ACI 224R-1990

ACI 224R-90

Control of Cracking in Concrete Structures Reported by ACI Committee 224

The principal causes of cracking in concrete and recommended crack control procedures are presented. The current state of knowledge in microcracking and fracture mechanics is discussed. The control of cracking due to drying shrinkage and crack control for flexural members, layered systems and mass concrete are covered in detail. Longterm effects on cracking are considered, and crack control procedures used in construction are presented. Information is provided to assist the engineer and the constructor in developing practical and effective crack control programs for concrete structures. Keywords: adiabatic conditions; aggregates: air entrainment; anchorage (structural); beams (supports); bridge decks; cement-aggregate reactions; cement content; cement types; compressive strength: computers; concrete construction; concrete pavements; concrete slabs; concretes; conductivity: consolidation; cooling; crack propagation; cracking (fracturing); crack width and spacing: creep properties; diffusivity; drying shrinkage; end blocks; expansive cement concretes; extensibility; failure; fibers; heat of hydration; insulation; joints (junctions); machine bases; mass concrete; microcracking; mix proportioning; modulus of elasticity; moisture content; Poisson ratio; polymer-portland cement concrete; pozzolans; prestressed concrete; reinforced concrete; reinforcing steels; restraints; shrinkage: specifications; specific heat; strain gages; strains; stresses; structural design; temperature; temperature rise (in concrete); tensile stress; tension; thermal expansion; volume change.

Contents Chapter 1 - Introduction, page 224R-2 Chapter 2 - Crack mechanisms in concrete, page 224R-2 2.1 - Introduction 2.2 - Microcracking 2.3 - Fracture

Chapter 3 - Control of cracking due to drying shrinkage, page 224R-9 3.1 3.2 3.3 3.4 3.5 3.6

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Introduction Crack formation Drying shrinkage Factors influencing drying shrinkage Control of shrinkage cracking Shrinkage-compensating concretes

Chapter 4 - Control of cracking in flexural members, page 224R-16 4.1 4.2 4.3 4.4

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Introduction Crack control equations for reinforced concrete beams Crack control in two-way slabs and plates Tolerable crack widths versus exposure conditions in reinforced concrete 4.5 - Flexural cracking in prestressed concrete 4.6 - Anchorage zone cracking in prestressed concrete 4.7 - Tension cracking

Cbapter 5 - Long-term effects on cracking, page 224R-21

ACI Committee Reports, Guides, Standard Practices , and Commentaries are Intended for guidance in designing, planning, executing, or inspecting construction, and in preparing specifications Reference to these documents shall not be made in the Project Documents. If items foun d in these documents are desired to be part of the Project Documents, they should be phrased in mandatory language and incorporated into the Project Documents.

5.1 5.2 5.3 5.4 5.5

Copyright 0 1990 , American Concrete Institute. All rights reserved including rights of reproduction and use in any form or by any means, including the making of copies by any photo process, or by any electronic or mechanical device, printed or

224R-1

-

Introduction Effects of long-term loading Environmental effects Aggregate and other effects Use of polymers in improving cracking characteristics

written or oral, or recording for sound or visual reproduction or for use in any knowledge or retrieval system or device, unless permission in writing is obtained from the copyright proprietors.

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Chapter 6 - Control of cracking in concrete layered systems, page 224R-23 6.1 6.2 6.3 6.4

-

Introduction Fiber reinforced concrete (FRC) overlays Latex modified concrete (LMC) overlays Polymer impregnated concrete (PIC) systems

Chapter 7 - Control of cracking in mass concrete, page 224R-26 7.1 7.2 7.3 7.4 7.5 7.6 7.7

-

Introduction Crack resistance Determination of temperatures and tensile strains Control of cracking Testing methods and typical data Artificial cooling by embedded pipe systems Summary - Basic considerations for construction controls and specifications

Chapter 8 - Control of cracking by correct construction practices, page 224R-36 8.1 8.2 8.3 8.4 8.5 8.6 8.7 -

Introduction Restraint Shrinkage Settlement Construction Specifications to minimize drying shrinkage Conclusion

Chapter 9 - References, page 224R-42 9.1- Specified and/or recommended references 9.2 - Cited references

Chapter 1 - Introduction Cracks in concrete structures can indicate major structural problems and can mar the appearance of monolithic construction. They can expose reinforcing steel to oxygen and moisture and make the steel more susceptible to corrosion. While the specific causes of cracking are manifold, cracks are normally caused by stresses that develop in concrete due to the restraint of volumetric change or to loads which are applied to the structure. Within each of these categories there are a number of factors at work. A successful crack control program must recognize these factors and deal with each of them, in turn. This report presents the principal causes of cracking and a detailed discussion of crack control procedures. The body of the report consists of seven chapters designed to help the engineer and the contractor in the development of effective crack control measures. This report is an update of a previous committee report, issued in 1972.1.1 The original report was supplemented by an ACI Bibliography on cracking,1 . 2 also issued by this committee. In the updating process, many portions of the report have undergone sizeable revision, and the entire document has been subjected to a detailed editorial review. Chapter 2, on crack mechanisms, has been completely rewritten to take into account the experimental and analytical work that has been done since the completion of the first committee report. Chapter 6, on crack control in concrete layered systems, is new to the report and deals with a form of concrete construction that was in its infancy at the time the first report was drafted. Individual chapters on crack control in re-

inforced and prestressed concrete members have been condensed into a single chapter, Chapter 4, on crack control in flexural members. The resulting presentation is more concise and, hopefully, more useful to the structural designer. Chapter 5, on long-term effects, details some interesting findings on the change of crack width with time. Chapters 3, 7, and 8, which consider drying shrinkage, mass concrete, and construction practices, respectively, have been expanded and updated to take into account the most recently developed procedures in these areas. In addition, new sections have been added to Chapters 7 and 8 which provide specific guidance for the development of crack control programs and specifications. The committee hopes that this report will serve as a useful reference to the causes of cracking and as a key tool in the development of practical crack control procedures in both the design and the construction of concrete structures. References 1.1. ACI Committee 224, “Control of Cracking in Concrete Structures,” ACI JOURNAL, Proceedings V. 69, N O . 12, Dec. 1972, pp. 717-753. 1.2. ACI Committee 224, “Causes, Mechanism, and Control of Cracking in Concrete,” ACI Bibliography No. 9, American Concrete Institute, Detroit, 1971, 92 pp.

Chapter 2 - Crack mechanisms in concrete* 2.1 - Introduction Beginning with the work at Cornell University in the early 1960s,2 .1 a great deal has been learned about the crack mechanisms in concrete, both at the microscopic and the macroscopic level. Of special interest during the early work was the realization that the behavior of concrete, under compressive as well as tensile loads, was closely related to the formation of cracks. Under increasing compressive stress, microscopic cracks (or microcracks) form at the mortarcoarse aggregate boundary and propagate through the surrounding mortar, as shown in Fig. 2.1. During the first decade of research, a picture developed that closely linked formation and propagation of these microcracks to the load-deformation behavior of concrete. Prior to load, volume changes in cement paste cause interfacial cracks to form at the mortar-coarse aggregate boundary.2.2,2.3 Under shortterm compressive load, no additional cracks form until the load reaches approximately 30 percent of the compressive strength of the concrete.2.1 Above this value, additional bond cracks initiate throughout the matrix. Bond cracking increases until the load reaches approximately 70 percent of the compressive strength, at which time microcracks begin to propagate through the mortar. Mortar cracking continues at an accelerated rate until the material ultimately fails. For concrete in uniaxial tension, experimental work indicates that major microcracking begins at about 60 percent of the ultimate tensile strength.2.4 ‘Principal author: David Darwin.

CONTROL OF CRACKING

Studies of the stress-strain behavior and volume change of concrete 2.5 indicate that the initiation of major mortar cracking corresponds with an observed increase in the Poisson’s ratio of concrete. The term “discontinuity stress” is used for the stress at which this change in material behavior occurs. In general, it has been agreed that the microcracking that occurs prior to loading has very little effect on the strength of concrete. However, work by Brooks and Neville 2.6 indicates that the effect of early volume change on microcracking of concrete may result in a reduction of both tensile and compressive strength as concrete dries out. Their study shows that upon drying, the strength of test specimens first increases and then decreases. They postulate that the initial increase is due to the increased strength of the drier cement paste and that the ultimate decrease in strength is due to the formation of shrinkage induced microcracks. Work by Meyers, Slate, and Winter 2.7 and Shah and Chandra2.8 demonstrates that microcracks increase under the effect of sustained and cyclic loading. Their work indicates that the total amount of microcracking is a function of the total compressive strain in the concrete and is independent of the method in which the strain is applied. Sturman, Shah, and Winter2.9 found that the total degree of microcracking is decreased and the total strain capacity in compression is increased when concrete is subjected to a strain gradient. At about the same time that the microcracking studies began, investigators began applying fracture mechanics to the studies of concrete under load. The field of fracture mechanics, originated by Griffith2.10 in 1920, serves as the primary tool for the study of brittle fracture and fatigue in metal structures. Since concrete has for many years been considered a brittle material in tension, fracture mechanics is considered to be a potentially useful analysis tool for concrete by many investigators. 2. .12 The field of fracture mechanics was first applied to concrete by Kaplan 2.11 in 1961. The classical theory serves to predict, the rapid propagation of a macrocrack through a homogeneous, isotropic, elastic material. The theory makes use of the stress intensity factor, KI , which is a function of crack geometry and stress. Failure occurs when KI reaches a critical value, K Ic , known as the critical stress-intensity factor under conditions of plane strain. KIc is thus a measure of the fracture toughness of the material. To properly measure KIc for a material, the test specimen must be of sufficient size to insure maximum constraint (plane strain) at the tip of the crack. For linear elastic fracture mechanics (LEFM) to be applicable, the value of KIc must be a material constant, independent of the specimen geometry (as are other material constants such as yield strength). The earliest experimental work utilized notched tension and beam specimens of mortar and con-

