Machining of Composite Materials

Machining of Composite Materials R. Teti University of Naples Federico II, Italy

Abstract Machining of composite materials is difficult to carry out due to the anisotropic and non-homogeneous structure of composites and to the high abrasiveness of their reinforcing constituents. This typically results in damage being introduced into the workpiece and in very rapid wear development in the cutting tool. Conventional machining processes such as turning, drilling or milling can be applied to composite materials, provided proper tool design and operating conditions are adopted. An overview of the various issues involved in the conventional machining of the main types of composite materials is presented in this paper.

Keywords: Machining, Composite Materials, Conventional Cutting Processes

ACKNOWLEDGEMENTS Acknowledgements are due for papers, contributions and correspondence received from Messrs (*CIRP members): Aspinwal, D.K., University of Birmingham UK; *Balazinski, M., Ecole Polytechnique de Montreal, Canada; *Byrne, G., University College Dublin, Ireland; *Brinksmeier, E., University of Bremen, Germany; Caprino, G., University of Naples Federico II, Italy; *Chandrasekaran, H., Swedish Institute for Metals Research, Stockholm, Sweden; Chen, L.J., Rotors Business Center, USA; *Dornfeld, D., University of California, Berkeley, USA; *Geiger, M., University of Erlangen-Nurnberg, Germany; *Inasaki, I., Keio University, Japan; *Jawahir, I.S., University of Kentucky, USA; *Klocke, F., Technical University of Aachen, Germany; *Komanduri, R., Oklahoma State University, USA; *Narutaki, N., Hiroshima University, Japan; *Pollmann, W., DaimlerCrysler AG, Stuttgart, Germany; *Spur, G., Technical University Berlin, Germany; Tomizuka, M., University of California, Berkeley, USA; *Uhlmann, F., Technical University Berlin, Germany; *Venkatesh, V.C., University of Technology Malaysia; *Weigl, E., HSC-Manufact. Engineering, Austria; *Weinert, K., University of Dortmund, Germany; *Wertheim, R., ISCAR Ltd., Israel. PhD student Doriana D'Addona, University of Naples Federico II, Italy, is gratefully thanked for her help and support in the preparation of the text.

1 INTRODUCTION Composite materials are used extensively as their higher specific properties (properties per unit weight) of strength and stiffness, when compared to metals, offer interesting opportunities for new product design. However, being nonhomogeneous, anisotropic and reinforced by very abrasive components, these materials are difficult to machine. Significant damage to the workpiece may be introduced and high wear rates of the cutting tools are experienced. Conventional machining practices, such as turning, drilling and milling, are widely applied to the machining of composite materials in view of the availability of equipment and experience in conventional machining. Although some of the materials used as reinforcement in composites,

such as glass, graphite, boron, alumina and silicon carbide, are highly abrasive and hard (sometimes as hard as or even harder than the tool material), conventional machining is considered for composites because their reinforcements are brittle and material separation is accomplished by brittle fracture rather than plastic deformation ahead of the tool. However, the cutting tool materials must be attentively chosen to minimize wear due to the hard abrasive constituents of the reinforcing phase in the composite representing the work material. Machining of a composite depends on the properties and relative content of the reinforcement and the matrix materials as well as on its response to the machining process. In addition, the choice of the specific process depends upon the following factors: type of machining operation, part geometry and size, finish and accuracy requirements, number of parts, diversity of parts (including the material of the parts), availability of appropriate machine and cutting tools, availability of in-house technology, current machining practice, manufacturing schedule, capital requirements and justification for new equipment, environmental and safety considerations, and overall costs. 2

COMPOSITE MATERIALS

Composite materials are formed from two or more materials producing properties that could not be obtained from any one material. One of the constituent materials acts as the matrix and at least one other constituent material acts as the reinforcement in the composite. The role of the matrix material comprises the following: - to protect the reinforcement materials; - to distribute the stress to the reinforcement material(s); - to provide for the final shape of the composite part. The role of the reinforcement material(s) is the following: - to provide the composite high mechanical properties; - to reinforce the matrix in preferential directions. The properties of a composite material depend on the nature of the reinforcement and the matrix, the form of the reinforcement (particles, fibres) and the relative content of reinforcement and matrix expressed as volume fraction: Vf = (reinforcement volume)/(composite volume) and Vm = (matrix volume)/(composite volume), where Vf+ Vm = 1.

Composite materials can be classified on the basis of the matrix material used for their fabrication: polymer matrix composites (PMC); metal matrix composites (MMC); ceramic matrix composites (CMC). Theoretically, a multitude of materials can come under these categories. In the following, a brief description of some of the PMC, MMC and CMC composites most commonly used for industrial applications is reported. 2.1 Polymer matrix composites (PMC) The most common types of reinforcement used for PMC are strong and brittle fibres incorporated into a soft and ductile polymeric matrix. In this case, PMC are referred to as fibre reinforced plastics (FRP). Capital letters G, C and A are placed before the acronym FRP to specify the nature of the reinforcing fibres: glass, carbon or aramid fibres. The fibres can be long (continuous) or short (discontinuous). Long fibres can be unidirectional (all fibres parallel to each other) or woven into a fabric or cloth. Unidirectional fibres provide for the highest mechanical properties in a composite. Glass fibre reinforced plastics (GFRP) are by far the most commonly used materials in view of their high specific mechanical properties and low cost. Carbon fibre reinforced plastics (CFRP) and aramid fibre reinforced plastics (AFRP) provide higher specific strength, higher specific stiffness and ligher weight. They are, however, expensive and are used only for those applications where performance and not cost is the major consideration. AFRP is used instead of CFRP where strength, lightness and toughness are major considerations, and stiffness and high temperature performance are not. The common matrix materials for FRP composites are: - thermoset polymers (e.g. polyester, epoxy) - thermoplastic polymers (e.g. polyamide, peek). Thermoset polymers remain rigid when heated and consist of a highly cross-linked three-dimensional network; they are quite strong and stiff and have poor ductility. Tensile strength, ail (MPa)

FRP material GFRP Unidirectional (Vr = 60 %) Woven cloth* Chopped . . roving* (short fibres) Sheet molding compound* (short fibres)

Elastic modulus, E (MPa)

I looo

45000

Strain to Density failure, d (g/cmS; Eu (%)

2.3

2.1

100-300 10000-20000

-

1.5-2.1

50-200

6000-12000

-

1.3-2.1

10-20

500-2000

1200

145000

0.9

1.6

800

220000

0.3

1.6

looo

75000

1.6

1.4

1.3-1.9

CFRP Unidirectional (Vr = 60 %) High strength Unidirectional (Vr = 60 %) High modulus AFRP Unidirectional (Vr = 60 %)

Table 1: Mechanical properties of FRP composite *For these materials: Vf = 20% - 50%.

Among the thermoset resins, polyester resins are lower in cost and are not as strong as epoxy resins. Polyester matrix composites are used in the fabrication of boat hulls, structural panels and parts for automobiles and aircrafts, building panels and beams, electrical appliances, water tanks, pressure vessels, etc. Epoxy resins, in addition, have a lower shrinkage after cure allowing for higher fabrication accuracy. Epoxy matrix is used commonly in CFRP and AFRP composites for aerospace applications, military equipments, satellite antennae, sports equipments, medical prostheses, etc. Thermoplastic polymers consists of flexible linear molecular chains that are tangled together and, as the name indicates, soften when heated; they have lower strength and modulus but quite high ductility. Among the thermoplastic resins, polyamide and peek resins are used as matrix materials in FRP composites for applications in the aerospace industry due to their superior mechanical properties and high glass transition temperature. Maximum service temperatures for FRP composites are relatively low, as the matrix material is prone to softening, chemical decomposition or degradation at moderate temperatures. The same temperature limitations apply to the machining of FRP composites. Table 1 reports the main mechanical properties of some FRP composite materials. 2.2 Metal matrix composites (MMC) MMC are used for applications requiring higher operating temperatures than are possible with PMC materials. Most of these composites are developed for the aerospace industry, but new applications are found in the automotive industry, such as in automobile engine parts, making use of continuous fibre, discontinuous fibre, or particle reinforced MMC. Continuous fibres provide for the highest stiffness and strength properties obtainable in MMC materials. Boron-aluminium composites are one of the earliest developed MMC material types. It is made by hot pressing layers of boron fibres between aluminium foils, so that the foils deform around the fibres and bond to each other [I]. By reinforcing with boron, the tensile strength can be increased by a factor of three to five while the elastic modulus can be tripled. Further reinforcing materials for MMC are silicon carbide, alumina and graphite in the form of particles, short fibres (whiskers) or long fibres. Aluminium, magnesium and titanium alloys are the most common matrix materials used in MMC materials. Table 2 reports the main mechanical properties of some MMC materials. Figure 1 is a plot of specific strength versus specific stiffness for various composites and conventional metal materials. It can be seen that composites, in general, have higher specific strength and specific modulus over conventional steel, Al, Ti, and Mg alloys, and MMC have properties superior to PMC composites. Applications of continuous fibre reinforced MMC include use of B-AI for the fuselage of the space shuttle orbiter, SIC-AI for the vertical tail section of advanced fighter planes, SIC-TiAI for hypersonic aircraft, etc. Discontinuous fibre and particle reinforced MMC are low cost MMC that provide higher strength and stiffness and better dimensional stability over the corresponding unreinforced alloys. Small additions of reinforcement (Vr = 20%) moderately increase the base alloy strength and stiffness. They also increase the wear resistance and contribute toward the difficulty in machining these materials. These MMC are used for sport equipments, automobile engine parts (pistons, cylinder liners, brake drums), missile guidance parts, etc.

