12.7 provides the outside mold dimensions along with some of the critical dimensions. The compressive stress in the mold is approximately:

s=

F Acompression

=

200 t × 9807 N t = 17 MPa 0.381 m × 0.302 m

This stress level corresponds to a strain of:

e=

17MPa s = = 8.3 ×10-5 E 205 GPa

The deflection across the entire mold during clamping is then:

dmold = e L = 8.3 ×10-5 × 403 mm = 0.03 mm In actual molding, the total mold deflection may be twice this amount since the rails on the sides of the ejector housing will exhibit a significantly higher state of stress. However, the above analysis provides a quick estimate on the correct order of magnitude.

Figure 12.7 Bezel mold dimensions for compressive stress analysis 

12.2 Analysis and Design of Plates

Example: Estimate the change in the height of the mold cavity due to the clamping of the surrounding A plate with 200 tons of force. The change in the height of the mold cavity is due to the compression of the plate surrounding the mold cavity. To provide an accurate analysis, the ­compressive stress in the plate surrounding the mold cavity is based on the area of the supporting material shown in Fig. 12.8. This area does not ­include the mold cavity, leader pins, and guide bushings since these components do not transmit any of the clamping force from the stationary to the moving side of the mold. As a result, the compressive stress in the surrounding B plate will be somewhat higher than the 17 MPa previously calculated. The projected area of the cavity retainer plate is approximately:

Acompression = 0.381 m × 0.302 m - 0.248 m × 0.168 m 2 2 æ p (0.020m) ÷÷ö çç p (0.032m) ÷÷ = 0.069 m2 -4 çç + ÷÷ 4 4 çè ø

Given 200 metric tons of clamping force, this projected area leads to a ­compressive stress in the plates surrounding the cavity of:

splate =

F 200 t × 9807N t = = 28.5 MPa A 0.069 m2

This stress level corresponds to a strain of:

e=

s 28.5 MPa = = 1.4 ×10-4 E 205 GPa

Figure 12.8 Support area around cavity

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Given a height of 12 mm, the deflection across the height of the mold cavity during clamping is then:

dmold = e L = 1.4 ×10-4 ×12 mm = 0.002 mm The change in the cavity height due to the mold clamping is very small. ­Furthermore, the melt pressure will be exerted on the surfaces of the mold cavity during molding, which will tend to counteract the mold clamping force. As such, the change in the cavity height during molding does not ­normally need further consideration. 

12.2.2 Plate Bending If the back face of the plate is not fully supported, then shear stresses will develop and cause the plate to bend. Plate bending is a typical issue for the plates located between the ejector housing and the mold cavity on the moving side of the mold. The shear stress, t, is defined as the force, F, per unit of area in shear, Ashear : t=

F Ashear

(12.8)

Figure 12.9 provides an example of a static force analysis of a portion of the bezel mold that is in shear. While the actual shear stresses will vary with the distribution of the melt pressure across the mold cavity, a reasonable estimate can be achieved by assuming a uniform distribution around the perimeter of the mold cavity. As such, the area in shear is:

Figure 12.9 Shear stresses around perimeter

12.2 Analysis and Design of Plates

Ashear = perimeter × height

Ashear = (2Wcavity + 2Lcavity )( H b_plate + Hsupport_plate )

(12.9)

Example: Calculate the shear stresses in the core insert and support plates if the melt pressure exerts a total force of 200 metric tons across the mold cavity. The dimensions from the mold design are provided in Fig. 12.10.

Figure 12.10 Bezel mold dimensions for shear stress analysis The area in shear is:

Ashear = (2× 0.248 m + 2× 0.168 m)(0.12 m - 0.012 m) = 0.090 m2 The shear stress is:

t=

F Ashear

=

200 t × 9807N t = 21.8 MPa 0.090 m2

This shear stress is very low relative to the limit stress of either aluminum or steel. 

The fundamental issue with plate bending in mold design is not the existence of shear stresses in the plates, but rather the development of large deflections across any long unsupported spans of the mold plates. Most molds utilize a moving ejector assembly, and so do not fully support the support plate between rails of the ejector housing. Accordingly, the mold plates behave like a beam in bending. The idealized case is represented in Fig. 12.11 in which the entire load, F, is assumed to be applied to the center of the mold section. This assumption is made to provide a conservative estimate of the maximum deflection.

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Figure 12.11 Plate bending modeled as a beam

The deflection of the plate is conservatively estimated assuming the beam bending equation with a central load as: dbending =

FL3 (12.10) 48 EI

where L is the length of the span and I is the moment of inertia. For a rectangular section, the moment of inertia is: I=

1 WH 3 (12.11) 12

where W is the width of the mold section in bending (in the direction normal to the section of Fig. 12.11) and H is the combined thickness of the core insert and the support plate. Example: Compute the deflection due to plate bending in the bezel mold assuming a ­loading of 200 metric tons from the melt pressure. The width of the mold that is in bending is conservatively assumed to be equal to the width of the mold cavity, which is shown as 248 mm in Fig. 12.10. The combined height of the core insert and support plate is 120 mm. Then, the moment of inertia is:

I=

3 1 0.248 m (0.120 m) = 3.6 ×10-5 m 4 12

The free span in bending is taken as the distance between the inside ­surfaces of the ejector housing, shown as 215.9 mm in Fig. 12.11. The ­deflection due to bending can then be estimated as: 3

dbending =

200 t × 9807N t × (0.2159 m) 48 × 205 GPa × 3.6 ×10-5 m 4

= 0.056 mm

12.2 Analysis and Design of Plates

This deflection is roughly twice the 0.024 mm deflection presented in Fig. 12.6 from the finite element analysis. The reason is that the finite element analysis assumed a uniformly distributed load, which reduces the predicted deflection by 60% compared to the single, centrally located load used above. The pre­ sented analysis methods will tend to over-predict the mold deflection due to bending, but is on the correct order of magnitude and should lead to robust mold designs. 

With respect to multiple-cavity molds, the analysis should be applied to separate portions of the mold cavity as appropriate. Fig. 12.12 provides a top and side view of a layout design for a six-cavity mold. One analysis approach is to lump the melt pressure across three cavities together to compute the applied force, F, which acts primarily on the effective width, W. It should also be noted that the effective plate thickness, H, should not include the thickness of the cores when the cores do not contribute significantly to the stiffness of the mold assembly.

Figure 12.12 Decomposition of separate bending areas

12.2.3 Support Pillars Plate deflection can be reduced significantly through the use of support pillars ­located between the rear clamp plate and the support plate. In general, support pillars are best placed directly under the portions of the mold cavity that generate significant force. By providing direct support of the mold plates, shear stresses and bending close to the support pillar are significantly reduced. A typical design is provided in Fig. 12.13. In this design, a clearance is provided through the ejector plate and the ejector retainer plate. The support pillar is then located using a dowel that mates the center of the support pillar to a hole is in the rear clamp plate. Since the support plate is secured to the rear clamp plate with socket head cap screws, the support pillar is fully secured upon mold assembly.

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Figure 12.13 Typical mold design with support pillar (preload not to scale)

Unfortunately, the location of support pillars can conflict with other components including the ejector pins and the ejector knock-out rod(s). For this reason, different layouts and sizes of support pillars should be analyzed. If mold deflection is a critical issue, then the ejector layout can be adjusted to provide space for several large support pillars at ideal locations. Three potential support pillar locations are provided in Fig. 12.14. At left, two smaller support pillars are located outside the ejector blades; the support pillars are fairly evenly spaced with regard to the span of the bezel. However, the support pillars cannot be placed directly under the bezel face without rearranging the ejector layout. Another design may call for one very large support pillar at the center of the mold so as to avoid interference with the ejector pin layout. However, this support pillar may not greatly reduce the deflection of the mold plates since significant plate bending can still occur due to the loading on the left and right sides of the molding. Furthermore, this design could conflict with the use of a centrally located ejector rod from the molding machine, which is quite common. As another alternative, the layout at right of Fig. 12.14 uses a single support pillar of intermediate size. This

12.2 Analysis and Design of Plates

design requires fewer support pillars than the first design, but has a larger span between the support pillar and the ejector rail and so can allow some deflection due to plate bending.

Figure 12.14 Different support pillar placements

The number, location, and size of the support pillars should be analyzed. One of the complexities of the analysis is that the support pillars are structural members of finite diameter and stiffness. This means that the support pillars will deflect under the compressive load. The core insert and support plate will also deflect with the support pillar. Furthermore, the core insert and the support plate will exhibit bending between the support pillar and the ejector rail. To estimate the total plate deflection, superposition is used to add the deflection due to compression and bending. This concept is shown in Fig. 12.15.

Figure 12.15 Superposition of compression and bending

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To perform the analysis, the forces across the mold must be converted to a set of load cases that is suitable for manual analysis. Figure 12.16 provides the conversion from the melt pressure imposed on the surface of the mold cavity to compression and bending load cases. The total force, F, is the integral of the melt pressure across the length and width of the cavity. To estimate the bending, this force is broken into two equal parts applied at the center of the span between the support pillar and the ejector rail. A force balance can then be applied to determine the forces that must be carried by the support pillar(s) and the ejector rails.

Figure 12.16 Compression and bending load cases

Once the load cases are developed for superposition, the previously presented ­compression and bending analyses may be performed to estimate the stress in the support pillar, s, the deflection due to compression, dcompression , and the deflection due to bending, dbending . The deflection across the surface of the mold cavity can be estimated as a function of the distance, x, from the centerline of the support pillar as: æ 3 L2 x - 4 x3 ÷ö æ xö ÷ (12.12) dtotal ( x) = dcompression çç1- ÷÷ + dbending ççç çè L ø÷ ÷÷ø çè L3

where L is the length of the span from the support pillar to the ejector rail and the range of x is restricted to one-half the length of the span. The maximum deflection of the mold will occur either at the center of the support pillar or half-way between the support pillar and the ejector rail, depending on the relative magnitudes of the deflections. This leads to the following formula for the maximum deflection given the superimposed compression and bending:

12.2 Analysis and Design of Plates

æ ö dcompression dmax = max çççdcompression , + dbending ÷÷÷ (12.13) 2 èç ø÷ Example: Design a support pillar for the bezel mold so that the total deflection is less than 0.1 mm. Analyze the stress in the pillar and resulting deflections. Since the support pillar will support the core insert and support plate underneath one side of the bezel, only this local area of the mold cavity is analyzed as shown in Fig. 12.17. The top and bottom sides of the bezel are close to the ejector rails and thus will not cause significant plate bending.

Figure 12.17 Area of mold cavity local to support pillar Assuming a melt pressure of 150 MPa, the total force, F, exerted by the plastic on this portion of the mold is:

F = P × A = 150 MPa × (168 mm ×13 mm) = 327,600 N = 33 t which is still a significant amount of force. A support pillar diameter of 37.5 mm is initially analyzed. A free-body diagram [4] could be drawn to ­assess the load carried by the rails of the ejector housing and any inter­ mediate support pillars. In the case of Fig. 12.15, a reasonable assumption is that the support pillar must convey one-half of the applied force. Then, the stress in the pillar is:

s=

F 2 327,600 N = = 297 MPa 2 A p (0.0375 m) 4

This stress is very close to the endurance stress for the support pillars if they are produced from SAE1040 steel. The resulting strain is:

e=

s 297MPa = = 0.014% E 205 GPa

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The length of the support pillar is 89 mm (3.5 in). The resulting deflection is:

dcompression = e L = 0.014% × 88.9 mm = 0.13 mm which is greater than the specified maximum deflection of 0.1 mm. To reduce the deflection, the support pillar diameter is increased to 50 mm. The stress is thereby reduced to 167 MPa while the deflection due to compression is ­reduced to 0.07 mm. The analysis of bending is very similar to the previous example. For the pur­ pose of estimating the moment of inertia and local plate bending, the effective width of the mold plate is conservatively estimated as just the width of the mold cavity (13 mm). The moment of inertia is then:

I=

3 1 0.013 m (0.120 m) = 1.9 ×10-6 m 4 12

With the use of the support pillar, the span has been reduced from 215.9 mm as shown in Fig. 12.11 to 108 mm as shown in Fig. 12.17. Applying the bending equation for a force of 33 tons in this area of the mold, the deflection due to bending is then estimated as: 3

dbending =

33 t × 9807 N t × (0.108 m)

48 × 205 GPa ×1.9 ×10-6 m 4

= 0.02 mm

Since the deflection due to bending is quite small, the maximum deflection will occur at the center of the mold due to compression of the support pillar. It should be noted that the thickness of the B plate and/or support plate could be slightly reduced while still meeting the deflection requirement. Further­ more, the deflection due to compression could be greatly reduced by preloading the support pillars. Specifically, support pillars of 88.97 mm (88.9 mm plus the 0.07 mm) can be used such that the pillars compress to their nominal 88.9 mm (3.5 in) length so that cavity becomes flat during molding. 

The mold designer should always consider the distribution of the forces applied by the polymer melt in the cavity to determine the location and size of the support pillars. Uniformly spaced support pillars may be sufficient, but are rarely optimal. Often, it is useful to decompose the mold into discrete areas for structural analysis and design. For example. Consider the structural design of a sixteen-cavity mold in Fig. 12.18. The structural behavior of each half of the mold for eight cavities could be analyzed with a span of 359 mm and a width of 178 mm. The structural design might include two large support pillars under each grouping of four cavities as shown at left, or six smaller support pillars spaced between pairs of cavities as shown at right. The mold designer could also consider one support pillar directly

12.2 Analysis and Design of Plates

under each cavity. Each of these designs can be analyzed. The final design will be determined not only by the structural performance, but also the ejection and maintenance considerations. Again, the simplest and most compact design is usually preferred.

Figure 12.18 Alternative layout designs for support pillars

The mold layout design of Fig. 12.18 is notable for two other reasons. First, this design shows 16 core inserts packed directly in a 4 × 4 grid in one large pocket in the core plate without any intervening mold steel. This design is much more compact and less costly to produce than a design with 16 individual pockets for the core inserts. However, the design will require careful machining, finishing, and assembly of the mold inserts since tolerance issues can cause positional errors and issues during molding such as coolant leaks, flashing at the parting line, or out of tolerance molded parts. ­Second, the design of Fig. 12.18 is a family mold in which each set of eight cavities is molded with different thicknesses and at different melt pressures. Designing the hot runner feed system would be challenging. A stacked “X” type manifold similar to the design of Fig. 6.13 could be used with different diameters for the primary runners to balance the melt pressures during filling. However, given the different part thicknesses and similar packing pressures provided to all the cavities, the mold designer should expect less shrinkage in the thinner lid moldings and ­specify a lower shrinkage value accordingly.

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The described analysis methods provide reasonable guidance for mold design. Alternatively, structural simulations using finite element analysis can be performed using the detailed mold geometry. Such structural analysis techniques [5, 6] are becoming increasingly integrated with computer flow simulations to provide high-fidelity predictions while also reducing significant barriers to routine implementation.

12.2.4 Shear Stress in Side Walls The foregoing analysis focused on plate deflection across the parting plane. However, the shear stresses in the side walls of the mold plates can also result in excess deflection and even failure. This concern becomes especially significant for molds with deep cavities. The cup mold represents a typical scenario with the load case shown in Fig. 12.19. Specifically, the deep cavity provides a tall side wall along which the melt pressure, P, is exerted. If the cavity is very deep, then significant shear stresses and bending deflections can develop. The width of the side wall, from the surface of the mold cavity to the side of the mold, is generally referred to as the “cheek.”

Figure 12.19 Shearing and bending of side walls

A common guideline in mold design is that the width of the cheek, Wcheek , should be equal to the height of the mold cavity, Hcavity . The maximum shear stress in the side wall can be estimated as a function of the height of the mold cavity, Hcavity , the width of the cheek, Wcheek , and the molding pressure, P: t=

F P × H cavity × Wcavity slimit (12.14) = < A Wcheek × Wcavity 2

12.2 Analysis and Design of Plates

To avoid failure, the shear stress should be less than one-half of the limit stress for the material. Applying this constraint to Eq. (12.14) and solving for the width of the cheek provides: Wcheek > 2H cavity ×

P (12.15) slimit

If the mold is made of SAE4140, then the endurance stress is 412 MPa. For a typical maximum injection pressure of 150 MPa, the width of the cheek should be specified as: Wcheek > 2H cavity ×

150MPa = 0.73 × H cavity (12.16) 412MPa

Accordingly, the rule of thumb that the width of the cheek should equal the thickness of the cavity provides a slight factor of safety under typical assumptions. Even though the shear stress may not exceed the specified limit, the mold designer should also verify the deflection of the side wall under load. Assuming that the side wall acts as a simply supported beam with a uniform load, then the deflection due to bending of the side wall can be estimated as: dbending =

4 3PH cavity 3 2EWcheek

(12.17)

Example: Estimate the shear stress and deflection of the side wall in the mold for the cup. Given the 3 mm wall thickness of the cup, a maximum melt pressure of 80 MPa is assumed. The height of the cavity from the parting plane to the top of the cup is 50 mm, and the width of the cheek is 45 mm. The shear stress in the side wall is approximately:

t = 80 MPa ×

50 mm = 89 MPa 45 mm

This shear stress is not problematic given the limit stress of 412 MPa. The maximum deflection of the side wall occurs at the parting plane, and will be approximately: 4

dbending =

3 × 80 MPa × (0.05 m)

3

2 × 205 GPa × (0.045 m)

= 0.04 mm

This stress and deflection should not be an issue so no changes are required to the mold design. 

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12.2.5 Interlocks In the previous example, the deflection of the side wall was not an issue. However, this issue would likely be significant if the melt pressures were higher, the mold cavity was deeper, or the molding tolerances were tighter. The mold designer could increase the width of the cheek to reduce the side wall deflection. However, this approach adds significant size and expense to the mold. Another alternative is to use interlocks on the parting plane near the edges of the mold to transfer part of the bending load from the stationary half of the mold to the moving half of the mold. Round and rectangular mold interlocks are shown in Fig. 12.20. Both types of ­interlocks should be placed on the parting plane and as close to the mold cavities as possible. In general, the rectangular interlock will provide greater resistance to deflection due to its larger size and cross-sectional area across the interlock. However, round interlocks are available in smaller sizes and are easier to install in a mold.

Figure 12.20 Round and rectangular interlocks

A detail view of a mold design incorporating a round interlock is shown in Fig. 12.21. In this design, the male interlock is fit into a through hole in the B plate of the mold. The female interlock is fit into a blind pocket in the deeper A plate of the mold. Both interlocks tightly fit into the surrounding plates, and are retained in the height direction with socket head cap screws. It is important that the mold designer does not jeopardize the structural integrity of the side wall by removing excess mold material when incorporating the interlocks. When the melt pressure is exerted on the side wall, the interlock will transfer part of the load from the A

12.2 Analysis and Design of Plates

half of the mold to the B half of the mold. The use of the interlock effectively doubles the stiffness of the side wall, resulting in a halving of the amount of the side wall deflection.

Figure 12.21 Mold design with round interlock

Since larger interlocks can carry higher loads, the largest interlock should be used that can be readily incorporated into the mold design. If the interlock is exposed to a lateral force, Flateral , exerted by the side wall, then the shear stress in the interlock, tinterlock , can be estimated as: t interlock =

Flateral (12.18) Ainterlock

where Ainterlock is the cross-sectional area of the interlock at the parting plane. If the interlock is made of S7 tool steel, then the design should provide a shear stress less than 300 MPa to avoid failure. It should be noted that temperature differences between the moving and stationary mold halves can cause misalignment of the interlocks and accelerated wear during mold operation. Regular inspection during mold maintenance is recommended since this wear can lead directly to core shifts and changes in part dimensions.

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Example: Estimate the shear stress in the 19 mm diameter interlock used to support the side wall of the cup cavity. The primary uncertainty in this analysis is the estimation of the lateral force applied to the interlock. This estimation of the lateral force is complicated by the round shape of the cup that provides a non-uniform cheek width ­between the guide pins. However, an estimate can be provided by assuming that the interlock is exposed to the lateral force from the nearby surface of the mold cavity. As shown by the hatched section of Fig. 12.22 that represents the nearby plastic, the effective area can be estimated as the product of the interlock width and the cavity height. Of course, the interlock will not be exposed to all of the lateral force from the melt pressure exerted on the side wall of the mold cavity. A conservative estimate is that half of the force will be carried by the interlock, so:

1 Flateral = Pmelt × finterlock × H cavity 2 1 Flateral = 40 MPa ×19.05 mm × 50 mm = 19,050 N 2 The shear stress in the interlock can then be estimated as:

t interlock =

Flateral 19,050 N = = 67 MPa Ainterlock p (0.019 m)2 4

Since this shear stress is less than the 300 MPa limit stress, the interlock is structurally sufficient to transfer half the loading from the side of the mold cavity to the moving half of the mold.

Figure 12.22 Projected view of interlock and cavity 

12.2 Analysis and Design of Plates

12.2.6 Stress Concentrations In mold plates, stress concentrations [7–10] will occur wherever material has been removed between the mold cavity and the supporting plates. Stress concentrations are especially common in injection molds due to the installation of water lines and ejector holes as well as the sharp corners that can be provided with electric ­discharge machining (EDM). The resulting stress distribution about the hole will be similar to that shown in Fig. 12.23. In this example, a hole has been provided in a mold plate at a distance of 1.5 times the hole’s diameter; a two-dimensional, plane strain model [11] was used with a fixed rear surface and symmetry conditions on the left and right sides. A pressure of 10 MPa has been applied to the top surface. The resulting maximum von Mises stress is 29.5 MPa, which corresponds to a stress concentration factor of about 3.

Figure 12.23 Stress concentration about hole

As the hole is moved further away from the mold cavity, the stress concentration is reduced. To evaluate the stress concentration factor, a series of finite element analyses were performed with varying mold geometries. The depth of the hole, from the top surface to the hole centerline, varied from 0.6 to 4 while the horizontal pitch between cooling lines was varied from 2 to 4 to 8. Fig. 12.24 plots the stress concentration as a function of the number of hole diameters from the cavity surface to the centerline of the hole. Fig. 12.24 also includes stress concentration data for different pitches, and indicates that cooling line pitch is not as significant a

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determinant of stress concentration compared to the hole depth. A model of the stress concentration factor, K, was fit to the data for a pitch of four diameters, providing a model for the stress concentration: æ f ö8.8 K = 2.75 + 0.056ççç hole ÷÷÷ (12.19) çè H hole ÷ø

where fhole is the diameter of the hole and Hhole is the distance from the cavity surface to the center of the hole. This model is also plotted as the dashed line in Fig. 12.24. Holes located close to the cavity surface obviously cause significant stress concentrations, and provide a dominating constraint regarding how close cooling lines and ejector holes can be located for mold materials with low yield stresses.

Figure 12.24 Stress concentration as a function of distance

Furthermore, it is observed that a stress concentration of 2.75 will occur even when a hole is located far from the cavity surface. This explains why many molds develop cracks emanating from the waterlines in molding applications with high melt pressures, even when the cooling lines are located far from the cavity surface. Cooling lines seem to cause more significant problems than ejector holes in practice. The reason is that cooling lines corrode, causing internal microscopic stress concentrations. Cracks then emanating from cooling lines will eventually leak and cause quality issues with the moldings. For this reason, molding applications with high melt pressures should be constructed of materials with high endurance stresses such as A6, D2, or H13.

12.2 Analysis and Design of Plates

By comparison, cracks emanating from ejector holes may not ever cause a catastrophic failure. The reason is that the deformation of the ejector hole under load can cause the plate around hole to be supported by the ejector, thereby reducing the stress around the hole. As such, cracks propagating from ejector holes will reach a critical length at which point the elastic deformation of the core insert onto the ejector prevents further crack propagation. Example: A thin wall molding application will utilize a filling pressure of 200 MPa with a core insert constructed of H13. Specify the closest allowable distance for a cooling line with a diameter of 9.5 mm. H13 has an endurance stress of 760 MPa. Since the melt pressure will ­provide a nominal compressive stress of 200 MPa, the allowable stress concentration factor is:

K=

sendurance 760 MPa = = 3.8 snominal 200 MPa

The distance may be evaluated using Fig. 12.24 or calculated using Eq. 12.19–. Solving Eq. 12.19 for the distance, Hhole , provides:

æ K - 2.75 ÷ö-1 8.8 æ 3.8 - 2.75 ÷ö-1 8.8 H hole = fhole çç = 9.5 mm × çç = 6.8 mm ÷÷ ÷ çè 0.056 ø çè 0.056 ÷ø 

Example: A prototype molding application will utilize a melt pressure of 100 MPa with 6160-T6. A mold designer is contemplating placing a 4 mm diameter ejector pin with 0.5 mm of aluminum between the edge of the cavity side wall and the edge of the ejector pin. Calculate the stress level and estimate the number of cycles to failure. The distance from the cavity side wall to the center of the ejector, Hhole , is 2.5 mm. The stress concentration factor, K, is:

æ 4 mm ö÷8.8 ÷ = 6.25 K = 2.75 + 0.056ççç çè 2.5 mm ÷÷ø Given a melt pressure of 100 MPa, the von Mises stress level at the pin will be approximately 625 MPa. This stress level is above the yield stress of 483 MPa provided in Fig. 12.5 so the material will plastically deform on the first molding cycle at this pressure. Since aluminum is quite ductile, the mold ­insert will likely deform around the ejector pin without cracking. The mold may continue to operate, but with significant resistance in the ­actuation of the ejectors and rapid wear at this ejector pin location. 

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„„12.3 Analysis and Design of Cores For the purpose of structural design, a core can be considered shallow when the height of the core is less than both the width and length of the core. Shallow cores, such as for the bezel mold, will not be subjected to excessive stress or deflection caused by the application of the melt pressure to the side walls of the core. As such, shallow cores can be designed according to the previously described analysis for mold plates. Deep cores need further consideration of stress and deflection as next discussed.

12.3.1 Axial Compression The vertical deflection of cores due to compression by the melt can be modeled as previously discussed in Section 12.2.1, though with the compressive forces and cross-sectional areas appropriate for the core as shown by example. Example: Estimate the vertical deflection of the core shown in Fig. 12.25 assuming a melt pressure of 80 MPa.