224R-3

$EGlrgyj m

0.0012 STRAIN

m

STRAIN

0. CKI

Fig. 2.1 - Cracking maps and stress-strain curves for concrete loaded in uniaxial compression. * *From S. P. Shah, and F. O. Slate, “Internal Microcracking, Mortar-Aggregate Bond and the Stress-Strain Curve of Concrete,” Proceedings, International Conference on the Structure of Concrete (London, Sept. 1965), Cement and Concrete Association, London, 1968, pp. 82-92.

crete. 2.11-2.14 The crack resistance was expressed in terms of the strain energy release rate at the onset of rapid crack growth, G, which is directly related to the fracture toughness of the material. Later investigations evaluated the crack resistance of paste, mortar and concrete in terms of the fracture toughness, itself.2.15 Work by Naus and Lott2.16 indicated that the fracture toughness of paste and mortar increased with decreasing water-cement ratio, but that the water-cement ratio had little effect on the fracture toughness of concrete. They found that KIc increased with age, and decreased with increasing air content for paste, mortar, and concrete. The effective fracture toughness of mortar increased with increasing sand content, and the fracture toughness of concrete increased with an increase in the maximum size of coarse aggregate. Additional work by Naus,2.17 presented just prior to the previous committee report,1.1 indicated that fracture toughness was not independent of specimen geometry for tensile specimens of paste, mortar and concrete and that fracture toughness was a function of the crack length. These observations lead to the possibly erroneous conclusion that fracture mechanics may not be applicable to concrete. Because certain size requirements must be met, before fracture mechanics is applicable, these results may only indicate that the test specimen did not satisfy all of the minimum size requirements of linear elastic fracture mechanics. The balance of this chapter describes some of the more recent studies of crack mechanisms in concrete and gives a somewhat different picture from that presented in the previous committee report.

224R-5

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Fig. 2.4 - Stress-strain curves as influenced by coating aggregates (Reference 2.36). seemed to indicate a very large effect, thus emphasizing the importance of interfacial strength on the behavior of concrete. These studies utilized relatively thick, soft coatings on the coarse aggregate to reduce the bond strength. Since these soft coatings isolated the aggregate from the surrounding mortar, the effect was more like inducing a large number of voids in the concrete matrix. Two other s t u d i e s 2 . 3 6 , 2 . 3 7 which did not isolate the coarse aggregate from the mortar indicate that the interfacial strength plays only a minor role in controlling the stress-strain behavior and ultimate strength of concrete. Darwin and Slate 2 . 3 6 used a thin coating of polystyrene on natural coarse aggregate. They found that a large reduction in interfacial bond strength causes no change in the initial stiffness of concrete under short-term compressive loads and results in approximately a 10 percent reduction in the compressive strength as compared to similar concrete made with aggregate with normal interfacial strength (see Fig. 2.4). They also found that the lower interfacial strength had no appreciable effect on the total amount of microcracking. However, in every case, the average amount of mortar cracking was slightly greater for the specimens made with coated aggregate. This small yet consistent difference may explain the differences in the stressstrain curves. Perry and Gillott 2 . 3 7 used glass spheres with different degrees of surface roughness as coarse aggregate. Their results indicate that reducing the interfacial strength of the aggregate decreases the initiation stress by about 20 percent, but has very little effect on the discontinuity stress. They also observed a 10 percent reduction in the compressive strength for specimens with low mortar-aggregate bond strength.

Work by Carino,2.38 using polymer impregnated concrete, seems to corroborate these two studies. Carino found that polymer impregnation did not increase the interfacial bond strength, but did increase the compressive strength of concrete. He attributed the increase in strength to the effect of the polymer on the strength of mortar, thus downgrading the importance of the interfacial bond. The importance of mortar, and ultimately cement paste, in controlling the stress-strain behavior of concrete is illustrated by the finite element work of Buyukozturk2.37 and Maher and Darwin. 2.31,2.32 Using a linear finite element representation of a physical model of concrete, Buyukozturk was able to simulate the overall crack patterns under uniaxial loading.

Mortar

Fig. 2.5 - Stress-strain curves for concrete model. * *From A. Maher. and D. Darwin, “Microscopic Finite Element Model of Concrete,” presented at the First International Conference on Mathematical Modeling (St. Louis. Aug.-Sept. 1 9 7 7 ) .

ACI COMMl=lTEE REPORT

224R-6 S t r e s s . PSI

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Fig. 2.6 - Stress-strain curve for finite element model of concrete with varying values of mortar-aggregate bond strength (Reference 2.32). However, his finite element model could not duplicate the nonlinear experimental behavior of the physical model using the formation of interfacial bond cracks and mortar cracks as the only nonlinear effect. Maher and Darwin2 .31,2.32 have shown that by using a nonlinear representation for the mortar constituent of the physical model, a very close representation of the actual behavior can be obtained. The results for Buyukozturk’s model are shown in Fig. 2.5 . The inability of linear elastic models 2.25,2.26,2.39 to duplicate the nonlinear behavior of concrete utilizing microcracking alone has been explained as being due to the fact that concrete is really a “statistical material.” When the proper statistical variation is selected, the nonlinear behavior of concrete can be

v MORTAR@21 v 0 CONCRETE 0

ti * OS

I I I I lJ4 l/2 314

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NOTCH DEPTH, INCHES (mm) Fig. 2.7 - Effect of notch depth on flexure strength (Reference 2.42).

duplicated 2.25 While the statistical variations undoubtedly play a part, the major nonlinear behavior can also be matched by considering the nonlinearities of the mortar constituent.2.31,2.32 Fig. 2.6 illustrates the results obtained for a highly simplified model of concrete under uniaxial compression using a nonlinear representation for mortar. The stressstrain curve for the model without cracking differs very little from that of models that have a normal, or above normal, amount of microcracking. Microcracks have a relatively minor effect on the primary stress-strain behavior of the models. The dominant effect of microcracking is to increase the lateral strain. In every case the failure of the model is governed by “crushing” of the mortar which occurs at an average strength below that of the mortar alone. Newman 2.5 s and Tasuji, Slate, and Nilson 2.40 lhave observed that the principal tensile strain in concrete at the “discontinuity stress” appears to be a function of the mean normal stress, 0, = (0,+0,+0,)/3. In their study of the biaxial strength of concrete, Tasuji, et al., observe that the final failure of their specimens consists of the formation of macroscopic tensile cracks. They also observe that the stress at discontinuity occurs at approximately 75 percent of the ultimate strength in compression and at about 60 percent of the ultimate strength for those cases involving tension, matching the levels at which mortar cracking begins. 2.3,2.4 l Their work seems to point very strongly toward a “limiting tensile strain” as the governing factor in the strength of concrete. Overall, the damage to cement paste seems to play an important role in controlling the primary stress-strain behavior of concrete under short-term axial load. In normal weight concrete, aggregate particles act as stress-raisers, increasing the initial stiffness and decreasing the strength of the paste. For cyclic and sustained loading, a great deal of the bond cracking results from load induced volume changes within the paste, but has no significant effect on strength. A number of investigators feel that the onset of mortar cracking marks the “true” ultil Whether mate strength of concrete. 2.6-2.8,2.33,2.34,2.41 mortar cracking itself controls the strength of concrete or whether it only signals intimate damage of the cement paste remains to be seen. Additional studies in this area are clearly warranted. 2.3 - Fracture Since the publication of the previous report, a number of investigations have shed additional light on the applicability of fracture mechanics to concrete and its constituent materials. Shah and McGarry utilized flexure specimens subjected to three-point loading.2.42 Their work indicates that while paste is notch sensitive, neither mortar nor concrete are affected by a notch (Fig. 2.7). Shah and McGarry also ran a series of tests using notched tensile specimens and determined that paste speci-