Tensile Elastic Strain to strength, modulus, failure, u,, (MPa) E (MPa) E,, (%)

MMC material IContinuous-fibre MMC IAI2124-T6(45% B) IAl6061-T6(51%B) IAl 6061-T6 (45% Sic)

I I I I

I I230000 I I 200000 I

1450

I220000

1410 1460

Matrix material

I 0.81 0.74 0.89

I I I

Discontinuous-fibre MMC Al 2124-T6 (2O%SiC)

650

125000

2.40

Al 6061-T6 (20% Sic)

480

120000

5.00

lparticle MMC IAl2124-T6(20%SiC)

I 500

Al 2124-F

450

700000

9.00

Al 6061-F

310

70000

12.00

550

7.00

5.50

No reinforcement

Table 2: Mechanical properties of MMC materials 0.6 h

0.45 B/Al 0

E

z

0.34 Gr/Mg

D

a

0.4

SlCfr

3j

5 m C

1

Conventional Steel, Al, Ti, Mg

I3

0.2

c

0.37 Gr/AI

0

0 0 0.37 Gr/AI 0.60IsIGr/Epoxy

NO.giCw/Al

u)

0

E 0

8 a v)

. 0 Be

0.50 Gr/Epoxy

0 0

50

100

Specific stiffness (10

150 Nrn/Kg)

Figure 1: Specific strength vs. specific stiffness for various MMC materials. Number in front of the composite is the reinforcement volume fraction [2]. 2.3

Ceramic matrix composites (CMC)

CMC materials are being developed mainly to improve the fracture toughness of unreinforced ceramics which already possess higher specific modulus and mechanical properties at high temperature superior to those of metals. Continuous fibres, discontinuous fibres (whiskers) or particles can be utilised as reinforcing constituents in CMC. The common reinforcement materials used in CMC are alumina and silicon carbide. A volume fraction Vf = 20% of S i c whiskers added to alumina can increase the fracture toughness from 25 to 50 MPa. Such an increase in toughness of a ceramic cutting tool will enable it to take heavy cuts or to perform without fracture in interrupted cutting. Conventional hot isostatic pressing techniques can be used to consolidate CMC composites. Other CMC include carbonkarbon composites in which high strength carbon fibres are embedded in a graphite matrix. The low density of carbon in combination with the very high strength of carbon fibres offers potential for the development of ultra high specific strength materials. Table 3 reports the main mechanical properties of some CMC materials.

Flexural strength, uf (MPa) 400-650 400-550 350-500

Fracture toughness, k (MPa) 30-45 40-60 45-65

400-500 500-800

45 45-55

20

Table 3: Mechanical properties of Sic whisker reinforced CMC materials at room temperature. 3

I

I 105000 I I 105000 I

IAl 6061-T6 (20% Sic)

I I

SiqNd

Sic whisker volume fraction, Vf (%) 0

I I

MACHINING APPLICATIONS

Machining of composite materials differs significantly in many aspects from machining of conventional metals and their alloys [3-51. In the machining of composites, the material behaviour is not only non-homogeneous and anisotropic, but it also depends on diverse reinforcement and matrix properties, and the volume fraction of matrix and reinforcement. The tool encounters alternatively matrix and reinforcement materials, whose response to machining can be entirely different. Thus, machining of composite materials imposes special demands on the geometry and wear resistance of the cutting tools. Accordingly, tool wear mechanisms and development must be attentively considered to establish correct cutting tool selection. In the following, applications of machining processes to composite materials are reviewed with reference to FRP materials and MMC materials. As regards the machining of CMC materials, the very small number of contributions received and the scarcity of information available in the open literature on this topic did not allow for the preparation of a dedicated section. 3.1 Machining of fibre reinforced plastic composites Orthogonal machining of FRP Investigations carried out in [6] by orthogonal cutting of FRP composites with different fibre orientations allowed for the clarification of the cutting mechanisms taking place in FRP (Figure 2). When machining is conducted at an angle of 0" to the fibre orientation, the laminate is subjected to stresses parallel to the fibres. In addition, the surface below the cutting edge is compressed. The material failure occurring in front of the cutting edge is due to delamination, matrix fracture or fibre-matrix interface failure, which is recognizable from the crack in the composite laminate ahead of the cutting edge. Individual fractures occurring in the fibres and in the matrix below the cutting edge are also visible and remain in the machined surface. As the angle between cutting direction and fibre orientation increases, fibres are compressed and bent in the direction opposite to the fibre orientation, ending up in fibre breakage as a result of bending and pressure load. This can result in fibre-matrix interface failure which extends into the unmachined surface. These load directions, which are the least favourable for FRP composites particularly at angles between 30" and 60" to the fibre direction, is reflected in a poor surface quality. In a composite machined at 90" to the fibre direction, the fibres are subjected to bending and are sheared off. In contrast to laminates with 0" fibres, each fibre has to be cut separately. The compressive strain normal to the fibres creates problems as interfacial fractures extend into the unmachined surface. More favourable conditions develop for fibre orientation 135". Fibres are subjected to bending and tensile stress and break in bundles. Problems arise, however, from the fact that individual fibres can be pulled out due to insufficient adhesion to the matrix.

Figure 3: Examples of turned FRP parts [15].

Figure 2: Cutting mechanisms for FRP composites [6] The machinability of CFRP and GFRP was deeply investigated in [7]. A model for cutting force prediction in orthogonal cutting operations was presented. Three parameters were initially varied during orthogonal cutting using HSS tools: tool rake angle, relief angle and depth of cut. Their effects on cutting forces were investigated and an optimal tool geometry was found. The effect of tool wear on cutting forces in the machining of unidirectional GFRP was investigated too [8]. In [9], cutting forces were found to increase with increasing depth of cut during orthogonal machining of unidirectional CFRP. The effect of fibre orientation on cutting forces and cutting quality in orthogonal machining of unidirectional CFRP was treated in [lo]. In [ I l l , the attention was focused on the mechanisms of chip generation. Because of the inferior surface quality of unidirectional CFRP after orthogonal machining for some fibre orientations, in [I21 the development of a new tool geometry to reduce work material surface damage was investigated. In [13, 141, tool wear development was studied and monitored using acoustic emission (AE) signal detection and analysis during orthogonal cutting of different types of composites: unidirectional GFRP and CFRP, and sheet moulding compound (SMC). Decision making on tool wear state was performed through graphical examination and neural network computation of AE spectrum features. Different results were obtained according to the composite material type: tool wear discrimination was reliably achieved for GFRP but not for CFRP and SMC. Turning of FRP A significant amount of research work has been carried out in applying turning processes to the various FRP composites with different cutting tools. Turning, together with drilling, milling and sawing, belongs to the most important cutting technologies for the machining of FRP [15]. Turning differs from milling and sawing mainly because an almost constant engagement of the tool exists. Apart from fluctuations in stress caused by the different cutting behaviour of the fibres and the matrix, a quasi-continuous cut exists during turning of FRP. The machinability of FRP is primarily determined by the physical properties of the fibres and the matrix as well as by the fibre orientation and volume fraction. While glass and carbon fibres break in a brittle manner under bending stresses, aramid fibres undergo shearing fracture under high deformation bending and tear under tensile loading. Moreover, the machining of short fibre reinforced composites is much easier than that of unidirectional FRP.

Although the cutting of FRP composite parts is rarely desired, it can be seldom avoided for the production of the final geometry, surface quality, and form accuracy of conventionally produced parts. Turning is applied to rotation-symmetric parts such as drag links, bearings, spindles, axles, rolls, or steering columns, etc. Figure 3 shows some typical FRP parts produced by turning. Particular attention was given by several authors to the aspects of tool wear mechanisms and development in turning of FRP composites with the aim of establishing correct cutting tool selection criteria. Among the possible wear mechanisms, which include abrasion, adhesion, tribo-oxidation and surface damage, only abrasion, surface damage and sometimes adhesion are of significance for FRP machining. Wear mechanisms are primarily related to the physical and mechanical characteristics of the different fibre-matrix systems. Glass and carbon fibres show a strongly abrasive behaviour because they are extremely abrasive by nature. Aramid fibres, on the other hand, impair the tool due to their low heat conductivity and ductile behaviour. Adhesive wear occurs when carbonised or molten matrix depositions settle on the tool surfaces. In [16], an analysis of tool wear during turning of GFRP and CFRP with diamond coated tools was carried out. The dominating wear mechanisms during cylindrical turning such as cutting edge blunting, elimination of the coating layer, retreat of the cutting edge, and crater wear formation were characterised and their development was explained. In [15], a survey on the possibilities and variants of application in turning of different types of FRP (polyesterglass, epoxy-glass, polyamide-carbon) with carbide, diamond coated, PCD and CBN tools was presented. The machinability of GFRP in precision turning by means of tools made of various materials and geometries was investigated experimentally in [17]. It was found that, by proper selection of the tool material and geometry, excellent machining of the workpiece is achieved and the surface quality relates closely to the feed rate and the tool. Flank wear as well as retreat and rounding of the cutting edge are the most frequently observed wear effects during cutting of FRP [18-201. Hereby, the wear speed is mainly related to the fibre content [18]. Furthermore, crater wear occurs only to a minor extent [19]. The cause for this wear behaviour results from the discontinuous chip formation during cutting of FRP. Hence, fracture on the face occurs only to a minor extent, whereas fracture on the flank is the main reason for the examined wear types [19, 22, 231. The selection of a large clearance angle is therefore recommended to improve the tool life. However, it must be noted that this causes a weakening of the tool that may promote cutting edge chipping. Thus, an optimum clearance angle has to be determined for every tool.

In [20, 211, investigations on cutting of CFRP showed that the hardness and microstructure of the cutting edge for various PCD tools exert a significant influence on the effectiveness of FRP machining. Coarse-grained PCD tools, in particular, reveal higher resistance to wear than medium- and fine-grained PCD types. The wear appears in the form of cutting edge rounding, chipping and crack formation on the different PCD types. Carbide tools also display a flank wear, yet more irregularly and with a tool life significantly shorter in comparison. The wear is characterised by scratches and chippings. A longer tool life is achieved by PCD and TIC or TaC free carbide types due to their higher thermal conductivity, pressure strength and wear resistance. Interrupted cutting during turning of CFRP causes a higher wear than continuous cutting with carbide tools under equal cutting conditions. In [15], tool wear was studied during turning of GFRP obtaining good results with carbide and PCD tools. Yet, the wear for these two tool materials is considerably different. Carbide tools exhibit mostly flank wear and rounding of the cutting edge. Crater wear does not occur in any significant way but carbonised chip material is deposited on the tool rake face during cutting. PCD tools also show flank wear. However, the rate of wear development is clearly slower in comparison with that for carbide tools. Furthermore, the cutting speeds attainable with PCD tools are much higher than those possible with carbide tools. Figure 4 shows the influence of cutting speed on tool life during turning of different GFRP composites: unidirectional glass cloth/epoxy with fibre volume fraction Vf = 55% (EPRU 5), bidirectional glass roving fabridepoxy with Vf = 45% (EPR 8), glass matlpolyester with Vf = 35% (UPM 72). For all GFRP materials, a decrease in tool performance was verified with increasing cutting speed. The EPRU 5 composite displays the worst tool life, the EPR 8 composite an intermediate tool life, and the UPM 72 composite the best tool life behaviour. This can be explained by the fact that lower glass fibre volume fractions result in lower thermomechanical stresses of the cutting edge and, consequently, higher tool life values. The influence of various cutting tool materials on tool wear is shown in Figure 5 during turning of the UPM 72 composite. The superiority of diamond-based cutting materials, PCD and diamond-coated carbide, over monolithic carbide tools is clearly seen. While the tool life trend is decreasing for carbide tools, a linear development is evident for diamond-coated tools and PCD tools.