Figure 12.25 Axial compression of hollow core It may appear that the cooling insert in the design of Fig. 12.25 will fully support the core insert. While this assumption may prove acceptable, a

12.3 Analysis and Design of Cores

more robust design is provided by assuming that the cooling insert provides no support. There are two reasons for making this assumption. First, the ­outer surfaces of the cooling insert may not tightly fit the inner surfaces of the core insert. Any clearance or gap greater than the deflection of the core will completely prevent the cooling insert from supporting the core. Second, the cooling insert may be made from a different material than the core ­insert, such as aluminum. Then, the cooling insert might not withstand the stresses imposed while supporting the core insert. The total deflection of the top surface of the core can be estimated by ­superimposing the compression of the side walls with the bending of the top surface. While the compressive stress distribution in the side walls of the core is not entirely uniform, the average stress is approximately the applied force from the top divided by the cross-sectional area of the side walls: 2

sside_wall =

80MPa × p (63mm) 4 Fvertical P p D22 4 = = = 216MPa 2 Aside_wall p D2 - D2 4 p (63mm)2 - (50mm)2 4 2 1

(

)

(

)

where P is the melt pressure applied to the top of the core, D2 is the outer diameter, and D1 is the inner diameter. This is a fairly high stress level indicating that a mild steel or aluminum should not be used in this application when considering the cyclic loading and possible fatigue. Assuming a steel core insert, the vertical compressive strain in the side walls is:

evertical =

sside_wall E

=

216 MPa = 0.11% 205 GPa

The total height of the core insert is 58 mm. With a strain of 0.11 %, the total vertical deflection at the top of the side walls equal to:

dvertical = evertical × H core = 0.11% × 58 mm = 0.06 mm Bending from the edge of the top surface to the center of the core can be calculated using beam or plate bending equations of Section 12.2.2 and ­added to the displacement due to compression to estimate the total vertical deflection. 

The structural design of deep cores is further complicated since deep cores can bend due to the application of lateral forces from the melt pressure against their sides. Core bending may be a significant problem when cores are slender and have a low stiffness associated with their cross-sectional area, especially when the cores are hollowed out to provide for mold coolant. Analysis and design of the core must ensure that the potential core deflection is not excessive and that the compressive stresses around the perimeter of the core are acceptable. These two concerns are next addressed.

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12 Structural System Design

12.3.2 Compressive Hoop Stresses When a core insert includes a hollow section for a cooling line or other purpose, the side walls must withstand the compressive forces imposed by the melt pressure. The load case is shown in Fig. 12.26.

Figure 12.26 Core insert loaded by melt pressure

The compressive stress, shoop , caused by the melt pressure is: shoop =

Pfcore (12.20) 2hcore

where P is the melt pressure, fcore is the outer diameter of the core insert, and hcore is the thickness of the core’s side wall. To avoid compressive failure in the side walls, the hoop stress must be less than the specified limit stress for the material. This leads to the following constraint on the thickness of the side wall: hcore >

Pfcore (12.21) 2slimit

The constraint can also be used to find the maximum inner core diameter, finner : æ P ö÷ ÷ (12.22) finner < fcore ççç1çè slimit ÷÷ø

A general guideline can be developed for inserts produced from P20 steel with an assumed melt pressure of 150 MPa. Since P20 has an endurance stress around 450 MPa, the thickness of the side wall should be greater than:

12.3 Analysis and Design of Cores

hcore >

f ×150 MPa f » (12.23) 2 × 456 MPa 6

and the maximum internal diameter of the core is: æ 150 MPa ö÷ 2 ÷÷ » fcore (12.24) finner < fcore ççç1èç 456 MPa ø÷ 3

In practice, the mold designer should customize the above analysis by utilizing the maximum melt pressure and endurance stress that are specific to the molding ­application. Example: Compute the compressive hoop stress in the core insert for the cup mold assuming that the outer diameter, fcore , is 60 mm, the wall thickness, hcore , is 10 mm, and the melt pressure is 80 MPa. Also recommend a maximum inner diameter if the core insert is made of aluminum 6160T6 for an ­application targeting 100,000 molding cycles. Given the above assumptions, the compressive hoop stress is:

shoop =

80 MPa × 60 mm = 240 MPa 2 ×10 mm

This is a safe but significant amount of stress. If the core insert is to be made of aluminum, then two different loadings might determine the allowable inner diameter of the core. First, the mold designer should consider the fatigue data provided in Fig. 12.5 for 6160-T6. The allowable stress for 100,000 cycles is 140 MPa. Given possible fatigue failure, the inner diameter should be no greater than:

æ 80 MPa ö÷ ÷ = 0.429fcore = 25.7 mm finner < fcore ççç1çè 140 MPa ÷÷ø Second, the mold designer should consider an overpressure situation wherein the molder accidentally injects the melt at the maximum pressure of the ­molding machine. A single cycle at too high a pressure could cause the core insert to fail. To check this, a melt pressure of 200 MPa can be used with 6160-T6’s yield stress of 483 MPa. This yield stress analysis indicates that the inner diameter should be:

æ 200 MPa ö÷ ÷ = 0.586fcore = 35.2 mm finner < fcore ççç1çè 483 MPa ø÷÷ Comparing the above two results indicates that cyclic fatigue is a more critical issue than yield in an overpressure situation. The maximum inner diameter when using 6160-T6 is 25.7 mm if 100,000 molding cycles are desired. 

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12 Structural System Design

12.3.3 Core Deflection Another common issue with deep cores is excessive deflection or “core bending” due to variations in the melt pressure around the periphery of the core. The variation in melt pressure is often due to side gating as shown in Fig. 12.27. However, slight variations in melt flow can cause significant bending in center gated designs when the cores are very slender (e. g., a core length on the order of 10 times the core diameter). The problem is compounded by the fact that the core bending effect is self-reinforcing, which means that a slight bending of the core facilitates more melt flow and pressure to the thicker portion of the cavity and further bending of the core.

Figure 12.27 Lateral loading of core insert

Core bending can be analyzed through appropriate use of bending equations. Typically, the core is held to the moving side of the mold, and the top of the core is free to bend. The deflection due to the pressure difference, DP, across the core is: dbending =

4 DPfcore H core (12.25) 8 EI

where I is the moment of inertia. For a hollow core with an outer diameter, fcore , and an inner diameter, finner , the moment of inertia is: I=

p 4 4 f - finner (12.26) 64 core

(

)

The magnitude of the pressure difference around the core will vary with the geo­ metry of the molding application. A shorter core, like that shown in Fig. 12.27, will have a pressure difference that is a significant fraction of the pressure required to

12.3 Analysis and Design of Cores

fill the mold (perhaps 50 % of the filling pressure). As the core becomes longer relative to its diameter, the pressure difference around the core will become less compared to the pressure difference along its height (perhaps 10 % of the filling pressure). However, the core deflection is proportional to the fourth power of the core height, so a small asymmetry of the melt pressure can cause a large deflection. Example: Estimate the magnitude of the deflection for the cup’s core assuming that the outer diameter is 60 mm, the inner diameter is 40 mm, the height is 58 mm, and the pressure difference around the core is 40 MPa. The moment of inertia for the core is:

I=

(

)

4 4 p (0.060 m) - (0.040 m) = 5.1×10-7 m 4 64

Assuming a steel core with a modulus of 205 GPa, the deflection is: 4

dbending =

40 ×106 Pa × 0.06 m × (0.058 m) 8 × 205 ×109 Pa × 5.1×10-7 m 4

= 0.03 mm

Since the deflection is small, core bending will likely not be an issue in this application even if the pressure difference around the core was significantly greater. 

Naturally, core bending becomes much more significant as the core becomes more slender. To minimize core bending, the mold designer should utilize solid cores with a minimal length to diameter ratio. When possible, slender core pins should be interlocked with the stationary side of the mold as shown in Fig. 12.28. Such interlocking of the core pin reduces the lateral deflection of the pin to approximately 10 % of the deflection for a pin that is supported on only one end. When ­interlocking or increasing the size of the core is constrained by the geometry requirements of the molding application, the mold designer should strongly ­ ­recommend using a center gate at the top of the corner or two opposing gates at the bottom of the core to minimize the pressure gradient exerted on the core.

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12 Structural System Design

Figure 12.28 Interlocking of slender core into cavity

Figure 12.30 shows the use of flow leaders as another approach to reduce the core deflection. In this design, the flow leaders will assist the melt to travel down the cavity with lower filling pressures. At the same time, the melt will propagate into the thinner adjacent sections of the cavity and partially freeze, thereby preventing the core from deflecting a significant amount even if significant pressure differences arise later in the filling stage. The flow leaders shown on the core in Fig. 12.29 may be undesirable as protrusions on the inner surface of the molded part, especially for applications involving contact with fluids. As such, the flow leaders may be integrated on the outside surface of the molding according to a variety of design configurations set into the cavity insert. Of course, core bending can be greatly reduced by using multiple gates on opposite sides of the core pin or by using a ring gate at the bottom of the core pin. In some mold designs, the core is offset so that then the melt is injected, the core pin deflects to produce a uniform wall thickness. This concept is similar to providing “windage” to compensate for warpage. The success of this strategy will depend on the quality of the mold design, but will remain sensitive to changes in the plastic resin and processing conditions that affect the melt pressure and loading on the core pin.

12.4 Fasteners

Figure 12.29 Use of flow leaders to minimize core deflection

„„12.4 Fasteners The mold design must also include fasteners to rigidly fasten the multiple components of the mold. There are three types of fasteners commonly used in molds. First, fits are used to tightly locate one component within another, such as a cavity or core insert being located within a retainer plate. Second, locating pins or dowels are used to locate one components above another, such as the ejector housing to the support plate. These first two fastening methods only provide fastening across the length and width directions of the mold. To fasten the mold components together in the height direction, socket head cap screws are used wherein the screw’s head is retained in a mold plate and the screw’s threads engage the component to be fastened. Each of these fastening methods is next analyzed.

12.4.1 Fits A “fit” refers to the mating of two components. A clearance fit refers to a mating in which a nominal clearance between the surfaces of the two components. While a clearance fit provides for easy assembly with minimal insertion forces, the clearance between the two components permits the precise location of components to remain unknown. Since tight tolerances are required in molds, interference fits are commonly used to locate the mold components.

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12 Structural System Design

Interference fits occur when the male component has a nominal dimension that is larger than the nominal dimension of the female component, as shown in Fig. 12.30 for a core insert and a retainer plate. Since metals have a high elastic modulus, a rigid interference fit can result when the difference between the nominal ­dimensions is very small, on the order of 0.01 % of the nominal dimension. The tightness and rigidity of the interference fit increases with the amount of inter­ ference between the two components. Unfortunately, the implementation of interference fits is impeded by the dimensional variations imposed in the components’ machining processes. For this reason, standard systems of fits have been developed to provide limits on the dimensions of the components.

Figure 12.30 Location-interference fit for inserts

The fits analyzed here are based on a unilateral hole basis [12] and have been ­converted from U.S. customary units to metric units. Two of the most common standards for fitting include “Preferred Limits and Fits for Cylindrical Parts,” ANSI B4.1-1967 (R1999), and “Preferred Metric Limits and Fits” ANSI B4.2-1978 (R1999). ANSI B4.1 is analyzed here due to its relative simplicity and broad ­applicability, though the mold designer may conform to whatever standard is most appropriate. In this method, rectangular members with width, W, and length, L, are modeled as a circular member with apparent diameter, D, computed as: D = W × L (12.27)

12.4 Fasteners

The tolerance limit, l, on a given dimension is then calculated according to a formula: 1 3

l = 0.001× C × D (12.28)

where C is a coefficient corresponding to the lower and upper limit for the male or female component provided by international standards. Table 12.1 provides coefficients for locational-interference fits (LN1 to LN3) and drive-interference fits (FN1 to FN3). Locational-interference fits are used when the accuracy of location is critical and the components require lateral rigidity. However, locational-interference fits do not provide significant retention force in the height direction, so the components must be secured in the height direction to another component via screws or other means. FN1 to FN3 correspond to drive fits with increasing interference and requiring increasing insertion forces. While drive fits provide semi-permanent assemblies, mold designs usually provide screws or other means for positively retaining the components in the height direction. Table 12.1 Location Tolerance Interference Coefficients [mm] Fit

Cinterference

Female (hole in plate)

Male (insert)

Lower limit

Upper limit

Lower limit

Upper limit

LN1

4.89

0.00

4.93

5.67

9.05

LN2

7.14

0.00

7.84

8.59

13.52

LN3

12.22

0.00

7.84

13.67

18.60

FN1

13.57

0.00

4.93

14.34

17.73

FN2

22.02

0.00

7.84

23.47

28.41

FN3

30.85

0.00

7.84

32.30

37.24

Example: The base of the core insert for the cup mold is 88.90 mm on each side. ­Specify the tolerances for a light drive (FN1) fit. The apparent diameter of the core insert is:

D = 88.90 mm × 88.90 mm = 88.90 mm The lower tolerance limit for the insert dimension is computed with C equal to 14.34: 1 3

lower linsert = 0.001×14.34 × 88.9 = 0.064 mm

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12 Structural System Design

The upper tolerance limit for the insert dimension is computed with C equal to 17.73: 1 3

upper linsert = 0.001×17.73 × 88.9 = 0.079 mm

The lower tolerance limit on the mating hole in the retainer plate is 0: 1 3

lower lplate = 0.001× 0.0 × 88.9 = 0.000 mm

The upper tolerance limit on the mating hole in the retainer plate is com­ puted with C equal to 4.93: upper lplate

1 3

= 0.001× 4.93 × 88.9 = 0.022 mm

The minimum and maximum dimensions on the insert are specified as 88.96 and 88.98 mm, respectively. The minimum and maximum dimensions on the hole in the plate are specified as 88.90 and 88.92 mm. These dimensional limits are shown in Fig. 12.31.

Figure 12.31 Insert and plate dimensions for an FN1 fit 

It may be of interest to estimate the insertion force required to achieve various interference fits, so that excessive insertion forces may be avoided. The insertion force may be estimated by the compressive stress required to strain the components during assembly. The expected amount of interference can be computed as the average male dimension minus the average female dimensions. Alternatively, the expected amount of interference, linterference , can be computed using the formula: 1 3

linterference = 0.001× Cinterference × D (12.29)

12.4 Fasteners

where Cinterference is a coefficient derived from the limit coefficients provided in Table 12.1. Assuming that the plate is much larger than the insert, the compressive stress, s, in the insert is estimated as: s=

linterference × E (12.30) 2D

where E is the modulus of the material. The factor of 2 in the above equation stems from the fact that the compressive stress in the insert will also drive a tensile stress in the plate. Accordingly, the interference causes equal strain in both the insert and the plate. The insertion force can then be estimated as the compressive stress multiplied by the contact area and the friction coefficient: Finsertion = f s (p DH ) (12.31)

where f is the friction coefficient and H is the height of the contact zone between the two components. Example: Estimate the insertion force for the core insert for the cup mold. Assume an FN1 fit with a contact height between the plate and the insert of 42 mm. The expected dimension for the core insert is 88.97 mm while the expected dimension for the hole in the retainer plate is 88.91 mm. The expected amount of interference, linterference , is 0.06 mm. The resulting stress in the steel components is:

s=

0.06 mm × 205 ×109 Pa = 69 MPa 2 × 88.9 mm

Assuming a coefficient of friction of 1.0, the resulting insertion force is:

Finsertion = 1.0 × 69 MPa (p × 88.9 mm × 42 mm) = 808 kN An insertion force of approximately 808 kN or 180,000 lbs is required to drive the core insert into the retainer plate. If a press is not available with this capacity, the mold designer can utilize a location-interference fit. Also, it is desirable to provide a slight taper along the leading edge of the core ­insert to assist in alignment and insertion during assembly. 

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12 Structural System Design

12.4.2 Socket Head Cap Screws Socket head cap screws are the most common fasteners used in molds. A ½"13 socket head cap screw is shown in Fig. 12.32. The primary reason is that socket head cap screws have been carefully designed [12] to optimize the performance of the head, threads, and bolt as a system. Accordingly, the socket head cap screw provides a standard and efficient method for retaining multiple components along the screw’s axis.

Figure 12.32 Typical socket head cap screw

The sizes and load carrying capability of socket head cap screws are related to their size, material, and treatments. Analysis of standard socket head screw designs indicates that the head height is equal to the thread diameter, and that the head diameter is approximately 150 % of the thread diameter. While the strength of the fastener varies somewhat with the coarseness of the thread, the tensile strength of standard DIN/ISO screws can be fairly well estimated by assuming an ultimate stress, sultimate , of 800 MPa multiplied by the cross-sectional area of the outer thread diameter: Ftensile = sultimate

2 p Dthread (12.32) 4

12.4 Fasteners

Example: Specify the size of the socket head cap screws used to fasten the stationary and moving halves of the laptop bezel mold shown in Fig. 12.7. Since this socket head cap screw is used in a critical application where ­failure may result in loss of equipment or life, a worst case scenario is ­assumed. First, the maximum mass of the mold is estimated assuming a ­solid block of steel according to the dimensions provided in Fig. 12.7. The maximum mass of the mold is:

M mold = rmold H mold LmoldWmold = 7800

kg × 0.403m × 0.381m × 0.302m = 362kg m3

Next, the worst case scenario is assumed. The worst case scenario occurs when the mold is clamped to only one side of the molding machine without the support of the moving platen, which may occur when the mold is being installed in the molding machine. Furthermore, the worst case scenario will assume that the entire mass of the mold must be supported by only one ­tightened screw, which may occur if the other cap screws are not tightened or tightened to lesser amounts. The resulting load case is shown in Fig. 12.33. The exerted force on the screw by the mold can be estimated by summing the moments about the locating ring to find:

Fscrew = M mold × ng × g ×

LCOG Lscrew

where g is the acceleration due to gravity (9.8 m/s2), LCOG is the distance ­between the platen and the mold’s center of gravity, and Lscrew is the ­distance from the locating ring to the screw. The coefficient ng relates to the number of gravities that may be exerted on the mold, and is usually set ­quite high for ­safety purposes. Due to the shock of a crane, ng is set equal to 10. Substituting the approximate values from Fig. 12.32 provides:

Fscrew = 362 kg ×10 × 9.8

m 0.2 m × = 47,000 N s2 0.15 m

Solving Eq. 12.32 for the diameter yields:

Dscrew =

4 Fscrew 4 × 47000 N 3 = = 8.65 mm ® 10 mm or " 6 psultimate 8 p × 800 ×10 Pa

The analysis indicates that a 3/8" or M10 socket head cap screw should be sufficient. For reference, the mold base selected for this application was ­ ­provided with ½" socket head cap screws. Failure of cap screws in this mold base is not expected.

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12 Structural System Design

Figure 12.33 Worst case analysis for screw loading 

12.4.3 Dowels Cap screws should not be relied upon to locate mold components given their relatively large radial clearances. As previously discussed, an interference fit should be used to locate one component within the pocket of another. For parallel plates or stacked components, however, dowels or other locating pins should be used as shown in Fig. 12.34. In this design, concentric holes are provided into the coplanar surfaces of the two plates. A dowel then mates with the two holes to locate the two components along the axis of the dowel.

Figure 12.34 Typical locating dowel design

12.4 Fasteners

Manufacturing variances in the holes’ location, diameter, and roundness limit the ability to precisely locate the two components relative to each other. Equation 12.28 can be used for various types of fits by varying the limit coefficients, C, for the dowel and holes according to international standards. Table 12.2 provides co­ efficients for a locational-clearance fit (LC1), locational-transitional fits (LT1 and LT3), as well as the loosest locational-interference fit (LN1). Locational clearance fits are intended for parts that are typically stationary but can be readily disassembled and reassembled. This fit provides the same order of tolerance as threaded fasteners, so is not recommended for injection molds since the large clearance can allow accelerated wear of sliding surfaces. Locational-transition fits provide for tighter control of location, but with the possibility of interference between the dowel and the hole which hinders the mold assembly. Table 12.2 Location Clearance and Transitional Coefficients [mm] Fit

CInterference

Female (hole in plate)

Male (dowel)

Lower limit

Upper limit

Lower limit

Upper limit

LC1

–4.16

0.00

4.93

–3.39

0.00

LT1

–6.38

0.00

7.84

–2.43

–2.51

LT3

–0.73

0.00

7.84

0.72

5.65

LN1

4.89

0.00

4.93

5.67

9.05

Example: A 12 mm dowel is to be used to mate the ejector housing to the support ­plate. Specify the dimensions for an LT3 fit. Estimate the expected clearance between the dowel and the hole, as well as the insertion force in the event of the worst case interference. The apparent diameter of the core insert is:

D = 88.90 mm × 88.90 mm = 88.90 mm The lower tolerance limit for the dowel diameter is computed with C equal to 0.72: 1 3

lower ldowel = 0.001× 0.72×12 = 0.002 mm

The upper tolerance limit for the dowel diameter is computed with C equal to 5.65: upper ldowel

1 3

= 0.001× 5.65 ×12 = 0.013 mm

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12 Structural System Design

The lower tolerance limit on the mating hole in the retainer plate is 0. The upper tolerance limit on the mating hole in the retainer plate is computed with C equal to 7.84: upper lplate

1 3

= 0.001× 7.84 ×12 = 0.018 mm

The minimum and maximum dimensions on the dowel are specified as 12.002 and 12.013 mm, respectively. The minimum and maximum dimen­ sions on the hole in the plate are specified as 12.000 and 12.018 mm. This design is shown above in Fig. 12.33. The average clearance between the two components is 0.0015 mm (or 1.5 mm, equal to the hole’s average diameter of 12.009 mm minus the dowel’s average diameter of 12.075 mm). Given that manufacturing varia­tion exists, it is important to check on the magnitude of the dowel’s insertion force when the hole and the dowel are at their specified limits. The worst case interference will occur when the hole’s diameter is 12.000 mm and the dowel’s diameter is 12.013 mm. The maximum amount of interference, linterference , is 0.013 mm. The resulting stress in the steel components is:

s=

0.013 mm × 205 ×109 Pa = 111 MPa 2 ×12 mm

Assuming an insertion length of 12 mm and a coefficient of friction of 1.0, the application of Eq. (12.31) results in an insertion force of:

Finsertion = 1.0 ×111 MPa (p ×12 mm ×12 mm) = 50 kN This magnitude of insertion force for a dowel is clearly undesirable since ­separation of the mold plates cannot be accomplished manually. The mold assembler would require grinding to reduce the pin diameter to avoid such excessive insertion forces. 

„„12.5 Review Molds are mechanical assemblies that must withstand high levels of stresses imposed by the pressure of the polymer melt. Several constraints drive the structural design of the mold. First, the mold must be designed to avoid yielding given a ­single overpressure situation. Second, the mold must be designed to avoid failure due to fatigue associated with the cyclic loads associated with molding thousands or millions of cycles. Third, the mold must be designed to avoid excessive deflection while molding, which would lead to flashing of the molded parts and accelerated wear on the mold’s parting line. Of these issues, fatigue and deflection tend to

12.5 Review

dominate though the relative importance is a function of the number of mold cavities, the molding pressures, the mold geometry, and the production quantity. Analyses were provided to model the compression of mold plates, cores, and support pillars as well as the bending of plates, side walls, and cores. Superposition of compression and bending can be used to estimate the total deflection of the cavity surfaces. Analyses were also developed for stress concentrations in mold plates. In general, all analyses indicate that increasing the amount of steel between the load and support points provides for lower levels of stress and deflection. As such, the mold designer must perform analysis to develop robust designs that are not un­ economical. The uses of support pillars, interlocks, and other designs were demonstrated to reduce deflection. Common fastening means were also analyzed including interference fits, socket head cap screws, and dowels with clearance and interference fits. The mold designer must remember to provide means for fastening the cavity and core inserts to the rest of the mold while providing tight control of location relative to other mold components. In practice, the provision of fasteners may interfere with other subsystems of the mold including part ejection and mold cooling. In such cases, iterative redesign of the mold may be required to efficiently locate all the mold’s subsystems without increasing the size and cost of the mold. After reading this chapter, you should be able to: ƒƒDescribe the flow of forces from the mold cavity to the machine tie bars; ƒƒState the relationship between modulus, stress, and strain; ƒƒState the relationship between ultimate stress, yield stress, and endurance stress; ƒƒSpecify the limit stress and maximum deflection based an application’s requirements; ƒƒEstimate the compressive, shear, and hoop stresses in various mold components; ƒƒEstimate the deflection of a plate, core, or support pillar due to compression; ƒƒEstimate the deflection of a plate, core, or side wall due to bending; ƒƒSpecify the plate thickness and use of support pillars to avoid excessive mold deflection; ƒƒSpecify the mold cheek and use of mold interlocks to avoid excessive stress or mold deflection; ƒƒSpecify the design of mold cores to avoid excessive hoop stress and core bending; ƒƒSpecify the distance between the mold cavity and stress concentrations (such as ejector holes and cooling lines) as a function of the material properties and application requirements; ƒƒSpecify the dimensional limits on male and female components to achieve clearance, transition, interference, and drive fits;

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ƒƒEstimate the expected clearance or insertion force for a specified fit; and ƒƒSpecify the use of socket head cap screws to securely fasten mold components. The analysis and design of each of the mold’s subsystems has been completed. The next chapter is intended to increase the mold designer’s awareness by providing a critical examination of available mold technologies.

„„12.6 References [1] Gąska, D. and C. Pypno, Strength and elastic stability of cranes in aspect of new and old design standards, Mechanics (2011) 17(3): pp. 226–231 [2] Forrest, P. G., Fatigue of metals, Elsevier (2013) [3] Yahr, G., Fatigue design curves for 6061-T6 aluminum, ASME-PUBLICATIONS-PVP (1993) 259: p. 43 [4] Bisplinghoff, R. L., J. W. Mar, and T. H. Pian, Statics of deformable solids, Courier Corporation (2014) [5] Guo, Z. Z., Y. Li, and X. J. Zhao, The analysis of core shift in injection mold and its optimization based on AMI, in Advanced Materials Research, Trans Tech Publ. (2011) [6] Choi, J., et al., Structural Analysis Examining the Mold Deformation Behavior for the Detection of the Flash in the Injection Mold, Int. Polym. Process. (2014) 29(4): pp. 489–494 [7] Peterson, R. E. and R. Plunkett, Stress concentration factors, J. Appl. Mech. (1975) 42: p. 248 [8] Timoshenko, S. and J. Goodier, Theory of elasticity, Engineering societies monographs, International Student edition, New York (1970) [9] Timoshenko, S. and S. Woinowsky-Krieger, Theory of plates and shells, Vol. 2, McGraw-Hill, New York (1959) [10] Timoshenko, S. P. and S. Woinowsky-Krieger, Symmetrical Bending of Circular Plates in Theory of Plates and Shells, McGraw-Hill, New York (1959) pp. 61–63 [11] Rybicki, E. F. and M. Kanninen, A finite element calculation of stress intensity factors by a modified crack closure integral, Eng. Fract. Mech. (1977) 9(4): pp. 931–938 [12] Shigley, J. E. and C. Mischke, Section 19, Limits and fits, Standard Handbook of Machine Design (1986) pp. 19.1–19.19

13

Mold Technologies

„„13.1 Introduction This book has sought to provide an engineering approach to mold design; the emphasis has been on the examination and modeling of fundamental mechanisms that govern the use and failure of injection molds. The examples have purposefully been made as simple and clear as possible, so that the practitioner can apply the design and analysis methods to more specific and advanced molding applications. There are many advanced molding process technologies that provide strategic advantages on an application-specific basis. Most advanced molding processes require specific mold designs. A flow chart has been provided in Fig. 13.1 to guide the selection of some of available mold technologies. Such mold technologies can be used to compete effectively by providing molded parts with higher quality in less time and at lower costs. Most of these technologies have been developed for specific purposes, such as to produce a molded part with unique properties, or to more economically produce large quantities of molded parts. Many molding technologies are interwoven. For instance, multishot molding (in which a molded part is made of two or more materials) has characteristics that are related to coinjection molding, insert molding, stack molding, and even injection blow molding. Regardless of the level of technology, the underlying physics and mold design fundamentals that have been previously provided still apply. As such, this chapter provides an overview of some available molding technologies, and discusses associated mold design issues. Examples of illustrative mold designs have been sourced from the U.S. patent literature. The objective here is not to provide an exhaustive survey of mold related technologies, or even to recommend specific mold designs. Rather, the intent is to show some interesting examples that will imbue the practitioner with specific insights into a range of mold technologies so that they may become better mold designers. In applications, mold and product designers should discuss the incorporation of mold technologies as part of a manufacturing strategy to either increase cus­tomer’s

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willingness to pay [1] or reduce costs. Breakeven analysis, such as presented in Section 3.4, can be used to assist the appropriateness of different mold technologies.