224R-7

CONTROL OF CRACKING

mens, and mortar specimens made with fine aggregate that passed the #30 sieve, are notch sensitive, but that mortar specimens containing larger sizes of aggregate are not notch sensitive. Brown utilized notched flexure specimens and double cantilever beam specimens of paste and mortar2.18 8 His tests show that the fracture toughness of cement paste is independent of crack length and is therefore a material constant. The fracture toughness of mortar, however, increases as the crack propagates, indicating that the addition of fine aggregate improves the toughness of paste. This behavior is similar to the behavior found in structural steels that exhibit a plane strain-plane stress transition. Because the plane strain-plane stress transition occurs beyond the limits of LEFM, the analysis is more complex. To re-establish the applicability of LEFM, larger test specimens must be used with tougher materials such as mortar. Mindess and Nadeau investigated the effect of notch width on KI for both mortar and concrete.2.20 Utilizing notched beam specimens of constant length and depth, with varying widths, they found that within the range studied, there was no dependence of fracture toughness upon the length of crack front. Since their work utilized small specimens with a depth of only about 50 mm (2 in.), there is some indication that rather than measuring the fracture toughness of the material, they were simply measuring the modulus of rupture. The applicability of these results, and much of the other fracture mechanics work, has been brought into perspective based on the experimental work by Walsh. In separate investigations of notched beam specimens2.21 ’ and beams with right angle re-entrant notches2.22 Walsh has demonstrated that specimen size has a marked influence on the applicability of linear elastic fracture mechanics to the failure of plain concrete specimens. As illustrated in Fig. 2.8, for specimens of similar geometry but below a certain critical size, the specimen capacity is governed by the modulus of rupture of concrete, calculated from the linear stress distribution. For specimens above this size, the strength is governed by the fracture toughness, which he approximated as a function of the square root of the compressive strength of the concrete. Walsh concluded that, for valid toughness testing of concrete, the depth of notched beams must be at least 230 mm (9 in.). This type of behavior is also observed in metals, i.e., for valid fracture mechanics test results, the test specimens must meet minimum size requirements (ASTM E 399). These size requirements are dependent upon the square of the toughness levels being measured. Thus a material whose toughness level is twice that of another material (all other properties being equal), must have specimen dimensions four times that of the first material for the test results to be equally valid. Gjorv, Sorensen and, Arnesen2.23 investigated the

0.10 L

I

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a/a0 (log scale) Fig. 2.8 - Relationship bet ween test results and theory for notched concrete beams (Reference 2.22).

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1

Fig. 2.9 - Effect of notch depth on flexural strength (Reference 2.23). notch sensitivity of paste, mortar and concrete using three-point bend specimens similar to those used by Shah and McGarry2.42 As shown in Fig. 2.9, they determined that both mortar and concrete are notch sensitive, but less sensitive than cement paste. They conclude that the disagreement with the earlier results is due in part to their improvement in the loading procedure. They feel that linear elastic fracture mechanics is applicable to the small specimens of

ACI COMMITTEE REPORT

paste, but not to the small size specimens of mortar and concrete. Even the small specimens of mortar and concrete, however, have some degree of notch sensitivity since the failure is not consistent with the modulus of rupture based on the net cross section. Citing Walsh’s earlier work,2.21 they agree that LEFM is applicable to large concrete specimens, but that it is not applicable to small specimens. Hillemeier and Hilsdorf 2.43 utilized wedge loaded, compact tension specimens to measure the fracture toughness of paste, aggregate and the paste-aggregate interface. They feel that, while the failure of concrete in tension and compression is controlled by many interacting cracks rather than by the propagation of a single crack, fracture mechanics does offer an important tool for evaluating the constituent materials of concrete. They found that paste is a notch sensitive material and that the addition of entrained air or soft particles has only a small affect on K I c . Their work indicates that the KIc values for interfacial strength between paste and aggregate is only about one-third of the KIc value for paste alone, and that the characteristic value of KIC for aggregate is approximately ten times that of paste. Swartz, Hu, and Jones2.24 used compliance measurement to monitor crack growth in notched concrete beams subjected to sinusodial loading. They conclude that this procedure is useful for monitoring crack growth in concrete due to fatigue. Based on the appearance of the fracture surface, which shows a combination of both aggregate fracture and bond failure, they feel that fracture toughness is not a pertinent material property. However, they state that an “effective” fracture toughness might be a significant material property if related to specific material and specimen variables such as aggregate size and gradation, and proportions of the mix, and if the calculation considers the nonlinear material response of concrete. A number of investigators do not feel that the Griffith theory of linear fracture mechanics is directly applicable to all concrete2.23, 2.24* 2.42 (ASTM E 399). Some like Swartz, et a1.2.24 feel that the theory has application when the limitations and specific nonhomogenous effects are taken into account. Clearly, specimen size requirements must be given more attention. Of key interest in future work are the observations by Walsh2.21’ 2.22 that show that if the specimens are large enough, the effects of heterogeneity are greatly reduced and that concrete may approximate a homogenous material to which the principles of fracture mechanics can be applied.

References 2.1. Hsu, Thomas T. C.; Slate, Floyd O.; Sturman, Gerald M.; and Winter, George, “Microcracking of Plain Concrete and the Shape of the Stress-Strain Curve,” ACI J OURNAL Proceedings V. 60, No. 2, Feb. 1963, pp. 209-224.

2.2. Hsu, Thomas, T. C., “Mathematical Analysis of Shrinkage Stresses in a Model of Hardened Concrete,” ACI JOURNAL , Proceedings V. 60, No. 3, Mar. 1963, pp. 371-390. 2.3. Slate, Floyd O., and Matheus, Ramon E., “Volume Changes on Setting and Curing of Cement Paste and Concrete from Zero to Seven Days,” ACI JO U R N A L, Proceedings V. 64, No. 1, Jan. 1967, pp. 34-39. 2.4. Evans, R. H., and Marathe, M. S., “Microcracking and Stress-Strain Curves for Concrete in Tension,” Materials and Structures, Research and Testing (Paris), V. 1, No. 1, Jan. 1968, pp. 61-64. 2.5. Newman, Kenneth, “Criteria for the Behavior of Plain Concrete Under Complex States of Stress,” Proceedings, International Conference on the Structure of Concrete (London, Sept. 1965), Cement and Concrete Association, London, 1968, pp. 255-274. 2.6. Brooks, J. J., and Neville, A. M., “A Comparison of Creep, Elasticity and Strength of Concrete in Tension and in Compression,” Magazine of Concrete Research (London), V. 29, No. 100, Sept. 1977, pp. 131-141. 2.7. Meyers, Bernard L.; Slate, Floyd O.; and Winter, George, “Relationship Between Time-Dependent Deformation and Microcracking of Plain Concrete,” ACI JOURNAL , Proceedings V. 66, No. 1, Jan. 1969, pp. 60-68. 2.8. Shah, Surendra P., and Chandra, Sushil, “Fracture of Concrete Subjected to Cyclic and Sustained Loading,” ACI JOURNAL, Proceedings V. 67, No. 10, Oct. 1970, pp. 816-824. 2.9. Sturman, Gerald M.; Shah, Surendra P.; and Winter, George, “Effects of Flexural Strain Gradients on Microcracking and Stress-Strain Behavior of Concrete,” ACI JOURNAL, Proceedings V. 62, No. 7, July 1965, pp. 805-822. 2.10. Griffith, A. A., “The Phenomena of Rupture and Flow in Solids,” Transactions, Royal Society of London, No. 221A, 1920, pp. 163-198. 2.11. Kaplan, M. F., “Crack Propagation and the Fracture of Concrete,” ACI JOURNAL , Proceedings V. 58, No. 5, Nov. 1961, pp. 591-610. 2.12. Glucklich, Joseph, “Static and Fatigue Fractures of Portland Cement Mortars in Flexure,” Proceedings, First International Conference on Fracture, Sendai, Japan, V. 2, 1965, pp. 1343-1382. 2.13. Romualdi, James P., and Batson, Gordon B., “Mechanics of Crack Arrest in Concrete,” Proceedings, ASCE, V. 89, EM3, June 1963, pp. 147-168. 2.14. Huang, T. S., “Crack Propagation Studies in Microconcrete,” MSc Thesis, Department of Civil Engineering, University of Colorado, Boulder, 1966. 2.15. Lott, James L., and Kesler, Clyde E., “Crack Propagation in Plain Concrete,” Symposium on Structure of Portland Cement Paste and Concrete, Special Report No. 90, Highway Research Board, Washington, D.C., 1966, pp. 204-218. 2.16. Naus, Dan J., and Lott, James L., “Fracture Toughness of Portland Cement Concretes,” ACI JOURNAL , Proceedings V. 66, No. 6, June 1969, pp. 481-489. 2.17. Naus, Dan J., “Applicability of Linear-Elastic Fracture Mechanics to Portland Cement Concretes,” PhD Thesis, University of Illinois, Urbana, Aug. 1971. 2.18. Brown, J. H., “Measuring the Fracture Toughness of Cement Paste and Mortar,” Magazine of Concrete Research (London), V. 24, No. 81, Dec. 1972, pp.185-196.