Figure 4: Tool life of diamond-coated carbide tools vs. cutting speed in turning of different GFRP composites [15].

Figure 5: Tool life for carbide, diamond-coated and PCD tools vs. cutting speed in turning of GFRP (Vf = 35%) [15]. A comparison in tool life of uncoated and diamond-coated carbide tools shows the protective effect of the diamond layer on the carbide substrate, granting protection against abrasive wear and thermal wear. At high cutting speed, the diamond-coated tool life is surpassed by PCD tools. The degradation of adhesion of the diamond layer to the carbide substrate is the main reason for this behaviour. Tool wear in CFRP machining is significantly different from tool wear in GFRP machining. As Figures 6 and 7 illustrate, a minor relation of tool life to cutting speed exists for CFRP machining in comparison with GFRP machining (Figure 4 and 5). The flatter trend of tool life versus cutting speed for CFRP is related to the lower temperature development during machining due to the much higher heat conductivity of carbon fibres. Thus, higher cutting speeds can be utilised during CFRP turning. A comparison between carbide and PCD tools in turning of thermoset matrix CFRP (Figure 6) indicates that much higher cutting speeds can be used with PCD tools. It must be noted that different criteria of tool life were adopted for the two tool materials. While during machining with carbide tools a VB = 0.2 mm was used as tool life criterion, the VB for PCD tools was reduced to 0.1 mm. If a standard VB = 0.2 mm were taken as a basis, the total volume of material removed for PCD would surpass the total volume of material removed for carbide by 250 times. Moreover, the tool life for both cutting tools shows that with increasing cutting speed the temperature influence on the tool life increases. The steeper rise in tool life for PCD tools demonstrates the extended range in cutting speed compared with that for carbide tools. Figure 7 shows a comparison between uncoated and diamond-coated carbide tools in turning of thermoplastic matrix CFRP. The tool wear during turning of polyamides matrix CFRP with diamond-coated tools is characterized by small chippings of the coating. The degree of chipping increases with the engagement time of the cutting edge up to little beyond the contact area between chip and rake face. The base carbide becomes smooth and the cutting edge rounded. The sharp-edged transition between carbide and coating layer on the tool face is also subject to the abrasive action of the carbon fibres and is removed in the direction of the chip flow. Independently of the cutting parameters, thermoplastic matrix deposits form on the tool face and flank, but are periodically removed during turning.

Figure 8: Mechanism of ultrasonic vibration cutting. vc = cutting speed; f = vibration frequency; IT = cutting distance during one period of tool vibration [24].

Figure 6: Tool life for carbide and PCD tools vs. cutting speed in turning of epoxy matrix CFRP ( V f = 40 %) [I51

This cutting speed is called "critical cutting speed" in the vibration cutting and is calculated by vc = 2naf, where a is the amplitude of the tool vibration and f is its frequency (in this study vc = 110 m/min). The mechanism of ultrasonic vibration cutting is shown in Figure 8. The performance of ultrasonic vibration cutting strongly depends on the cutting distance during one period of the tool vibration IT = vJf. It was experimentally confirmed that IT must be smaller than the fibre diameter (7 pm in this study) to take advantage of ultrasonic vibration cutting. By making IT smaller than the fibre diameter, the matrix and the fibre, which have different mechanical properties, can be sheared separately. Hereby, the fibres do not prevent the shearing of the plastic matrix and, consequently, the surface quality is improved even if the angle between fibre orientation and cutting direction is 90". A comparison of surface roughness between conventional and ultrasonic vibration cutting is shown in Figure 9. When IT is larger than the fibre diameter, the surface roughness in ultrasonic vibration cutting is similar to that of conventional cutting (Figure 9a). On the contrary, when IT is smaller than the fibre diameter, the roughness in ultrasonic vibration cutting becomes smaller than that of conventional cutting (Figure 9b).

Figure 7: Tool life of diamond-coated and uncoated carbide tools vs. cutting speed in turning of polyamide matrix CFRP (Vf = 40%) [ I 51. As regards cutting parameters, speed and feed primarily influence the life of the cutting edge (Figure 7). The tool life of both uncoated and diamond-coated carbide tools reveals that wear is reduced by the diamond layer. An increase in thermal stress of the cutting edge is connected with the increase in cutting speed. Due to the high thermal conductivity of the diamond layer, an increase in thermal load capacity is available and, accordingly, higher cutting speeds are allowed for. Because of the difficulty in machining CFRP composites with high efficiency, in [24] it was proposed to apply ultrasonic vibrations in turning of CFRP pipes using a diamond-coated tool. The performance of the ultrasonic vibration cutting was evaluated in terms of cutting force, burr formation and surface roughness. Ultrasonic vibration cutting allows to obtain good surface quality when machining difficult-to-cut materials. This is due to the fact that the ultrasonic vibration avoids the continuous contact between the tool rake face and the chip. As reported in [25], when the cutting speed becomes faster than the speed of the tool vibration, the tool rake face is not separated from the chip and consequently ultrasonic vibration cutting loses its effectiveness.

Figure 9: Surface roughness for conventional and ultrasonic vibration cutting [24].

Fibre orientation (deg)

Fibre orientation (deg)

Figure 10: Cutting forces for conventional and vibration cutting for a cutting speed = 4 m/min corresponding to IT = 3.6 pm [24].

Figure 13: Typical thrust force vs. time plot for a single drilling operation on CFRP [26].

Figure 12: Microscopic observation of the edge when fibre orientation is 90" [24]. Cutting forces were also investigated to confirm that the critical limit of IT in CFRP cutting was correct. The cutting forces for different fibre orientations in conventional and ultrasonic vibration cutting are shown in Figure 10. In cutting of CFRP the thrust force is higher than the principal force. The average thrust force in ultrasonic vibration cutting becomes less than half of the thrust force in conventional cutting if IT < 7 pn. The improvement of the surface quality was confirmed by microscopic photographs (Figures 11 and 12). Arrows in the figure show the cutting and the feed direction. For fibre orientation O", in conventional cutting lots of fibres are pulled out (Figure I l a ) . In ultrasonic vibration cutting, those fibres are absent (Figure 11b). In addition, for fibre orientation go", in conventional cutting fibres are not cut at the edge of the surface (Figure 12a). In ultrasonic vibration cutting, however, those fibres are not visible (Figure 12b). Drilling of FRP A number of research workers have investigated the drilling of different FRP composite materials using various cutting tool materials. In [26], the use of high performance carbide drills in drilling CFRP epoxy matrix composites was studied. To reduce the high wear rate of the carbide drills, speciality coatings including titanium nitride (TIN) and diamond-like-carbon (DLC) can be used. The performance of the coatings was analysed in terms of damage to the composite and thrust force and torque produced during drilling. Figures 13 and 14 show typical thrust force and torque profiles, in the case of an uncoated tool. The general form of the thrust force and torque profile, comprises six main stages. Initially, there is a sharp increase in thrust force and torque due to the initial entry of the drill into the composite. This is followed by a further increase in the force and torque as the second cutting edge enters the workpiece. The maximum force and torque occur as the tip of the tool breaks through the bottom ply of the laminate.

Figure 14: Typical torque vs. time plot for a single drilling operation on CFRP [26]. This is followed by a sharp reduction of the force and a slight drop of the torque due to the fact that the tip of the tool has broken through the back face of the workpiece. When the first chisel edge breaks through the back face of the laminate, the reduction in force becomes more gradual and the torque is seen to slightly increase. Finally, the force and the torque drop to zero as reaming takes place. As the number of holes drilled increases, so does the magnitude of both the maximum torque and the thrust force values. Similar profiles were noted for both the uncoated and the coated tools. Figure 15 shows a combination of the maximum thrust force and torque for uncoated and coated drills. Also included are the flank wear results, which show wear in the order of 0.07 mm after 32 drilling operations. The maximum thrust force, maximum torque and flank wear curves for the three drill types exhibit similar trends. Both the thrust force and torque curves rise sharply in the initial stages after which the subsequent rate of increase is seen to reduce. A change in form of both these curves is apparent in the region 5 < n < 10, drilled holes.

(a) n = 1

(b) n = 1 0 0 0

Figure 17: Hole exit in drilled GFRP (Vf = 60%). Work thickness: 10 mm; drill: fish tail carbide; drill diam.: 10 mm; feed: 0.1 mm; cutting speed: 163 m h i n [30]. Figure 15: Variation of maximum thrust force, torque and flank wear with number of drilled holes [26]. x Uncoated tool, DLC tool, TIN coated tool.

+

In [27], the tool life of uncoated and diamond-coated carbide tools in drilling of GFRP composites was studied (Figure 16). The comparison of the tool life of the different types of tools illustrates the protective effect of the diamond layer. In addition to the protection against abrasive wear, the diamond layer also protects against thermal wear. A shift of the tool life line towards higher values is obtained for higher cutting speeds. Nevertheless, the tool life curve of the diamond-coated carbide bends at high cutting speeds. This indicates the thermal failure of the substrate material. An increase in cutting speed is connected with an increase in cutting temperature. On the one hand, this can be explained by a crater wear development and, on the other, by the decrease in tool life for uncoated carbide tools due to the insufficient heat resistance of the substrate [28]. In [29] an overview of the potential uses of PCD in FRP composite drilling was shown and PCD tools were compared to carbide tools in terms of both economics and quality. It was found that drilling processes performed on FRP composites are strongly dependent on the tools implemented. PCD is an economical alternative to carbide despite the higher cost because tool life is longer and higher processing speeds can be used.

Cutting speed vc ( r n h i n )

Figure 16: Tool life of carbide and diamond-coated carbide tools vs. cutting speed in drilling of GFRP (Vf = 55%). Drill diameter: 10 mm; workthickness: 18 mm; feed: 0.08 mm [27].