Figure 13.1 Mold technology selection flow chart

13.2 Coinjection Molds

„„13.2 Coinjection Molds Coinjection molding is a process in which two materials are sequentially injected into a mold cavity, typically through the same gate. Since the first material forms a skin and the second material forms the core of the molded part, it is possible to use coinjection molding to produce plastic parts with unique aesthetic or structural properties with potentially lower costs than injection molding. Some typical co­ injection molding applications include: ƒƒThe use of a first virgin material having preferred cosmetic properties followed by a second material having different structural properties and/or recycled content, as in the fascia of a car bumper; ƒƒThe use of a first material followed by a second foaming material to produce a cosmetic part with lower density, as in structural foam applications; ƒƒThe use of a first material followed by a second fluid, such as air or water, to produce a hollow part like a door handle. While this last example (commonly known as gas assist or water assist or fluid ­assist molding) may not seem a coinjection process, the molding process and mold designs are sufficiently similar to warrant a joint discussion.

13.2.1 Coinjection Process In coinjection molding, two materials are sequentially injected, often similar to the sequence provided in Fig. 13.2 [2]. As shown, a first melt is partially injected into the mold through a sprue 6 or some other feed system. After a desired volume of the first material 7 has been injected, a second material 8 is injected at the same location. If the volume of the first material is too small, then the second material may “blow through” the first material. Conversely, too large an initial charge of the first material may leave too small a volume for the injection of the second material. Since the first material is adjacent to the mold wall, and may have partially solidified, the second material will tend to flow through the core of the first material. After the second material has been injected, it is fairly common to then inject a small amount of the first material 9. This latter injection of the first material serves to purge the feed system of any undesired amount of the second material, which might otherwise contaminate the subsequent molding cycle.

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Figure 13.2 Coinjection molding process

It is observed in Fig. 13.2 that the mold core 2 is moving in and out of the mold cavity 1 during the injection of the materials into the mold to thereby adjust the wall thickness of the cavity 3. This injection compression serves at least two purposes. First, in foam molding, the compression and subsequent expansion of the mold cavity can be used to delay and subsequently encourage the nucleation of gas cells, thereby controlling the distribution and density of the injected foam. Second, in nonfoam molding, the compression of the cavity can be used to control the pack pressure throughout the mold and thereby control the shrinkage characteristics of part features molded of the first material while injecting the second material. The control of the cavity wall thickness can be accomplished by profiling the displacement of the molding machine’s platen during the filling stage, or alternatively ­profiling the clamp tonnage profile. The mold uses a sliding fit (refer to Section 12.4.1) along the vertical sides where the mold core mates with the mold cavity. While not discussed in this reference [2], the sliding fit can be assisted through the use guide pins, interlocks, or k ­ eyways to mate the cavity and core inserts to avoid accelerated wear on the sliding ­surfaces.

13.2 Coinjection Molds

13.2.2 Coinjection Mold Design A schematic for a coinjection mold and feed system is shown in Fig. 13.3 [3]. As shown, material is delivered to the mold from the barrels of two injection units 8 and 9 to the mold 10 via corresponding flow channels 15 and 16. These two channels converge at a control valve 17 prior to the sprue 11. The control valve uses a valve pin 18 with two skewed flow channels. By rotating the pin, one of the two flow channels in the pin will register with the channels 15 or 16 to allow material to flow from the corresponding barrels 8 or 9 into the mold while also preventing the materials from flowing between the barrels. A control system is required to coordinate the actuation of the valve pin 18 with the injection of material from the barrels 8 and 9. Given this feed system design, a first and last injection of the first material is warranted to avoid contamination of any material residing between the valve pin 18 and the sprue 11 as previously discussed with respect to Fig. 13.2.

Figure 13.3 Coinjection mold and process

For the most part, design of coinjection molds is very similar to that of conventional molds; many conventional molds can be successfully used in a coinjection process since the mechanisms for coinjection are mostly integrated with the molding machine and not the mold itself. However, the mold designer should modify the analyses for coinjection. With regard to mold filling, the mold designer should ­ensure that the mold cavity is designed to achieve the desired filling patterns at

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reasonable pressures. Analytical solutions and simulations have been developed for the coinjection of two materials with dissimilar viscosities into a mold [4, 5]; however, in many coinjection applications, the mold will operate successfully if the mold is designed to fill completely with only the more viscous material. Analysis of cooling, shrinkage, and ejection should also be modified to consider the melt ­temperatures and thermal properties of the two materials. Given the multilayered structure of the coinjected molding, a reasonable approach is to derive a “meta-­ material” that has material properties in proportion to the layer thickness of the two constitutive materials.

„„13.3 Gas Assist/Water Assist Molding Gas and water are both fluids, so both gas assist molding and water assist molding can be considered as types of fluid assisted molding processes. Since these assisting fluids are injected inside of a first material, all fluid assist molding processes can be considered a type of coinjection molding process. Compared to traditional coinjection with polymer melts, fluid assisted molding has two distinct differences. First, the second injected fluid (such as nitrogen or water) has a very low viscosity compared to the previously injected polymer melt. This low viscosity provides for a very low pressure drop along the flow path, and thus gives excellent pressure transmission for packing out the previously injected polymer melt. Second, the assisting fluid is later removed from the molded part so as to hollow out the inside of the molded part. With careful mold design, fluid assisted moldings can have ­increased strength, lighter weight, and reduced cycle times compared to conventional or coinjected molds. Fluid assisted molding is a fairly old process having served as an alternative to blow molding [6]. Two variations of a more modern gas assist process with injection decompression are shown in Fig. 13.4 [7]. In the first method, two mold halves 10 and 11 form a mold cavity 12 into which the plastic melt will flow. The molding machine’s nozzle 13 has an internal core 16 with a sliding valve 19 that is actuated by compressed gas alternately introduced through gas lines 20 and 21 through three-way valves 22 and 23. At the beginning of the process, the plastic melt 24 is introduced into the mold cavity through the machine nozzle 13; at the same time, the sliding valve 19 is in a position which blocks the gas inlet tube 18 compressing the gas through line 21 decompressing the gas through line 20. After the mold cavity is partially filled, gas line 20 is pressurized while gas line 21 is depressurized. This causes the sliding valve to assume the position as shown so that gas inlet tube 18 is opens and delivers compressed gas to the mold cavity. Once the gas has been injected, the sliding valve is then actuated to prevent the undesired flow of

13.3 Gas Assist/Water Assist Molding

the plastic melt. After the molded part cools, the opening of the mold causes the sprue to break and the release of any compressed gas to the ambient atmosphere.

Figure 13.4 Gas assist with injection decompression

A second method is also shown in Fig. 13.4 in which the reverse of injection ­compression, injection decompression, is used to form a hollow part with a very large cavity. In this design, the plastic melt flows into a cavity formed between two mold halves 30 and 32. While not shown, these mold halves can have fine details like bosses that are fully formed by the initial filling of the cavity with the polymer melt. The compressed gas is then injected into a thicker portion of the mold cavity. At the same time, the mold core 32 is retracted from its opposing mold half 30 to enlarge the cavity. In this manner, moldings with very large internal gaps (for ­example, 50 mm) can be formed while preserving fine features on the exterior surfaces of the molded part. The mold design of fluid assisted molds needs to vary considerably from that of injection and coinjection molds. In particular, the mold designer needs to consider the location for the injection of the gas or water. As demonstrated in Fig. 13.4, both the nozzle and cavity are common locations. As importantly, the mold designer must carefully design the mold to have appropriate flow channels to strategically direct the gas or water through the mold cavity. In most mold designs, the cavity wall thickness is made as uniform as possible to avoid nonuniform cooling and shrinkage. However, such a mold design will not lead to an effective mold for fluid assist. The reason is that the gas or water will permeate or “finger” in random directions through a uniformly thick mold cavity, thereby weakening the molding without significantly reducing the part weight. As such, thick flow channels as

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shown in Fig. 13.5 are commonly added to the mold cavity to direct the gas or water through the mold cavity [8]. All the gas channels will exhibit some ­ ­irregularity regardless of the magnitude of penetration. In general, it is desirable to develop a gas channel to provide as uniform a molded wall as possible while providing the necessary fluid flow and part stiffness. For this reason, the top right gas channel in Fig. 13.5 is least preferred. Since the other flow channels are cored out by the fluid, the cooling and shrinkage is made relatively uniform without ­extended cycle times.

Figure 13.5 Flow channel sections for fluid assisted molding

Water assist molding seems to have received increased interest lately [9–12]. Compared to gas assist, water assist provides at least three key benefits. First, water has a very high specific heat and so can be injected to reduce the cycle time compared with gas assist molding application. In fact, in some water assist molding applications, the flow channels are designed with inlets and outlets, such that the water can be circulated within the molded part and thereby greatly reduce heat transfer via heat convection. Second, water is incompressible compared to gas, and so can be used to provide higher melt pressures in the cavity with less energy and risk than gas. Third, it has been shown that water assist provides more uniform and smooth surface in the inside of the molded parts. With these advantages, however, water assist does bring two significant disadvantages. First, the water must be removed from the interior of the molded parts; ­various schemes have been developed to remove the water internal to the molding prior to the mold opening [13]. Second, the use of water in the molding environment tends to increase humidity and corrosion, so a corrosion resistant mold ­material such as SS420 is recommended.

13.4 Insert Molds

„„13.4 Insert Molds Insert molding refers to a process in which a discrete component is placed within a mold, and then at least partially encapsulated by the subsequently injected ­plastic melt. Some commonly inserted components include electrical devices, nuts or other fasteners, stiffening members, and other plastic components. After the insert molding process, the inserted component is usually permanently joined with the molded plastic.

13.4.1 Low Pressure Compression Molding One common method for encapsulation of delicate components is compression molding as shown in Fig. 13.6 for the production of a tantalum capacitor [14]. In this process, the capacitor 19 is placed as an insert between two layers of plastic 17 and 18 prior to the mold closure. The mold design provides matching cavities 16 and 22 to receive the plastic as well as grooves 21 to receive the lead wires 20. In any molding process, the mold designer should explicitly consider the handling of the molded parts upon de-molding. In the design of Fig. 13.6, the lower cavity 22 is deeper than the upper cavity 16. In addition, the lower mold half 11 is provided with a flash well 23 for the collection of any plastic that flows out of the cavity ­during the compression molding process. As a result of these design elements, the molded part will remain on the lower mold half when the mold opens. In this compression molding process, the plastic layers 17 and 18 were cut from sheet stock in a form to fit into their corresponding cavities while also supporting the capacitor 19 and lead wire, 20. It is desirable that the plastic fully contacts the rear surface of the mold cavities to facilitate heat transfer and plastic forming. Prior to mold closure, cartridge heaters in the two mold halves 10 and 11 bring the temperature of the mold and plastic layers to above the glass transition temperature of the plastic. Once the plastic is softened, the mold is slowly closed with low force. As the mold slowly closes, the plastic slowly flows around the capacitor until it is fully encapsulated – any excess plastic in the cavity will flow out the flash surface, 24, and into the flash well, 23. Once the inserted component is fully encapsulated, full clamping force may be applied to the two mold halves to compensate for shrinkage and achieve the desired dimensions while the mold is cooled.

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Figure 13.6 Compression molding with inserted component

The mold design of Fig. 13.6 has some unique features. First, the grooves secure the insert component in the mold to avoid undesired movement caused by the movement of the mold or the flow of the plastic melt during the molding process. Second, this process was specifically designed to impart low stress on the inserted component by the controlled heating of the mold and softening of the plastic ­followed by the clamping and cooling of the mold. Given this heating and cooling cycle, the mold should be carefully designed to minimize the size and thickness of the plates so that energy consumption and cycle time are minimized. Third, this process used a flash surface and reservoir to control flashing of excess plastic; it is clearly desirable to select an amount of plastic stock that minimizes the amount of flashing while ensuring a fully filled cavity. In this design, the mating flash s­ urface, 24, on the lower mold requires a relatively large clearance with mating surface on

13.4 Insert Molds

the upper mold half. If this clearance is too small, then the rate of the compression molding process can be limited by the flow of the plastic melt out of the cavity. The filling and cooling analyses of Chapters 5 and 9 can provide useful design support.

13.4.2 Insert Mold with Wall Temperature Control Another example of insert molding is provided in Fig. 13.7 [15], which is particularly directed to the control and improvement of weld lines around an inserted component for the production of a water faucet handle. The mold design consist of two separable mold halves 30 and 31 having recesses 32 and 33 that together form a mold cavity. The inserted component 35 is held in position by two opposing pins 36. After mold closure and prior to mold filling, a substantially uniform cavity thickness exists between the inserted component 35 and the mold halves 30 and 31. In this mold design, the mold wall temperature of the mold is locally controlled by the flow of a controlled fluid through channels 40 and 42. Different fluids such as water, oil, or steam can be provided to different portions of the mold at different temperatures. In the molding process, plastic melt is fed through the gate 38 and will follow the path of least resistance through the mold cavity. Upon entering the cavity, the plastic melt will divide into two streams 44 and 45 flowing around the insert. In Fig. 13.7, the recess 33 in the lower mold half is controlled at lower temperature compared to the recess 32 in the upper mold half. As a result, the upper melt stream 44 will advance more rapidly than the lower melt stream 45. Given the importance of aesthetics in this molding application, the melt front advancement and knit-line location 50 can then be adjusted by specifying the difference between the mold wall temperatures in each zone. Furthermore, a heating element 46 is used to ­locally heat the mold wall to a temperature above the plastic’s glass transition ­temperature to melt and fuse the area around the knit-line ensuring desirable ­aesthetic and structural properties.

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Figure 13.7 Insert molding with mold temperature control

The design of the multiple temperature control channels seems quite advanced, especially for 1937 when this patent application was filed. To facilitate the implementation of the cooling channels, the recesses 32 and 33 are themselves provided as mold plates 48 and 49 that are placed into the cavities in the two mold halves 30 and 31. This design is quite similar to the bezel example of Fig. 9.16 in which the cooling lines have been milled into the rear surface of the core insert. The cooling and structural analysis of Chapters 9 and 12 should be applied to determine the

13.4 Insert Molds

cooling channel’s hydraulic diameter and layout, as well as the required amount of plate stock require to avoid excessive stresses. The resistive means of heating, as here for element 46, is discussed in more detail in Section 9.4.2.

13.4.3 Lost Core Molding Lost core molding refers to a process in which a mold core is inserted into a mold cavity to form the interior of a molded plastic part. After the molding with the core insert is ejected, the core is melted out of the molded part to leave a complex ­interior cavity. One lost core molding application is shown in Fig. 13.8 [16]. This particular application molds a valve housing with internal threads and an internal cavity containing a spring and ball check.

Figure 13.8 Lost core molding with internal components

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The lost core molding process requires two sets of molds. The first mold design consists of two mold halves 4 and 6 which meet at a parting line 8. The mold cavity 10 includes threaded ends 12, a central bore 14, and a valve seat 16. Prior to mold closure, a ball check 26 and a compression spring 28 are inserted into the cavity 10 from the parting plane. After these components are inserted into the cavity and the mold is closed, a pin 22 is lowered into the mold cavity in opposition to the spring to lift the ball check from its seat. A low melting point material 30 is then injected through an opening 20 to completely fill the mold cavity. Upon solidification, the low melting point material has locked the ball check and spring in position. The mold is then opened, and the molded part 32, is removed with the pin 22. The molded part 32 is then inserted into the second mold’s cavity 46 formed by mold halves 40 and 42 for use as a core piece to form the inside surfaces of the valve housing. From the previous molding process, the core includes externally threaded ends 34 and a central section 36 that encloses the ball check 26, spring 28, and a conical surface 38, which is used to form the contour of the valve seat of the valve housing. The sliding fit of the pin 22 with the hole 52 serves to center the core 32 properly within the mold cavity 46. Plastic 56 is then injected through the sprue 50 to fill the mold cavity and surround the core 32. Upon removal of the ­solidified molding, the core 32 is melted away to leave the final structure 60. This housing 60 has internally threaded ends 62 and a central chamber 64, which contains the ball check 26 and the spring 28. It may seem unlikely that such a lost core molding process is feasible. To melt the core from the housing, after all, the first material 30 that makes up the core must have a lower melting point than the second material 56, which makes up the housing. Then, why does the core not melt during the injection of the second material? The reason is that the first material 30 has sufficient mass to act as a heat sink and absorb heat from the second material 56 without melting. For example, the first material is suggested to be a metallic alloy of 58 % bismuth and 42 % tin. Such an alloy melts at about 138°C. For the second material, various plastics with a wide range of melting points have been used, including acetal and polycarbonate. For molding with higher temperature plastics, both the mold and the lost core can be cooled to minimize melting of the core. Another more recent example [17] describes the use of a molded POM core with overmolding of a ceramic-polymer powder that is subsequently debinded and sintered to produce a ceramic part. The author expects the use of lost core processes to increase with the availability of 3D printing processes by which such cores can be made.

13.5 Injection Blow Molds

„„13.5 Injection Blow Molds Injection blow molding is a process by which complex, thin-walled parts can be made including large internal cavities. In this section, two different blow molding processes are presented. The first design utilizes a conventional injection molding machine with a rotary mold while the second design uses a four position indexing system with multiple molding stations.

13.5.1 Injection Blow Molding An injection blow mold design is shown in Fig. 13.9 [18]. The design includes two injection molds 48 and 50 which are positioned at equal radial distances from the axially located main sprue. The injection molds each include gating 54 and 56, an injection molding cavity 58 and 60, and other common mold components. The ­design also includes two sets of split cavity blow molds 140 a/b and 142 a/b so that undercut parts may be readily ejected after inflation. The injection blow molds are located diametrically opposite each other on the same radius as the injection molds 48 and 50. All four molds in this design spaced at 90 degree angles. The mold cores are spaced on the same radius as the mold cavities so that the four cores can engage the four mold cavities simultaneously upon mold closure. A manifold delivers melt from the nozzle of the molding machine to the mold ­cavities. The manifold rotator 18 is designed to oscillate between two orthogonal positions through actuation of a hydraulic drive cylinder 124 pivoted at one end by a pin 126 to the bearing block 78. A piston 128 having a clevis end 130 pivotally engages a crank arm 132, which is secured by screws or other suitable means to the circumference of the manifold axle 15. Stops 134 and 136 are fastened to the bearing block to limit the travel of bellcrank arm to a 90 degree sweep. This design ensures that the mold cores 114 and 116 engage the injection mold cavities 58 and 60 at one position of the manifold unit, while the other two mold cores engage the injection blow molding cavities 65 and 67.

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Figure 13.9 Injection blow mold with rotating cavities and reciprocating core

During the molding process, the machine clamps the mold cores against the mold cavities. The injection unit of the molding machine provides plastic through the main sprue 42 and runners 44 and 46 of the manifold 14 to the injection mold cavities 58 and 60, where the blow molding parisons or preforms are molded. ­Afterwards, the cores carrying the hot parisons are reciprocated out of the cavities by action of the clamp. The manifold assembly then rotates 90 degrees to align the injection blow mold cavities with the molded parisons. The machine then clamps the mold and injects compressed air through the bores of the parisons to inflate the parisons and form the blow molded products while the injection unit fills the two injection mold cavities to form the next set of parisons. Upon mold opening, the split blow molds open and the finished parts are ejected.

13.5 Injection Blow Molds

This design utilized a rotating set of mold cavities with a reciprocating set of mold cores. A clear alternative would be to utilize a stationary set of cavities with a rotating and reciprocating set of mold cores. Either design strategy provides a method for compactly and economically performing injection blow molding through the modification of a conventional molding machine. The design may be guided by filling analysis to ensure appropriate runner system and cavity design, cooling analysis to control the temperature of the hot parisons, and structural analysis to minimize the size and stress of the injection and blow molds.

13.5.2 Multilayer Injection Blow Molding A different machine configuration for injection blow molding is shown in Fig. 13.10 for the molding of a two layered product [19]. The inner and outer layers are c­ hosen for particular reasons related to the use of the product. For example, the inner layer 44 may be made of material which resists reaction with the contents of the container while the outer layer 44’ may be made of a material of substantially greater strength than the inner layer. In this design, the injection molding system includes a first injection station 10, a second injection station 12, a blowing station 14, and a stripper station 16. The system has an indexing head 18 with four faces. A set of core rods 22 extends from each of the four faces. The indexing head rotates intermittently about a center shaft 20 to sequentially index each set of cores with the different stations. Due to its configuration, this design is known as a “four ­position machine.”

Figure 13.10 Two layer injection blow molding

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At the first injection station 10, the injection mold 26 is supplied with plastic melt from an injection unit 28 via a runner system 30. The injection mold is a split ­cavity design with a stationary lower section 32 and a movable upper section 34. When these mold sections are closed together, each of the core rods 22 extends into a cavity. The neck portion 38 of each core rod is firmly gripped by the wall of the opening 36. The plastic is injected into the cavity at an opening opposite the top of the core rod, and flows around and down the length of the core rod to form a parison 44. As the plastic flows down the length of the cavity, the melt loses some of its heat to the mold 26 and the core rod 22. Because the core is slender and has a large thermal resistance as analyzed in Section 9.3.5.5, the highest temperature of the parison will occur at the end of the core rod 40 near the gate and runner 30. After the molding of the parison in the first station 10, the mold 26 opens and the core rods 22 are lifted clear of the mold cavity in the lower section 32 by rising movement of the indexing head. The indexing head turns 90 degrees in the ­counterclockwise direction and brings the core rods with the parisons 44 on them over the lower section of a second injection mold 48 at the second injection station 12. The second injection mold closes on the neck portion 38 to form mold cavities 42a, which are larger in cross-section than the corresponding cavity 42 if the first molding station. As previously noted, the tip of the first parison 40 is at the highest temperature and so most easily deformed. If the second layer of the parison were gated from the same location, then the direct impingement and high flow rates of the second layer could wash some or all of the first parison off the core rod and thereby lessen the value of the molded product. For this reason, the plastic for the second layer 44’ is introduced into the cavity 42 from a manifold 50 and individual runners to the neck end of the cavity. The plastic flows around the first layer to form a parison with two layers. The second mold is then opened, and the indexing head moves the core rods to a blow mold, 58, at the blowing station 14. The blow mold holds the neck portion of the laminated parison. The mold cavity 42b is in the form of the desired article, which is to be blown from the laminated parison. The blowing operation then fully inflates the two layer parison 60. The blown products 60 next advance to the ­stripper station 16 where a stripper plate 64 pushes the molded products 60 off the core rods 22. With the next rotation of the indexing head, the bare core rods are presented to the first forming station to begin the next round of moldings. There are three suggested benefits of this mold design. First, the witness mark formed by the gating at the tip of the core is wiped away by the flow of the second layer. Second, the design reduces the cycle time associated with the molding of the first layer by gating the second layer into a location away from the hottest portion of the first layer. Third, the improved consistency of the first layer facilitates the molding of a thinner first layer and with it a lower cost product.

13.6 Multishot Molds

„„13.6 Multishot Molds Multishot molding sequentially injects different types of plastic, to mold a part with distinct regions. There are several potential advantages for the use of a multishot molding process. These include the use of multiple shots with: ƒƒdifferent shot capacities to sequentially mold very large parts; ƒƒdifferent colors to mold multicolor parts, such as automotive taillights; ƒƒdifferent structural properties to mold parts with improved living hinges, tactile feel, etc.; ƒƒand others. There are many different methods to accomplish multishot molding. Perhaps the oldest and simplest is overmolding, which can be considered a variant of insert molding. Another approach is the core-back method, which lends itself to a relatively simple mold design. However, multi-station mold designs are the most ­common method in use today due to the capability of this process to economically produce more complex part geometries. Each of these mold design strategies will be discussed. Regardless of the type of multishot mold design, the provided mold design and analysis methods generally apply with a few special considerations. First, multishot molds may require extended cooling times. The reason is that the first layer will not be at the mold coolant temperature when the second layer is injected. Furthermore, the first layer will largely prohibit the transfer of heat from the second layer to the mold. For these reasons, the mold designer should consider multishot molding using the analysis for one-sided heat flow as discussed in Section 9.3.5.6; the second layer should be 40% thinner than the first layer to avoid extending the cycle time. Second, multishot molds provide the mold designer the opportunity to utilize the second injection of plastic to melt and wipe out small imperfections or witness lines on the first layer of plastic. Because of this effect, however, the mold designer should avoid the placement of fine details on the some surfaces of the first molding that may be degraded by the second injection of the plastic melt.

13.6.1 Overmolding A multishot mold design using an overmolding approach is shown in Fig. 13.11 [20]. In this design, a separate mold has used a branching runner 4 to fill two ­lateral runners 2 and a series of mold cavities for the production of key caps, B. Each key cap has a window 7 molded into its top surface 6 in the form of that key’s desired character, such as the number “5.” This particular design is directed to the bonding of incompatible plastic materials through the use of a solvent such as

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a­ cetone applied to the key caps’ rear faces 5. Accordingly, the branching runner 4 is intended to be used as a handle by the operator during the application of the solvent; more modern designs may use bosses or shoulders to assist automated part handling systems.

Figure 13.11 Two layer injection blow molding

Once the solvent has been applied to the key caps’ rear faces 5, the key caps are placed in the cavities 1 of the lower half of a mold, A. The upper half of the mold, C, provides cores 12 that mate with the cavities in the upper half of the mold for the molding of the keys. After the mold is closed, the second plastic material is ­injected through the primary runner 13 and secondary runner 14 into the mold cavities. A portion of the material will fill the back 9 and window 7 of the key cap as well as the key’s boss 17 for later assembly with other components.