CONTROL OF CRACKING

2.19. Evans, A. G.; Clifton, J. R.; and Anderson, E., “The Fracture Mechanics of Mortars,” Cement and Concrete Research, V. 6, No. 4. July 1976, pp. 535-547. 2.20. Mindess, Sidney, and Nadeau, John S., “Effect of Notch Width of KIC for Mortar and Concrete,” Cement and Concrete Research, V. 6, No. 4, July 1976, pp. 529-534. 2.21. Walsh, P. F., “Fracture of Plain Concrete,” Indian Concrete Journal (Bombay), V. 46, No. 11, Nov. 1972, pp. 469-470, 476. 2.22. Walsh, P. F., “Crack Initiation in Plain Concrete,” Magazine of Concrete Research (London), V. 28, No. 94, Mar. 1976, pp. 37-41. 2.23. Gjorv, O. E.; Sorensen, S. I.; and Arnesen, A., “Notch Sensitivity and Fracture Toughness of Concrete,” Cement and Concrete Research, V. 7, No. 3, May 1977, pp. 333-344. 2.24. Swartz, Stuart E.; Hu, Kuo-Kuang; and Jones, Gary L., “Compliance Monitoring of Crack Growth in Concrete,” Proceedings, ASCE, V. 104, EM4, Aug. 1978, pp. 789-800. 2.25. Shah, Surendra P., and Winter, George, “Inelastic Behavior and Fracture of Concrete,” ACI JOURNAL , P r o ceedings V. 63, No. 9, Sept. 1966, pp. 925-930. 2.26. Testa, Rene B., and Stubbs, Norris, “Bond Failure and Inelastic Response of Concrete,” Proceedings, ASCE, V. 103, EM2, Apr. 1977, pp. 296-310. 2.27. Darwin, David, Discussion of “Bond Failure and Inelastic Response of Concrete,” by Rene B. Testa and Norris Stubbs, Proceedings, ASCE, V. 104, EM2, Apr. 1978, pp. 507-509. 2.28. Spooner, D. C., “The Stress-Strain Relationship for Hardened Cement Pastes in Compression,” Magazine of Concrete Research (London), V. 24, No. 79, June 1972, pp. 85-92. 2.29. Spooner, D. C., and Dougill, J. W., “A Quantitative Assessment of Damage Sustained in Concrete During Compressive Loading,” Magazine of Concrete Research (London), V. 27, No. 92, Sept. 1975, pp. 151-160. 2.30. Spooner, D. C.; Pomeroy, C. D.; and Dougill, J. W., “Damage and Energy Dissipation in Cement Pastes in Compression,” Magazine of Concrete Research (London), V. 28, No. 94, Mar. 1976, pp. 21-29. 2.31. Maher, Ataullah, and Darwin, David, “A Finite Element Model to Study the Microscopic Behavior of Plain Concrete,” CRINC Report-SL-76-02, The University of Kansas Center for Research, Lawrence, Nov. 1976, 83 pp. 2.32. Maher, Ataullah, and Darwin, David, “Microscopic Finite Element Model of Concrete,” Proceedings, First International Conference on Mathematical Modeling (St. Louis, Aug.-Sept. 1977), University of Missouri-Rolla, 1977, v. III, pp. 1705-1714. 2.33. Karsan, I. Demir, and Jirsa, James 0.. “Behavior of Concrete under Compressive Loadings,” Proceedings, ASCE, V. 95, ST12, Dec. 1969, pp. 2543-2563. 2.34. Neville, A. M., and Hirst, G. A., “Mechanism of Cyclic Creep of Concrete,” Douglas McHenry Symposium on Concrete and Concrete Structures, SP-55, American Concrete Institute, Detroit, 1978, pp. 83-101. 2.35. Nepper-Christensen, Palle, and Nielsen, Tommy P. H., “Modal Determination of the Effect of Bond Between Coarse Aggregate a n d M o r t a r o n t h e C o m p r e s s i v e Strength of Concrete,” ACI JO U R N A L , Proceedings V. 66, No. 1, Jan. 1969, pp. 69-72.

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2.36. Darwin, David, and’ Slate, F. O., “Effect of PasteAggregate Bond Strength on Behavior Concrete,” J o u r nal of Materials, V. 5, No. 1, Mar. 1970, pp. 86-98. 2.37. Perry, C., and Gillott, J. E., “The Influence of Mortar-Aggregate Bond Strength on the Behavior of Concrete in Uniaxial Compression,” Cement and Concrete Research, V. 7, No. 5, Sept. 1977, pp. 553-564. 2.38. Carino, Nicholas J., “Effects of Polymer Impregnation on Mortar-Aggregate Bond Strength,” Cement and Concrete Research, V. 7, No. 4, July 1977, pp. 439-447. 2.39. Buyukozturk, Oral, “Stress-Strain Response and Fracture of a Model of Concrete in Biaxial Loading,” PhD Thesis, Cornell University, Ithaca, June 1970. 2.40. Tasuju, M. Ebrahim; Slate, Floyd 0.; and Nilson, Arthur H., “Stress-Strain Response and Fracture of Concrete in Biaxial Loading,” ACI JO U R N A L , Proceedings V . 75, No. 7, July 1978, pp. 306-312. 2.41. Shah, Surendra P., and Chandra, Sushil, “Critical Stress, Volume Change, and Microcracking of Concrete,” ACI JO U R N A L, Proceedings V. 65, No. 9, Sept. 1968, pp. 770-781. 2.42. Shah, Surendra P., and McGarry, Fred J., “Griffith Fracture Criterion and Concrete,” Proceedings, ASCE, V. 97, EM6, Dec. 1971, pp. 1663-1676. 2.43. Hillemeier, B., and Hilsdorf, H. K., “Fracture Mechanics Studies of Concrete Compounds,” Cement and Concrete Research, V. 7, No. 5, Sept. 1977, pp. 523-535.

Chapter 3 - Control of cracking due to drying shrinkage* 3.1 - Introduction Cracking of concrete due to drying shrinkage is a subject which has received more attention by architects, engineers, and contractors than any other characteristic or property of concrete. It is one of the most serious problems encountered in concrete construction. Good design and construction practice can minimize the amount of cracking and eliminate the visible large cracks by the use of adequate reinforcement and contraction joints. Although drying shrinkage is one of the principal causes of cracking, temperature stresses, chemical reactions, frost action, as well as excessive tensile stresses due to loads on the structure, are frequently responsible for cracking of hardened concrete. Cracking may also develop in the concrete prior to hardening due to plastic shrinkage. Information presented in this chapter concerns only the subjects of cracking of hardened concrete due to drying shrinkage; factors influencing shrinkage; control of cracking; and the use of expansive cements to minimize cracking. The subject of construction practices and specifications to minimize drying shrinkage is covered in Chapter 8 (Sections 8.3 and 8.6) of this report.

*Principal author: Miloss Polivka.

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ACI COMMITTEE REPORT

3.2 - Crack formation Why does concrete crack due to shrinkage? If the shrinkage of concrete caused by drying could take place without any restraint, the concrete would not crack. However, in a structure the concrete is always subject to some degree of restraint by either the foundation or another part of the structure or by the reinforcing steel embedded in the concrete. This combination of shrinkage and restraint develops tensile stresses. When this tensile stress reaches the tensile strength, the concrete will crack. This is illustrated in Fig. 3.1. Another type of restraint is developed by the difference in shrinkage at the surface and in the interior of a concrete member, especially at early ages. Since the drying shrinkage is always larger at the exposed surface, the interior portion of the member restrains the shrinkage of the surface concrete, thus developing tensile stresses. This may cause surface cracking, which are cracks that do not penetrate deep into the concrete. These surface cracks may with time penetrate deeper into the concrete member as the interior portion of the concrete is subject to additional drying.

ORIGINAL LENGTH

I

I

UNRESTRAINED SHRINKAGE

t-

RESTRAINED SHRINKAGE DEVELOPS TENSILE STRESS

IF TENSILE STRESS IS GREATER THAN TENSILE STRENGTH, CONCRETE CRACKS Fig. 3.1 - Cracking of concrete due to drying shrinkage.

The magnitude of tensile stress developed during drying of the concrete is influenced by a combination of factors, such as (a) the amount of shrinkage, (b) the degree of restraint, (c) the modulus of elasticity of the concrete, and (d) the creep or relaxation of the concrete. Thus, the amount of shrinkage is only o n e factor governing the cracking. As far as cracking is concerned, a low modulus of elasticity and high creep characteristics of the concrete are desirable since they reduce the magnitude of tensile stresses. Thus, to minimize cracking, the concrete should have low drying shrinkage characteristics and a high degree of extensibility (low modulus and high creep) as well as a high tensile strength. However, a large extensibility of a concrete member subjected to bending will cause larger deflections.