In [30], the problem of burr generation in drilling of GFRP composites with different cutting tools is studied. The fish tail drill is found to be very effective in suppressing the generation of burr. Several grades of carbide materials were tested as fish tail drills. Among the tested carbides, KO1 and K10 showed the highest cutting performance and drill wear depended only on the fibre type and volume fraction. Even with these drills, tool wear causes the generation of burr after a certain length of drilling. Figure 17 shows an example of burr after drilling 1000 holes on a 10 mm thick GFRP laminate with Vf = 60 %. The burr is mainly caused by the outer corner wear of the drill. To get a longer tool life, it is necessary to use higher wear resistant tool materials and diamond is the most suitable. A trial diamond endmill with sintered diamond blades brazed on a carbide substrate was used for drilling GFRP. Compared with carbide drills, the wear of diamond endmills was very small and after drilling 1000 holes the burr was scarcely generated. In addition, the torque and thrust for diamond endmills was less than a half of that for carbide drills. The roughness of the hole wall drilled with carbide drills and diamond endmills was compared. Holes drilled with fish tail carbide drills have a high roughness with ,R = 30 pm after drilling 1000 holes. Holes drilled with diamond endmills have a low roughness with ,R Carbon fibers Short fibers -> Long fibers then: Carbide tools -> PCD tools

Number of holes

Figure 21: Tool wear when drilling the aluminium/CFRP/titanium sandwich structure for different tool specifications [34]. Milling of FRP Milling operations conducted on FRP parts, as opposed to metal parts, are characterized by a low ratio of material removed to total part volume. Milling is used, as a rule, as a corrective end-machining operation or to produce defined, high quality surfaces. The fibre type, reinforcement architecture and matrix volume fraction are the most important factors governing tool selection and machining parameter setting. In the case of glass and carbon fibre reinforcement, it is the cutting tool material, that dominates the tool selection. In the case of aramid fibre reinforcement, it is the cutting tool geometry that dictates the choice of the cutting tool. The behaviour in machining operations is determined mainly by the characteristics of the fibres reinforcing the composite. This exerts a major influence on process parameter selection or on the suitability of tool concepts. The fibres are characterized by their high tensile strength, modulus of elasticity (higher than those of the plastic matrix) and low strain at failure (lower than that of the plastic matrix). Additionally, there is a variety of thermal characteristics, depending on the fibre type, which differ considerably from those of the plastic matrix. Tool selection The hardness of the glass and, more especially, of the carbon fibres results in a high level of wear during machining. Since this wear manifests itself above all in tool cutting edge rounding, the cutting edge should possess a high degree of resistance to abrasion and chipping. Fine grain carbide from the K 10 group or, better, PCD are, therefore, suitable as tool materials. Ceramic materials are unsuitable because their low strength and high brittleness make them very sensitive to shocks, resulting in tool cutting edge spalling, and their low heat conductivity does not allow for the dissipation of the heat generated during FRP composite machining. Due to its low wear resistance, CBN, which is as expensive as PCD, presents no advantage over the latter. In order to ensure that the glass and carbon fibres are severed in a clean cut, it is very important to ensure a high cutting edge sharpness. As regards cutting edge geometry, care should be taken to ensure that cutting edge serration and radius are as small as possible. Due to the pronounced susceptibility of these fibres to brittle fracture, tool geometries correspond approximately to those of the tools used in metal working. These requirements are summarised in Figure 22. Distinctions in tool selection can be drawn between areas of application on the basis of the fibre type, fibre length and fibre volume fraction in the FRP composite material.

Figure 22: Tool requirements for milling of GFRP and CFRP composite materials. The standard tool geometries cannot be applied to machining of AFRP composites because the individual aramid fibres can be separated in a clean cut only under simultaneous prestress. Accordingly, the tool geometry must allow for prestressing of the aramid fibres before the cutting process begins [37-391. A high cutting edge sharpness and a small cutting edge radius are further requirements. To minimize the friction at the tool rake face and, therefore, the tendency to build-up, the tool must satisfy high requirements in terms of face and flank surface quality due to the high friction coefficient of the tough aramid fibres. Tools made of fine grain carbide prove successful in milling of AFRP composites [6]. The tool requirements for AFRP milling reported in Figure 23 illustrate the differences in tool characteristics compared with tools used for GFRP or CFRP milling (Figure 22). Delamination and top layer fraying can only be avoided in milling of AFRP composites by using tools with a counterclockwise spiral. The tool used depends on the thickness of the part to be milled. For thin laminates, opposed helical tools prove effective. The forces thus released point from the top and bottom layers towards the middle of the workpiece. This variant requires accurate axial alignment of the workpiece. When thicker parts are milled, the tool becomes clogged up in the middle with fibre material. Split helix milling cutters should be used for such parts. The constantly alternating stress prevents the fibres from avoiding the cutting edge. The high dynamic stress, which can result in strong vibration and chatter, is a disadvantage for these tools. PCD tipped tools can only be used in special cases as it is extremely difficult to grind the complex geometry of such tools. Only tools having PCD cutting edges set at opposed helix angles soldered onto solid carbide shanks have found general acceptance for certain applications.

0

Special macroscopic geometry of tool cutting edge

0

Very high resistance to wear

0

Low cutting edge radius, r = 10-15 pm

0

Very high surface quality of rake and flank, Ra ~ 0 . pm 8

0

Going from: Low thickness (c 3 mm) -> High thickness then: Carbide tools -> PCD tools Figure 23: Tool requirements for milling of AFRP composite materials.

Deburring short fibre reinforced composites, such as sheet moulding compound or glass mat reinforced composites, is a frequently required milling operation [30-411. Abrasive pencils and multi-sided cutting tools such as diamond interlocked carbide milling cutters are generally used for this application. Milling cutters with very sharp double cutting edges should be used to mill composite parts reinforced with unidirectional long fibres and a high fibre volume fraction since this is the only means of separating the fibres in a clean cut. Milling cutters with PCD cutting edges are superior to carbide milling cutters in terms of surface quality and tool life. Tool wedge angles of approximately 75", with a rake angle of 0" to 7", were proven favourable in [41]. A reduction in the wedge angle offers only a short-term advantage which is off-set by increased wear. When the rake angle selected is too small, grinding-in occurs at the flank as a result of the tendency of the polymer matrix to deform. It can be assumed that the material demonstrates elastic lateral yield and bounces back against the flank after the cutting operation, an effect which occurs more frequently when fibre reinforced thermoplastic composite materials are machined, due to the high strain at failure of the matrix. Evidence of this is also provided by deposits of matrix material on the tool face. The use of wear-minimizing coatings is one means of reducing the wear of carbide tools [6]. TIN, TiCN, TiWN and CrCN coatings can be used to mill FRP composites. PCD tools treated with ion implantation offer an interesting alternative means of reducing tool wear. The hardness of the bonding phase and the friction coefficient are enhanced by the introduction of ions into the surface. This, in turn, should result in increased wear resistance. In [42] the tool design was studied with particular reference to the milling of CFRP composites. In CFRP milling, the boundary conditions for the selection of the most suitable tool are the mechanical load, caused by the abrasive carbon fibres, and thermal stress due to the insufficient thermal characteristics of the plastic matrix. Analysis of the wear mechanisms was used to reveal potential means of optimising the tool, taking into account the process-specific parameters at the tool design stage. Milling processes performed on FRP composites are strongly dependent on the tools implemented, both in terms of quality and economics [29]. Carbide tools and, particularly relevant for CFRP, those made of PCD proved successful. In the latter case, the diamond cutter is soldered onto a carbide support that allows for small tool diameters even with multiple-toothed tools. This is particularly important when milling slots or grooves with small internal radii. Larger tool diameters are achieved by attaching the cutters to a tool holding system as illustrated in the examples of Figure 24.

-

50 prn

Figure 25: Wear curves for PCD and carbide tools during milling of CFRP [29]. The design and modification of tools should follow an analysis of the wear mechanisms, mainly characterised by the abrasive effects of the reinforcing fibres. Carbon fibres, in particular, can lead to ridging along the cutting edge, depending on the fibre orientation. In comparison to carbides, the diamond grains are affected more by the cobalt bonding phase dissolving than by the friction between the workpiece and the cutter [43]. This causes less wear of the tool flank than observed with carbide tools. What is more obvious with PCD tools is that the cutting edge begins to round after only a few metres of milled path: this can be partly attributed to manufacturing related defects. After further milling, the wear behaviour of the PCD tool stabilises and results in a tool life which is significantly longer than for carbide lools (Figure 25). PCD materials with a grain size of 1 - 2 pm can be used for milling FRP composites [44]. The cutting edge radius of a PCD cutter is between 6 and 8 pm depending on the grind. Tools with coarser grains, and thus with larger cutting edge radii, produce slightly poorer quality and have less wear resistance (Figure 26). This can be mainly attributed to the superior quality of fine-grained PCD cutters which is also reflected in superior cutter surface properties and lower notching. Despite the coarser material being harder, these factors have a crucial effect on the machined surface quality. Figure 26 shows that the surface quality achieved using a finer grain tool is 25% better than that for a coarser grain tool over the same milled path. Compared to carbides, the smaller cutting edge radius and notching as well as the hardness have a positive effect on tool wear.

-

PCD arain size 2 um edge radius 6.8 pm

width of wear mark [mm] 0 10

.-.

0 06

--

PCD grain size 6-10 pm edge radius 8 pm

/ Slot cutter head

Inserted cutter head

blade

coredrill bit

A

\

Figure 24: PCD milling heads for FRP composites [29].

Y

n

1

10 20 100 milled path [m]

18 roughness R, [pml

------

Carbide K10 grain size 1-2 pm edge radius 9 pm

x)\

0,02

I0

-

-.-- ------

---

2 20

60 100 milled path [m]

Figure 26: Comparison of different grain sizes and tool materials in milling of CFRP [29].

Quality-optimised cut edge

Burnt matrix material

Figure 28: Quality achieved when milling thermoplastic matrix GFRP composite laminates [29]. 150 machining force F,, F, [N] 0

100 50

m

m

n

u

u

0

Figure 27: Surface quality in milling of thermoset and thermoplastic matrix CFRP [6]. Surface quality as determined by the fracture behaviour of the FRP composites was studied by SEM microscopy in [6] (Figure 27). The 0" orientation fibres pulled out of the thermoset matrix composites and the heavily structured surface are clearly recognizable (Figure 27a). In contrast, the machining of thermoplastic matrix composites is more defined (Figure 27b). The surface structure is much more uniform due to the considerably higher strain at failure of the thermoplastic matrix which is between 60 and 100%. The fibres are also held firmly in place in the deformed matrix, which results in an altogether more cohesive chip. In [29], the quality of the cut surface during milling of thermoplastic matrix GFRP composites was studied. In this case, the quality indicators are influenced by thermal stresses which, in turn, are determined by the plastic matrix. The lower abrasive properties of the glass fibres imply that fraying is not as problematic here as in the case of carbon fibres. High cutting velocities combined with low feeds lead to thermoplastic matrix melting. Molten matrix can end up sticking to the surface of the composite laminate or to the chips which have been already cut off. If thicker materials or larger cutting depths cause this thermal stress to increase, the thermoplastic matrix could even burn (Figure 28). A significant reduction in cutting velocity does remedy this problem to a certain extent. Process parameters As regards process parameters selection in the milling of FRP composites, due to the multiplicity of possible composite structures, recommendations can only be given for specific machining applications. In most cases, however, a high cutting speed ranging from 800 to 1200 m h i n with a low to moderate feed per tooth proved advantageous. This is shown in Figure 29 for milling of GFRP using PCD tools [6]. The slight improvement in surface quality with increasing cutting speed beyond the range 800-1200 m/min is offset by a rise in feed force Ff. The latter increases notably, particularly at high feed rates per tooth. The result of this, especially with smaller tool diameters, is that a high level of thermal stress is exerted on the tool, in addition to purely mechanical stress imposed on the cutting edge. The thermal stress is attributable to the process heat which must be dissipated almost solely by the tool. Despite a highly dynamic content, the forces remain below those typical for metal cutting due to the nonhomogeneous material structure. Further major influences, with effects similar to the effect of feed rate, are the depth of tool engagement and the workpiece thickness.