13.6 Multishot Molds

This type of mold design is quite common for production of keys and signs to avoid noticeable wear. Specifically, the number “5” is formed by two materials, each with the same thickness as the window 7. The key cap’s entire top face 6 will have to be worn away before the character disappears. On a side note, this mold design has two features that may be useful in other multishot molding applications. First, the projections 11 increase the surface area and therefore the bond strength ­between the two materials; these same projections will also tend to increase the lateral strength of the molded parts. Second, a rib 10 is placed below the window 7 to undercut the second molded material and ensure that the two pieces are not ­separable.

13.6.2 Core-Back Molding Core-back molding refers to a multishot molding process in which a portion of the mold, typically a core, is moved to create or reveal a mold cavity into which a second plastic melt can be injected adjacent to a previously molded first plastic melt. A design for a mold with core-back capabilities used to make front or tail ­indicator lights for vehicles is shown in Fig. 13.12 and Fig. 13.13 [21]; the molded piece in this application may consist of three different colors such as red, clear, and amber. The mold consists of an upper mold half 4 and a lower mold half 5 that together form a cavity. The cavity is split into three different portions 11, 12, and 13 through the use of four blades 14, 15, 16, and 17 that are independently actuated by pistons 18, 19, 20, and 21 integral with cylinders 22, 23, 24, and 25. Each cavity is fed polymer melt through runners 7, 8, and 9.

Figure 13.12 Plane view of mold with core-back

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Figure 13.13 Section view of mold with core-back

In the molding operation, the mold and the blades are closed to isolate the different portions of the mold cavity so that the different plastic materials can be sequentially injected. To reduce cycle time, it is preferable to concurrently inject polymer melts from nozzles 1 and 2 through the runners 7 and 8 into the areas 11 and 12. Once these materials are sufficiently solidified, the cylinders are actuated to retract the blades 14, 15, 16, and 17. The third plastic may then be ejected through runner 9 into the third area 13 of the mold cavity. In this manner, a molded part can be produced consisting of multiple materials without ever opening the mold. There are two items of note regarding this core-back design. First, it is possible to have designed a mold utilizing a single set of blades 16 and 17 to reduce the size and complexity of the mold. Indeed, for strength reasons a preferred design would use a single set of blades that interlock with a slot on the opposing face of the mold cavity. However, one possible reason for the design of Fig. 13.12 and Fig. 13.13 is that the retraction of the blades into both sides of the mold provides a means for the molding of protruding ribs into the upper and lower sides of the mold 4 and 5. In any case, structural analysis should be used to ensure that the blades are of sufficient thickness to avoid excessive shear stress and bending in the blades given the thickness of the cavity and the operating melt pressure. The second item of note concerns the use of two sets of blades as opposed to a ­single central core that could be withdrawn. An alternative mold design could avoid the use of blades altogether by making the entire center section 13b of Fig. 13.13 a single actuated member. In this alternative design, the center section 13b would be in a forward position upon mold closure, closing the cavity area 13 of

13.6 Multishot Molds

Fig. 13.12 from the melt and providing the same cavity side walls effectively provided by the blades 14, 15, 16, and 17 in the previous design. After the left and right areas 11 and 12 were filled with plastic, the center section 13b could be retracted and the cavity area 13 filled with a third plastic via runner 9. It may seem that this alternative design would require extremely high actuation forces for the center section 13b given its large projected area subjected to melt pressure. However, this is not the case since the center section will not see significant pressure when molding areas 11 and 12, and can be supported by a shoulder or other mold components when retracted and exposed to high melt pressures.

13.6.3 Multi-station Mold Parts consisting of multiple materials are also often molded in multi-station molds. One such design is shown in Fig. 13.14 to produce a replica of the Canadian ­national flag [22]. In this design, a first mold is composed of a transfer plate 47, a cavity plate 48, and a runner plate 49. The cavity plate 48 defines a cavity 51 ­including lateral panel segments 52, a central maple leaf image segment 53, and bridges 54 that connects all of the segments together. In operation, the first plastic melt is injected from a runner 58, into a sprue 56, through a gate 57, and into the cavity 51. Once this plastic solidifies, the mold is opened and the transfer plate is removed from the first mold. Because of the mold design, the solidified runner 68 and molding 61 remain with the transfer plate. The transfer plate 47 with the solidified runner 68 and molding 61 is then transferred to a second mold. In this design, the transfer plate inserts the solidified molding 61 into another cavity 70 in a second mold cavity plate 69. The transfer plate is then moved laterally to strip the feed system from the molding 68. Additional plates 71 and 75 are then positioned with the mold cavity plate 69 to completely form a second mold. The second material is then injected adjacent and over the first material to form a single part integrating the two plastic materials.

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Figure 13.14 Multi-station mold design

Compared to a core-back mold design, the multi-station mold provides for greater flexibility in the molded part design. Specifically, multi-station molds allow for complex parts to be molded, and then inserted into other arbitrarily complex cavities for injection of additional plastic materials adjacent to, above, and around the prior molding(s). Accordingly, several different mold and machine designs have been developed to support multi-station molds. These include the transfer of parts via rotating mold sections similar to the designs shown in Figs. 13.9 and 13.10. More recently, dedicated two-shot molding machines have been developed as shown in Fig. 13.15 [23]. In this type of design, the injection units provide material to two sets of mold cavities mounted on two opposing platens. Because the platens oppose each other, a single clamping mechanism can be used to provide the mold closure force for both sets of cavities, very similar to stack molds as discussed in Section 13.6.2. Different drive mechanisms have been developed to index the cores including turret drives as shown in Fig. 13.15, rack and pinion arrangements [24], and others.

13.7 In-Mold Labeling

Figure 13.15 Turret style molding machine and mold design

„„13.7 In-Mold Labeling Injection molding generally provides highly functional and economical parts. In many applications, however, there is a requirement for decorating, detailing, or otherwise labeling the molded parts according to the consumer’s needs. In some cases, the decorating may be provided by post-molding processes such as hot stamping, pad printing, painting, screen printing, and others. In other cases, these techniques are not feasible since some plastic resins such as polypropylene, polystyrene, polyethylene, and others are inherently resistant to dyes, inks, and paints. One approach might be to utilize two-shot molding with different types of plastics to achieve the necessary detail. However, this approach is relatively expensive to implement. In addition, two-shot molding does not provide for the level of drawing detail or the number of colors that may be desired. As such, molds can be designed to incorporate printed labels that are secured to the molded part during the molding process. There are several advantages related to such in-mold labeling. First, the labels themselves can have very fine graphic details with multiple colors ­produced by screen printing or offset lithography. Second, the labels can provide for wear resistance through the use of an acrylic or polycarbonate layer laminated

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over the printed surface. Third, the cost of in-mold labeling is relatively small with little additional tooling cost and only slight extensions of the molding process, and requires no post-molding processes.

13.7.1 Statically Charged Film In this design, labels are typically printed on a film with a thickness on the order of 0.15 mm (5 mils), and of a polymer (such as polypropylene or polystyrene) that is compatible with the plastic being molded. Since a thin film is made of flexible plastic, the thin film can be placed onto curved surfaces. Different approaches may  be used to secure the film in place during the molding process, including adhesives, compressed air, vacuum, and static charge. Figure 13.16 shows a ­ method for in-mold labeling using a film that is statically charged prior to its placement in the mold [25]. In this design, two mold halves 17 and 19 form a cavity 23. Prior to molding, a film 11 is placed in the mold cavity and secured by static charge applied to either the film or the mold block; most often, the film is charged by ­ionizing the air around the film with a high voltage from a nearby electrode. The film 11 is made of the same material as the molded part 25, and has a printed ­design 13 facing the mold cavity.

Figure 13.16 In-mold labeling with static charge

Once the film is placed in the mold, the molding process continues as normal. The heat of the polymer melt causes the film 11 to melt and fuse with the part 25, such that the printed design 13 appears without any evidence of the film 11. Although the printed areas 13 may not fuse into the plastic, these areas can be adequately bonded by the fusion of the surrounding nonprinted areas. If necessary, the printing may be imperceptibly dithered to facilitate fusion between the molded part and the film.

13.7 In-Mold Labeling

A few comments are warranted about the film thickness and the processing con­ ditions. The mold designer should recognize that the film must withstand both thermal and structural loadings. The structural loading is driven by the shear stress applied by rate of the polymer melt flowing across the film and not the magnitude of the melt pressure; the analysis of Section 5.3.1 can be used to estimate the shear stress as a function of the polymer properties, part geometry, and processing conditions. The thermal loading is related to the heating and melting of the film by the polymer melt. If the film is too thin, then the printed design may be destroyed by the complete melting and flow of the film during the molding process. Analysis and experimentation may be required to optimize the film and process.

13.7.2 Indexed Film Statically charged films are quite common with in-mold labeling. One potential ­issue, however, is the placement of the films into the mold by either human or ­robotic operation. If a human operator is used, then issues may arise pertaining to processing delays, safety, or repeatability. If a robot is used, then issues may arise related to processing delays and cost. For molding applications with higher production volumes, it may be better to resolve these issues through the use of a mold design that automatically indexes the printed film through the mold with each molding cycle. One such design is shown in Fig. 13.17 [26] for the production of bottle cap with a retaining ring. In this design, the mold 25 includes a top clamp plate 26, a movable cavity plate 28, and a conventional moving half of the mold 30. The cavity plate 28 is retained to the top clamp plate 16 via fasteners 37. However, a counterbore is provided in the top clamp plate to allow the fasteners to slide such that the cavity plate is moved away from the top clamp plate by compression springs 39 when the mold opens. The resulting gap is approximately 0.5 mm in height, and provides clearance for the carrier ribbon 55 or 90 to slide between the cavity plate and the top clamp plate. The carrier ribbon is supplied from roll 58, around guide rolls 59, through the gap 41, to the mold cavities 50. When the printed film 60 is properly positioned, the mold closes and the film is secured by the clamping of the carrier ribbon between the mold cavity plate and the top clamp  plate. With the mold filling, the printed design is transferred from the ­carrier ­ribbon to the molded part 72 which subsequently solidifies. After the opening of the mold and the removal of the molded part, the used carrier ribbon is ­indexed by the drive motor 64 and the feed rolls 62 and 66, then directed to a ­suction tube 68 where it is recycled or discarded. As such, this mold is designed to operate very rapidly.

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Figure 13.17 In-mold labeling with indexed film

More recently, advanced mold and process designs have been developed to allow for very complex in-mold labels that may be shaped and contain cut-outs. These designs may utilize an indexing mechanism to consecutively handle thicker printed sheets. In such an operation, each sheet is thermoformed to conform to the surface of a complex mold cavity. Then, one or more punches may provide holes or otherwise size the formed and printed sheet to the shape of the mold cavity. ­Finally, each completed label is placed into the mold where it is bonded to the molded part. As such, in-mold labeling can be used to provide nearly any appearance (such as marble, national flags, etc.) to nearly any part (such as cell phones, etc.)

„„13.8 Review There are many technologies that can be incorporated into a mold’s design to: ƒƒEnable extremely complex molded part geometries, ƒƒMake molded parts containing multiple materials,

13.9 References

ƒƒProduce a hollow part with simple or complex internal features, ƒƒProvide a molded part with aesthetic or decorative surfaces, ƒƒControl the flow of the melt in the molded part, ƒƒIncrease the consistency of the molded parts, ƒƒIncrease the molder’s productivity by increasing the number of cavities in a mold, ƒƒIncrease the molder’s productivity by decreasing the required clamp tonnage, ƒƒDecrease the cost of the plastic consumed in molding the product, ƒƒAnd other objectives. There seems to be almost no limit to what the injection molding process can ­accomplish with advanced mold designs. For many molding applications, however, the issue to be deliberated is not what can be done but rather what should be done for a specific application. The decision as to how to develop a mold design is for the mold designer, who must strive to serve the needs of the molder and end-user. For this reason, critical decisions about the mold design and related technologies should be approved and documented between all the involved parties with a ­common understanding of the costs, benefits, and risks.

„„13.9 References [1] Makadok, R. and D. G. Ross, Taking industry structuring seriously: A strategic perspective on product differentiation, Strategic Management Journal (2013) 34(5): pp. 509–532 [2] Barrie, I. T., Injection Moulding Apparatus, in U.S. Patent No. 3,909,169 (1975) [3] Garner, P. J., Injection Molding Machines, in U.S. Patent No. 3,733,156 (1973) [4] Himasekhar, K., et al., Current trends in cae: Simulations of latest innovations in injection molding, Adv. Polym. Technol. (1993) 12(3): pp. 233–241 [5] Li, C. T. and A. I. Isayev, Interface Evolution and Penetration Behavior During Two-Component Transfer Molding, Part I: Modeling and Formulation, Polym. Eng. Sci. (2004) 44(4): pp. 687–696 [6] Hobson, J. R., Method and Apparatus for Making Hollow Articles of Plastic Material, in U.S. Patent No. 2,331,688 (1943) [7] Friederich, E., Method for injection molding of hollow shaped bodies from thermoplastic resins, in U.S. Patent No. 4,101,617 (1978) [8] Chen, S. C., et al., Investigation of Gas-Assisted Injection Molding. Part III: Effect of Gas Channel ­Design on Part Bending Strength, Polym. Eng. Sci. (1998) 38(7): p. 1085 [9] Knights, M., Water injection molding makes hollow parts faster, lighter, Plast. Technol. (2002) 48(4): pp. 42–47 [10] Mapleston, P., Water-assist molding bears fruit, homes in on automotive applications, Mod. Plast. (2004) 81(3): pp. 38–39 [11] Kuang, T., et al., Analysis of Air Bubble and Secondary Penetration of Water in Water Assisted Injection Molding Process, Polym. Mater. Sci. Eng. (2014) 11: p. 024

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[12] Park, H., B.-S. Cha, and B. Rhee, Experimental and Numerical Investigation of the Effect of Process Conditions on Residual Wall Thickness and Cooling and Surface Characteristics of Water-Assisted ­Injection Molded Hollow Products, Adv. Mater. Sci. Eng. (2015) [13] Kirkland, C., Let’s just call it AIM–assisted injection molding, Injection Molding (2002) 10(5): pp. 63–64 [14] Williams, G. H., Molding of Thermoplastic Materials, in U.S. Patent No. 2,975,487 (1957) [15] Reid, J. S., Method and Means for Molding Thermoplastic Articles, in U.S. Patent No. 2,182,389 (1937) [16] Ryder, F. E., Method of molding utilizing meltable core, in U.S. Patent No. 3,882,220 (1975) [17] Attia, U. M. and J. R. Alcock, Fabrication of hollow, 3D, micro-scale metallic structures by micro-­ powder injection moulding, J. Mat. Process.Technol. (2012) 212(10): pp. 2148–2153 [18] Manen, D. T.V., Rotary Paraxial-Cavity Reciprocable-Core Injection Mold/Injection Blow Mold System, in U.S. Patent No. 3,752,625 (1973) [19] Rainville, D., Double injection mold with neck gating, in U.S. Patent No. 2,975,487 (1976) [20] Dofsen, F. J., Process for Bonding Thermoplastic Materials, in U.S. Patent No. 2,492,973 (1950) [21] Mares, P., Injection molding articles of more than one resin component, in U.S. Patent No. 4,275,030 (1981) [22] Burry, A., Method of multi-material injection moulding, in U.S. Patent No. 4,073,854 (1978) [23] Rees, H., P. Brown, and M. Grund, Turret-type injection-molding machine, in U.S. Patent No. 4,330,257 (1982) [24] Brown, P., Support for an intermediate platen of a stack mold, in U.S. Patent No. 4,408,981 (1983) [25] Jardine, P. R., et al., Method of using static charge to decorate molded thermoplastic srticles, in U.S. Patent No. 3,270,101 (1966) [26] Ruhl, G. F., Production of molded plastic articles with artwork thereon, in U.S. Patent No. 4,537,739 (1985)

14

Mold Commissioning

Injection molding is a preferred manufacturing process given its ability to quickly and efficiently make complex products to high quality. However, it is quite common for problems to be encountered during mold commissioning given the challenge of delivering stringent yet diverse key product characteristics (KPCs) while also managing significant uncertainty related to material properties and start-up processing conditions. When problems occur, it is important to assess the root cause and associated corrective remedy. Typically there issues will arise from one of four sources: 1) material properties, 2) processing conditions, 3) product design, or 4) mold design. Multiple tuning loops are often required to develop a mold design and molding process that provide acceptable quality levels. A significant issue with mold ­commissioning is that the root cause(s) and potential remedies can be subject to debate. Different decision makers may strongly advocate different remedies based on their prior experiences and financial interests. Fortunately, most companies are self-interested in long-term financial stability and so will work cooperatively as partners to resolve issues and develop more strategic partnerships. While each application is governed by the specifics of the negotiated mold purchase agreement, there are some well-known customs set forth by the Society of the Plastics Industry and other industry organizations. This chapter provides an overview of some of the most important concepts with some practical guidance.

„„14.1 Mold Commissioning Objectives 14.1.1 Certify Mold Acceptability As described in Section 2.2, it is common for the mold purchase cost to not be fully paid until the mold has been found acceptable and the customer signs off on the mold acceptance. The mold designer and mold maker appreciate prompt payment

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for any balance due, so the molder and end-user of the molded products should strive to certify mold acceptance within 30 days after the mold has been delivered. Longer delays can cause financial distress with the mold maker. Furthermore, very long delays can impede corrective remedies as the mold designer and mold maker will move on to other applications and may, eventually, forget or discard details related to the mold’s development such as sketches, drawings, CNC programs, patterns, and cutting tools. For this reason, molders should plan to trial received molds within a week of their arrival. Given the potential for conflict during mold commissioning, parties in the molded product supply chain need to be reasonable with respect to mold acceptance and the implementation of corrective remedies as needed. The molder often serves as an intermediary between the mold designer/maker and the end-user of the molded parts. As such, the molder will try to balance the interests of all parties and seek the most cost- and time-effective solutions. Molders will often try to resolve molding issues first through process changes, then material changes, and finally mold design changes. Since molders routinely maintain their inventoried molds, many molders are able to quickly perform many of the changes to the mold design. However, the molder should contact the mold designer and mold maker prior to making these changes, since modification of the mold without permission can constitute acceptance of the mold by contract. In a best case scenario, the mold will be found acceptable as shipped. In most cases where significant mold rework is required, the mold is typically shipped back to the mold maker. The cost of the mold rework can be significant and is dependent on the needed remedy as well as the expertise of the mold designer and mold maker. The owner of the mold should budget approximately 50 % or more of the initial purchase price of the mold for mold rework and maintenance. Indeed, some companies employ a mold procurement strategy of purchasing multiple copies of the cheapest molds possible, then budgeting an amount for rework equal to the full purchase cost of the molds. In a worst case scenario, the molds are not found acceptable and the cooperating parties dispute the best course of action. In some cases, the contractual obligations may not be clear or reparations cannot be made. Then, the final payment to the mold maker is never made and the molder/end-user will seek out a third party to implement corrective remedies. The original parties may simply let the matter drop or seek legal remedy regarding financial remuneration and property ownership.

14.1 Mold Commissioning Objectives

14.1.2 Optimize Molding Process and Quality Once a mold has been found acceptable, the mold commissioning process turns to optimizing the molding process and the quality of the molded products. This ­optimization process is typically performed by the molder with the support and approval of the end-user. The molder is motivated to maximize their profit by maximizing the yield of acceptable products while also minimizing material consumption and cycle time. Meanwhile, the end-user of the molded parts is motivated to ensure the product quality and so needs to provide strict guidance as to acceptable quality levels during mold commissioning. Often, purchase agreements for molded products assume annual productivity gains in injection molding. The end-user should assume that the molder will attempt to continue to improve their molding processes. Accordingly, such process optimization is best conducted in early production runs, before reference process settings and quality levels are established. Minor mold design changes are often made to facilitate process optimization. The mold designer and mold maker may or may not be involved and, if so, may charge for their services on a “cost plus” basis that accounts for their time and related expenses.

14.1.3 Develop Mold Operation and Maintenance Plans Molders will typically work with many end-users to develop mold operation and maintenance plans. During initial mold commissioning, these “plans” can be fairly rough with significant uncertainty that needs to be resolved on an application-­ specific basis. The reason is that each molding application has its own molding behavior with a unique set of requirements that must be fulfilled. Indeed, each mold should be considered a custom-designed machine with distinct components, operation, and maintenance requirements. Hundreds or thousands of parts will typically be molded during the mold commissioning process, leading to valuable experience with the operation of the mold. The molder should strive to leverage these molding trials to validate and customize the acceptance and maintenance plans. Subsequently, the mold designer and mold maker are rarely involved unless replacement parts or mold rebuilding is planned on an intermittent basis. In such cases, it may be advantageous to purchase replacement parts (e. g. pins, spare cavities/cores) with the mold. Similarly, it is standard practice to order standard mold components (for example, ejector pins, cooling plugs, nozzle heaters, etc.) that may shut-down production if ­damaged.

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„„14.2 Commissioning Process Figure 14.1 provides a flow chart of the mold commissioning process, where the parties that are typically involved are shown on the left. The mold designer and mold maker are usually responsible for an internal inspection and test before they ship the mold to the molder. The mold designer and molder should work together to determine the molding process conditions such as temperature, pressures, and timings; many of these process conditions should have been estimated early in the mold specification and design. Both these parties usually work together during the initial molding trial where the mold operation is verified. Any significant defects in the mold design or workmanship are often revealed at this time, and engineering change orders (ECOs) are issued to the mold designer/maker as needed.

Figure 14.1 Mold commissioning process

14.2 Commissioning Process

Once the initial mold verification is complete, the mold designer and mold maker have fulfilled their obligations and should be paid though they are still liable for warranty costs according to the mold purchase agreement. The molder will perform a first article inspection to fully characterize the quality of the moldings. Process capability studies are often performed to optimize the molding process, perhaps with the use of scientific molding techniques and design of experiments [1]. Engineering change orders for the mold may be requested to remedy defects in the mold design or workmanship, increase the molded product quality, or otherwise improve molding productivity. The cost of these ECOs should be paid by the party responsible for the root cause: ƒƒMold design change due to product design change: end-user (original equipment manufacturer, OEM) ƒƒMold design change due to change in the mold specification: end-user or molder ƒƒMold design change due to defect in mold design or making: mold designer or maker Once the mold is fully qualified with acceptable operation and molded product quality, the standard operating procedures should be recorded with a maintenance plan. Each of these foregoing steps is described in greater detail in subsequent sections.

14.2.1 Mold Design Checklist Figure 14.2 provides a checklist for the completed mold design. At the top of the list is a set of design documents that the mold designer should provide to the molder/owner. The mold designer might begin by reviewing the mold purchase agreement and mold specification to verify that all requirements are fulfilled in the implemented mold design. The design documentation is typically specified ­relative to the bill of materials (BOM). Every mold component should be listed in the bill of materials along with that component’s supplier and drawing number if custom. A full set of drawings should be delivered, including completed title blocks with material, tolerances, and finishes. The design documentation should include a mold design report or manual describing the rationale for the mold design including analysis and simulation. Layout drawings for the feed system, water lines, and ejector systems should be provided. This mold manual should also provide layout drawings of the assembled mold from every slide and views from the parting plane; these drawings can be helpful with respect to mold maintenance. The mold manual should also provide a basic process setup sheet with the estimates used for mold design. D ­ rawings of the molded parts, both isometric and orthogonal views, should be ­provided with critical to quality

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attributes indicated. If the mold includes a hot runner system, then the hot runner drawings and instructions should also be provided with the mold manual. All this information should be provided in native electronic CAD format unless otherwise agreed to.

14.2.2 Component Verification Mold designers/makers often take pride in their work and will typically fully ­assemble the mold prior to inspection by the molder/owner. Molders are often tempted to immediately take the fully assembled mold and begin molding trials. However, if the mold is to be used for long-term production, then a thorough inspection of the mold components and their assembly is warranted. The component verification items identified in Fig. 14.2 can be performed at the mold maker prior to assembly, or at the molder/owner’s location after the mold has been disassembled. Each component in the mold’s bill of materials (BOM) should be verified with respect to its materials, finishes, treatments, and quantity. For complex molds, it is standard practice to number cores, cavities, ejector pins, etc. according to the mold drawings to facilitate assembly and maintenance of the mold. The core and cavity inserts should be carefully inspected with respect to finish, texture, and critical dimensions against the design drawings. During mold assembly, the molder should verify that the mold is fully marked to their satisfaction. Each plate can be marked at its top corner with a “0” or the plate number (from 1 to the number of plates in the stack) to facilitate mold reassembly. Each mold plate should have its external edges chamfered, and eyebolt holes ­centered on its side(s). Each water line circuit should be labeled, with water line connectors per the molder specification. To interface with the molder’s machinery, the molder should verify the appropriateness of the mold’s locating ring, sprue bushing, and ejector rod knock-out pattern.

14.2 Commissioning Process

Figure 14.2 Sample mold inspection checklist

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14.2.3 Mold Assembly The quality of the mold design and workmanship often becomes apparent during the mold assembly when inspecting the various items listed in Fig. 14.2. Core and cavity inserts should fit tightly into their respective mold pockets. Interchangeability of the cavity and core inserts may be verified during mold assembly, but best practices (to obtain tight tolerances) are to qualify and operate the mold with a constant layout per the mold drawings. The mold’s plates and other components should assemble without excessive force. When securing the mold plates with bolts, the support pillars should preload the cavity/core plates. No shims, however thin or thick, should be necessary to fine-tune the mold fittings since shims may move or become lost. For molds with hot runners or in-mold sensors, electrical connectors should be securely mounted with wires retained in wire channels. After the mold is assembled, the mold should readily separate at the parting plane(s) without any inadvertent contact on molding surfaces. The actuation of the ejector system(s), core pulls, hot runner valve gates, and any other mechanisms should be verified to move easily through their range of travel.

14.2.4 Mold Final Test Prior to actual molding, the mold’s subsystems should have a final dry cycle test as indicated in Fig. 14.2. Both the water lines and feed system should be pressure tested with shop air pressure to verify that there are no leaks; leaks in these systems are quite common due to minor design or assembly issues but can present major issues during molding trials. The function of the electrical circuits can also be tested by measuring the resistance across the pins for each circuit: a thermocouple should read on the order of an ohm while a heater is typically on the order of 10 to 100 ohms. If the resistance is zero, then the circuit has likely shorted with the mold. If the resistance is infinite, then the circuit is likely open due to a cut wire. Once the mold seems acceptable, the mold is mounted in the molding machine. Standard die set procedures are used to set the die height, ejector travel, interface signals, etc. with low initial speed and force settings. Prussian blue or another ­temporary ink can be placed on one surface of the mold at the parting lines and vents. After clamping, the ink should uniformly transfer to the other side of the mold: ƒƒIf no ink is transferred, then the parting line has not sealed and the mold may need rework. ƒƒIf the ink is transferred in a nonuniform pattern, then portions of the mold shutoff surfaces may be under too much compression and will wear.