3.3 - Drying shrinkage When concrete dries, it contracts or shrinks, and when it is wetted again, it expands. These volume changes, with changes in moisture content, are an inherent characteristic of hydraulic cement concretes. It is the change in moisture content of the cement paste that causes the shrinkage or swelling of concrete, while the aggregate provides an internal restraint which significantly reduces the magnitude of these volume changes. When cement is mixed with water, several chemical reactions take place. These reactions, commonly called “hydration,” produce a hydration product consisting essentially of some crystalline materials (principally calcium hydroxide) and a large amount of hardened calcium silicate gel called “tobermorite gel.” This rigid gel consists of colloidal size particles and has an extremely high surface area. In a hardened cement paste, some of the water is in the capillary pores of the paste, but a significant amount is in the tobermorite gel. Shrinkage is due to the loss of adsorbed water from the gel. On drying the first water lost is that which occupies the relatively large size capillaries in the cement paste. This loss of water causes very little, if any, shrinkage. It is the loss of the adsorbed and inter-layer water from the hydrated gel that causes the shrinkage of the paste. When a concrete is exposed to drying conditions, moisture slowly diffuses from the interior mass of the concrete to the surface where it is lost by evaporation. On wetting this process is reversed, causing an expansion of the concrete. In addition to drying shrinkage, the cement paste is also subject to carbonation shrinkage. The action of carbon dioxide, CO2, present in the atmosphere on the hydration products of the cement, principally calcium hydroxide, Ca(OH)2, results in the formation of calcium carbonate, CaCO,, which is accompanied by a decrease in volume. Since carbon dioxide does not penetrate deep into the mass of concrete, shrinkage due to carbonation is of minor importance in the overall shrinkage of a concrete structure. However,

CONTROL OF CRACKING

carbonation does play an important role in the shrinkage of small laboratory test specimens, particularly when subjected to long-term exposure to drying. Thus, the amount of shrinkage observed on a small laboratory specimen will be greater than the shrinkage of the concrete in the structure. The subject of shrinkage due to carbonation is discussed in detail by Verbeck.3.1 3.4 - Factors influencing drying shrinkage The major factors influencing shrinkage include the composition of cement, type of aggregate, water content, and mix proportions. The rate of moisture loss or the shrinkage of a given concrete is greatly influenced by the size and shape of the concrete member, the environment, and the time of drying exposure. These and other factors influencing magnitude and rate of shrinkage are herein discussed. 3.4.1 Effect of cement - Results of an extensive study made by Blaine, Arni, and Evans,3.2 of the National Bureau of Standards on a large number of portland cements indicate that it is not possible to say that a cement, because it conforms to the requirements of one of the standard types of cements, will have greater or less shrinkage than a cement meeting requirements for some other type of cement. Their results on neat cement pastes showed a wide distribution of shrinkage values especially for the Type I cements. The 6 month drying shrinkage strain of the neat pastes ranged from about 0.0015 to more than 0.0060 with an average for the 182 cements tested of about 0.0030. They found that lower shrinkage of pastes was associated with: 1. lower C3A/SO3 ratios, 2. lower Na2O and K2O contents, and 3. higher C4AF contents of the cement. Tests by Brunauer. Skalny, and Yudenfreund3.3 show that for short curing periods Type II cement pastes exhibited considerably less shrinkage than Type I pastes. However, the shrinkage of pastes cured for 28 days was about the same for the two types of cements. Tests made by the California Division of Highways 3.4 on mortar or paste as a measure of behavior in concrete indicate that Type II cements generally produce lower shrinkage than Type I cements, and much lower than Type III cements. Tests by Lerch1.5 show that the proportion of gypsum in the cement has a major effect on shrinkage. Cement producers moderate the differences in shrinkage due to cement composition by optimizing its gypsum content. The fineness of a cement can have some influence on drying shrinkage. Tests by Carlson3.6 showed that finer cements generally result in greater concrete shrinkage, but the increase in shrinkage with increasing fineness is not large. His results show that the composition of the cement is a factor and thus for some cements an increase in fineness may show little change and in some cases even a lower concrete shrinkage.

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TABLE 3.1 - Effect of type of aggregate on shrinkage of concrete3.6

Aggregate Sandstone Slate Granite Limestone Quartz

t

Specific gravity 2.47 2.75 2.67 2.74 2.66

Absorption, percent

l-year shrinkage, percent

5.0 1.3 0 .8 0.2 0.3

0.116 0.068 0.047 0.041 0.032

3.4.2 Influence of type of aggregate - Coarse and fine aggregates, which occupy between 65 and 75 percent of the total concrete volume, have a major influence on shrinkage. Concrete may be considered to consist of a framework of cement paste whose large potential shrinkage is being restrained by the aggregate. The drying shrinkage of a concrete will be only a fraction (about l/4 to l/6) of that of the cement paste. The factors which influence the ability of the aggregate particles to restrain shrinkage include (a) the compressibility of aggregate and the extensibility of paste, (b) the bond between paste and aggregate, (c) the degree of cracking of cement paste, and (d) the contraction of the aggregate particles due to drying. Of these several factors, compressibility of the aggregate has the greatest influence on the magnitude of drying shrinkage of concrete. The higher the stiffness or modulus of elasticity of an aggregate, the more effective it is in reducing the shrinkage of concrete. The absorption of an aggregate, which is a measure of porosity, influences its modulus or compressibility. A low modulus is usually associated with high absorption. The large influence of type of aggregate on drying shrinkage of concrete was shown by Carlson.3.6 As an example some of his shrinkage data for concretes with identical cements and identical water-cement ratios are given in Table 3.1. Quartz, limestone, dolomite, granite, feldspar, and some basalts can be generally classified as lowshrinkage producing types of aggregates. Highshrinkage concretes often contain sandstone, slate, hornblende and some types of basalts. Since the rigidity of certain aggregates, such as granite, limestone or dolomite, can vary over a wide range, their effectiveness in restraining drying shrinkage will vary accordingly. Although the compressibility is the most important single property of aggregate governing concrete shrinkage, the aggregate itself may contract an appreciable amount upon drying. This is true for sandstone and other aggregates of high absorption capacity. Thus, in general, aggregate of high modulus of elasticity and low absorption will produce a lowshrinkage concrete. However, some structural grade lightweight aggregates, such as expanded shales,

ACI COMMITTEE REPORT

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+

119

142

166

190 kg/m3

240

280

320 Ib/yd3

5 0.060 u % 0.050

I

," 0.020 z

zis

0.010

200

WATER CONTENT OF CONCRETE

Fig. 3.2 - Typical effect of water content of concrete on drying shrinkage (Reference 3.8).

clays, and slates which have high absorptions, produced concretes exhibiting low shrinkage characteristics.3.7 Maximum size of aggregate has a significant effect on drying shrinkage. Not only does a large aggregate size permit a lower water content of the concrete, but it is more effective in resisting the shrinkage of the cement paste. Aggregate gradation also has some effect on shrinkage. The use of a poorly graded fine or coarse aggregate may result in an oversanded mix, in order to obtain desired workability, and thus prevent the use of the maximum amount of coarse aggregate resulting in increased shrinkage. 3.4.3 Effect of water content and mix proportions The water content of a concrete mix is another very important factor influencing drying shrinkage. The large increase in shrinkage with increase in water content was demonstrated in tests made by the U.S. Bureau of Reclamation.3.8 A typical relationship between water content and’drying shrinkage is shown in Fig. 3.2. An increase in water content also reduces the volume of restraining aggregate and thus results in higher shrinkage. The shrinkage of a con-

400’ 2.5

19.0

37.5

75

1 1/2

3

150 m m

(237)

350 (208)

300 (178)

250 (148)

200 3/8

(119)

3/4

6 in.

MAXIMUM SIZE OF AGGREGATE

Fig. 3.3 - Effect of aggregate size on water requirement of non-air-entrained concrete (ACI 211.1).

crete can be minimized by keeping the water con-

tent of the paste as low as possible and the total aggregate content of the concrete as high as possible. This will result in a lower water content per unit volume of concrete and thus lower shrinkage. The total volume of coarse aggregate is a significant factor in drying shrinkage. Concrete proportioned for pump placement with excessively high sand contents will exhibit significantly greater shrinkage than will similar mixes with normal sand contents. Tests reported by Tremper and Spellman3.4 show that the cement factor has little effect on shrinkage of concrete. Their data show that as the cement factor was increased from 470 to 752 lb/yd 3 (279 to 446 kg/m3) the water content remained nearly constant, while percentage of fine aggregate was reduced. The amount of mixing water required for concrete of a given slump is greatly dependent on the maximum size of aggregate. The surface area of aggregate, which must be coated by cement paste, decreases with increase in size of aggregate. The large effect that the maximum size of aggregate has on the water requirement of concrete is shown in Fig. 3.3. The data plotted in this figure, taken from ACI 211.1 shows, for example, that for a 3 to 4 in. (75 to 100 mm) slump concrete, increasing the aggregate size from 3/4 in. (19 mm) to 11/2 in. (38 mm) decreases the water requirement from 340 to 300 lb/yd3 (202 to 178 kg/m3). This 40 lb (24 kg) reduction in water content would reduce the 1 year drying shrinkage by about 15 percent. Also shown in Fig. 3.3 is the effect of slump on water requirement. For example, the water requirement of a concrete made with 3/4 in. (19 mm) size aggregate is 340 lb/yd3 (202 kg/m3) for a 3 to 4 in. slump, but only 310 lb/yd3 (184 kg/m31 for a 1 to 2 in. slump (25 to 50 mm). This substantial reduction in water content would result in a lower drying shrinkage. Another important factor which influences the water requirement of a concrete, and thus its shrinkage, is the temperature of the fresh concrete. This effect of temperature on water requirement as given by the U.S. Bureau of Reclamation 3. is shown in Fig. 3.4. For example, if the temperature of fresh concrete were reduced from 100 to 50 F (38 to 10 C), it would permit a reduction of the water content by 33 Ib/yd 3 (20 kg/m 3) and still maintain the same slump. This substantial reduction in water content would significantly reduce the drying shrinkage. From the above discussion it must be concluded that, to minimize the drying shrinkage of concrete, the water content of a mix should be kept to a minimum. Any practice that increases the water requirement, such as the use of high slumps, high temperatures of the fresh concrete or the use of smaller size coarse aggregate, will substantially increase shrinkage and thus cracking of the concrete.