-

0 , l mm

0

1000 2000 cutting speed v, [tdmin]

Figure 29: Cutting forces and surface roughness in milling of GFRP composites [6]. As in the case of the process parameters for GFRP and CFRP milling, those for AFRP milling can be specified only in very general terms. When carbide milling cutters are used, it is advisable to cut at a speed between 200 and 400 m h i n (Figure 30). Here too, it emerges that, like tool engagement depth, feed rate per tooth exerts a considerably greater influence on the forces and on the machining result than cutting speed. In contrast to milling of GFRP and CFRP, carbide tools coated with TIN allow for successful milling of AFRP. The coating reduces the coefficient of friction at the cutting area which, in conjunction with the enhanced abrasion resistance of the tools, results in longer service life [36].

15

machining force Ff, Ffn [N]

5 U

0

10 20 30 40 cutting speed vc [ m h i n ]

Figure 30: Influence of process parameters on cutting forces in milling of AFRP [6].

In addition to the fibre type, the fibre orientation with respect to the machining direction exerts a major influence on the cut surface quality. This has a particularly unfavourable effect on contour machining operations performed on unidirectionally reinforced parts. In such cases, a variety of surface qualities and machining forces are obtained depending on the orientation of the fibres to the machining direction [38, 45, 461. The type of polymer matrix exerts a further influence on the process, the greatest difference being between thermoset and thermoplastic matrix materials. In the case of fibre reinforced thermoset polymers, the best machining quality is achieved when milling is performed parallel to the fibre orientation (orientation angle 0"). At angles between 30" and 45", the fibre is cut by a combined compressive and bending strain which produces the highest forces and lowest surface qualities. Surface quality improves again as the fibre orientation angle increases. Fibre reinforced thermoplastic polymers display a somewhat different behaviour when milled. Tests carried out on unidirectional composite laminates with various types of fibre, show that the best surface quality is achieved when machining is conducted at an angle of 90" to fibre orientation [40]. When machining is performed at 0" to the fibre orientation, the composite laminate is cut at the fibre-matrix interface since cutting is carried out parallel to the fibres. In this case, individual fibres are torn out of the matrix. Milling operations carried out at 90" to the fibre orientation are characterized by a high level of plastic deformation of the matrix material, extending over the entire cut surface. The individual fibre cut edges are thus hardly recognizable. This partially melted and plasticised matrix material is typical of the machining result of cutting operations performed on thermoplastic matrix FRP composites. Additionally, it has been shown that tool wear, particularly in the case of carbide tools, is higher when thermoplastic matrix composites are milled than when the same operation is conducted on thermoset matrix composites. Satisfactory results are obtained only when PCD is used as a tool material. Machine tool reauirements If FRP composites are to be machined with a reliable level of high quality, both material-specific process control and careful machine tool selection are vital [6]. Fundamental requirements which must be met by any machine tool used to machine FRP composites can be formulated on the basis of the specific process technology (Figure 31). The high cutting speeds in conjunction with the relatively small tool diameters demand high spindle speeds. High frequency spindles permitting the rotational speed to be set within a wide range are, therefore, commonly used. This, in turn, imposes special requirements regarding axial run-out and concentricity accuracy of the spindle, the tool and the tool holding fixture. Accordingly, high feed rates are essential to ensure that the feed rates per tooth required at high cutting speeds are achieved. Due to the complex geometry of FRP composite parts and the generally small batch sizes, the machine tool should, ideally, have five degrees of freedom in motion and a flexible clamping fixture Due to the sensitivity of FRP composites to compressive stress, the clamping device should transmit the clamping pressure required to the workpiece in planar form. If required, the top layers can be supported in order to avoid damage to the FRP part. Since the material removed is highly abrasive and, in some cases, electrically conductive, guideways and bearings should be covered and electrical components should be separately ventilated.

Figure 31 : Machine tool requirements for reliable, highquality machining of FRP composites [6]. Processina and tool costs PCD tools are expensive but this can be balanced through the machining process and tool life. If a double-bladed, straight-grooved, 8 mm diam. PCD finger milling tool is considered, the PCD version is 12 times more expensive than the carbide tool. On the other hand, the wear resistance of PCD tools and their capability to withstand higher cutting speeds makes them a very attractive option. The following comparison between processing costs for PCD and carbide tools in milling two FRP composites is based on the cost per milled metre [6]. The following costs are considered: i) interest and repayments; ii) depreciation; iii) wages; iv) rent; v) energy costs; vi) maintenance costs; vii) tool costs. A cost of 250.000 Euro for the milling machine, including the casing around the machine and an extraction unit for the dust emitted, is assumed. The operating rate in a one-shift factory is 80%. The plant is to be written off in 6 years and the interest rate is 10% per year. The wages are set at 30 Euro per hour which corresponds to rate for a machine tool operator [47]. The overheads are set at 220%. The use of PCD tools for GFRP milling resulted in a tool life of 1900 m at a vc = 800 m/min. The tool life of the carbide tool was only 100 m at a vc = 400 m/min. Despite the PCD tool being much more expensive (325 Euro versus 30 Euro for the carbide tool), the more favourable processing conditions reduce the machining costs: PCD tools provide over 50% saving in comparison with carbide tools. This comparison is less positive in the case of CFRP for which the tool life is shorter: at slightly lower cutting speeds, PCD tools only provide an 18% saving in comparison with carbide tools. The effect of various factors on tool life and, thus, on the process cost should be taken into consideration when machining FRP parts. The example in Figure 32 shows that an increase in tool life leads to an almost 9% cost reduction. An increase in cutting parameters will, however, lead to a reduction in machining costs of up to 18%. 100 costina 1%1

I 80 1

15% increase in cutting parameters

70

1850

2000

2150

2300

2450 2600 tool life [m]

Figure 32: Cost reductions achieved by adjusting cutting parameters. Material: thermoplastic matrix GFRP; tool: PCD; cutting speed: 800 m/min [6].

3.2 Machining of metal matrix composites A large number of parts made of MMC are produced by machining for the aerospace industry, the sport equipment industry and the automotive industry (Figure 33). The main problem in machining MMC is the high tool wear, which under certain circumstances, leads to an uneconomical production process or makes the process impossible [48].

Figure 33: Applications of MMC materials in the automotive industry [48].

I

counterbody (workpiece)

Tribosystem

1

4 . -.. .:_

counterbody (chip)

surrounding (air, vacuum)

(abraded particles, reaction particles, lubricant)

body

The extensive tool wear is caused by the very hard and abrasive reinforcements. This may be explained by looking at the tribological system (Figure 34). The main reason for wear is the direct contact between the reinforcing particles or fibres and the cutting edge, which causes both a mechanical and a thermal load on the cutting edge. The dominant wear mechanism is abrasion, which is generated by impacts at the cutting edge and by the sliding motion of the particles relative to the rake and clearance face [49]. Additionally, a thermal load stresses the cutting edge. This thermal load results from hot spots which are generated by microcontacts between the cutting edge and the reinforcement. Despite the relatively low process temperature, this thermal load is limited by the melting temperature. Different wear mechanisms are responsible for the abrasive tool wear. These are known as microploughing, microfatigue, microcutting and microcracking [50]. A further important aspect regarding tool wear is the microstructural composition of the cutting material. Commonly used cutting materials like cemented carbide are also composites consisting of carbides as hard phase (WC) and cobalt as binder. Due to this, depending on the ratio of the properties of both the cutting material and the reinforcements (e.g. grain size to particle size), one of the different wear mechanisms predominates. When high performance MMC cutting is sought for, PCD tools characterized by a wear behaviour superior to the one of carbide tools should be used. The wear behaviour of PCD tools, however, is not as high as required and, because of their high cost, they do not seem to foster a real break-through in MMC machining. An interesting alternative is represented by cutting tools coated by chemical vapour deposition (CVD) of diamond. The development of this coating technology over the last decade has led to the growth of millimetre-thick self supporting layers over larger areas. Because CVD diamond is essentially pure carbon with no binder phase, the hardness of CVD diamond (I0000 HV) is much higher than the hardness of PCD (6000 HV). In combination with the gradient structure, CVD diamond thick layers utilized as cutting material are highly suitable for machining MMC. Tests with this new cutting material were carried out: when turning brake drums made of aluminium alloy AISi9Cu3 reinforced with Vf = 20% Sic particles, the flank wear of the cutting inserts could be reduced by using CVD diamond thick layers [51]. Moreover, boring tests of cylinder faces of motor units were carried out with similar results. The tool wear of the cutting inserts with a CVD diamond thick layer is also much lower in comparison with the wear of the PCD tools [51]. The high machining costs caused by the extreme abrasive tool wear can be reduced in the future by using CVD diamond thick layers as a cutting material. The CVD diamond technology might lead to a better productivity in MMC machining and thereby to a break-through of MMC material applications in the future. Turning of MMC In [52], the authors worked on turning of MMC with CBN, PDC, W C and DCC tools, compared the tool wear, and studied the effects of turning parameters. Moreover, in [5], an attempt to model the tool wear due to cumulative machining was made. In particular, the proposed equations model the cumulative wear in turning of MMC. In [53], it was first noted that tool wear of W C tools was independent of cutting speed and temperature, then a selfpropelled rotary WC tool was used to rough turn a Sic whisker reinforced Al alloy. The distribution of wear prolonged the WC tool to such an extent that its tool life was comparable to that of a PCD tool and the subsurface damage was comparable to that produced by a PCD tool.

n Stress on the cutting edge when machining metal matrix Mechanical loads: - considerable abrasion caused by contact with fibres or particles - high dynamic loads caused by reinforcement impacts at the cutting edge - alternating stress resulting from inhomogeneity of the workpiece material Thermal loads: - relatively low cutting temperature (limited by the melting temperature of the workpiece matrix material) - high local temperature generated by intensive microcontact between cutting edge and reinforcement Figure 34: Tribological system in the machining of MMC materials [48].