14.2 Commissioning Process

The clamp force and actuation speeds are incrementally increased to reasonable values. The inked pattern at the parting line should remain with increasing clamp tonnage and cycles. When full clamp force is reached, the inked vents should be checked to verify that no ink has transferred across the parting plane. Ink transfer from the vent channels to the other side of the mold indicates crushing of the vents during mold compression. Engineering change orders should be issued to remedy any deficiencies with the mold’s components or assembly. If the issues are very minor, then the mold’s initial molding trials may proceed with permission of the molder/owner.

14.2.5 Preliminary Molding Recommendations Figure 14.3 provides a preliminary process setup sheet that facilitates the mold start-up. The top section of includes the part name, wall thickness, shot weight, an isometric view of the molding(s), and view of the A and B side of the mold from the parting plane. The material data includes the type of material, specific grade name, melt flow index, and melt/mold temperatures as well as drying conditions. While most molders can start up a mold without any documentation, they usually appreciate knowing the mold’s physical characteristics as well as any guidance on the filling, packing, and cooling stages. To develop an initial estimate of the filling profile, the shot volume, Vshot , is first calculated as: Vshot =

mshot (1 + fpacking + fcushion ) (14.1) rmelt

where mshot is the shot weight of the moldings (including any cold runner segments) available from the mold design, ρmelt is the polymer’s density at mid-range melt temperature provided from Appendix A, fpacking is the percentage of the ­injection volume reserved for compensation of the volumetric shrinkage during packing, fcushion is the percentage of the injection volume to remain at the end of the injection stage. These percentages will vary by application, but 5 % is a reasonable starting point. The injection stroke may then be calculated by dividing the 2 . If the screw diameter is not shot volume by the projected area of the screw, p Rscrew known, then the mold designer can still provide the filling time as well as the flow rate (relative to maximum) as a function of injection volume. The middle section of Fig. 14.3 provides an example for the cup/lid mold.

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Figure 14.3 Preliminary molding process conditions

The filling velocity profile in Fig. 14.3 is developed to try to maintain a uniform melt front velocity during filling of the mold cavity. In this mold design, the runner volume is 5.6 g, or 8.1 % of the shot volume. A lower flow rate is thus used during the first 10 % of the injection stage to avoid excessive shear heating and potential problems in melt flow at the gates (such as jetting and splay); the reason for the

14.2 Commissioning Process

use of 10 % rather than the 8.1 % of the shot volume is to allow for melt compressibility in the barrel as well as slower melt velocity in the initial filling of the cavities where the projected area of the melt front is very small. The magnitude of the flow rate can be adjusted, and hopefully increased, during the molding trials once it is verified that there are no issues with the runners/gates. At Point 3 in the filling profile of Fig. 14.3, the flow rates are increased as the melt fills the cavity and the projected area of the melt front increases. The use of v­ elocity ramps, rather than step changes, is recommended for two reasons. First, a sudden step change in flow rate is likely to induce a step change in the melt velocities and shear stresses, with potential defects such as hesitation marks, etc. in the moldings. Second, the molding machines can induce process instabilities related to closed loop control of step changes. Specifically, the process controllers will have proportional and integral terms that can cause overshoot with significant step changes if the controller gains have been set aggressively. Conversely, conservative controller gains can cause a relatively sluggish response. In both cases, the resulting process will not be as repeatable as a well-tuned machine using a simple velocity profile. The injection flow rate is ramped down towards the end of the filling stage (at ­filling profile Point 5 in Fig. 14.3) as the polymer melt converges on the end of the mold cavity. Since the projected area of the melt front decreases as the polymer melt contacts the mold walls at the end of flow, the reduction in flow rates is helpful to try and maintain a consistent melt velocity. Furthermore, the reduction in flow rates is also useful to avoid excessive injection pressures as well as a smooth transition to the packing stage. The total filling time is based on the mold filling analysis described in Chapter 5 and/or recommendations from injection molding simulation. The magnitude of the velo­city profile can then be calculated accordingly as a function of the screw position. The suggested packing pressure should be around 75 % of the peak filling pressure, with the magnitude adjusted upward or downward according to the shrinkage analysis of Chapter 10 and, subsequently, the dimensional requirements of the molded parts. The pack time is related to the gate freeze-off (unless hot runner valve gates are used). The gate freeze time can be estimated according to Section 7.3.5. As discussed therein, the actual gate freeze time will be longer than that predicted assuming pure heat conduction since the convection of hot melt during packing will tend to keep the gate from freezing. Accordingly, a reasonable first assumption is that the pack time will equal 25 % of the total cooling time for the part (in which the cooling time here is defined to include the packing time). As such, a pack time of 6 s is initially predicted for the cup/lid mold as shown in the bottom section of Fig. 14.3. In this profile, a ramped packing pressure decreasing to the plastication pressure is provided during the last 2 s to assist the gate freezeoff and avoid over-packing of the sprue.

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„„14.3 Molding Trials Molders generally know how to set up and operate molds, and this section is not intended to provide a treatise on injection molding. However, the mold designer can assist the molder by providing molding information that was used in the mold design. The section is also intended to provide some practical information that the author has found consistently reliable in industry application. These guidelines are generally consistent with scientific molding principles and the mold design methodology presented throughout the book. The preliminary process conditions provided in Fig. 14.3 are based on the mold design and related scientific principles. However, no mold designer or molder should expect the preliminary process conditions originating from the mold design to be optimal. Accordingly, the molding trial methodology provided in Fig. 14.4 is recommended. After the operation of the mold has been verified with dry cycles, the molding trials proceed in two major stages. First, the filling and packing stages are set up as subsequently described to produce moldings that “appear” to be acceptable. The word “appear” is purposefully used here since it is often impossible to fully confirm the acceptability of the moldings by visual or other direct ­inspection alongside the molding machine. A delay of hours or days is often needed to allow the moldings to equilibrate prior to metrology, and such detailed metrology is often performed off-line given the time and effort required to fully characterize the molded part quality. Accordingly, the optimization stage is described in Section 14.4.

Figure 14.4 Molding trial methodology

14.3 Molding Trials

14.3.1 Filling Stage After the operation of the mold has been verified using dry cycles, the filling stage of the mold can be set using short shot studies [2] in which the packing time is set to the expected packing time but the packing pressure is set to 0. Figure 14.5 provides a continuation of the cup/lid example that focuses on the ram velocity profile during the filling stage. Given the initial process conditions from the mold design, profile Ê uses a simple profile with a large shot volume along with a low screw velocity and small injection stroke. The reasons for this profile are 1) to ensure that the mold is structurally sound, 2) to verify that there are no missing components or leaks in the mold, and 3) to determine that the feed system and moldings appear to be working as designed. If the mold appears to be operating as intended, then the shot size, injection ­velocity and maximum injection pressure are iteratively tuned based on the ­observed short shots until the moldings are nearly fully formed without any filling-related defects. Profile Ë illustrates a simple ram velocity profile to provide nearly full cup/lid moldings. However, the molder would likely realize that this profile does not provide adequate cushion given further ram displacement during the packing stage. Furthermore, it may be possible to improve the quality of the moldings by using a lower screw velocity at the start of injection as well as improved molding productivity with a higher screw velocity during the majority of the filling stage. Profile Ì provides a final profile that corresponds to an injection time of 1.5 s.

Figure 14.5 Evolution of the injection velocity profiling

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There is some disagreement over the value and practicality of injection velocity profiling [3, 4]. However, velocity profiling can provide the degrees of freedom ­necessary to improve molded product quality and molding productivity [5, 6]. Very complex profiles that have steep ramp rates and oscillating velocities should not be used since the compressibility of the polymer melt and delays related to the ­controller response do not allow direct control of the polymer melt velocity in the cavity. For example, an “optimal” velocity profile suggested by computer simulation may have 15 points requiring a decrease in the ram velocity in 0.01 s as the polymer melt passes through the gate(s). Such a profile is unlikely to be achievable and should be simplified. There is also some disagreement about the V/P switchover condition for transfer from the filling stage to the packing stage; the “V/P” acronym refers to the Velocity/Pressure feedback control corresponding to the filling and packing stages. Generally, switchover based on screw position is preferred unless feedback is available from an in-mold sensor located at the end of the cavity to mark the arrival of the polymer melt [7]. Switchover based on injection pressure and time have been shown to reduce molded product consistency due to extrinsic variation of the material and other processing states. A position-based switchover at 95–99 % of the volume is the norm, with the final V/P position dependent on the mold geometry, ending screw velocity, and the packing pressure as next described.

14.3.2 Packing Stage The objective of the packing stage is to achieve the desired product dimensions and aesthetics by compensating for volumetric shrinkage as the polymer melt cools. Figure 14.6 provides an example for packing pressure profiling of the cup/ lid mold. Section 6.4.6 indicates that the total pressure drop was designed to be ~65 MPa. Profile Ê corresponds to an initial pack time of 8 s with the packing pressure set to 55 MPa, approximately 85 % of the injection pressure at the end of the filling stage. A gate freeze study should be conducted in which the packing time is first set to a long time, e. g., 12 s, to ensure that the packing pressure is maintained until the gate has frozen and no more material can be packed into the cavity as the moldings are cooled. The packing time can then be successively reduced across molding cycles until the weight of the moldings is found to decrease. Generally, setting the pack time equal to the gate freeze time will result in consistent molded product dimensions.

14.3 Molding Trials

Figure 14.6 Evolution of the packing stage profiling

The magnitude of the packing pressure and time can be guided by the analysis of Chapter 10 to tune the volumetric shrinkage and part dimensions. Profile Ë uses a packing pressure of 65 MPa, an increase made to reduce shrinkage/sink, and a packing time of 6.4 s to ensure adequate packing prior to gate freeze. The use of a packing time equal to or longer than the longest gate freeze time is generally not mandated, though some knowledgeable and demanding end-users may, in fact, stipulate such. Pack pressure profiling can provide the ability to reduce the packing (and cycle) time by using an artificially high pack pressure with an artificially long ramp at the end of packing. This combination can provide for higher melt pressures in the cavity during packing. Then, since the gate is not quite but almost frozen, the decreasing pressure ramp serves to equilibrate the melt pressures on opposite sides of the gate. With no pressure imbalance, there is no melt flow through the gate and the gate can freeze more rapidly than otherwise possible without a pressure ramp. Such a strategy is exemplified by profile Ì in Fig. 14.6.

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Figure 14.7 Dynamics of velocity to pressure (V/P) switchover

The injection velocity and pack pressure should be smooth (technically, C1 continuous) at the changeover from velocity to pressure control. Figure 14.7 provides an example for the cup/lid mold. At the ideal transition, the melt pressure at the end of the filling stage equals the pack pressure as shown at a time of 1.6 s. Then, the screw velocity will decay as the polymer melt completes filling of the mold and the melt throughout the cavity is pressurized. If the changeover occurs too late, the melt pressure will follow the ideal case until the melt reaches the end of the cavity. Afterwards, the melt pressure will begin to increase significantly as the machine is under velocity control so the screw will ram more material into the cavity and cause a potential overpacking situation with high melt pressures as observed at 1.65 s. When the machine does switch over to a lower packing pressure, the pressure differential acting on the screw will cause it to “kick back,” literally with a negative velocity. Conversely, if switchover occurs too early, e. g., at 1.55 s, then the melt pressure is significantly less than the set packing pressure, which will cause

14.3 Molding Trials

the screw to “surge forward.” Both of these events are undesirable since they can cause defects such as back flow, flash, nonuniform shrinkage, etc. Note how in both of these events, the rapid change in the screw velocity is the result of a mismatch between the melt pressure at the end of the filling stage and the set packing pressure. For a smooth transition, these two pressures should be equal; the switchover (V/P) position and the injection velocity specified in the last step in the filling stage should be adjusted so that the pressure at the end of the filling stage equals the specified packing pressure.

14.3.3 Cooling Stage The cooling stage is typically set to allow the moldings to cool to the point where they may be ejected from the mold. There are generally three benefits to longer cooling times. First, longer cooling provides more volumetric shrinkage, such that parts may be more readily ejected from deep and narrow ribs where there is a ­tendency to stick. Second, cooler parts tend to be stiffer given that modulus is a function of temperature, and so longer cooling times will facilitate the transmission of the ejection forces across the part. Third, the parts have a lower temperature difference relative to room temperature, and so will have less post-molding shrinkage. That said, increased cooling times are highly undesirable since they will directly reduce molding productivity. Also, longer cooling times can cause ­excessive in-mold shrinkage, which will actually increase the necessary ejection forces per the analysis of Section 11.2.2. If ejection forces are too high, then defects such as push-pin may result; this defect is often resolved by increasing the packing pressure and reducing the cooling time. When asked why a particular cooling time has been selected, molders often indicate that the cooling time is needed to obtain the necessary product quality. However, sometimes the rationale for a specific cooling time is unknown. As stated in the bottom left block in Fig. 14.4, it is always a good idea to successively reduce the cooling time until the dominating constraint is identified. In many cases, the cooling time will be limited by the plastication time. In other cases, the cooling time will be limited by the ejection of the runner system or drool from nozzle(s). The molder and mold designer should experiment with a variety of cooling times and mold coolant temperatures, then weigh the cost and benefit of potential strategies for further reducing the cooling and cycle time. As suggested in Chapter 3, the ­return on investment is a strong function of the expected production quantity and lifetime of the mold.

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„„14.4 Production Part Approval The objective of the molding setup procedure of Section 14.3 was to provide a set of operating conditions that produces acceptable product. Further molding trials are needed to validate that the developed process can be used as the standard process. Every molded part design has its own set of quality requirements, and many end-users of injection molds will have their own product development and approval processes. A formal production part approval process (PPAP) has been developed by the Automotive Industry Action Group (AIAG) to qualify materials, processes, molds, and molded parts for the automotive industry. This joint initiative was founded in 1982 by North American automotive manufacturers (Chrysler, Ford, and General Motors) whose membership has grown to include Daimler, Toyota, Honda, and Nissan, and many part suppliers and service providers. The PPAP is quite detailed and requires significant investment, yet can provide evidence that the design documents are “properly understood” and the manufacturing process provides consistent product at production rates. This section provides an overview of the fundamental concepts useful for first article inspection.

14.4.1 Quality Assurance Quality assurance is vital to ensuring the consistency of incoming received materials, internal manufacturing processes, and outgoing manufactured products. A flow chart of a quality control methodology is shown in Fig. 14.8. The materials received by the facility for use in their manufacturing processes typically undergo acceptance sampling. If the materials do not meet specification, the lot should be rejected. During production, key product characteristics can be collected with cameras, scales, or other on-line metrology systems. In theory, this quality data can be used with a feedback controller to automatically adjust the process settings to ­regulate the product quality. However, most injection molders do not use on-line metrology systems and so these elements are indicated with dashed lines. Regardless of the use of on-line metrology, statistical process control is often applied with the process data and/or quality data to ensure process consistency. As a final check, acceptance sampling of the manufactured product can be conducted prior to shipment. If either the statistical process control or the final acceptance sampling identify quality issues, then material, process, or mold design changes may be necessary to achieve acceptable quality levels.

14.4 Production Part Approval

Figure 14.8 Quality assurance methodology

Quality control methods vary widely by the molding application requirements, and not every molder or molding facility will use these same techniques. However, ­acceptance sampling, on-line metrology, and statistical process control are quite common and provide a complementary basis for measuring and controlling quality. Prior to the use of any of these techniques, however, gauge repeatability and reproducibility is necessary to ensure the statistical significance of whatever measurements are being taken.

14.4.2 Gauge and Process Repeatability & Reproducibility The measured product quality is the result of many factors including the mold ­design, polymeric material, molding process, environmental conditions, and the capability of the measurement processes themselves. The purpose of a gauge repeatability and reproducibility (gauge R&R) study is to determine the consistency of the measurement system [8, 9]. Given the number of potential sources for variations in quality, gauge R & R studies should be performed prior to using the measurement data for process modeling, optimization, control, or product acceptance/ rejection. Gauge R & R is classically applied to dimensional metrology, but the ­concept is also applicable to other process/product measurements. In a gauge R & R study, the same measurement instrument or “gauge” is used by the same trained operator to measure the same dimension on the same part ­several times. The number of repeated measurements is typically on the order of 10 to 40.

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The “repeatability” of the measurement is the variation of the same measurement obtained by one operator with one gauge. The operator “reproducibility” is the variation of the average measurement obtained by different operators with the same gauge. Similarly, the process repeatability is the variation of the average measurement obtained by one operator with the same gauge for different parts. For example, suppose that the thickness of one molded part is measured at the same location with the same micrometer 10 times by three different operators resulting in the data plotted at left in Fig. 14.9. It is apparent that operators 1 and 2 provide better measurement repeatability with the gauge than operator 3. The ­statistics for each operator are computed independently. The average thickness for operators 1, 2, and 3 are respectively 1.4001, 1.3982, and 1.4022 mm. The gauge reproducibility, taken as the standard deviation of these averages, is equal to 0.003 mm. The gauge repeatability, taken as the standard deviation of the measurements for each of the three operators, are 0.0023, 0.0016, and 0.0063 mm. Note that there are three gauge repeatability measurements, one for each of the three operators using the same gauge. Using these statistics, the normal probability ­density functions can be plotted as shown at right in Fig. 14.9.

Figure 14.9 Gauge repeatability & reproducibility results

14.4 Production Part Approval

The horizontal dashed lines represent the lower specification limit, LSL, and upper specification limit, USL, on the thickness. The process capability index, CP , is a standard measurement of process capability defined as [10] the breadth of the ­specification limits divided by six standard deviations, σ, of observed variation: CP =

USL - LSL (14.2) 6s

The process capability index was defined such that a value of 1 is considered standard. If the process is centered between the upper and lower specification limits, then three standard deviations of process variation exist between the mean and each specification limit, such that the yield is 99.7 %. An issue with the process capability index is that it assumes that the process is centered with the mean, m, between the upper and lower specification limits. If the process is not centered, then the fraction of acceptable products (referred to as the “yield”) will be much lower than suggested by the process capability index. In such cases, the ­asymmetric process capability index, CPK , should be used: é m - LSL USL - m ù ú (14.3) CPK = min ê , êë 3s 3s úû

Applying the operators’ measurement statistics of the data in Fig. 14.1 with a LSL/ USL of 1.39/1.41 mm provides CP values of 1.45, 2.08, and 0.53, respectively, for the three operators. These results indicate that the gauge repeatability for operator 3 is a significant issue in quality control since the measurement of the parts will lead to a high portion of falsely rejected product. Such repeatability issues are quite common in practice for reasons such as operator fatigue, limited dexterity, lack of care, or improper instruction. Whatever the reason, the manufacturer should definitely avoid assigning operator 3 to this measurement task. Frankly, none of the operators really provide measurement fidelity sufficient to develop highly capable processes relative to the specification limit, and the molder should strive to develop better testing methods since the quality of the measurement itself defines the limit of the quality that can be achieved.

14.4.3 Image-Based Dimensional Metrology As implied by the need for the above section on gauge R & R, a recurring issue in production part approval is the quantification of part quality. Recent advances in imaging techniques and computational processing is supporting part metrology by optical and X-ray scanning techniques to provide significant metrology data. Both these techniques provide for faster, more complete, and more accurate metrology

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compared to coordinate measurement machines and manual measurement with calipers or micrometers. Although digital cameras, optical comparators, and other methods have been used in dimensional metrology, rapid progress in optoelectronic components has ­significantly increased the speed and resolution at which images can be ­acquired [11]. Con­currently, increased computational power and improved optical image recog­nition algorithms have enabled automatic conversion of point clouds to parametric features [12]. Multiple cameras at different angles can be used to ­recognize part features in three dimensions. The resolution is fundamentally ­related to the number of pixels in the image and potential errors in alignment. For this reason, it is fairly common to place gauge features of known absolute dimension in the ­imaging ­system’s field of view for calibration. A 10 megapixel image, composed of a 1000 by 1000 grid of pixels, would then have a raw accuracy of 0.1 mm across a 100 mm image length. The actual measurement accuracy could be significantly improved if multiple i­ mages were taken at slightly different positions and analyzed with ­feature recognition. Such optical techniques are nearly instantaneous, and are ­being applied in online metrology systems. Future systems based on time of light laser systems (LIDAR, [13]) when coupled with adaptive coarse-fine controller [14] promise ­potential micron-level measurement of part ­features. Another image-based method for dimensional metrology is industrial computer ­tomography (CT), which is already being used in medical and many other critical applications. A typical system is shown in Fig. 14.10 being used for the measurement of a turbine blade [15]. In application, an X-ray source projects an X-ray fan beam against a detector on the other side of an object to be measured. A moving stage holding the object with gauge pins (made of known materials) is moved through the fan beam in various directions to scan different areas of the blade. The computed tomography scan can be performed at varying resolutions and X-ray ­intensities. The inclusion of multiple gauge pins made of known materials provides for calibration with respect to both size and material properties (e. g., density). The controller analyzes the acquired point clouds in a modeling program to extract vector information (parametric features) about the geometry that can then be directly compared to specifications from the object’s original computer aided design [16, 17]. Comparing these two techniques, optical imaging is faster while CT scanning provides insight to internal features including voids, internal geometry or components, and density variations. As these techniques become commoditized, they are likely to be used not only for first article inspection, but also for intermittent sampling and ultimately 100 % quality assurance [18].

14.4 Production Part Approval

Figure 14.10 Computed tomography system

14.4.4 Process Capability Evaluation The yield and process capability indices are directly related through the normal cumulative distribution function. If the process is centered, then the yield can be estimated from the process capability index as: æ 3 × CP ö÷ ÷ (14.4) yield = erf çç èç 2 ø÷÷

where erf is the Gaussian error function. The expected yields for processes of varying capability are listed in Table 14.1. It is observed that yield increases significantly as the process capability increases from zero and then approaches 100 % at process capabilities above 1. Given normal statistics, the yield provides 99.9999 %, sometimes referred to as “six nines” at a process capability index of 2. This quality level is the goal of many Six Sigma initiatives and in theory should result in only 2 defects per million opportunities. Of course, a process capability greater than 2 does not preclude defects due to random, uncontrolled events such as a power ­outage. Molders and end-users need to determine acceptable quality levels on an application-specific basis by weighing the value and cost of quality [19–21].

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Table 14.1 Process Capability Levels and Their Corresponding Yield and Defects Per Million Opportunities (DPMO) Process Capability, CP

Yield (%)

DPMO

0

0

1,000,000

0.167

38.29

617,100

0.333

68.27

317,300

0.5

86.64

133,610

0.667

95.45

45,500

0.833

98.76

12,420

1.0

99.73

2700

1.333

99.95

63

1.667

99.99

6

2.0

99.9998

2

2.5

99.99999999

0.0001

3.0

~1

~0

During mold commissioning, molders will require a process capability index in which parts are continuously molded at the developed operating conditions. The automotive production part approval process [22] requires a minimum of 400 ­consecutive molding cycles with a process capability index of 1.67. Other molding applications, such as some for medical or military or aerospace, may require a process capability index of two according to “Six Sigma” guidelines. Six Sigma is a quality control initiative, originally developed by Motorola [23], to ensure that six standard deviations of process variation should be maintained between the process mean and the closest specification limit. Statistically, only two defects per million opportunities (DPMO) would occur if the Six Sigma criterion was satisfied. Since its inception, Six Sigma has drawn upon many previously established methodologies including the design, measure, analyze, improve, control (DMAIC) process [24]. Another significant underpinning of Six Sigma is the capability roll-up for manufacturing processes that have multiple quality requirements, each with its own process capability index. In a capability roll-up, the defects per million ­opportunities are individually calculated for each specification according to the previously defined normal statistics. The defects per million are then summed to provide the total number of expected defects per million opportunities, DPMO. The calculation of the total DPMO is expressed by the equation: m æ æ 3 × CP,i ö÷ö÷ ÷÷÷÷ (14.5) DPMO = å1×106 × ççç1-erf ççç èç 2 ø÷ø÷÷ èç i=1

14.4 Production Part Approval

where CP,i is the process capability associated for the i-th of m specifications. After the total DPMO is calculated, the final yield considering all quality requirements is estimated as: æ DPMO ö÷ yield = çç1÷ (14.6) èç 1×106 ø÷

The rolled up process capability index is then calculated as: CP =

2 -1 erf (yield) (14.7) 3

For a process to qualify as Six Sigma, the rolled-up process capability index must be 2 or greater. Molders will often find that the process capability is limited by a couple of key quality attributes, and that improving one of these attributes by changing the molding conditions can push other quality attributes out of specifi­ cation. The interplay between the process and quality constraints defines the bounds of the “process window” as shown in Fig. 14.11. Here, two processing ­parameters, x1 and x2 , are allowed to vary between lower and upper processing limits, LPLi and UPLi . Only a portion of the processing space will yield acceptable moldings, because there are multiple quality attributes, each of which must be between its lower and upper specification limits, LSLj and USLj .

Figure 14.11 Process window for multiple process and quality constraints

Figure 14.11 illustrates how some of the constraints are more demanding or “­dominate” the boundaries of the process window. In this example, suppose that:

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ƒƒx1 represents the barrel temperature, with minimum and maximum limits of 200 and 250°C, respectively; ƒƒx2 represents the injection velocity with minimum and maximum limit of 10 and 200 mm/s, respectively; ƒƒy1 represents the thickness, with lower and upper specifications of 1.39 and 1.41 mm, respectively; ƒƒy2 represents the knit-line strength, with only a lower limit of 30 MPa; ƒƒy3 represents the flash, with a maximum limit of 100 mm. It can be understood from the diagram that the thickness, knit-line strength, and flash can each be problematic at different combinations of barrel temperature and injection velocity. However, the upper constraint on the thickness is unlikely to be the governing constraint given that flash and knit-line strength will become ­problematic first. The nominal process window is defined by the intersection of the feasible region associated with each constraint. In theory, any operating point within the process window will provide a manufactured product that meets all the process and quality constraints. In practice, however, operating close to the boundary of the process window will tend to yield poor results. The reason is that there is variation in the processing conditions, xi , as well as uncertainty in the product quality, yj . For this reason, a smaller but more “robust” process window is sought within which the process is insensitive to variation. An analytical method has been developed for solving the process window in high dimensional space for many variables [25]. However, the algorithm and data ­structures are complex and still subject to uncertainty in the models, yj = f(xi), that relate the quality attributes to the processing conditions [26]. A practical approach to confirming the robustness of a newly commissioned mold is to perform a design of experiments (DOE) centered around the proposed operating conditions [27, 28]. In the DOE, each processing condition is set to the proposed set-point plus or ­minus a small perturbation representative of the variation that is expected during longterm production. As indicated in Fig. 14.11, the process is robust if the quality for all the moldings made across the DOE runs is acceptable. Mold commissioning according to these formal methodologies represents a fairly significant investment, and engineering change orders are frequently requested during these process capability studies to develop an acceptable molding process. If the quality attributes are close to specification, requests can be made to slightly broaden or shift the product specification to achieve the required process capa­ bility levels. If the request is denied or the molded product quality is far from specification, then changes to the mold design are recommended.