CONTROL OF CRACilNG

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4.4 10.0 15.6 21.1 26.7 32.2 378 OC 310 084)

300 (I 78) 290 (I 72)

280 (166)

0

0

w iii

270 (160)

260 (154140

a a 50

60

70

80

90

100 OF

TEMPERATURE OF FRESH CONCRETE

-0 w

4

8

I2

16

20

24

28 in.

DEPTH BELOW CONCRETE SURFACE

Fig. 3.4 - Effect of temperature of fresh concrete on its water requirement (Reference 3.8).

Fig. 3.5 - Rates of drying of concrete exposed to 50 percent relative humidity (Reference 3.9).

3.4.4 Effect of chemical admixtures - Chemical admixtures are used to impart certain desirable properties to the concrete. Those most commonly used include air-entraining admixtures, water-reducing admixtures, set-retarding admixtures, and accelerators. It would be expected that when using an air-entraining admixture, the increase in the amount of air voids would increase drying shrinkage. However, because entrainment of air permits a reduction in water content with no reduction in slump, the shrinkage is not appreciably affected by air contents up to about 5 percent.3.8 Some air-entraining agents are strong retarders and contain accelerators which may increase drying shrinkage by 5 to 10 percent. Although the use of water-reducing and set-retarding admixtures will permit a reduction in the water content of a concrete mix, it will usually not result in a decrease in drying shrinkage. Actually some of these admixtures may even increase the shrinkage at early ages of drying, although the later age shrinkage of these concretes will be about the same as that of corresponding mixes with no admixtures. The use of calcium chloride, a common accelerator, will result in a substantial increase in drying shrinkage, especially at the early ages of drying. Tests made by the California Department of Transportation 3.44showed that the 7 day shrinkage of a concrete containing 1.0 percent of calcium chloride was about double that obtained for the control mix without admixture. However, after 28 days of drying, the shrinkage of the concrete containing calcium chloride was only about 40 percent greater than that of the control mix. 3.4.5 Effect of pozzolans - Fly ash and a number of natural materials such as opaline cherts and shales, diatomaceous earth, tuffs and pumicites are pozzolans used in portland cement concrete. The use of some natural pozzolans can increase the water de-

mand as well as the drying shrinkage of the concrete. Also, it was observed that the use of some of these pozzolans increased drying shrinkage although they had little effect on the water content of the concrete. Some fly ashes have little effect on drying shrinkage, while others may increase the shrinkage of the concrete. All of these observations are based on results of tests made on laboratory size specimens. However, as noted in Section 3.4.7 and Fig. 3.6, the larger the concrete member, the lower the shrinkage. This may explain the negligible difference in shrinkage cracking of field structures, with and without pozzolan, despite clearly greater shrinkage of the concretes with pozzolans in laboratory tests on small size specimens. 3.4.6 Effect of duration of moist curing - Car1son3.6 reported that the duration of moist curing of concrete does not have much effect on drying shrinkage. This is substantiated by the test results of the California Department of Transportation3.’ which show substantially the same shrinkage in concrete that was moist cured for 7, 14, and 28 days before drying was started. As far as the cracking tendency of the concrete is concerned, prolonged moist curing may not necessarily be beneficial. Although the strength increases with age, the modulus of elasticity also increases by almost as large a percentage, and the net result is only a slight increase in the tensile strain which the concrete can withstand. Steam curing at atmospheric pressure, which is commonly used in the manufacture of precast structural elements, will reduce drying shrinkage (AC1 517). Also, because stream curing will produce a high early-age strength of the concrete, it will reduce its tendency to crack, since the pre’cast members are unrestrained. 3.4.7 Influence of size of member - The size of a concrete member will influence the rate at which moisture moves from the concrete and thus influence the rate of shrinkage. Carlson3*’ has shown

ACI COMMITTEE REPORT

224R-14

that for a concrete exposed to a relative humidity of 50 percent, drying will penetrate only about 3 in. (75 mm) in 1 month and about 2 ft (0.6 m) in 10 years. Fig. 3.5 shows his theoretical curves for the drying of slabs. Hansen and Mattock3.10 made an extensive investigation of the influence of size and shape of member on the shrinkage and creep of concrete. They found that both the rate and the final values of shrinkage and creep decrease as the member becomes larger. This significant effect of size of member on drying shrinkage of concrete must be considered when evaluating the potential shrinkage of concrete in structures based on the shrinkage of concrete specimens in the laboratory. The rate and magnitude of shrinkage of a small laboratory specimen will be much greater than that of the concrete in the structures. Test results of several studies carried out to compare the shrinkage of concrete in walls and slabs in the field with the shrinkage of small laboratory specimens have shown, as expected, that the shrinkage of the concrete in a field structure is only a fraction of that obtained on the laboratory specimens. Even in laboratory tests the size of the specimen used has a significant influence on shrinkage. As an example of the effect of specimen size on shrinkage is the data presented in Fig. 3.6, giving the results of shrinkage tests obtained on four different size concrete prisms. It will be noted that the shrinkage of the prisms having a cross section of 3 x 3 in. (7.5 x 7.5 cm) was more than 50 percent greater than that of the concrete prism having a cross section of 5 x 6 in. (12.5 x 15 cm). 3.5 - Control of shrinkage cracking Concrete tends to shrink due to drying whenever its surfaces are exposed to air of low relative humidity. Since various kinds of restraint prevent the con7.5x7.5

I

3x3

10 x 10

10x12 5

I

I 4x4

4x5

12.5x 15 cm

I 5x6 in

AVERAGE END AREA DIMENSION OF CONCRETE PRISM ( LOG SCALE )

Fig. 3.6 - Effect of specimen size on drying shrinkage of concrete (Principal author’s data).

crete from contracting freely, the possibility of cracking must be expected unless the ambient relative humidity is kept at 100 percent or the concrete surfaces are sealed to prevent loss of moisture. The control of cracking consists of reducing the cracking tendency to a minimum, using adequate and properly positioned reinforcement, and using control joints. The CEB-FIP Code give quantitative recommendations on the control of cracking due to shrinkage, listing various coefficients to determine the shrinkage levels that can be expected. Control of cracking by correct construction practices is covered in Chapter 8 of this report, which includes specifications to minimize drying shrinkage (Section 8.6).

Cracking can also be minimized by the use of expansive cements to produce shrinkage-compensating concretes. Shrinkage-compensating concretes are discussed in Section 3.6. 3.5.1 Reduction of cracking tendency - As mentioned previously, the cracking tendency is due not only to the amount of shrinkage, but also to the degree of restraint, the modulus of elasticity, and the creep or relaxation of the concrete. Some factors which reduce the shrinkage at the same time decrease the creep or relaxation and increase the modulus of elasticity, thus offering little or no help to the cracking tendency. Emphasis should be placed, therefore, on modifying those factors which produce a net reduction in the cracking tendency. Any measure that can be taken to reduce the shrinkage of the concrete will also reduce the cracking tendency. Drying shrinkage can be reduced by using less water in the mix and larger aggregate size. A lower water content can be achieved by using a well-graded aggregate, stiffer consistency, and lower initial temperature of the concrete. As discussed in Section 3.4.4, however, the reduction of water content by the use of water-reducing admixtures will not usually reduce shrinkage. Another way to reduce the cracking tendency is to use a larger aggregate size. A larger aggregate size allows an increase in aggregate volume and a reduction in the total water required to obtain a given slump. The larger aggregate also tends to restrain the concrete more, and although this may result in internal microcracking, such internal cracking is not necessarily harmful. A third way to reduce the cracking tendency is to apply a surface coating to the concrete, which will prevent the rapid loss of moisture from within. This means of controlling cracking has not been used to its full potential and should be given better consideration. However, many surface coatings such as allpurpose paints are ineffective, because they permit the moisture to escape almost as fast as it reaches the surface. Chlorinated rubber and waxy or resinous materials are effective coatings, but there are probably many other materials which will slow the evaporation enough to be beneficial. Any slowing of

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CONTROL OF CRACKING

the rate of shrinkage will be beneficial, because concrete has a remarkable quality of relaxing under sustained stress. Thus, concrete may be able to withstand two or three times as much slowly applied shrinkage as it can rapid shrinkage. 3.5.2 R e i n f o r c e m e n t - P r o p e r l y p l a c e d r e inforcement, used in adequate amounts, will not only reduce the amount of cracking but prevent unsightly cracking. By distributing the shrinkage strains along the reinforcement through bond stresses, the cracks are distributed in such a way that a larger number of very fine cracks will occur instead of a few wide cracks. Although the use of such reinforcement to control cracking in a relatively thin concrete section is practical, it is not needed in massive structures such as dams due to the low drying shrinkage of these mass concrete structures. The minimum amount and spacing of reinforcement to be used in floors, roof slabs, and walls is given in AC1 318. 3.6.3 Joints - The use of joints is the most effective method of preventing formation of unsightly cracking. If a sizable length or expanse of concrete, such as walls, slabs or pavements, is not provided with adequate joints to accommodate shrinkage, it will make its own “joints” by cracking. Contraction joints in walls are made, for example, by fastening to the forms wood or rubber strips which leave narrow vertical grooves in the concrete on the inside and outside of the wall. Cracking of the wall due to shrinkage should occur at the grooves, relieving the stress in the wall and thus preventing formation of unsightly cracks. These grooves should be sealed on the outside of the wall to prevent penetration of moisture. Sawed joints are commonly used in pavements, slabs and floors. Joint location depends on the particulars of placement. Each job must be studied individually to determine where joints should be placed.*

STEEL\ _---_B--ORIGINAL LENGTH

T A b T ___++IC~*___ t

EXPANSION PUTS STEEL IN TENSION AND CONCRETE IN COMPRESSION M

STRESS LOSS DUE TO SHRINKAGE AND CREEP RESIDUAL EXPANSION OR, -+j SMALL CONTRACTION

. Qr

3. 7 - Basic concept of shrinkage-compensating concretes. l

r/ .