The independence of tool wear from cutting speed was shared among [53] and other authors (e.g. [54]). They found that cutting speeds in turning have negligible effects on tool life for both WC and PCD tools. The low temperature rise due to the increase in cutting speed when machining aluminium was thought to be the reason. This finding, however, contradicts the fact that abrasion damage is related to the rubbing speed, and abrasion is the major mechanism causing tool wear when machining MMC reinforced with abrasive particles such as Sic. In [54], however, an empirical relationship between tool wear and milling speed was found. The issue was more controversial when relationships with opposite trends were found in [55]. Here, a S i c particle reinforced Al alloy (A1 1050/SiC/25p) was turned with different types of W C tools and other ceramic tools. The wear rate of SIN tools was negative with cutting speed, but positive for other tools. In [56], a Sic particle reinforced Al alloy (AISiMg/SiC/14p) was rough turned with WC and coated WC tools under dry conditions while the same MMC was fine turned with PCD tools under wet conditions. Higher feed rates reduced tool wear due to thermal-induced softening of the MMC. The measured surface roughness values were far from the theoretical values. Again, different coatings on WC tools failed to protect the tools from abrasive wear. In [57], magnesium-based MMC (ZC71 reinforced with Vf = 20% S i c particles) was turned under dry conditions using TiN-coated, PCD-coated and PCD-tipped tools. While TiN-coated tools were destroyed immediately, PCDcoatings showed a good resistance against abrasion until the PCD film was torn. Chipping off of the layer could then be observed. A further improvement of the tool life of PCD-coated tools could be achieved by increasing the film thickness. In spite of the remarkable performance of the PCD layer, PCD-tipped tools showed the best results. In [58,591, blocky shaped Sic particles were burnished into the surface of a Mg alloy to add a reinforced top layer to the workpiece surface. By comparing the surface roughness of the turned unreinforced work material (R, = 1.5 mm) and of the same work material reinforced with S i c particles of different size, the achievable surface roughness was improved for small particles ( R a = 1.0 pm) and got worse for coarse particles (R, = 2.2 pm). However, for small particles, plastic deformation could be observed in the indentation tracks of a tribological testing and the reinforced layer was destroyed, whereas the layer with coarser particles showed a good wear resistance. In [49], the tool wear development during turning of MMC materials with carbide tools and PCD tools was studied. The work materials consisted of an Al alloy matrix reinforced with A1203 short fibres (Vf = 20%), SIC particles (Vf = 20%) or B4C particles (Vf = 10%). The main reason for tool wear was identified in the direct contact between the particles or fibres and the cutting edge and their motion relative to the rake and clearance face. Thus, the reinforcement hardness was considered a dominant factor for tool wear. In Figure 35, the flank wear of a medium grain carbide tool is reported vs. cutting time when turning the three different MMC. Although the volume fraction of B4C particles is only l o % , the wear rate is very high. When turning the A1203 short fibre reinforced Al alloy, the wear is relatively low. As the wear of carbide tools is very large when turning S i c or B4C reinforced Al alloy even with Vf as low as 10% or 20%, PCD tools appear particularly suitable for the turning of these MMC. In Figure 36, the flank wear of PCD tools is reported versus cutting time during turning of different Al alloy matrix MMC. Although the cutting speed increases from 100 m/min with carbide tools to 500 m/min with PCD tools, the tool wear decreases. The influence of reinforcement

hardness on the wear rate of PCD tools corresponds to that of carbide tools, but the wear rate increase due to the reinforcement hardness is not as high. In [30], the machinability of MMC materials was evaluated in turning by examining tool wear, cutting forces and surface finish. The examined MMC consisted in a 6061 Al alloy matrix reinforced with S i c whiskers (SiCw/6061 with Vf = 15%, 25%) or with A1203 fibres (A1203/6061 with Vf = 7.5%, 15%). Triangular inserts of WC, ceramic and sintered diamond were used as cutting tools.

Cutting time t,

-

Figure 35: Tool wear of carbide tools when turning Al alloy matrix MMC materials at 100 m/min cutting speed [49]. 600

1

w400

9 300 L

m

$

200

Y C

m

ii 100 0 200

0

400

600 800 Cutting time t,

1000

s

1400

I

Figure 36: Tool wear of PCD tools when turning Al alloy matrix MMC materials at 500 m/min cutting speed [49].

d

0.05

-

0

Figure 37: Effect of cutting speed on tool wear in turning of MMC materials with different tools [30].

Figure 37 shows the relation between cutting speed and tool wear of various cutting tools. Among the tested tools, the diamond tool has the smallest wear. Moreover, for carbide or diamond tools the wear is independent of cutting speed, while for ceramic tools the wear increases with increasing cutting speed. This phenomenon is observed only when the work material contains S i c whiskers, whereas in the other cases the wear of the ceramic tool is independent of cutting speed. The application of cutting fluid did not affect the tool wear. Also, when the feed rate was varied, the tool wear was not affected by feed rate and simply increased in proportion to the cutting length. From the above results, the tool wear in turning of MMC is considered to be caused only by mechanical abrasive wear. In the case of ceramic tools, Sic whisker reinforcement seems to have some effect on the cutting speed dependence of the tool wear, but the mechanism was not clarified. Figure 38 shows the effect of the reinforcement volume fraction on the tool wear. The tool wear and wear rate depend largely on the fibre type and volume fraction. Namely, tool wear increases with increasing fibre volume fraction and fibre hardness. This is exactly the same tendency as for FRP composite materials. Figure 39 shows the relation between surface roughness and cutting speed. The roughness of the matrix material is very large and varies with cutting speed, while that of the MMC is small and independent of cutting speed.

a) Diamond coated carbide

b) Polycrystalline-diamond

Figure 40: Roughness profile obtained during turning of GrA-Ni 10S.4G MMC material [60].

Figure 41: Pareto chart for average roughness analysis [60].

3

‘5

Cutting length L (m)

Figure 38: Effect of reinforcement volume fraction on tool wear during turning of MMC. Work: SiCw/6061; tool: carbide KIO; cutting speed: 60m/min; depth of cut: 0.5 mm; feed: 0.2 mm [30].

B

Cutting speed vc (mlmin)

Figure 39: Surface roughness vs. cutting speed in turning of MMC. Work: A1~03/6061,SiCw/6061; tool: cermet X407; depth of cut: 0.5 mm; feed: 0.15 mm [30].

In [60], the machinability of an Al alloy reinforced with S i c particles (Vf = 10%) and Ni coated graphite particles (Vf = 5%) (GrA-Ni 10S.4G) was investigated. This MMC material is designed for high wear resistance applications (cylinder liners and brake drums) as replacement for grey cast iron, where lower weight and high thermal properties are desired. Two diamond tool materials were used: PCD and diamond coated carbide. Turning tests were conducted to assess the surface finish of turned MMC parts. Figure 40 displays the roughness profiles of turned GrA-Ni 10S.4G parts obtained with different tool materials. The average roughness, Ra, was extracted and used for analysis. A Pareto chart was used to compare the effect of cutting parameter changes (feed, speed and depth of cut) on surface finish (Figure 41). For both PCD and DCC tools, the feed rate had a significant influence on the Ra values. As for metals, cutting speed and depth of cut did not show significant influences on Ra for both PCD and DCC tools. The cutting speed did not influence the Ra significantly because the range of cutting speeds was higher than the speed at which BUE are formed. At higher cutting speeds, unless chatter or vibration occurs, the surface finish is independent of speed. In most tests, PCD tools generated better finishes than DCC tools. Based on the results of statistical analysis, the average roughness Ra of GrA-Ni 10S.4G parts can be predicted using equations 1 and 2. Ra( DCC) = 12.1 (f’ 166/r);

R2 = 98.8%

Ra(PCD) = 7.23* (f’ 177/r);

R2 = 99.2%

(2)

where r is the tool nose radius (0.8 mm), f is the feed rate, and R2 is the coefficient of correlation. Figure 42 shows the Ra values obtainable when turning GrA-Ni 10S.4G at different feed rates using DCC and PCD tools. PCD tools generated a better finish than DCC tools because the PCD edge is sharper and its friction coefficient is lower. At very low feeds (- 0.1 mm/rev), Ra

values for surface finish generated using both PCD and DCC tools are similar. In most work on surface finish obtained during machining of composite materials [56], the average roughness is compared to that obtained using the theoretical model described by equation 3: Ra = 0.0321* (f?)

(3)

where Ra is the average roughness, r is the tool nose radius, and f is the feed rate. As noticed by [56], at lower feed rates, the experimental data obtained (Figure 41) during machining of composites is higher than the values predicted by the theoretical model (equation 3). Ra values are low because the tool nose radius becomes larger as the tool wears out. In general, the surface finish obtained with PCD tool is better than that obtained with DCC tools. In fact, the active faces of a PCD tool are smoother than those of a DCC tool and the rougher tool generates a poorer part finish. Drilling of MMC The results of machinability tests on a wrought spray deposited Al alloy MMC containing S i c particles with Vf = 15% were presented in [61]. The experimental work concentrated on conventional speed drilling and high speed drilling operations involving PCD tipped tvvlst drills. In [62], the authors, worked on the development and application of PCD tools used for drilling of MMC; they also presented the results from drilling tests with HSS, diamond-plated HSS, W C and TIN coated carbide tools. The drilling tests carried out on Al alloy A42618 reinforced by S i c particles with Vf = 15% showed that PCD tipped drilling tools were effective over a wide range of cutting conditions, offer the only realistic option in large batch and mass production, and cutting speed was not a significant factor affecting the tool life. Blind hole drilling is difficult if chips are not removed properly. When this is overcome, the wear of the drill edge is another problem when drilling MMC materials. In [54], it was found that carbide drills were not suitable for Al alloy 2618 reinforced with S i c particles (2618/SiC/15p), but PCD tipped drills were. The drilling speed did not affect the drill life, but higher feed rates allowed for a higher number of holes per drill [63]. In [64, 651, it was found that increasing the drilling speed accelerated the wear of WC drills, but had no effect on the wear of PCD tvvlst drills in the range of 15-300 m/min. The Al alloy 2014 reinforced with S i c particles was used in a comparative drilling test in [66]. The thrust force decreased at higher speeds and, therefore, with higher drill wear. This contradicts the trend found by others (e.g. [64, 651). An attempt to increase the thrust force by drilling with different tools was made by [67], but no significant improvement on drill life was achieved.

Figure 42: Average roughness Ra vs. feed for DCC and PCD tool materials [60].