14.5 Mold Maintenance

„„14.5 Mold Maintenance Menges [29] provides an interesting industry survey of industry practices based on “fire brigade” and preventive maintenance strategies: only about 30 % of injection molding facilities implement formal preventative maintenance approaches. Practitioners agree that molds should be well-maintained for at least two reasons: ƒƒFirst, molds are expensive and regular preventive maintenance can significantly extend their lifetime. For example, wear on parting lines will allow the formation of flash. If the parting line is not well maintained, then the amount of flash can successively increase which thereby increases the projected area of the polymer melt and the force it exerts on the parting plane of the injection mold. Across additional cycles, the growing flash will act like a wedge that wears the parting plane and deforms the parting plane until the mold must be reworked. ƒƒSecond, a well-maintained mold will increase molding productivity and avoid ­unscheduled delays associated with unforeseen repairs. For example, poorly maintained cooling lines will develop internal corrosion that reduces the flow rate and heat transfer from the mold cavities. If the corrosion is never addressed, the product quality can become inconsistent across the lifetime of the mold. Molders may compensate by extending the cooling time; however, such changes in the process conditions will reduce productivity and may impact product ­quality in unforeseen ways. To ensure proper maintenance, a mold operating log should be kept that includes a record of maintenance as well as any issues and/or molding defects. Figure 14.12 provide a sample mold operating log. There are essentially three levels of mold maintenance that are recommended at different time intervals: 1. Pre-molding and post-molding maintenance: these simple tasks are performed every time a mold is run to respectively ensure that the mold is fit to operate and will not be damaged during storage. The mold map and operating log are simple methods to track the mold performance and record any issues. 2. Regular maintenance: these items are performed on a regular basis to ensure that the mold is kept in good working order with consistent performance. The required maintenance interval is application dependent, but usually on the order of 50,000 cycles. Most molds will undergo several regular maintenance rounds throughout their lifetime. 3. Rebuilding: intermittent reworking of the most critical mold components but is necessary to ensure continued mold operation. Mold rebuilding represents a ­significant expense that should be planned when the mold is not needed in pro­ duction. The number of molding cycles before major maintenance is application dependent, but will typically vary between 100,000 and 500,000 cycles.

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Figure 14.12 Sample mold log and maintenance checklist

If the mold is not well maintained, then the performance of the mold will decline over time. In high volume production with multicavity molds, some molders routinely shut off damaged or otherwise inoperable cavities and continue production by adjusting the molding process conditions (and, most often, by using an earlier VP switchover in proportion to the number of working cavities). This practice should be avoided, and is indicative of very loose product/process/quality development methodologies. Reducing the number of mold cavities not only reduces molding productivity, but also unbalances each leg of the runner system, and with it the proportion of melt flow and pressure delivered to each cavity. The parts may appear to be fine, even with the same part weights, but the fact is that the process is not the same as the approved standard. Accordingly, the molder is likely liable if the moldings are deficient. Simple, regular maintenance can avoid such issues and maximize molding productivity.

14.5 Mold Maintenance

14.5.1 Pre-Molding Maintenance The exact tasks and their sequence for pre-molding maintenance will vary with each molder and application, and the checklist provided near the bottom left of Fig. 14.12 provides as simple a process as possible. Prior to molding, the date and time that each mold is taken from the mold inventory should be recorded along with the work order number that authorizes the mold to be put into production. The standard process conditions (a more detailed and fully validated version of Fig. 14.3) should be verified; most modern molding machines have a computer-based, graphical user interface that allows molding recipes to be stored and reloaded. As the machine is warming up, the operator should inspect the mold for damage. The most common components that are damaged include the cooling line plugs, ­electrical cables and connectors, external cooling hoses, gates, slides, and pins. Inspection of the external components should be performed prior to clamping the mold in the molding machine; many minor repairs can be made quickly without production delays. Once the mold is clamped in the molding machine, the mold should be operated on a slow, dry cycle and the internal components of the mold further inspected for any damage. The mold and ejectors should operate silently, without any sound when the guide pins insert into the guide bushings upon mold closure or any grating noises when the ejector plate and any slides are actuated. The water lines (and hot runner) can then be connected and started to bring the mold up to temperature. Any leaks or other abnormalities should be resolved prior to mold operation. Any grease or residue on the mold parting plane, molding surfaces, and vents should be removed, taking care to never damage the molding surfaces. Any damage or discoloration or haze on the molding surfaces should be recorded in the mold operating log and assessed by quality control for authorization to proceed. The molding machine should then be operated in a semiautomatic mode to ensure that the moldings are being consistently produced and ejected prior to switching to a fully automatic cycle. If all seems well, then the operator should operate the molding process in a fully automatic cycle for several cycles prior to having the quality control group inspect the new moldings and authorize the production run. If the production run is authorized, then the operator can check and initial the box for the pre-molding checklist. Otherwise, a comment (and possibly engineering change order) should be entered with a plan to correct any issues.

487

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14 Mold Commissioning

14.5.2 Molding Observation and Mold Map Dealey [30] provides a detailed of discussion of mold maintenance and introduces the concept of the mold map for tracking the condition of the mold. During molding, the operator should intermittently observe the molding process and visually inspect the quality of the moldings. Any issues can be documented in the mold operating log in the comment section with a mapping of the problematic areas on the views of A and B sides of the mold from the parting plane. Table 14.2 provides an example of some mold defect codes that can be used, with others added by the molder on an application-specific basis. The top right of Fig. 14.12 provides an ­example of a mold map wherein the designated mold defect codes correspond to the observed defects as further described in the comment section of the mold ­operating log. Table 14.2 Some Mold Defect Codes for Marking a Mold Map Mold defect codes: problem CL: cooling-related problem

PD: plastic defect

CR: cracked

 PD1: flow lines  PD2: burn marks  PD3: jetting  PDx: etc.

DT: dirt or dirty DM: dimension-related issue EJ: ejector-related problems FL: flash

PO: polish or surface finish

GT: gate-related problems

RN: runner

LP: lubricant problem

RT: rust

MB: mold base issue

SC: scratch

OL: oil leak

SL: slides VN: vent WL: water leak

The mold map at the top right of Fig. 14.12 provides examples of indicated mold-­ related defects. During mold commissioning, two issues were identified with EJ and RN mold defect codes and comments in the mold operating log. First, the RN code indicates that the primary runner to cavity #2 was restricting melt flow and causing the cavity to fill too late; an engineering change order (ECO) requests increasing the runner diameter from 2.5 to 2.78 mm. Second, the EJ code indicates that the runners system was intermittently hanging up upon ejection, so an ECO requests keying the sprue puller pin so the undercut that catches the sprue faces downward. Subsequently, at run 5, the CL(8) code indicates that the cooling line plugs do not match the hoses at press #12 per comment 2. As shown, the mold map and operating log provide a very simple method to track the operation and performance of the mold across its lifetime.

14.5 Mold Maintenance

14.5.3 Post-Molding Maintenance The post-molding maintenance is also intended to be a simple, standard procedure that operators can perform without annoyance. A basic procedure is provided at the center of Fig. 14.12, and begins with purging and cooling of the hot runner system (if the mold has a hot runner system and purging is needed). The mold’s water lines should then be blown out and the mold dried; corrosion of the mold’s water lines and surfaces are significant issues that can be avoided with a little effort. The parting plane and any vents/slides are often easily cleaned with the mold open while still in the press. The time spent cleaning also provides an opportunity to inspect the mold for any damage. Subsequently, mold preservative may be ­applied if the mold is to be out of service for an extended period. Finally, the mold is removed from the molding machine and secured with safety straps. The post-molding maintenance is not completed, however. The last molding should be stored as a record of the mold’s performance and molding/material consistency. Some molders, especially in medical and other industries with high warranty costs, save molding samples from each production run. This practice allows the molder to trace changes in the mold, material, and process that may adversely ­affect the molded product quality relative to that which was accepted during mold commissioning. A green (good) or red (bad) tag indicating the current state of the mold should also be attached, and the mold stored at its designated inventory location. Afterwards, the mold operation log can be completed, along with any needed comments or engineering change orders.

14.5.4 Scheduled Regular Maintenance The objective of the regular maintenance is to keep the mold operating in a consistent manner. Molders often develop their own maintenance schedules according to their application requirements and experiences. The plans provided at the bottom right of Fig. 14.12 provide some basic guidance. These maintenance items should be performed regularly, though the number of molding cycles between maintenance will vary significantly with the type of polymer, quality of the mold, and the ­implemented molding process. Most medium-sized and larger molders are capable of completing all the listed tasks completely in their facility within a day or two. During maintenance, the molded parts should be fully inspected with the cause of any significant changes in the quality ascertained. The highest priority is given to moving components that may wear and cause ­significant damage to the mold should they break. All core pulls, lifters, slides, stripper plate, unscrewing mechanisms, etc. should be disassembled, cleaned, and lubricated to ensure that they operate freely with minimal friction. Angle pins,

489

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14 Mold Commissioning

cams, and gears should be closely inspected. If any of these components are worn or bent, they should be replaced before the mold is put back into service. All molding surfaces should be carefully cleaned with a cotton or felt pad with polishing compound or mold cleaners as appropriate. Scratches as well as poorly polished and textured surfaces should be repaired. The gates, shut-offs, and vents should be inspected against the mold drawings to characterize wear with changes above 10 % of the nominal dimensions noted in the mold operating log. The progression of wear should be tracked in subsequent maintenance cycles and addressed before any mold cavities need to be removed from service. Johnson [31] suggests close inspection of flash as an indicator of mold wear, with rebuilding scheduled when the length of flash exceeds one-half of the allowable limit. The cooling lines should be serviced to minimize corrosion and maintain a high level of performance. First, all cooling line plugs should be removed to give access to the cooling lines. Mineral, rust, and other materials will tend to form on the ­internal surfaces of the cooling lines, causing the water flow and cooling rates to progressively decline. Most of this scale can be removed with a brass bristle brush sized to the cooling line diameter. The cooling plugs can then be reinstalled with any external hoses/clamps. The cooling lines can then be treated with a chemical cleaner such as hydrogen chloride to remove remaining scale. If corrosion is a ­persistent problem, then the molder should consider water treatment systems to minimize acidity (pH) and chlorine levels.

14.5.5 Mold Rebuilding Scheduled major maintenance represents a significant investment, and should be scheduled when production will be off-line. Most molds will only rebuilt a couple times across their life, and a well-designed mold used at low pressure with a non­ abrasive, noncorrosive polymer (like polypropylene) may never need to be rebuilt. Most molders do not have the resources to rebuild their molds, and ship the mold back to a mold maker for this service. The mold maker should also be provided the updated mold design drawings, manual, and operating log. Here, a well-maintained set of documents will help to efficiently rebuild the mold while also helping to ­improve future mold designs by informing the mold designer/maker about the mold’s past performance. During mold rebuilding, the mold should be fully disassembled and every component cleaned and inspected as discussed in Section 14.5.4. However, the disassembly and cleaning should now include the hot runner and all other subassemblies. The dimensions and surface finish of every component should be checked against the mold’s updated design drawings. Any discrepancies between the mold assembly and documentation should be discussed with the molder to update the design and documentation.

14.6 Summary

Sections of cracked plates and mold inserts can be welded or completely replaced if the crack length is long relative to the thickness. Similarly, the mold surface and features should be restored. This rework may require machining and welding one or more inserts with subsequent machining and surface refinishing. Alternatively, chrome and nickel plating allow the deposition of metal for repair of damaged or worn molding surfaces. All of the parting line, shut-offs, and vents should be rebuilt and checked with Prussian blue or another temporary ink as discussed in Section 14.2.4. The mold can then be reassembled, with cavities/cores/pins returned to their documented location. Even though all of these components have been restored to their original condition and location, it is recommended that the mold be formally recommissioned as described in Sections 14.2–14.4. The reason is that the mold repairs may change the behavior of the mold while fit and tolerance variations may cause small dimensional errors. Mold recommissioning should be a rapid process given the depth of documentation and prior molding experience. The molded part characteristics should be fully assessed and approved, with any needed mold changes, prior to the rebuilt mold entering the next round of mass production.

„„14.6 Summary This chapter provided practical guidance for mold commissioning and maintenance. The key concepts are: ƒƒIssues with mold design can cause reduced molding productivity and molded part quality. For each issue, the root cause should be identified and remedied prior to the mold entering production. The cost of the corrective action(s) should be paid by the party responsible for the end-user. ƒƒTwo sets of documentation are critical to develop during mold commissioning: 1. The mold design documentation should include a full description of the mold and its intended use. A report should describe the analysis and rationale for its design relative to the specifications stated in the mold purchase agreement. A mold manual should include a set of useful drawings for the molder as well as preliminary processing conditions that were used in the mold ­design. 2. The mold processing documentation should include a setup sheet with all ­required process conditions. If the mold is to be operated on different makes/ models/sizes of molding presses, then the molding process should be validated on each molding press with its own setup sheet. At the same time, a mold operating log should be developed to track the mold’s operation and

491

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performance. Any issues with the mold should be indicated on a mold map and corrected as appropriate. ƒƒMold commissioning often requires cooperation between the mold designer/ maker, molder, and end-user of the molded parts. Prior to shipping the mold to the molder, the mold maker should perform a detailed inspection of the mold. Prior to molding, the molder should similarly inspect the mold and perform dry cycles to verify its operations. ƒƒThe molder develops a process for appropriate filling, packing, cooling, and mold resetting. The mold designer/maker and end-user are frequently involved with this process. The mold designer/maker helps to assess and correct any mold-related issues. The end-user coordinates with the molder’s quality control team to assess and ultimately approve the molded part quality. ƒƒOnce a mold is commissioned with an approved process, the molder should not vary from the standard operating conditions. The molder should not operate the mold with a reduced number of cavities or otherwise try to correct for material or mold issues by changing the molding process conditions. Any change may adversely affect the molded part quality and should be approved by the end-user of the molded parts. ƒƒMany molders perform (and many end-users require) process capability assessment. The purpose is to ensure that the molding process provides consistent moldings that will meet product specifications in production. Gauge repeatability and reproducibility should be verified prior to process capability studies. ƒƒThe process capability, CP, is usually dominated by a few critical quality attributes that constrain the molding process window. A CP value of 1 is often sufficient, but many molders and end-users strive for a CP of 2 as per Six Sigma guidelines. Engineering change orders may be required to change the mold design or quality specification to achieve such high process capability values. ƒƒRoutine maintenance will help to sustain molding productivity, molded part quality, and the lifetime of the mold. Simple care prior to and after each molding run can ensure that the mold remains in good condition; the date, number of cycles, and any issues encountered during molding should be recorded. Molders should regularly schedule more extensive maintenance to address wear, cooling, damage, or other issues to keep the mold operating in a consistent manner. ƒƒMold rebuilding may be necessary in which every component in the mold is restored to an almost new condition. Major rework of the mold may change the mold’s behavior and molded part quality, so the mold should be formally recommissioned prior to re-entering mass production. Injection molds are complex systems, and their design and operation requires the experience and cooperation of the mold designer, mold maker, molder, and enduser. Mold design should not be performed in a vacuum; feedback from the mold’s

14.7 References

operation and maintenance will help to educate all parties for continuous improvement and the realization of more creative and competitive designs in the future.

„„14.7 References [1] Kulkarni, Suhas, Design of Experiments for Injection Molding, In Robust Process Development and Scientific Molding: Theory and Practice, Hanser Publishers (2010) [2] Lord, H. and G. Williams, Mold-filling studies for the injection molding of thermoplastic materials, Part II: The transient flow of plastic materials in the cavities of injection-molding dies, Polym. Eng. Sci. (1975) 15(8): pp. 569–582 [3] Bozzelli, J. W. Systematic molding for accurate comparisons: part to part, resin to resin, plant to plant, and lot to lot, in SPE Annu. Tech. Conf. (1991) [4] Bozzelli, J. W. and J. Cardinal, Process Control on Injection Molding Machines: What Is the Correct Hydraulic Response during Switchover from First to Second Stage?, SPE ANTEC (1996) pp. 584–587 [5] Chen, X., et al., Automatic velocity profile determination for uniform filling in injection molding, Polym. Eng. Sci. (2010) 50(7): p. 1358–1371 [6] Kazmer, D., V. Kudchadkar, and R. Nageri, Validation of moulding productivity with two self-­ regulating melt pressure valves, Plast., Rubber Compos. (2004) 33(9-10): pp. 446–451 [7] Kazmer, D. O., et al., A comparison of seven filling to packing switchover methods for injection molding, Polym. Eng. Sci. (2010) 50(10): pp. 2031–2043 [8] Berger, R. W., The Certified Quality Engineer Handbook, American Society for Quality (2002) [9] Barrentine, L. B., Concepts for R & R Studies, American Society for Quality (2003) [10] Kane, V. E., Process Capability Indices, J. Qual. Technol. (1986) 18(1): pp. 41–52 [11] Cuypers, W., et al., Optical measurement techniques for mobile and large-scale dimensional metro­ logy, Optics and Lasers in Engineering (2009) 47(3): pp. 292–300 [12] Weckenmann, A., et al., Multisensor data fusion in dimensional metrology, CIRP Annals-Manu­ facturing Technology (2009) 58(2): pp. 701–721 [13] Peggs, G., et al., Recent developments in large-scale dimensional metrology, Proc. Inst. Mech. Eng., Part B: J. Eng. Manuf. (2009) 223(6): pp. 571–595 [14] Li, A., et al., Laser coarse–fine coupling scanning method by steering double prisms, Appl. Optics (2012) 51(3): pp. 356–364 [15] Michaels, D. J. and R. H. Warner, U.S. Patent 8,861,673, Component aperture location using com­ puted tomography, Oct. 14 (2014) [16] Kruth, J.-P., et al., Computed tomography for dimensional metrology, CIRP Ann. Manuf. Technol. (2011) 60(2): p. 821-842 [17] Carmignato, S., Accuracy of industrial computed tomography measurements: experimental results from an international comparison. CIRP Ann. Manuf. Technol. (2012) 61(1): pp. 491–494 [18] Maropoulos, P. G. and D. Ceglarek, Design verification and validation in product lifecycle, CIRP Ann. Manuf. Technol. (2010) 59(2): pp. 740–759 [19] Huang, M.-F., Y.-R. Zhong, and Z.-G. Xu, Concurrent process tolerance design based on minimum product manufacturing cost and quality loss, Int. J. Adv. Manuf. Technol. (2005) 25(7-8): pp. 714–722 [20] Evans, J. R. and W. M. Lindsay, The management and control of quality, South-Western, Mason, OH (2005)

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[21] Garvin, D. A., What does product quality really mean, Sloan management review, (1984) 26(1) [22] Production Part Approval Process, Automotive Industry Action Group (2006) [23] Harry, M. J., The Nature of Six Sigma Quality, Motorola University Press (1988) [24] Snee, R. D., Six-Sigma: the evolution of 100 years of business improvement methodology, Int. J. Six Sigma and Compet. Advantage (2004) 1(1): pp. 4–20 [25] Zhu, L. and D. Kazmer, An Extensive Simplex Method Mapping the Global Feasibility, J. Eng. Optim. (2003) 35(2): pp. 165–176 [26] Kazmer, D., S. Westerdale, and D. Hazen, A Comparison of Statistical Process Control (SPC) and OnLine Multivariate Analyses (MVA) for Injection Molding, Int. Polym. Process. (2008) 23(5): p. 447–458 [27] Kazmer, D. and C. Roser, Evaluation of product and process design robustness, Res. Eng. Des. (1999) 11(1): pp. 20–30 [28] Kazmer, D., L. Zhu, and D. Hatch, Process Window Derivation With an Application to Optical Media Manufacturing, ASME J. Manuf. Sci. (2001) 123: pp. 303–314 [29] Menges, G., W. Michaeli, and P. Mohren, How to Make Injection Molds, Hanser, Munich (2001) pp. 527–533 [30] Dealey, R. W., Plastics Technician’s Toolbox, in Plastics Technician’s Toolbox (2004) pp. 31–45 [31] Johnson, S., Creating a Mold-Repair Plan, Plast. Technol. (2007)

Appendix

496

  Appendix

„„Appendix A: Plastic Material Properties Source: Autodesk Inc. Moldflow (Waltham, MA). Used with permission. Plastic material

ABS

Acetal

PA6

PA66

PA6633%GF

PC

PE, LD

Trade name

Cycolac

Delrin

Capron

Leona

Leona

Lexan

Lupolen

Grade

MG47

500

8202

1402

14G33

121

1800H

Supplier

Sabic

DuPont

BASF

Asahi

Asahi

Sabic

Basell

Description

Multi-purpose, injection molding ABS pro­ viding a favorable balance of engineering properties

Low friction, wear resistant grade with high viscosity for bushings and engineering applications

Non-impact modified, general purpose mold­ ing grade of Nylon 6

Unfilled Nylon 6/6 with low warpage, aging resistance, and fast cycle times

33% glass fiber re­inforced Nylon 6/6 for underhood automotive parts

Medium viscosity polycarbonate with good impact properties

Low density polyethylene with low modulus and high elon­ gation for injection molding

Approximate cost ($/kg)

2.52

3.09

3.20

3.64

4.18

4.08

2.34

Approximate cost ($/m3)

2340

3550

3107

3510

5058

4295

1755

Modulus (MPa)

2280

2800

970

1200

4400

900

110

Poisson’s ratio

0.35

Yield strength (MPa)

44

50

36

67

132

62

9

Strain to yield (%)

2

12

16

35

3

7

15

DTUL (°C, 0.45 MPa, ASTM D648)

96.7

160

160

240

250

138

41

No flow melt temperature (°C)

132

175

216

263

254

143

106

Minimum melt temperature (°C)

218

180

230

260

275

280

205

Maximum melt temperature (°C)

260

230

300

300

300

305

245

Minimum coolant ­temperature (°C)

49

50

70

40

60

70

20

Maximum coolant ­temperature (°C)

71

105

110

80

120

95

60

Appendix A: Plastic Material Properties

PE, HD

PET

PMMA

PP

PP30%GF

PS

PS30%GF

PVC, rigid

PVC, soft

Alathon

Eastapak

Plexiglas

Inspire

Nepol

Styron

Questra

Geon

Geon

5370

9921

V052

702

GB303HP

478

WA212

M3000

C9000

Equistar

Eastman

Atohaas

Dow

Borealis

Dow

Dow

PolyOne

PolyOne

High density polyethylene with good impact strength, crack resistance, and color

Unfilled polyester for packaging applications

A general purpose acrylic with good light transmission and high heat ­stability

Unfilled, high flow, high impact grade for consumer products and automotive applications

30 % long glass fiber reinforced poly­ propylene to provide high strength and impact resistance

High impact poly­ styrene resin typically used in toys, housewares, and appliances

Syndio­ tactic polystyrene with toughness and heat, moisture, and chemical resistance

A medium flow, rigid PVC with normal impact often used with potable water fittings

A general purpose, medium gloss, ­flexible PVC

2.03

2.03

2.87

2.10

2.54

2.65

3.11

2.47

2.97

1470

2355

3061

1637

2353

2537

3426

3040

3132

400

1400

2740

1740

7400

2110

7600

3200

570

28

55

70

21

150

23

92

51

9

7

4

2

6

2

1.4

1.2

1.6

350

72

68

98

80

160

87

240

73

55

142

74

215

176

170

121

248

77

130

230

270

220

200

220

180

290

190

170

260

300

260

240

280

220

330

220

210

18

16

50

20

20

40

50

20

40

30

40

90

50

60

60

80

40

60

497

498

  Appendix

Plastic material

ABS

Acetal

PA6

PA66

PA6633%GF

PC

PE, LD

Density at 20°C (kg/m3)

1044

1435

1153

1151

1426

1192

894

Density at melt temp (kg/m3)

930

1149

971

964

1210

1052

750

Specific heat (J/kg°C)

2340

2020

2630

1670

3100

1260

3180

Thermal conductivity (W/m°C)

0.19

0.23

0.28

0.19

0.31

0.25

0.23

Thermal ­diffusivity (m2/s)

8.73E–08

9.91E–08

1.10E–07

1.18E–07

8.26E–08

1.89E–07

9.64E–08

Thermal expansion (m/m°C)

8.82E–05

1.00E–04

7.80E–05

8.00E–05

5.60E–05

6.80E–05

2.30E–04

Minimum shrinkage (% m/m)

0.4%

1.4%

0.2%

0.8%

0.3%

0.4%

1.4%

Maximum shrinkage (% m/m)

0.8%

2.6%

1.7%

1.6%

1.0%

0.8%

3.0%

Parallel shrinkage (% m/m)

0.6%

2.1%

0.9%

0.9%

0.4%

0.6%

1.6%

Perpendicular shrinkage (% m/m)

0.6%

1.5%

0.9%

0.9%

0.9%

0.6%

1.6%

Maximum shear rate (1/s)

50000

40000

100000

60000

60000

40000

40000

Mid-range melt temperature (°C)

239

205

265

280

288

293

225

Power law viscosity (Pa sn)

3.50E+04

8.25E+04

8.62E+03

1.02E+03

1.19E+04

4.05E+04

1.10E+04

Power law index, n

0.256

0.197

0.355

0.680

0.431

0.325

0.295

WLF: n

0.247

0.122

0.191

0.679

0.400

0.211

0.291

WLF: τ* (Pa)

9.97E+04

4.43E+05

2.54E+05

3.81E+00

1.19E+05

6.97E+05

2.35E+04

WLF: D1 (Pa s)

1.93E+13

1.20E+11

1.43E+08

7.16E+19

1.08E+18

5.06E+12

2.64E+19

WLF: D2 (K)

373.15

263.15

323.15

323.15

323.15

417.15

233.15

WLF: D3 (K/Pa)

0

0

0

0

0

0

0

WLF: A1

31.4

24.1

18.025

44.24

43.215

31.41

44.335

WLF: A2 (K)

51.6

51.6

51.6

51.6

51.6

51.6

51.6

Appendix A: Plastic Material Properties

PE, HD

PET

PMMA

PP

PP30%GF

PS

PS30%GF

PVC, rigid

PVC, soft

952

1336

1186

929

1127

1036

1265

1350

1198

724

1160

1067

781

927

958

1101

1230

1056

2890

1980

2093

2890

1969

1820

2400

1630

1580

0.33

0.23

0.157

0.184

0.13

0.133

0.25

0.185

0.22

1.58E–07

1.00E–07

7.03E–08

8.15E–08

7.12E–08

7.63E–08

9.46E–08

9.23E–08

1.32E–07

1.50E–04

7.50E–05

7.00E–05

9.50E–05

4.40E–05

9.00E–05

3.00E–05

7.50E–05

2.40E–04

1.2%

0.5%

0.2%

1.2%

0.4%

0.3%

0.1%

0.2%

1.0%

2.6%

1.0%

0.6%

2.2%

0.9%

0.7%

1.1%

0.6%

3.0%

1.6%

0.6%

0.4%

1.5%

0.5%

0.5%

0.2%

0.4%

1.7%

1.6%

0.6%

0.4%

1.5%

0.8%

0.5%

0.7%

0.4%

1.7%

40000

50000

40000

100000

80000

40000

80000

32000

32000

245

285

240

220

250

200

310

205

190

1.09E+04

8.53E+04

3.68E+04

3.51E+03

9.52E+03

1.31E+04

8.72E+03

3.88E+04

1.36E+04

0.389

0.237

0.239

0.379

0.332

0.346

0.361

0.311

0.294

0.360

0.100

0.237

0.377

0.329

0.346

0.358

0.306

0.291

8.85E+04

8.46E+05

5.81E+04

3.90E+03

1.69E+04

1.03E+04

1.30E+04

9.20E+04

2.33E+04

3.75E+17

1.26E+16

3.99E+20

1.99E+14

1.19E+15

1.00E+12

2.20E+15

1.02E+19

2.42E+13

153.15

351

377.15

263.15

263.15

373.15

373.15

353.15

360.15

0

0

0

0

0

0

0

0

0

38.78

39.12

51.858

30.02

33.317

25.29

34.19

49.42

33.79

51.6

51.6

51.6

51.6

51.6

51.6

51.6

51.6

51.6

499

500

  Appendix

Plastic material

ABS

Acetal

PA6

PA66

PA6633%GF

PC

PE, LD

Apparent zero shear rate viscosity, η0 (Pa s)

1330

745

134

964

890

1060

1720

b1m (m3/kg)

9.83E–04

8.49E–04

1.01E–03

1.03E–03

8.12E–04

8.66E–04

1.23E–03

b2m (m3/kg K)

6.51E–07

4.21E–07

5.05E–07

6.94E–07

4.14E–07

5.67E–07

8.78E–07

b3m (Pa)

1.35E+08

1.33E+08

1.78E+08

1.37E+08

1.65E+08

1.69E+08

1.05E+08

b4m (1/K)

4.38E–03

3.37E–03

5.09E–03

3.51E–03

3.77E–03

4.23E–03

4.12E–03

b1s (m3/kg)

9.83E–04

7.33E–04

9.41E–04

9.80E–04

7.58E–04

8.65E–04

1.17E–03

b2s (m3/kg K)

3.47E–07

2.41E–07

3.83E–07

4.66E–07

2.49E–07

2.16E–07

6.60E–07

b3s (Pa)

1.70E+08

3.59E+08

2.47E+08

1.68E+08

2.98E+08

2.58E+08

1.71E+08

b4s (1/K)

4.21E–03

2.75E–03

3.67E–03

3.16E–03

2.57E–03

3.02E–03

2.42E–03

b5 (K)

370.6

448.15

489.16

535.66

527.15

416.23

379

b6 (K/Pa)

2.30E–07

2.00E–08

6.05E–08

6.24E–08

6.40E–08

3.34E–07

2.39E–07

b7 (m3/kg)

0

1.16E–04

6.50E–05

4.57E–05

5.41E–05

0

5.44E–05

b8 (1/K)

0

1.24E–01

5.20E–02

1.13E–01

3.32E–02

0

7.37E–02

b9 (1/Pa)

0

4.65E–09

4.73E–09

7.96E–09

4.71E–09

0

2.28E–08

Material prices are as of November, 2015 from Plastics News (Detroit, MI) for ­representative generic grades in annual volumes of 1 million kg.