CURING

p-

DRYING

SHRINKAGE- COMPENSATING CONCRETE, p = 0.16Ym

PORTLAND CEMENT CONCRETE ;

3.6 - Shrinkage-compensating concretes Shrinkage-compensating concretes made with expansive cements can be used to minimize or eliminate shrinkage cracking. The properties and use of expansive cement concretes is published in numerous papers and reports.3-11* 3*12 Of the several types of expansive cements produced, the Type K shrinkage-compensating expansive cement is most commonly used in the United States. In a reinforced concrete, the expansion of the cement paste during the first few days of curing will develop a low level of prestress inducing compressive stresses in the concrete and tensile stresses in the steel. The level of compressive stresses developed in the shrinkage-compensating concretes ranges from 25 to 100 psi (0.2 to 0.7 MPal. When subjected to drying shrinkage, the contraction of the concrete will result in a reduction or elimination of its precompression. The initial precompression of the

, 0

I 50

I 100

I 150

I 2oc

AGE OF CONCRETE, DAYS

Fig. 3.8 - Length change characteristics of shrinkage-compensating and portland cement concretes (Relative humidity = 50 percent). concrete minimizes the magnitude of any tensile stress that may ultimately develop due to shrinkage, and thus reduce or eliminate the tendency to cracking. This basic concept of the use of expansive cement to produce a shrinkage-compensating concrete is illustrated in Fig. 3.7. A typical length change history of a shrinkagecompensating concrete is compared to that of a portland cement concrete in Fig. 3.8. The amount of reinforcing steel normally used in reinforced concrete *Guidance on joint sealants and control joint location in slabs is available in ACI 504 and in ACI 302, respectively.

224R-16

ACI COMMITTEE REPORT

made with portland cements is usually more than adequate to provide the elastic restraint needed for shrinkage-compensating concrete. To take full advantage of the expansive potential of shrinkage-compensating concrete in minimizing or preventing shrinkage cracking of unformed concrete surfaces, it is important that positive and uninterrupted water curing (wet covering or ponding) be started immediately after final finishing. For slabs on well saturated subgrades, curing by sprayed-on membranes or moisture-proof covers have been successfully utilized. Inadequate curing of shrinkage-compensating concrete may result in an insufficient expansion to elongate the steel and thus subsequent cracking during drying shrinkage. Specific recommendations and information on the use of shrinkage-compensating concrete are contained in ACI 223. References 3.1. Verbeck, George J., “Carbonation of Hydrated Portland Cement,” Cement and Concrete, STP-205, American Society for Testing and Materials, Philadelphia, 1958, pp. 17-36. 3.2. Blaine, R. L.; Arni, H. T.; and Evans, D. N., “Interrelations Between Cement and Concrete Properties: Part 4 - Shrinkage of Hardened Portland Cement Pastes and Concrete,” Building Science Series No. 15, National Bureau of Standards, Washington, D.C., Mar. 1969, 77 pp. 3.3. Brunauer, S.; Skalny, J.: and Yudenfreund, H., “Hardened Cement Pastes of Low Porosity: Dimensional Changes,” Research Report No. 69-8, Engineering Research and Development Bureau, New York State Department of Transportation, Albany, Nov. 1969, 12 pp. 3.4. Tremper, Bailey, and Spellman, Donald L., “Shrinkage of Concrete - Comparison of Laboratory and Field Performance,” Highway Research Record. Highway Research Board, No. 3, 1963, pp. 30-61. 3.5. Lerch, William, “The Influence of Gypsum on the

Hydration and Properties of Portland Cement Pastes,” Proceedings, ASTM, V. 46, 1946, pp. 1252-1297. 3.6. Carlson, Roy W., “Drying Shrinkage of Concrete as Affected by Many Factors,” Proceedings, ASTM, V. 38, Part II, 1938, pp. 419-437. 3.7. Reichard, T. W., “Creep and Drying Shrinkage of Lightweight and Normal Weight Concrete, ” M o n o g r a p h 74, National Bureau of Standards, Washington, D.C., 1964, 30 pp. 3.8. Concrete Manual, 8th Edition, U.S. Bureau of Reclamation, Denver, 1975, 627 pp. 3.9. Carlson, Roy W., “Drying Shrinkage of Large Concrete Members,” ACI JOURNAL , Proceedings V. 33, No. 3, Jan.-Feb. 1937, pp. 327-336. 3.10. Hansen, Torben C., and Mattock, Alan H., “Influence of Size and Shape of Member on the Shrinkage and Creep of Concrete,” ACI JOURNAL, Proceedings V. 63, No. 2, Feb. 1966, pp. 267-290. 3 . 1 1 . ACI C o m m i t t e e 2 2 3 , “ E x p a n s i v e C e m e n t Concretes-Present State of Knowledge,” ACI J O U R N A L , Proceedings V. 67, No. 8, Aug. 1970, pp. 583-610. 3.12. Klein Symposium on Expansive Cement Concretes, SP-38, American Concrete Institute, Detroit, 1973, 491 pp.

Chapter 4 - Control of cracking in flexural

members*

4.1 - Introduction With the regular use of high strength reinforcing steel and the strength design approach for reinforced concrete, and higher allowable stresses in prestressed concrete design, the control of cracking may be as important as the control of deflection in flexural members. Internal cracking in concrete can start at stress levels as low as 3000 psi (20.7 MPa) in the reinforcement. Crack control is important to promote the aesthetic appearance of structures, and for many structures, crack control plays an important role in the control of corrosion by limiting the possibilities for entry of moisture and salts which, together with oxygen, can set the stage for corrosion. This chapter is concerned primarily with cracks caused by flexural and tensile stresses, but temperature, shrinkage, shear and torsion may also lead to cracking.”4.1 Cracking in certain specialized structures, such as reinforced concrete tanks, bins and silos, is not covered in this report. For information on cracking concrete in these structures, see Reference 4.2 and ACI 313. Extensive research studies on the cracking behavior of beams have been conducted over the last 5 0 y e a r s . M o s t o f t h e m a r e r e p o r t e d i n ACI Bibliography No. 9 on crack control.4.3 Others are referenced in this chapter. Reference 4.1 contains an extensive review of cracking in reinforced concrete structures. Several of the most important crack prediction equations are reviewed in the previous committee report. 1.1’ Additional work presented in the CEB-FIP Model Code for Concrete Structure gives the European approach to crack width evaluation and permissible crack widths. Recently, fiber glass rods have been used as a reinforcing material.4.4To date, experience is limited, and crack control in structures reinforced with fiber glass rods is not addressed in this report. It is expected, however, that future committee documents will address crack control in structures using this and other new systems as they come into use. 4.2 - Crack control equations for reinforced concrete beams A number of equations have been proposed for the prediction of crack widths in flexural members; most of them are reviewed in the previous committee report 1.1Pand in key publications listed in the references. Most equations predict the probable maximum crack width, which usually means that about 90 percent of the crack widths in the member are below the calculated value. However, research has shown that isolated cracks in beams in excess of twice the width of the computed maximum can

*Principal authors: Edward G. Nawy and Peter Gergely.