In [68], a new type of microdrill made of fine carbide grains (2 pm) was used to drill an Al alloy reinforced with Sic particles (Vf = 12% and 20%). Drilling was carried out to study the performance of the new drill types. High speed drilling conditions were: spindle speed 20,000 rpm, feed rate 5 - 8 mm/s, coolant applied, drill diam. 0.5 - 1.0 mm. Holes drilled with the lower feed rate had better peripheral waviness for both types of MMC. Results showed that the MMC with higher Vf has a higher tool wear. The best life prediction was Taylor's tool life equation VT" = C. Ex erimental data were fitted to get VTo7' = 4969 and V?" = 2028 for Vf = 12% and 20%, respectively. Both n and C are higher than for carbide tools and correspond to ceramic tools, possibly due to the higher performance and lower tool life. The difference may also be due to the uneven distribution of hard S i c particles in the MMC causing impact and the fact that the heat produced softens the matrix and makes impact even more serious. In [30], the effect of machining parameters on drill wear during drilling of SiCw/6061 and A1203/6061 MMC materials was studied. Figure 43 shows the effect of feed rate on HSS drill wear in drilling a through hole in a 6 mm thick board. The flank wear decreases with increasing feed rate due to the decrease of total cutting length. The drill wear was also independent of spindle speed and decided mainly by the kind and content of fibres in the MMC. The effect of tool material on drill wear was also examined showing that, compared with HSS drills, the wear of carbide drills is very small (Figure 44).

Figure 43: Effect of feed on drill wear in drilling of SiCw/6061 and A1~03/6061with different fibre contents. Tool: standard HSS drill; drill diam.: 6 mm; thickness of board: 6 mm; numbers of hole: 1 [30].

Figure 44: Tool material effect on wear in drilling of SiCw/ 6061 (Vf= 25%). Workthickn.: 6 mm; feed: 0.2 mm; spindle speed: 415 rpm; drill diam.: 6 mm [30].

In [69], dry drilling was used to evaluate the machinability of a new family of MMC materials consisting of an Al alloy reinforced with Ni coated graphite particles and ceramic particles (GrA-Ni@). Machinability was assessed on the basis of cutting force when drilling with HSS tvvlst drills. The better machinability of GrA-Ni@verified over existing Sic particle reinforced MMC, such as AI-9Si.20SiC, can be explained by the following points-[70]: a Low Sic content in GrA-NiePcompared to Sic content in AI-SSiC.20SiC, although the total volume fraction of particles in the two MMC composites is roughly equivalent (GrA-Ni@:10% Sic, 6% AkNi, 4%Gr). a AhNi and graphite particles found in GrA-Ni@are not as hard as SIC or A1203 particles. a The hardness of GrA-Ni@alloy is lower (-70 HRB) than that of AI-9Si.20SiC (-75 HRB) [71]. Low values of hardness and strength are favourable for machinability, except in very ductile materials, as they may generate BUE, burrs, and poor finish. a AhNi present in graphitic MMC is a brittle compound and leads to a reduction of material ductility. The lower the ductility, the lower the energy required to shear the metal and the easier the chip breakage. a Graphite flakes present in the aluminium matrix act as a solid lubricant [72, 731. To demonstrate the wav in which different wear mechanisms cause tool wear, SEM photographs were taken to show worn carbide cutting edges used for drilling of fibre and particle reinforced Mg alloys (Figure 45). In both cases, a very regular flank wear is typically formed [74]. When drilling &A1203 short fibre reinforced Mg alloy, the dominant wear mechanisms are microcracking and fatigue. Because of the low hardness of &A1203 short fibres of approximately 800 HV, it is not possible for the short fibre to abrade the carbide by microcutting. The hard phases of the cutting tool are abraded due to microcracking and fatigue at the cutting edge. These particles, which are separated from the substrate, contribute to the amount of wear because they slide over the rake and clearance face. They act like a polishing compound because of their small grain size of approximately 1-3 pm and cause a very smooth topography of the worn surface. Some larger cracks on the cutting edge could also be observed.

(a)

(b)

Matrix material: AZ 91 Reinforcement: (a) 20% &A1203 short fibres (b) 20% Sic-particles Tool: Cemented carbide K10 Drill hole length j = 20 rnrn Drill length: Lf = 400 rnrn Cutting speed vc = 75 rn/rnin Feed: f = 0.1 rnrn Coolant: Dry machining

Figure 45: SEM photographs of worn cutting edges [48]

m

> L

m

Y C

m LL

Drilling length

L,

Figure 46: Tool wear when drilling two Mg alloy MMC reinforced with either &A1203 fibres or SIC particles [48] The tool wear which results from drilling MMC materials containing harder reinforcements is different. The topography of the worn surface is characterised by grooves. These grooves are mainly oriented parallel to the cutting direction. Since the hardness of cemented carbide is much lower than the hardness of S i c (2400 HV), these particles abrade the carbide tool by microcutting as shown by the grooves on the worn clearance face. This wear mechanism leads to extensive tool wear. The only possibility of limiting tool wear is to use harder cutting tool materials, such as PCD. But considering the high cost of PCD drills, alternatives such as drills with coatings based on TiAlN or diamond should be considered. To examine the efficiency of these different tool materials, drill tests were carried out. The results are reported in Figure 46, showing the development of wear of the coated and uncoated tools as a function of drilling length when machining Mg alloy containing different reinforcements. Short &A1203 fibres cause only a slight deterioration of the tool depending on the hardness of the reinforcement. The carbide tool shows the highest wear because of its low hardness compared to the reinforcement. The best results were obtained with TiAIN-coated tools. Surprisingly, diamond-coated tools are unfavourable. The latter tools show more wear if tool wear is predominantly due to micro-cracking and fatigue, which occurs when the &A1203 fibres hit the cutting edge. In this case, the TiAlN-coating shows the best results. When drilling the ZC 63 Mg alloy reinforced with 12% S i c particles, the wear of the uncoated and the TiAlN coated drills increases progressively due to the high hardness of the Sic particles and causes an early end of the tests. In this case, the diamond-coated tools show the best results: after a drilling length of L = 400 mm, no tool wear could be measured. Thus, it can be confirmed that when tool wear is due to microcutting, diamond coated tools are favourable. Besides tool wear, a further important factor is the surface integrity of the MMC material. Damage of the reinforcements, caused by the machining process, can lead to a decrease in the properties of the MMC parts. The cutting and accompanying deformation mechanism affect the surface quality in a multitude of ways. Examples of this are damage in the form of microcracks, pores, microstructural transformation, plastic deformation, residual stresses and changes in hardness.

+I Matrix material: AZ 91 (MgAI9Znl) Reinforcement: 20% A1203 short fibres Cutting tool: cemented carbide drill, 0 6 rnrn Drill hole length: I = 20 rnrn Cutting speed: vc = 75 rn/rnin

Figure 47: Surface integrity of a drilled A1203 fibre reinforced Mg alloy [48].

Drill hole depth: I = 20 rnrn Cutting speed: vc = 75 rn/rnin Feed: f = 0.1 rnrn

Figure 48: Surface integrity of a drilled Sic particle reinforced Mg alloy [48]. To analyse the influence of machining on the subsurface zone of MMC materials, SEM and optical microscope analyses of the machined surfaces and of special cross sections were carried out. Figure 47 shows a SEM photograph of the subsurface zone of a Mg alloy reinforced with short A1203 fibres. A carbide drill was used at a cutting speed of vc = 75 mm/min and a feed rate o f f = 0.1 mm. Fibres lying up to approximately 50 pm below the surface are shown to be fractured. Figure 48 shows SEM photographs of the subsurface zone of a Mg alloy reinforced with S i c particles. Carbide and PCD drills were used at a cutting speed of v c = 75 mm/min and a feed rate of f = 0.1 mm. In the case of the subsurface machined with the carbide drill, the S i c particles lying up to approx. 40 pm below the surface are fractured as a result of plastic deformation. On the other hand, the S i c particles of subsurface zones machined with the PCD drill are almost cleanly cut and nearly undamaged because of the very sharp cutting edge of the PCD tool. Milling of MMC In [54], face milling of an Al alloy reinforced with Sic particles (2618/SiC/15p) using carbide and PCD tools was studied. It was found that cutting speed contributed to faster flank wear rate in milling, but the speed had little effect in turning and drilling. Milling of particle and fibre reinforced MMC materials was performed in [65] with PCD tools. BUE were observed when the feed rate was above 0.6 mmhooth, and chipping occurred when these BUE were sheared. Coolant was recommended at high feed rate at the expense of tool life.

Tool wear was found to increase with cutting speed, though chipping occurred at speeds below 250 m/min. The fact that no chipping was reported in [54] when milling at 114 m/min indicates that chipping is tool dependent rather than speed dependent. Flank wear was lower for the coarse-grain PCD at feed rates above 0.25 mmhooth, but the reverse trend was found at lower feed rates. Surface finish worsened with higher feed rate as expected, but was constant for milling speeds between 250 and 1000 m/min. In [63], cast MMC materials were end-milled and facemilled with different tools. HSS and coated tools were not economically justifiable for the task. Carbide tools had limited tool life even at the low cutting speed of 30 m/min. TIC or TIN coatings offered little advantage. End-milling of MMC was studied by [75], finding that 100 pm wears on PCD tools were established in minutes during the "break-in" period regardless of the matrix and reinforcement compositions. Surface roughness measurement was tried as alternative tool wear criterion, but inconsistency of measurement was found instead. In [75], the wear of PCD tools was reported to be high when face-milling between 600-800 m/min and the wear reduced when tools were used outside that speed range. In [69], the machinability of a new family of MMC materials consisting of an Al alloy reinforced with Ni-coated graphite particles and ceramic particles (GrA-Ni@)was assessed in terms of tool life and chip form. The work materials consisted of an Al alloy matrix with the following types and contents of reinforcement: 0 GrA-Ni 5A.4G: 5 vol% A1203, 4 vol% Ni-Gr. 10 vol% Sic, 4 vol% Ni-Gr. 0 GrA-Ni 10S.4G: 6 vol% Sic, 2.5 vol% Ni-Gr. 0 GrA-Ni 6s-2.5G: 0 AI-9Si.20SiC: 20 vol% Sic. Dry milling tests were conducted using TiCN coated carbide inserts. Figure 49 displays the wear curves observed when milling at a cutting speed of 61 m/min. GrA-Ni 5A.4G showed the lowest wear rate compared to all the other MMC materials. The incorporation of Nicoated graphite particles reduces the tool wear rate. When the reinforcing particles are alumina and Ni-coated graphite, the wear rate is quite similar to that observed during machining of high silicon content Al alloy, such as Al 380, but the wear progresses quicker at the beginning when cutting alumina graphitic MMC. A material with high wear rates such as AI-9Si.20SiC is more abrasive since the wear mechanism is pure abrasion due to the harder particles that grind the flank face of the tool. The higher the content of harder particles (e.g. Sic), the higher the wear rate and thus the poorer the machinability. On the other hand, the higher the content of graphite, the better the machinability since the graphite will lubricate the cutting process.

h

E E

v

m

m

> .. L

m

$

Y

C

m LL

0

50

100 150 200 Cutting time (rnin)

250

300

Figure 49: Wear progression during milling of various MMC materials at 61 m/min cutting speed [69].