Appendix A: Plastic Material Properties

PE, HD

PET

PMMA

PP

PP30%GF

PS

PS30%GF

PVC, rigid

PVC, soft

660

370

7400

680

1010

6880

3060

5720

4090

1.27E–03

7.54E–04

8.61E–04

1.24E–03

1.02E–03

9.97E–04

8.78E–04

7.51E–04

9.02E–04

1.03E–06

5.104E–07

6.00E–07

8.72E–07

7.209E–07

5.98E–07

4.98E–07

4.95E–07

7.72E–07

9.26E+07

2.46E+08

2.07E+08

8.05E+07

8.69E+07

1.56E+08

7.29E+07

2.19E+08

1.15E+08

4.94E–03

3.80E–03

5.45E–03

4.93E–03

5.45E–03

4.58E–03

1.94E–03

5.05E–03

5.36E–03

1.08E–03

7.53E–04

8.62E–04

1.17E–03

9.33E–04

9.94E–04

8.17E–04

7.51E–04

9.01E–04

2.08E–07

1.00E–07

1.00E–07

6.14E–07

3.10E–07

2.96E–07

1.18E–07

2.32E–07

6.28E–07 1.15E+08

3.32E+08

3.77E+08

2.77E+08

1.10E+08

1.50E+08

1.92E+08

2.09E+08

2.67E+08

2.46E–06

5.87E–04

4.23E–03

5.12E–03

5.93E–03

4.96E–03

1.00E–04

4.96E–03

5.03E–03

414.5

346.88

386.15

449.15

443.15

394.25

521

350.15

403.15

2.39E–07

1.75E–07

2.00E–07

6.25E–08

6.90E–08

8.10E–08

2.90E–07

1.33E–07

6.56E–08

1.87E–04

0

0

7.44E–05

8.15E–05

0

6.14E–05

0

0

5.16E–02

0

0

1.02E–01

1.42E–01

0

1.90E–02

0

0

1.02E–08

0

0

8.67E–09

1.18E–08

0

1.22E–08

0

0

501

502

  Appendix

„„Appendix B: Mold Material Properties B.1 Nonferrous Metals Mold material

Al 7075-T6

Al QC-10 Alloy

Cu 17200

Description

Aircraft grade aluminum alloy with high strength and corrosion resistance

Aluminum alloy developed for molds with higher strength, hardness, and conductivity

Beryllium-free copper alloy with high strength and thermal conductivity

26.9

20.7

61.2

Cost ($/cm )

0.076

0.058

0.531

Ultimate strength (MPa)

565

579

689

Modulus (MPa)

71,000

70,000

120,000

Yield stress (MPa)

421

525

517

Fatigue limit stress (MPa)

149

166

290

Hardness, Brinell (HB)

150

160

210

Feed per tooth (mm)

0.0762

0.0762

0.0762

Cutting speed (m/h)

23,600

23,600

3,600

Volume machine rate (m3/h)

0.0091

0.0091

0.0014

Area machine rate (m2/h)

0.225

0.225

0.034

Thermal expansion (°m/m°C)

24

24.7

18

Thermal conductivity (W/m°C)

130

160

259

Cost ($/kg) 3

Specific heat (J/kg°C)

960

879

506

Density (kg/m3)

2,810

2,850

8,690

Thermal diffusivity (m2/s)

4.82E–05

6.39E–05

5.89E–05

Appendix B: Mold Material Properties

B.2 Common Mold Steels Mold material

1045

4140

P20

Description

High strength carbon steel, low cost but poor corrosion and wear resistance

Chrome alloyed steel with good fatigue, abrasion, and impact resistance

Common mold steel with good fatigue, abrasion, and impact resistance

Cost ($/kg)

4.8

10.3

11.6

3

Cost ($/m )

0.037

0.081

0.091

Ultimate strength (MPa)

752

778

965

Modulus (MPa)

207,000

200,000

205,000

Yield stress (MPa)

647

669

830

Fatigue limit stress (MPa)

291

412

456

Hardness, Brinell

225

259

300

Feed per tooth (mm)

0.0762

0.0508

0.0508

Cutting speed (m/h)

5600

4700

3800

Volume machine rate (m3/h)

0.0021

0.0012

0.001

Area machine rate (m2/h)

0.053

0.03

0.024

Thermal expansion (°m/m°C)

12.2

12.2

12.8

Thermal conductivity (W/m°C)

49.8

42.7

32

Specific heat (J/kg°C)

515

523

500

Density (kg/m3)

7850

7850

7820

Thermal diffusivity (m2/s)

1.23E–05

1.04E–05

8.18E–06

503

504

  Appendix

B.3 Other Mold Steels Mold ­material

A6

D2

H13

S7

SS420

Description

Heat treatable to be very hard with good wear resistance and fatigue life

High carbon/ chrome steel for wear and abrasion resistance

Heavily alloyed, hard steel with excellent ­temperature and wear resistance

Excellent toughness and high strength but lower wear resistance

Excellent polishability and corrosion resistance with good hardness

Cost ($/kg)

20.6

19.7

46.7

25.5

63.6

Cost ($/m ) 3

0.165

0.151

0.364

0.199

0.496

Ultimate 2380 strength (MPa)

2200

1990

1620

655

Modulus (MPa)

203,000

210,000

210,000

207,000

207,000

Yield stress (MPa)

2100

1929

1650

1380

345

Fatigue limit stress (MPa)

834

755

760

528

190

Hardness, ­Brinell

650

685

528

369

195

Feed per tooth (mm)

0.0508

0.0508

0.0508

0.0508

0.0508

Cutting speed (m/h)

2900

2700

700

3900

5000

Volume machine rate (m3/h)

0.0007

0.0007

0.0002

0.001

0.0013

Area machine rate (m2/h)

0.018

0.017

0.004

0.025

0.032

Thermal expansion (°m/m°C)

11.8

11.8

11.5

12.1

10.8

Thermal ­conductivity (W/m°C)

27

21

24.3

29

24.9

Specific heat (J/kg°C)

460

460

460

460

460

Density (kg/m3)

8030

7670

7800

7810

7800

Thermal diffusivity (m2/s)

7.31E–06

5.95E–06

6.77E–06

8.07E–06

6.94E–06

Appendix C: Properties of Coolants

ƒƒCost data was produced from commodity pricing for rectangular plates approximately 30 cm × 30 cm × 4 cm. ƒƒThe fatigue endurance stress was analyzed using empirically fit S-N coefficients at 1,000,000 cycles with a safety factor of 1.0. ƒƒThe volumetric removal rate assumes a carbide, two fluted, ¾ inch diameter end mill with a depth of cut of 0.125 inches. The surface area removal rate assumes a carbide, four fluted, ¼ inch diameter end mill operating at half the nominal feed rate. ƒƒThermal properties were evaluated as the average of room temperature and 200°C if data was available, and at room temperature otherwise.

„„Appendix C: Properties of Coolants Coolant material

Water

Ethylene glycol

Oil

Formulation

100 % H2O

100 % C2H6O2

(CH4)n

Description

Typical plant water, possibly contaminated with corroded metals

Undiluted ethylene ­glycol with corrosion inhibitors

ISO grade 32 oil, a lower viscosity oil appropriate for ­circulating systems

Cost ($/L)

0.0

7.0

3.0

Lower use temperature 1 (°C)

–56

32

Upper use temperature 100 (°C)

134

288

Density (kg/m3)

800

900

1000

Specific heat (J/kg°C)

4187

2261

1842

Thermal conductivity (W/m°C)

0.6

0.18

0.16

Thermal diffusivity (m2/s)

1.43E–07

9.95E–08

9.65E–08

Viscosity (Pa s, at 50 °C)

0.0010

4.8

23.5

Viscosity (Pa s, at 100 °C)

0.0002

3.4

4.6

505

506

  Appendix

„„Appendix D: Statistical Labor Data D.1 United States Occupational Labor Rates The average wages for various occupations related to mold design, mold making, and molding are listed in Table D1. These data are from the U.S. Department of Labor, Bureau of Labor Statistics’ Occupational Employment Survey (OES) as of May, 2014. These data do not include the cost of benefits (estimated in Appendix D.2 by country), other indirect costs, or profit. Table D.1 United States Wage Data Position

Average wage (USD)

Mechanical engineer

41.89

Tool and die maker

26.61

Mold assembler/coremaker

20.65

Machinist

19.97

Tool and die maker apprentice

18.07

Lathe setup operator

17.96

CNC programmer

24.13

Milling machine operator

16.82

Lathe operator

15.88

Drilling operator

14.21

Machinist apprentice

16.55

Buffing and polishing setter

14.41

Molding machine operator

14.71

D.2 International Labor Rate Comparison The average manufacturing costs rates for different countries are listed in Table D.2 with average hourly compensation listed in U.S. dollars. Source: U.S. Department of Labor Statistics’ International Labor Comparison dated August, 2013. The occupational labor rates listed in Table D.1 can be proportioned by the international average manufacturing rates to estimate occupational labor rates internationally.

Appendix D: Statistical Labor Data

Table D.2 International Manufacturing Cost Data Country

Employee cost

Paid benefits

Worker hourly pay

Norway

63.36

11.33

52.03

Switzerland

57.79

19.60

38.19

Belgium

52.19

27.33

24.86

Sweden

49.80

21.77

28.03

Denmark

48.47

11.93

36.54

Australia

47.68

13.94

33.73

Germany

45.79

19.34

26.45

Finland

42.60

17.57

25.03

Austria

41.53

19.57

21.96

France

39.81

19.30

20.51

Netherlands

39.62

16.96

22.66

Ireland

38.17

12.39

25.78

Canada

36.59

10.92

25.68

United States

35.67

11.74

23.93

Japan

35.34

15.22

20.11

Italy

34.18

14.87

19.31

United Kingdom

31.23

8.96

22.27

Spain

26.83

12.17

14.67

New Zealand

24.77

3.90

20.87

Singapore

24.16

8.46

15.71

Korea, Republic of

20.72

4.45

16.27

Israel

20.14

4.92

15.22

Greece

19.41

8.12

11.29

Argentina

18.87

5.93

12.94

Portugal

12.10

4.71

7.39

Czech Republic

11.95

4.72

7.23

Slovakia

11.30

5.17

6.12

Brazil

11.20

5.25

5.95

Estonia

10.41

3.66

6.75

Taiwan

9.46

1.37

8.08

Hungary

8.95

3.87

5.08

Poland

8.25

3.16

5.09

Mexico

6.36

1.92

4.45

India

2.85

0.92

1.93

China

2.52

0.77

1.75

Philippines

2.24

0.54

1.70

507

508

  Appendix

„„Appendix E: Unit Conversions This book uses the following system of units: ƒƒTemperature: degree Celsius, °C ƒƒLength: meter, m ƒƒMass: kilogram, kg ƒƒForce: Newton, N ƒƒPressure: Pascal, Pa ƒƒFlow Rate: cubic meters per second, m3/s ƒƒViscosity: Pascal seconds, Pa s ƒƒEnergy: Joules, J Conversions from each of these units to other common systems are next provided to four significant digits. To convert from degree Celsius to degree Fahrenheit, multiply the temperature by 1.8 and add 32: T (°F) = 1.8 × T (°C) + 32 (E.1) T (°K) = T (°C) + 273.1

E.1 Length Conversions Table E.1 Length Conversion Factors To convert from

to

Multiply by

Meter, m

Millimeter, mm

1000

Meter, m

Centimeter, cm

100

Meter, m

Micrometer, µm

1,000,000

Meter, m

Inch, in

39.37

Meter, m

Feet, ft

3.281

Appendix E: Unit Conversions

E.2 Mass/Force Conversions Table E.2 Mass/force Conversion Factors To convert from

to

Multiply by

Kilogram, kg

Newton weight, N

9.807

Kilogram, kg

Gram, m

1000

Kilogram, kg

Pound force, lbf

2.205

Kilogram, kg

Ounce, oz

35.27

Kilogram, kg

Metric ton, t

0.001

Kilogram, kg

U.S. short ton, t (short)

0.0011023

Kilogram, kg

U.K. long ton, t (long)

0.0009842

E.3 Pressure Conversions Table E.3 Pressure Conversion Factors To convert from

to

Multiply by 2

Megapascal, MPa

Dyne per sq. centimeter, dyn/cm

10,000,000

Megapascal, MPa

Pascal, Pa

1,000,000

Megapascal, MPa

Kilopascal, kPa

1000

Megapascal, MPa

Pounds per sq. inch, lb/in2

145.04

Megapascal, MPa

Bar, bar

10

Megapascal, MPa

Standard atmosphere, atm

9.86

E.4 Flow Rate Conversions Table E.4 Flow Rate Conversion Factors To convert from

to

Multiply by

Cubic meter per second, m3/s

Cubic centimeter per second, cc/s

1,000,000

Cubic meter per second, m /s

Liter per minute, L/min

60,000

Cubic meter per second, m3/s

Gallon per minute, gal/min

15,840

Gallon per hour, gal/h

950,400

3

3

Cubic meter per second, m /s

509

510

  Appendix

E.5 Viscosity Conversions Table E.5 Viscosity Conversion Factors To convert from

to

Multiply by

Pascal second, Pa s

Poise, P

10

Pascal second, Pa s

Centipoise, cP

1000

Pascal second, Pa s

Gram per centimeter second, g/(cm s)

10

E.6 Energy Conversions Table E.6 Energy Conversion Factors To convert from

to

Multiply by

Joule, J

Watt second, W s

1

Joule, J

Newton meter, N m

1

Joule, J

Foot pound, ft lbf

0.7376

Joule, J

British thermal unit, Btu

0.000948452

Joule, J

Kilowatt hour, kW h

0.000000278

Joule, J

Ton hours of refrigeration, ton h

0.000000079

Appendix F: Estimation of Melt Velocity

„„Appendix F: Estimation of Melt Velocity The following derivation provides an estimate of the linear melt velocity required to balance the heat lost from the melt to the mold with the heat internally generated due to shear heating. As such, the plastic melt should maintain a uniform melt temperature throughout filling if the suggested velocity is maintained. A power law viscosity model is assumed in which the viscosity is modeled as: h = k g n-1 (F.1)

where I is the “consistency index” indicative of the viscosity at a shear rate of 1 s-1, and n is the power law index. For this viscosity model, the velocity, v, as a function of thickness, z, is: 1ü ì æ 1 + 1 n ö÷ïïïï æ 2 z ö1+ n ïïï ç ÷ ç ÷í1- ç ÷ ý v (F.2) v ( z ) = 2çç çè 2 + 1 n ø÷÷ïï èç H ø÷ ïï ïîï ïþï where v is the average linear (bulk) melt velocity that we seek to define. The shear

rate as a function of the thickness is: 1

æ 1 öæ 2 z ön v (F.3) g ( z ) = 2çç2 + ÷÷çç ÷÷ çè n ÷øèç H ÷ø H

The one-dimensional, steady heat equation with internal shear heating is: d 2T q ( z ) + = 0 (F.4) k dz 2

The internal heating, q(z), due to viscous dissipation is: 2

q ( z ) = h (g ( z )) g ( z ) (F.5)

Substituting the above terms for the viscosity and the shear rate, the shear heating energy, q, as a function of the thickness is:

2+

2 q(z) =

2 æ n k ç2 +

ç èç

1 ö÷-1+n 2 2/ n æ 1 ö n ÷ 1 ö÷ 2 æç z ö÷ ççç 1+ n æç 1 öæ z ÷ v ç ÷÷ ç2 ç2 + ÷÷÷çç ÷÷÷ v H ÷÷÷ ÷ ç ç ç ç èHø ç è nø n øè H ø ÷÷ø çè (F.6) H2

511

512

  Appendix

Substituting this shear heating term into the energy equation and integrating once provides: 1 æ 1 ö÷1+n çç 1+ æ öæ ö n ÷ 1 z k (1 + 2n) z çç2 n çç2 + ÷÷çç ÷÷ v H ÷÷÷ ÷ç H ÷ø çç ÷ èç n øè dT ( z ) èç ø÷ =+ C1 (F.7) æ ö dz çç2 + 1 ÷÷(1 + 2n)k èç n ÷ø

where C1 is a constant of integration. To find this constant, the temperature profile through the thickness is assumed to be symmetric with respect to the centerline so that dT/dz equals 0. The coefficient C1 is then found to be 0. Integrating Eq. F.7 again, 1 æ 1 ö÷1+n çç 1+ æ öæ ö n ÷ 1 z kz 2 çç2 n çç2 + ÷÷çç ÷÷ v H ÷÷÷ ÷ ÷ çè ç çç ÷ n øè H ø èç ø÷ + C2 (F.8) T (z) = æ öæ ö çç2 + 1 ÷÷çç3 + 1 ÷÷k èç n ÷øèç n ÷ø

To find C2 , the temperature at the mold wall is assumed constant. Setting T(z=H/2)=Twall . The coefficient C2 is found to be:

C2 = Twall +

ææ ö1+n çççç2 + 1 ö÷÷ v ÷÷ çèç n ÷ø ÷÷÷ 2-1+n H 2k ççç ÷ çç H ÷÷÷ ÷÷ çç è ø æ öæ ö çç2 + 1 ÷÷çç3 + 1 ÷÷k ÷ èç n øèç n ø÷

(F.9)

Accordingly, the temperature profile through the thickness is

-1+n

2 T ( z ) = Twall +

1 æ 1 ö÷1+n çç 1+ æ ææ ö÷1+n ö ö æ ö n ÷ 1 1 z 2 H k ççççç2 + ÷÷÷ v H ÷÷ - kz çç2 n çç2 + ÷÷÷ v çç ÷÷÷ H ÷÷÷ ç ç ç ÷ ç ÷ è ø è ø è ø n n H è ø èçç ø÷ (F.10) æ öæ ö çç2 + 1 ÷÷çç3 + 1 ÷÷k èç n ÷øèç n ÷ø 2

The goal in the selection of the linear melt velocity is to force the bulk melt temperature during filling to equal the melt temperature. The average temperature can be

Appendix F: Estimation of Melt Velocity

evaluated by integrating the melt temperature profile from the centerline to the mold wall and dividing by the half thickness:

ò T= 0

H 2

ò0

T ( z ) v ( z ) dz

H 2

v ( z ) dz

(F.11)

which evaluates to:

T = Twall +

æ1 + 2n ö÷n 2n Hkn (1 + 3n) v çç v÷ çè Hn ÷ø

(2 +13n + 20n2 )k

(F.12)

Setting the average temperature equal to the melt temperature and solving for the linear melt velocity provides: 1

-1+ -1/ n ö æ ÷ 1+n ç(1 + 2n) çæ 2-n (1 + 4n)(2 + 5n)(Tmelt - Twall )k ÷ö ÷ ç ÷÷ ÷÷ ç v = çç (F.13) ÷÷ ÷÷ çç Hn çççè Hkn (1 + 3n) ø è ø÷

For a Newtonian material with n equal to 1 and the consistency index, k, equal to some apparent viscosity, m, the linear melt velocity simplifies to: v=

35(Tmelt - Twall )k 24m

(F.14)

It is a good idea to verify the solution by checking the units. For the Newtonian material: 1

1

1

æ ö2 æ é W ù ö÷2 æ é N × m ù ö÷2 1 çç( éC ù ) éê W ùú ÷÷ çç ê ú ÷ çç ê ú÷ æ é m2 ù ö÷2 é m ù çç ë û ê mC ú ÷÷ çç ê m ú ÷÷ çç ê s × m ú ÷÷ çê ë û ÷÷ = ç ë û ÷÷ = ç ë û÷ ú÷ v = çç çç é Pa × s ù ÷÷ çç é N ù ÷÷÷ = ççç ê 2 ú÷÷÷ = êê s úú (F.15) çç ëé Pa × s ûù ÷÷÷ s ûú ø è ê ë û ÷ ÷ ç ç ë û ë ê ú ÷ ÷ çç çç çç ê 2 s ú ÷÷ ÷ø ÷ø è è è ëm û ø

It may appear that the velocity is independent of the wall thickness, but the role of wall thickness enters through the effect of the shear rate. When evaluating Eq. F.14 using a Newtonian model, an iterative approach is needed to ensure that the ­apparent shear rate and viscosity match the predicted velocity. See Section 5.5.1 for usage guidelines with an example.

513

514

  Appendix

Many polymers have a power law index, n, around 1/3. For this case, the optimal velocity is: 3

v = 3.81

4

(Tmelt - Twall ) H 2k 3

k3

(F.16)

It is observed that the velocity is more sensitive to the difference between the melt and mold wall temperature, thermal conductivity, and consistency index since these terms are raised to the ¾ power in Eq. F.16 rather than the ½ power in Eq. F.14. This behavior is further exaggerated for shear sensitive materials as the power law index, n, approaches 0 with the exponent of (∆T × κ/k) approaching 1.

The Author

David Kazmer (Ph. D. Mechanical Engineering, Stanford University) is the Dandeneau Professor for Sustainable Manufacturing in the Department of Plastics Engineering at the University of Massachusetts Lowell. He performs research and teaches courses related to plastics product and process development. Prior to his current appointment, he was an Applications Engineer at General ­Electric and Director of Research and Development at Synventive HotRunners. He is a licensed professional manufacturing engineer, and is a fellow of the American Society of Mechanical Engineers and the Society of Plastics Engineers. He is also the recipient of the ASME Kos Ishii-Toshiba Award for sustained, meritorious ­contributions to the field of design for manufacturing.