1000

0

GrA-Ni5A.4G

1

Figure 52 compares the resultant cutting forces obtained during milling with both tool materials. As expected, DCC inserts (active face rougher than that of PCD) generated higher cutting forces. The cutting forces are also higher during milling with DCC tools because the cutting edge is slightly honed prior to the deposition of the diamond film.

0 0

I - 1 30

40

50

60

70 80 90100

Cutting Speed

Vc

200

(rnlrnin)

Figure 50: Tool life vs. cutting speed during milling of various MMC of the GrA-Ni@family [69]. The tool life was obtained graphically from the wear curves. The cutting speed vs. tool life relationship for each material is reported in Figure 50 and shows that: 0 The wear mechanism of Sic-reinforced GrA-Ni@ materials is similar (abrasion of the clearance face). Thus, the slope of the Taylor relation is quite comparable. 0 All these MMC have a high Taylor exponent (0.69 < n < 0.95) which denotes the low influence of cutting speed on tool life. 0 Alumina reinforced graphitic MMC, such as GrA-Ni 5A.4G, can be machined at higher metal removal rates than other Sic reinforced graphitic MMC and Al alloy MMC reinforced with S i c particles only. This can be explained by the fact that alumina particles are softer than Sic particles, thus generating a low abrasive load on cutting tools.

Figure 52: Cutting force vs. feed rate and tool material during milling of GrA-Ni 10S.4G MMC material with PCD and DCC tools [60].

E E m

v

> In [60], the machinability of GrA-Ni 10S.4G with two diamond tool materials (PCD and diamond coated carbide) was studied, finding that DCC tools with good adhesion properties outperform PCD tools in terms of productivity, while PCD tools present a more predictable performance. Figure 51 compares the performance of DCC inserts that did not fail prematurely with that of PCD inserts. The first DCC insert (#I) had a high initial wear rate at the beginning of the cut, but the wear stabilised at about 0.2 mm flank wear. From this point, the wear did not increase for a long period of cut. This behaviour can either be attributed to the fact that the peaks usually found on the DCC tool active face wear off or that the layer beneath the coating is stronger than the surface. The second DCC insert (#2) performed better than the PCD insert at the beginning of the cut, then the flank wear increased and stabilised at about 0.12 mm.

Milling duration, t (s) Figure 53: Wear of carbide tool when milling a range of CeC alloys at two feed levels (f = 0.19 mm and f = 0.38 mm) and a cutting speed of vc = 15 m/min [77].

h

E E

Y

m

>

0

Figure 51: Tool wear vs. volume of metal removed during milling of GrA-Ni 10S.4G MMC material with PCD and DCC tools [60].

60

120 180 Milling duration, t (s)

240

Figure 54: Wear of PCD tool when milling a range of CeC alloys at two feed levels (f = 0.19 mm and f = 0.38 mm) and a cutting speed of vc = 1178 m/min [77].

In [77], symmetric single tooth dry face milling tests was carried out on centrifugal cast tubes made of an Al alloy reinforced with Sic particles (CeC material). Milling forces were measured and the quality of the milled surface was evaluated in addition to normal tool wear measurements. Cutting speeds of 15 and 1178 m/min were used respectively with carbide tools and PCD tools, using two feed rate values and a constant depth of cut of 1.5 mm. The evolution of flank wear during milling of the CeC material is shown in Figures 53 and 54. The flank wear was most severe in both cases at the higher feed for the MMC material with 38% S i c reinforcement. Furthermore, with PCD tools (Figure 54) the least wear was obtained at a feed of 0.18 mm/tooth for the MMC with Vf = 23% Sic, indicating the possibility of an MCC with optimal particle content for improved machinability. 4 CONCLUSIONS Machining of composite materials differs significantly in many aspects from machining of conventional metals and their alloys. In the machining of composites, the material behaviour is not only non-homogeneous, but it also depends on diverse reinforcement and matrix properties, reinforcement architecture and orientation, and relative content of matrix and reinforcement. The tool encounters continuously alternate matrix and reinforcement materials, whose response to machining can be entirely different. Accordingly, the machining of composite materials imposes special demands on the geometry and the abrasion resistance of the cutting tool. As regards plastic matrix composites (PMC), and in particular fiber reinforced plastics (FRP), tool wear mechanisms and development during machining need to be attentively considered to establish correct cutting tool selection criteria. Among the possible wear mechanisms, only abrasion, surface damage and sometimes adhesion are of significance for the machining of FRP materials. Tool wear mechanisms are primarily related to the physical and mechanical characteristics of the different fiber matrix systems. Glass and carbon fibers show a strongly abrasive behaviour because they are extremely abrasive by nature. Aramid fibers, on the other hand, impair the tool because of their low heat conductivity and their ductile characteristics. Adhesive wear occurs when carbonized or molten matrix depositions settle on the tool surfaces. The fiber type, the reinforcement architecture and the matrix content are the most important factors governing tool selection and machining parameters setting. In the case of glass fiber reinforced plastics (GFRP) and carbon fiber reinforced plastics (CFRP), it is the cutting tool material that dominates the tool selection. In the case of aramid fiber reinforced plastics (AFRP), it is the tool geometry that dictates the choice of the cutting tool. The hardness of the glass and, more especially, of the carbon fibers results in a high level of tool wear in machining operations. Since this wear manifests itself above all in tool cutting edge rounding, the cutting edge should possess a high degree of resistance to abrasion and chipping. Moreover, to guarantee that the fibers are severed in a clean cut, it is very important to ensure a very high cutting edge sharpness. Carbide tools, coated carbide tools and PCD tools yield good results in terms of tool wear and tool life during the machining of GFRP and CFRP, although the wear for these tool materials is considerably different. Moreover, the superiority of diamond-based cutting materials, PCD and diamond-coated carbide, over monolithic carbide tools is clearly demonstrated. Ceramic materials are unsuitable because their low strength and high brittleness make them very sensitive to shocks and their low heat conductivity

does not allow for the dissipation of the machining heat. Due to its low wear resistance, CBN, which is as expensive as PCD, presents no advantage over the latter. A comparison between carbide, DCC and PCD tools reveals that much higher cutting speeds can be used with PCD tools. Furthermore, at higher cutting speeds, the life of DCC tools is surpassed by PCD tools, the degradation in adhesive properties of the diamond layer to the carbide substrate being the main reason. Finally, by comparing PCD tools to carbide tools in terms of both economics and quality, PCD can be an economical alternative to carbide in the machining of FRP despite the much higher purchase cost because tool life is longer and higher processing speeds can be used. Due to the pronounced susceptibility of glass and carbon fibres to brittle fracture, tool geometries in the machining of GFRP and CFRP composites correspond approximately to those of the tools used in metal working. In contrast, the standard tool geometries cannot be applied to machining of AFRP because the individual aramid fibres can be separated in a clean cut only under simultaneous prestress. Accordingly, a special tool geometry must be designed to allow for prestressing the aramid fibres before the cutting process begins. Also the type of polymer matrix exerts a significant influence on the process, the greatest difference being between thermoset and thermoplastic matrix materials. For the same machining operation, tool wear is higher for thermoplastic matrix FRP rather than for thermoset matrix FRP. In the machining of thermoplastic matrix FRP composites, satisfactory results are obtained only when PCD is used as tool material. As regards metal matrix composites (MMC), though challenging to manufacturing engineers and research workers at the beginning, MMC are machinable using conventional cutting processes, although particular care must be exercised. PCD tools can be used effectively to machine MMC in turning, drilling, milling, etc. The less expensive diamond-coated tools emerge as a promising alternative to solid diamond tools if adequate coating adhesion is assured. Tool wear and tool life are rather independent of cutting parameters and mostly dependent on tool geometry, tool material and work material. MMC machinability is critically affected by the reinforcement and matrix hardness: higher hardness shortens the tool life critically. CBN and PCD tools are one and two orders of magnitude better than carbide tools in terms of wear resistance. While carbide tools could be used economically for roughing operations, PCD tools should be used for finishing operations because of their longer tool life. TIN coating of carbide tools does not improve tool life but proper diamond coating on complex geometry tools is by far the most promising tool technology. Besides tool wear, a further critical factor is the surface integrity of machined MMC. Damage of the reinforcement in the subsurface zone, caused by the machining process, can lead to a reduction in the properties of the finished MMC part. When using PCD tools, the reinforcement in the subsurface zone is generally almost cleanly cut and nearly undamaged because of the very sharp cutting edge of the PCD tools. In the case of carbide tools, the reinforcement in the machined subsurface can be fractured as a result of plastic deformation. In any case, fractured and pulled out particles are practically unavoidable in MMC machining, and the subsurface damage is more substantial in whisker or fibre reinforced MMC than in particle reinforced MMC materials. Water based cutting fluids help reducing the built-up-edge formation but fail to significantly improve tool life. Inadequate flushing of the chips reduces tool life because the abrasive chips tend to grind the tool flank.

Figure 55: Principle of laser assisted machining [78]. Finally, as regards the machining of ceramic matrix composites (CMC), as mentioned in section 3, the very small number of contributions received and the scarcity of information available in the open literature did not allow for the preparation of a section dedicated to this topic. Among the very few contributions received, it is worth mentioning the research activity on laser-assisted machining of ceramics and CMC materials reported in [78] (Figure 55). Experimentations were carried out on unreinforced SIN ceramics and CMC materials. The conclusions on the effectiveness of laser-assisted machining report that SIN ceramics demonstrate excellent potential for benefiting from the advantages of laserassisted machining. However, investigations on CMC materials lead to the conclusion that, in spite of some advantages in tool wear development in comparison with conventional turning, the overall cutting length is too short for an efficient machining of these materials. In other words, the CMC material was very difficult to machine and practically remained so even when the laser-assisted machining technology was applied. The reason for the scarcity of information on the topic of CMC materials machining could, most probably, be attributed to the fact that these materials are actually very difficult to machine, making research activities hard to carry out and the few experimental results obtained, especially in industrial research context, are not easily publicised. The need for quantitative data and information on CMC machining is, however, decidedly felt, calling for an increased research effort in the area of conventional machining of CMC materials. 5 REFERENCES [ I ] Garlasco, F.M., 1969, Machining of Boron Fibrous Composites, American Society of Manufacturing Engineering, SME, TP EM: 69-74. [2] Maclean, B.J., Misra, M.S., 1983, Mechanical Behavior of Metal Matrix Composites, Eds. Hack, J.E., Amateau, M.F., TMS-AIME 301. [3] Konig, W., Wulf, C., Grap, P., Willerscheid, H., 1989, Machining of Fibre-Reinforced Plastics, Annals of CIRP, 3412: 537-548. [4] Komanduri, R., 1997, Machining of Fibre-Reinforced Composites, Int. J. of Machining Science and

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