Index

A

B

acceptable quality levels  461, 476, 481 acceptance sampling  476 actuation  466 actuation force  363 additional draft  38 aesthetic defect  111 aesthetics  29, 439 aesthetic surface  377 air channel  277 allowance  89 aluminum  261 aluminum 6061-T6  386 aluminum tooling  72 amorphous  313 amortized cost  46 angle pins  366, 371, 489 anisotropic shrinkage  301 anisotropy  314 annealing  103 anodizing  104 A plate  6, 94, 149 apparent shear rate  119 AQLs, See  acceptable quality levels artificial balancing  159 automatic de-gating  13, 197, 209 automatic molding  374 Automotive Industry Action Group (AIAG)  476 auxiliaries  50 auxiliary equipment  14 auxiliary systems  72 avoid uneven filling  111 axial compression –– of cores  410 axial mold opening direction  82

baffles  274 banana gate  209 barrel temperature  312 beam bending  394 bending  397 BHN, See  Brinell Hardness Number bill of materials  463, 464 blush  52 bolt strength –– ultimate stress  422 BOM, See  bill of materials bore diameter  365 boss  34, 355 boss design  34 B plate  6, 94 branched runners  188 breakeven analysis  69 Brinell Hardness Number  20, 101 bronze gib  366 bubbler  275, 373 buckling  320, 323, 350 buckling constraint  352 bulk temperature  123 burn marks  52, 125, 135, 227 business development  23 C CAD, computer aided design  18 cam  372, 374 carbon black  314 carburizing  104 case hardening  103 cashew gate  209 cavities –– shutting off during molding  486 cavity complexity  59 cavity cost estimation  55

518

Index

cavity discount factor  62 cavity filling analysis  109 cavity finishing cost  63 cavity insert  79, 80, 87 cavity insert retainer plate  6 cavity layout  91 cavity machining cost  58 cavity materials cost  56 cavity retainer plate  372 cavity set cost  56 chamfer  348 chamfers  35 changeover, See  switchover change-over times  14 checklist –– for mold design inspection  463 –– for mold layout design  106 cheek  88, 372, 402 circular layout  92 clamp force  186 clamp tonnage  50, 79, 97, 110, 130 class  72 Class 101 mold  19, 99 Class 103 mold  20 clearance  356 closed loop control  190 coefficient of friction (COF)  104 coefficient of linear thermal expansion  296 coefficient of thermal expansion  291, 314 coefficient of volumetric thermal expansion  295 coinjection  431 coinjection mold design  433 coinjection molding  431 cold runner  49, 50, 142, 144, 180, 185 collapsible cores  373 color change  14, 71, 175, 211 color matching  29 color streaking  211 common defects  52 complexity factor  58 compressibility  291, 294, 296 compression  397 compression molding  437 compression spring  368 compressive stress  340, 389, 421 –– on cores  411 computed tomography system  481 computer aided design  18 computer simulation  317, 322, 472 concurrent engineering  17 conduction heating  284

conductive inserts  270 conductive pin  276 conformal cooling  269 constraints  426 contamination  52 continuous improvement  493 contoured ejector pins  347 convective boundary  251 coolant  11 coolant flow rate  253 coolant manifolds  266 coolant temperature  311 cooling  297 –– air channel  277 –– baffle  274 –– complexity  245 –– conductive pin  276 –– cooling line depth  257 –– cooling line pitch  260 –– cooling line routing  262 –– cooling power  252 –– cooling time estimate  250 –– heat pipe  275 –– heat transfer  243 –– heat transfer coefficient  250 –– insulating layer  280 –– internal manifold  267 –– minimum time  248 –– mold-making cost  244 –– parallel setup  267 –– post-mold  292 –– reliability  245 –– required coolant flow rate  253 –– series setup  266 –– shrinkage  292 –– system design  243, 246, 266 –– temperature distribution  265 –– temperature gradient  271 –– turbulent flow  255 –– wall temperature  244 cooling circuit  266 cooling insert  269, 273 cooling line  11, 87 –– layout  91 –– maintenance  490 –– networks  266 cooling plugs  256 cooling stage  3 cooling system  49, 65, 243 cooling system cost  66 cooling time  3, 11, 173, 243, 251, 284, 311, 447, 475 –– estimate  246

Index 519

copper  261 core  11, 410 –– minimum wall thickness  412 –– slender  414 core back  447 core-back molding  449 core bending  414 core deflection  414 core height  415 core insert  79, 80, 87 core insert retainer plate  6 core inserts –– with stripper plate  357 core pull  330, 361 –– actuators  330 corner design  34 corrosion –– in cooling lines  490 cost drivers  45 cost estimates  22 cost plus  43, 461 CP, See  process capability index CPK, See  process capability index cracks  385, 408 –– in molds  491 critical milestones  27 critical stress  116 Cross-WLF model  115 CTE, See coefficient of thermal expansion Cu 940  270 cycle efficiency  50 cycle efficiency factor  50, 51 cycle time  3, 16, 27, 49, 243, 284, 446, 450 –– reduction  473, 475 cycle time estimate  251 cyclic stresses  385 D daylight  13, 96, 153 dead pockets  231, 239 deep cores  273, 410 defect –– race-tracking  135 defect cost per part  52 defects –– burn marks  125 –– flash  125 –– hesitation  111 –– jetting  111 –– short shot  111, 138 –– warpage  111 defects per million opportunities  481, 482

deflection  372, 389, 394 –– side walls  403 deflection temperature under load  248 degradation  97 delivery terms  44 density  246, 314 design changes  323 design for assembly  23, 30 design for injection molding  31 design for manufacturing  23, 30 design for manufacturing and assembly  291 design iterations  17 design of experiments  306, 484 design requirements  25 design standards  28 detailed design  23 development time  16 diaphragm  205 diaphragm gate  206 dieseling  227 die set for mold stack height  466 differential shrinkage  31, 243, 265, 318 dimensional adjustments  89 dimensional metrology  479 –– computer tomography (CT)  480 –– coordinate measurement machines  480 –– optical image recognition  480 dimensions  28 direct metal laser sintering  269 discount factor  56 dispute –– during mold commissioning  460 DMLS, See direct metal laser sintering documentation –– of mold design  463 DOE, See  design of experiments double domain  294 dowels  417, 424 DPMO, See  defects per million opportunities draft angle  38 drawings –– layout of subsystems  463 –– of mold design  463 drive-interference fit  419 drops  153 dry cycle  466, 470, 487 DTUL, See deflection temperature under load dynamic melt control  190 E early ejector return  369 ECOs, See  engineering change orders

520

Index

edge gate  202 EDM, See  electric discharge machining effective area  336, 337 efficiency  50 ejection  38, 292 –– coefficient of friction  335 –– internal stresses  334 –– molding machine setup  330 –– normal force  334 –– part removal system  331 –– surface roughness  335 ejection force  334, 345 –– hoop stress  337 –– pin-to-pin variations  352 –– undercuts  359 ejection forces –– unbalanced  360 ejection stage  330 ejection system  327, 371 –– cooling interference  332 –– cost  333 –– ejection forces  330 –– mold opening  330 –– part aesthetics  332 –– part distortion  331 –– positive return  371 –– speed  331 ejection temperature  248, 336 ejector –– layout  345 ejector assembly  327 ejector blade  353 –– buckling  354 ejector housing  6, 95, 377 ejector knock-out rod  328, 369, 396 ejector locations  110 ejector pad  346 ejector pin  181, 230, 238, 327, 350, 396 –– clearance  238 –– contoured  39 –– stepped  352 ejector plate  11, 146, 327, 369, 377 ejector retainer plate  327, 349 ejectors –– alignment  349 –– buckling  351 –– clearance  349 –– compressive stresses  341 –– detailing  348 –– interference  347 –– number  343 –– placement  345 –– push area  341

–– push pin  342 –– shear stress  342 –– size  343 –– sliding bearing  348 –– stripper plate  356 –– total required perimeter  342 ejector sleeve  346, 355 ejector system  65 –– design process  333 –– design strategies  343 ejector system cost  66 ejector travel  94, 95 elastic deformation  359 elastic limit  359 elastic modulus  383 electrical connectors  466 electric discharge machining  407 encapsulated  437 endurance limit  100 endurance stress  258, 385, 505 energy efficiency  20 engineering change orders  463, 467, 484 ethylene glycol  257 excessive deflection  426 external undercuts  371 F factor of safety  384 family mold  9, 248 fan gate  204 fasteners  417 fatal flaws  17 fatigue  100, 385, 413, 426 –– in cores  413 FDM, See fused depositon modeling feed system  48, 141, 185 –– artificially balanced  145, 170 –– branched layout  160 –– comparison  15 –– cooling time  173 –– cost  66 –– cross-sections  176 –– custom layout  162 –– diaphragm  160 –– dynamic feed control  191 –– fill times  172 –– hybrid layout  161 –– imbalances in naturally balanced  188 –– insulated runner  185 –– layouts  159 –– maximum pressure drop  143 –– maximum volume  143

Index 521

–– naturally balanced  160, 170 –– number of turns  175 –– objectives  156 –– optimization  166 –– pressure drop specification  167 –– primary runners  147 –– radial layout  160 –– residence time  175 –– secondary runners  147 –– self-regulating valves  192 –– stack mold  186 –– standard runner diameters  183 –– steel safe design  184 –– sub-runners  216 –– tertiary runners  147 –– volume  165 –– waste factor  49 fidelity –– of quality measurements  479 fillers –– carbon black  314 –– glass bead  314 –– glass fiber  314 –– mica  314 –– rubber  314 fillet  35 filling  297 –– complete cavity  110 filling patterns  133 filling pressure  128 filling profile –– of injection velocity  122, 467, 472 filling stage  2, 10 filling time  3, 109, 125, 469 finger  435 finishing method  36 finishing rates  63 finishing time  63 first article inspection  17, 463, 476 fit  417 fit for purpose  1 fits  417 –– apparent diameter  418 –– clearance  417 –– insertion force  420 –– interference  418 –– locational-clearance  425 –– locational-interference  425 –– locational-transitional  425 –– retention force  419 –– unilateral hole basis  418 –– using dowels  424 fixed core pin  355

flash  52, 83, 125, 228, 475 flash gate  205 flashing  387 flow channel  127, 436 flow leaders  135, 138, 416 flow length  213 flow rate  142 fluid assist  431 fluid assisted molding  434 foam  432 freeze-off  11 fully automatic  50, 182 fully automatic molding  13, See also injection ­molding fused deposition modeling  72, 269 G gantry robots  331 gas assist  431 gas assist molding  434 gas trap  134, 145 gas traps  229, 239 gate  11, 146, 197 –– ring  416 gate freeze time  221, 223, 469, 473 gate types  216 gate well  201 gating –– automatic de-gating  197 –– comparison  216 –– design recommendations  218 –– diaphragm gate  205 –– direct sprue  200 –– edge gate  202 –– fan gate  204 –– film gate  205 –– fine-tuning  224 –– flash gate  205 –– freeze time  222 –– gating location  213 –– no-flow temperature  222 –– objectives  197 –– pack time  199 –– pin-point gate  201 –– pressure drops  219 –– shear rates  198, 217 –– submarine gate  209 –– tab gate  203 –– thermal gate  209 –– thermal sprue gate  211 –– tunnel gate  206 –– vestige  198

522

Index

gating design  197, 213 gating flexibility  12, 14, 16 gating location  109, 127 gauge repeatability and reproducibility  309, 477, 478 gauge R & R, See  gauge repeatability and ­reproducibility geometric complexity  58 geometric distortion  31 gibs  371 glass bead  314 glass fiber  314 glass filled  301 gloss  281, 285, 286 gloss level  29 grid layout  92 guides –– for ejector blades  354 gusset  34 H H13 steel  102 Hagen-Poiseuille  163, 257 hardness  101 HDT  248 hear stresses  114 heat conduction  246 heat content –– of moldings  252 heat deflection or distortion temperature  248 heater resistance  466 heat flux  260 heating element  439 heat load  284 heat pipes  275 heat transfer  11 –– insulating layer  287 heat transfer coefficient  250 heel block  368 height allowance  88 height dimension  87 helix  375 hesitation  111 hoop stress –– in cores  412 hot runner  14, 46, 64, 142, 153, 185, 209 –– color change  144 –– maintenance  489, 490 –– residence time  144 –– turn-over  144 hot runner mold  14, 16

hot runner system –– configurations  155 –– H manifold  155 –– stacked manifolds  155 –– straight-bar  155 –– X manifold  155 hot spots  270 hot sprue bushing  14, 153 hourly rate  49 hourly wage  62 hybrid layout  93 hydraulic actuators  364 hydraulic diameter  177 I improper color match  52 increased molding productivity  14 indexing head  445 indirect costs  45 induction heating  285 initial investment  16 injection blow molding  443 injection blow molds  443 injection compression  373, 432, 435 injection decompression  435 injection mold  4, 6 injection molding –– cooling stage  475 –– filling stage  471 –– fully automatic molding  487 –– packing stage  472 –– process capability  481 –– semiautomatic mode  487 injection molding process  1, 2, 16 injection molding process timings  3 injection pressure  97 –– maximum  110, 471 injection velocity  471 injection velocity profiling  471 ink –– after mold rebuilding  491 –– to check fits  466 in-mold film –– indexed  455 –– statically charged  454 in-mold labeling  453 in-mold sensors  306 insert creation  87 insertion force  417 insert mold  437, 439 insert sizing guidelines  87 inspections  47

Index 523

insulated runner  185 intellectual property  45 interlock  362 interlocking  415 interlocking core  277 interlocking features  85 internal corners  271 internal thread  41 internal threads  374 internal voids  33 isothermal boundary  249 isotropic  300 iterative mold development  16 J jetting  125, 199, 468 K Kentucky windage  322, 323 key product characteristics  459, 476 keyway  362 knit-line  145 knit-line location  439 KPCs, See  key product characteristics L laminar flow  163 laser sintering  72 lay flat  110, 129, 133 layout design –– conflict  93 lean manufacturing  72, 267 length dimension  88 liability –– mold designer/maker  463 –– molder  486 lifter  40 limit stress  100, 384, 386 limit switches  366 linear flow velocity  119 linear melt flow  204 linear melt velocity  157 linear shrinkage  300 linear velocity  112 locating dowel  357 locating pins  417 locating ring  7 locational-interference fit  419 lofted surfaces  85 lost core molding  41, 441

LPL, See  processing limits LSL, See  specification limits lubricity  104 M machine capability factor  50 machining and wear performance  101 machining efficiency factor  61 machining factor  58 machining labor rate  58 machining rate  101 machining time  58 maintenance  44, 228 –– venting  228 maintenance cost  16, 48 maintenance plan  461 managed heat transfer  286 manifold  153, 187, 443 –– cooling  267 manufacturing strategies  72 manufacturing strategy –– for purchasing molds  460 marginal cost  69 material consumption  16 material cost per part  46, 48 material removal rate  59 materials cost  45 material supplier  316 material waste  49 maximum cavity pressure  110 maximum deflection  398 maximum diameter  255 maximum shear stress  402 maximum stroke  353 mechanisms  466, 489 melt flipper  160, 189 melt flow –– pressure drop  121 –– velocity profile  120 melt front advancement  110 melt front velocity  468 melt pressure  109, 142 –– injection limit  143 –– maximum, due to endurance stress  258 melt temperature  294 metrology  476 MFI, melt flow index  115 mica  314 microfinish  36 minimum cooling line diameter  255 minimum draft angle  38 minimum wall thickness  129

524

Index

mirror finish  36 modulus  336 mold acceptance  459 mold base  17, 18, 53, 79, 93, 97 mold base cost estimation  54 mold base selection  91 mold base sizing  93 mold base suppliers  97 mold cavity  2, 80 mold commissioning  17, 29, 459 –– component verification  464 –– general process  462 –– mold assembly  466 –– mold map  488 –– mold verification  463 –– process  462 –– recommissioning  491 –– saving the last molding  489 mold cost estimation  53 mold cost per part  47 mold customer  43 mold customization  64 mold defect codes  488 mold design  16, 21, 24 mold development process  17, 25 mold dimensions  54 molded-in stresses  292 Moldex3D  303 mold filling simulation  121, 317 –– lay-flat analysis  128 mold filling simulations  110 Moldflow Plastics Insight  122, 303 mold functions  4 molding cycle  297 molding machine  50 molding machine capability  51 molding machine compatibility  95 molding processes –– strategic advantages  429 molding process instabilities  469 molding process setup sheet  468 molding productivity  15 molding trial methodology  470 molding trials  17, 44, 470 mold insert  361 mold inspection checklist  465 mold interlocks  404 mold layout design  54, 79 mold log and maintenance checklist  486 mold maintenance  37, 97, 485 –– post-molding  485, 489 –– pre-molding  485, 487

–– rebuilding  485 –– regular preventive  485, 489 mold manual  463 mold map  488 mold materials –– A2  100, 102 –– aluminum 6061-T6  99 –– aluminum 7075-T6   99 –– aluminum QC10  98, 99 –– C-18200  100 –– D2  100 –– digital ABS  102 –– grain structure  103 –– H13  100 –– SS420  99 –– Ultem (PEI)  102 mold material selection  98, 107 mold opening direction  79, 80 mold opening distance  150 mold opening height  13 mold operating log  488 mold prototyping  73 mold purchase agreement  44 mold quoting  24, 43 mold rebuilding  47, 490 mold reset time  3 mold setup time  267 mold structures  6 mold supplier  43 mold technology  429 mold technology selection  430 mold temperature controllers  253 mold texturing  37 mold wall temperature control  281 –– conduction heating  284 –– induction heating  285 –– insert mold  439 –– managed heat transfer  286 –– passive heating  286 –– pulsed cooling  281 moment of inertia  351, 354, 394, 414 moving cavity inserts  371 moving core  330, 362 moving half  11, 404 moving platen  11, 96 moving side  332 multicavity molds  9, 91 multigated  14 multilayer injection blow molding  445 multishot molding  280 multishot molds  447 multi-station  447

Index 525

multi-station mold  451 multivariate optimization  166 N naturally balanced  154, 162 naturally balanced feed system  92 net shape manufacturing  1 Newtonian  219 Newtonian limit  116 Newtonian model  117, 163 nitriding  104 nominal dimensions  311 nominal shrinkage rate  312 nonuniform shrinkage  475 normal probability  478 nozzles  153 O oil, for cooling  257 one-sided heat flow  278, 447 opening time  153 open loop control  191 operating cost  15 orientation  112 orifice diameter  95 original design manufacturer  30 original equipment manufacturer  30 over-filling  145 overmolding  280, 447 over-packing  111, 315, 469 overpressure  413, 426 P P20 steel  98, 100, 102, 383 packing  297 packing pressure  294, 469, 473 packing stage  2 –– gate freeze time  222 packing stage profiling  473 packing time  3, 199, 311, 473 pack pressure profiling  312, 323, 473 parison  444, 446 part cost  46 part dimensions  473 parting line  83, 84, 86 parting plane  9, 79, 80, 84, 93, 146, 228, 236, 404 parting surfaces  334 part interior  231 payment terms  27

peak clamp tonnage  131 physical vapor deposition  104 pilot production  24 pin length  352 pin-point gate  201 planetary gears  376 plastication  297 plastication stage  2 plastication time  3 plastic part design  21 plate bending  382, 392 plate compression  389 platen deflection  387 platens –– bending  381 plating  104, 491 polyjet printing  73, 269 polymer –– amorphous  295 –– compressibility  296 –– semicrystalline  295 poor gloss  52 positive return  369 power-law  163, 167, 219 power law index  116, 119, 120 power law model  119 power law regime  116 PPAP, See  production part approval process preliminary quote  16 preloading  400 pressure difference  414 pressure drop  110, 113, 142, 163, 219, 255 –– annulus  179 –– channel flow  119 –– gates  198 –– in vents  234 –– tube flow  163 pressure test –– of water lines and feed system  466 pressure transmission  14 pressure-volume-temperature  309 preventive maintenance  47 process capability, See  injection molding process capability index  479 –– rolled-up  483 processing conditions  124 –– robust  484 processing cost  44 processing cost per part  46, 49 processing limits  483 process optimization –– of injection molding  461

526

Index

process simulation  121, See also simulation process window  483 process window development  483 product definition  22 product design  16, 23 product development process  21, 24 production data  27 production flexibility  72 production part approval process  476, 482 production planning  23, 27 projected area  51 projections  449 prototype mold  316 prototype molding  29 pulsed cooling  281 purchase agreement –– for injection molds  459, 463 –– warranties for injection molds  463 purchase agreements –– for molded products  461 purchase cost  15 purge  175 purging  14 push area  340 push-pin  331, 342 –– defect  475 PVT, pressure-volume-temperature behavior  294 Q QC7 aluminum  383 QC10 aluminum  270 quality assurance  476 quality assurance methodology  477 quick ship  98 quoting process  43, 158 R race-tracking  134, 135 radial flow  204 radial mold opening direction  81 rails  6, 327 rear clamp plate  6, 327, 395 recommended melt velocity  125, 127 reduced material consumption  14 reduce setup times  72 regulatory agencies  25 replacement parts  461 requests for quotes  43 required heat transfer rate  252 residence time  175

residual stress  111, 292 retainer plate  88, 366 return pins  327 reverse ejection  333, 377 rework –– cost of  460 Reynolds number  163, 255 RFQs, request for quote  43 rheology  115 rib design  33 root cause analysis  459 rotating cores  375 rubber  314 rule of thumb  251 runner  10, 142, 146, 149, See also feed system –– annulus  179 –– efficiency  178 –– full round  176 –– half-round  176 –– hydraulic diameter  177 –– round-bottom  176 –– shut-offs  182 –– standard sizes  183 –– trapezoidal  176 runner volume  165 S safety margin  109 scientific molding  123, 463, 470 selective laser sintering  269 self-regulating valve  191 self-threading screws  34 semiautomatic  50 semiautomatic mode, See  injection molding semicrystalline  313 sensor –– cavity pressure  306 –– cavity temperature  306 sensor stack  306 series layout  91, 159 setup sheet –– for molding  467 sharp corners  34 shear heating  468 shear rate  112, 115, 217 shear rates –– maximum  217 shear stress  112, 340, 372, 392, 405, 455 shear thinning  120 shims  466 short shot  52, 111, 125, 142, 199 short shot studies  471

Index 527

shot size  97, 471 shot volume  97 shot weight stability studies  3 shrinkage  11, 112, 291, 292, 432 –– anisotropic  302 –– contractual obligation  317 –– in-mold  310, 475 –– isotropic  303 –– linear  292, 300 –– lower limit  314 –– negative  315 –– nonuniform  304 –– pack pressure profiling  312 –– post-mold  310 –– post-molding  475 –– processing dependence  311 –– recommendations  315 –– uncertainty  316 –– uniformity  323 –– upper limit  315 –– validation  306 –– volumetric  300 shrinkage analysis  293 shrinkage behavior  29 shrinkage data  316 shrinkage range  314 shut-offs  86 shut-off surface  230 side action  361 side wall –– deflection  402 side walls  372 –– bending due to shear  402 Sigmasoft  303 simulation –– Moldex3D  122 –– mold filling  121 –– Moldflow  122 –– shrinkage  303 –– Sigmasoft  122 –– Simpoe  122 single cavity  14 single cavity mold  91 sink  33, 203 sink marks  286 sintered vent  240 slender  415 slender core  272, 278 slides  366 slideways  371 sliding cores  366 sliding fit  362, 432 SLS, selective laser sintering  269

snap beam  39 snap finger  39 S-N, stress-number fatigue curve  385, 505 Society of the Plastics Industry  19, 36 socket head cap screws  6, 417, 422 solidification temperature  336 solidified plug  209 solidified skin  281 solvent  448 specification limits  479 specific heat  246 specific volume  313 –– relation to shrinkage  298 SPI  See Society of the Plastics Industry SPI finish  36 splay  52, 199, 468 split cavity  443, 446 split cavity design  82 split cavity mold  82, 334, 371 sprue  95, 142 sprue break  148 sprue bushing  10, 146, 149 sprue gate  200 sprue knock-out pin  147 sprue pickers  331 sprue pullers  12, 150, 180 SS420 steel  102 stack height  94, 96, 153, 187, 388 stack molds  186, 452 staged deployment  311 stagnant material  210 standards  19 start-up times  16 stationary half  11, 404 statistical process control  476 steady flow  113 steel safe  143, 218, 310, 316, 347 steel safe designs  184 stereolithography  72, 269 stop pins  327 strain  359, 383, 389 strength  100 stress  383 –– during ejection  359 stress concentrations –– due to cooling lines  257 –– ejector holes  407 –– water lines  407 stress-strain behavior  383 stripper bolt  12, 150 stripper plate  12, 149, 356 structural and thermal performance  100

528

Index

structural design  67, 381 –– minimize stress  382 –– mold deflection  387 –– mold size  388 –– safety factor  384 structural integrity  245 structural system design  381 –– cost  67 structured development  21 submarine gate  209 sub-runners  216 sucker pins  150, 180, 209 superposition  397 supply chain  19, 23, 44, 158 support pillars  388, 395 support plate  6, 94, 327, 372, 395 surface area  59 surface area removal rate  505 surface finish  36 surface refinishing  491 surface roughness  36, 38 surface striations  52 surface texture  36, 37 surface treatments  103 switchover  486 –– dynamics of velocity and pressure  474 –– position  472, 475 –– surge forward  475 switchover condition  472 T tab gate  203 Tait equation  294 technical feasibility  27 temperature differences  286 temperature differential  266 temperature fluctuations  281 temperature gradient  243, 260, 265 temperature variation  262 tensile stress  336 test mold  316 thermal conductivity  243, 246 thermal contact resistance  264, 332 thermal contraction  291 thermal diffusivity  246 thermal expansion  291, 296 thermal gate  14, 209 thermal sprue gate  211 thermal strain  336 thermocouple  466 thermoplastic elastomer  280 thermoreactive diffusion  104

thickness  49 thin wall  125, 143, 281, 409 three-dimensional printing  72 three-plate  142, 148, 153, 180 three-plate mold  12, 13, 16, 185 thrust pads  154 tie bar  95 –– tension  381 tie bar spacing  95 tight tolerance  29, 281, 305, 313, 323, 387, 389 tight tolerances  293 tolerance  28, 311 –– stack-up  347, 348 tolerance limit  419 tolerances –– tight  293 –– typical  293 tolerance specifications  29 tolerance stack-up  356 toll-gate process  21 top clamp plate  6 torpedo  209 total cost  68 TPE, thermoplastic elastomer  280 tuning loops  459 tunnel gate  206 turbulent flow  255 turret drives  452 two cavity  9 two-plate  142, 146, 153 two-plate mold  7, 11, 16 two-shot molding  280 type of gate  110 typical tolerance  29 U ultimate stress  100, 383 undercut  39, 208, 359, 361, 449 –– horizontal boss  39 –– internal thread  41 –– overhang  39 –– side window  39 –– snap finger  39 undercuts  334 undercutting  373 uniformly distributed  345 uniform wall thickness  31 unsupported spans  393 UPL, See  processing limits USL, See  specification limits

Index 529

V

W

valve gate  187, 212 valve pin  179, 212 velocity to pressure switchover  474 vent channel  236 venting  227, 348 –– analysis  228 –– dead pockets  239 –– defects  227 –– design  229, 236 –– dimensions  232 –– ejector pins  238 –– flashing  228 –– locations  110, 229 –– maintenance  228 –– pressure drop  234 –– relief  236 –– thickness –– maximum  235 –– minimum  233 vertically integrated molders  44 viscosity  112, 115 –– Arrhenius temperature dependence  116 –– Cross-WLF model  115 –– Newtonian model  118 –– Newtonian plateau  117 –– no-flow temperature  222 –– power law model  119 –– power law regime  116 –– WLF temperature dependence  116 viscous flow  112 volumetric flow rate  118, 120, 171 volumetric removal rate  505 volumetric shrinkage  33, 297, 473 von Mises stress  382, 407 V/P, See  switchover condition

wall thickness  32, 109 –– minimum  127 warpage  52, 111, 243, 291, 317 –– avoidance strategies  323 –– differential shrinkage  318 –– Kentucky windage  322 –– out of plane deflection  318 –– pressure gradient  319 –– radius of curvature  318 –– sources  318 –– temperature gradient  318 water assist  431 water assist molding  434 water lines  487 –– maintenance  489 wear –– maintenance of  490 wear plates  372 weld line  134 width dimension  88 windage  322 window  86 witness line  83, 86, 333, 358, 375 witness mark  198, 332, 377, 446 worst case scenario  384, 423 Y yield  46, 53, 481 yield estimates  52 yield stress  100, 383 Z zero shear viscosity  116, 118