CONTROL OF CRACKING

sometimes occur,*-4 though generally the coefficient of variation of crack width is about 40 percent.4-1 Evidence also exists indicating that this range in crack width randomness may increase with the size of the member? Besides limiting the computed maximum crack width to a given value, the designer should estimate the percentage of cracks above this value which can be tolerated. Crack control equations recommended by ACI Committee 224 and the Comite Euro-International du Beton (CEB) are presented below. 4.21 ACI Committee 224 recommendations - Requirements for crack control in beams and thick oneway slabs in the ACI Building Code (ACI 318) are based on the statistical analysis4-6 of maximum crack width data from a number of sources. Based on the analysis, the following general conclusions were reached: 1. The steel stress is the most important variable. 2. The thickness of the concrete cover is an important variable, but not the only geometric consideration. 3. The area of concrete surrounding each reinforcing bar is also an important geometric variable. 4. The bar diameter is not a major variable. 5. The size of the bottom crack width is influenced by the amount of strain gradient from the level of the steel to the tension face of the beam. The equations that were considered to best p r e diet the most probable maximum bottom and side crack widths are: W* =

0 . 0 9 1 v-a p (f, - 5) x 10-3

w, =

0.091 rt,, A . (f, - 5) x 1 0 - 3 -1-G. t,l&

(4.la)

l

(4.lb)

where W* = most probable maximum crack width at bottom of beam, in. w, = most probable maximum crack width at level of reinforcement, in. = reinforcing steel stress, ksi fJ = area of concrete symmetric with reinforcing A steel divided by number of bars, in.2 tb = bottom cover to center of bar, in. t, = side cover to center of bar, in. = ratio of distance between neutral axis and P tension face to distance between neutral axis and centroid of reinforcing steel = 1.20 in

224R-17

where W = most probable maximum crack width, in. d c thickness of cover from tension fiber to center of bar closest thereto, in. =

When the strain, Ed, in the steel reinforcement is used instead of stress, f,, Eq. (4.2) becomes w = 2.2 p L, V-JX E,

=

(4.3)

strain in the reinforcement

Eq. (4.3) is valid in any system of measurement. The cracking behavior in thick one-way slabs is similar to that in shallow beams. For one-way slabs having a clear concrete cover in excess of 1 in. (25.4 mm), Eq. (4.2) can be adequately applied if p = 1.25 to 1.35 is used. AC1 318 Section 10.6 uses Eq. (4.2) with p = 1.2 in the following form 2 =

f,cQi-

(4.2a)

Using the specified cover in AC1 318, maximium allowable z = 175 kips per in. for interior exposure corresponds to a limiting crack width of 0.016 in. (0.41 mm). The Code allows a value of z = 145 kips per in. for exterior exposure based on a crack width value of 0.013 in., (0.33 mm), which may be excessive based on Table 4.1. While application of Eq. (10.4) of AC1 318-771 to beams gives adequate crack control values, its application to one-way slabs with standard 3/4 in. (19 mm) cover and reinforced with steel of 60 ksi (414 MPa) or lower yield strength results in large reinforcement spacings. However, the provisions of Code Section 7.6.5 indirectly limit the spacing of such reinforcement in one-way slabs. AC1 340.1R contains design aids for the application of Eq. (4.2a). 4.2.2 CEB recommendations - Crack control recommendations proposed in the European Model Code for Concrete Structures apply to prestressed as well as reinforced concrete and can be summarized as follows: The mean crack width, wm in beams is expressed in terms of the mean crack spacing, srm such that Kn = L&n

(4.4)

where (4.5)

beams h1

=

distance from neutral axis to the reinforcing steel, in.

Simplification of Eq. (4.la) yielded the following equation w = 0.076~fs ~AX D3

(4.2)

and represents the average strain in the steel. fs steel stress at the crack f II = steel stress at the crack due to forces causing

cracking at the tensile strength of concrete

K

= bond coefficient, 1.0 for ribbed bars, reflecting

influence of load repetitions and load duration

ACI COMMITTEE REPORT

224~018

The mean crack spacing is S

(4.6)

rm

where

c = clear concrete cover S bar spacing, limited to 15d, x2 = 0.4 for ribbed bars x3 = depends on the shape of the stress diagram, 0.125 for bending QR = A, /A, A t = effective area in tension, depending on arrangement of bars and type of external forces; it is limited by a line c + 7d, from the tension face for beams; in the case of slabs, not more than halfway to the neutral axis A simplified formula canbe derived for the mean crack width in beams with ribbed bars, d

f

w, = 0.7 _“- 3c + 0.05 -! E QR

(4.7)

A characteristic value of the crack width, presumably equivalent to the probable maximum value, is given as 1.7~~. 4.3 - Crack control in two-way slabs and plates Crack control equations for beams underestimate the crack widths developed in two-way slabs and plates4.7 and do not tell the designer how to space the reinforcement. The cracking mechanism in twoway slabs and plates is controlled primarily by the steel stress level and the spacing of the reinforcement in the two perpendicular directions. In addition, the clear concrete cover in two-way slabs and plates is nearly constant [3/4 in. (19 mm) for interior exposure], whereas it is a major variable in the crack control equations for beams. Analysis of data in the only major work on cracking in two-way slabs and plates4s7 has provided the following equation for predicting the maximum crack width:

w

s

&,sI:

(4.8)

n

where the radical rl = db,s21et, is termed the grid index, and can be transformed into

] k

=

fs

=

d b1

=

s1

=

s2

=

46

1" =

Qrl

S

w=

P

= fracture coefficient, having a value k = 2.8 x lO-5 for uniformly loaded restrained two-way

= =

action square slabs and plates. For concentrated loads or reactions, or when the ratio of short to long span is less than 0.75 but larger than 0.5, a value of k = 2.1 x 1O-5 is applicable. For span aspect ratios 0.5, k = 1.6 x 1O-s (as defined in Section 4.2.1) 1.25 (chosen to simplify calculations though varies between 1.20 and 1.35) actual average service load stress level, or 40 percent of the design yield strength fy, ksi diameter of the reinforcement in direction “1” closest to the concrete outer fibers, in. spacing of the reinforcement in direction “l”, in. spacing of the reinforcement in perpendicular direction “2”, in. direction of reinforcement closest to the outer concrete fibers; this is the direction for which crack control check is to be made active steel ratio Area _ - ofV steel - -A, Pper - ft -width - 12 (dbt + 2CJ

where Cl is clear concrete cover measured from the tensile face of concrete to the nearest edge of the reinforcing bar in direction b& 1VW w = crack width at face of concrete, in., caused by flexural load Subscripts 1 and 2 pertain to the directions of reinforcement. For simply supported slabs, the value of k should be multiplied by 1.5. Interpolated k values apply for partial restraint at the boundaries. For zones of flat plates where transverse steel is not used or when its spacing s2 exceeds 12 in., use s2 = 12 in. in the equation. If strain is used instead of stress, Eq. (4.8) becomes (4.9) where values of the kl = 29 x 100~ times the k values previously listed. References 4.8 and 340.1R contain design aids for the application of these recommendations. 4.4 - Tolerable crack widths versus exposure conditions in reinforced concrete Table 4.1 is a general guide for tolerable crack widths at the tensile face of reinforced concrete structures for typical conditions and is presented as an aid to be used during the design process. The table is based primarily on Reference 4.9. It is important to note that these values of crack width are

CONTROL OF CRACKING

TABLE 4.1 - Tolerable crack widths, reinforced concrete Exposure condition Dry air or protective membrane Humidity, moist air, soil Deicing chemicals Seawater and seawater spray: wetting and drying Water retaining structures*

Tolerable crack width, in.

(mm)

0.016 0.012 0.007

(0.41) (0.30) (0.18)

0.006 0.004

(0.15) (0.10)

*Excluding nonpressure pipes

not always a reliable indication of the corrosion and deterioration to be expected. In particular, a larger cover, even if it leads to a larger surface crack width, may sometimes b e preferable for corrosion control in certain environments. Thus, the designer must exercise engineering judgment on the extent of crack control to be used. When used in conjunction with the recommendations presented in Sections 4.2.1 and 4.2.3 to limit crack width, it should be expected that a portion of the cracks in the structure will exceed these values by a significant amount. 4.5 - Flexural cracking in prestressed concrete Partially prestressed members, in which cracks may appear under working loads, are used extensively. Cracks form in these members when the tensile stress exceeds the modulus of rupture of the concrete (Sfl to 90 under short-term conditions). The control of these cracks is necessary mainly for esthetic reasons. The residual crack width, after removal of the major portion of the live load, is small [about 0.001 in. to 0.003 in. (0.03 to 0.08 mm)] and therefore, crack control is usually not necessary if the live load is transitory. The prediction of crack widths in prestressed concrete members has received far less attention than in reinforced concrete members. The available experimental data are limited and, at the same time, the number of variables is greater in prestressed members. 4.5.1 Crack prediction equations - One approach to crack prediction, w h i c h r e l a t e s i t t o t h e nonprestressed case, has two steps. First the decompression moment is calculated, at which the stress at the tension face is zero. Then the member is treated as a reinforced concrete member and the increase in stress in the steel is calculated for the additional loading. The expressions given for crack prediction in nonprestressed beams may be used to estimate the cracks for the load increase above the decompression moment. A multiplication factor of about 1.5 is needed when strands, rather than deformed bars, are used nearest to the beam surface in the

224R-19

prestressed member to account for the differences in bond properties. The difficulty with this approach is the complexity of calculations. The determination of the decompression moment and, especially, the stress in the steel is complicated and unreliable unless elaborate methods are used.4.10 For this reason, approximate methods for crack width prediction are attractive. These are not much less accurate than the more complicated methods, and the lack of sufficient data, covering large variations in the variables, precludes further refinements at this date. The CEB Model Code has the same equation for the prediction of the crack width in prestressed members as in nonprestressed members (see Section 4.2.2). The increase in steel strain is calculated from the decompression stage. Several other equations have been proposed.4.11-4.“0 Limited evidence seems to indicate that unbonded members develop larger cracks than bonded members. Nonprestressed deformed bars may be used to reduce the width of the cracks to acceptable levels. The cracks in bonded post-tensioned members are not much different from cracks in pretensioned beams. 4.5.2 Allowable crack widths - Some authors state that corrosion is a greater problem in prestressed concrete members because of the smaller area of steel used. However, recent research results4.“’ indicate that there is no general relationship between cracking and corrosion in most circumstances. Furt h e r m o r e cracks close upon removal of the load, and the use of crack width limits should depend on the fluctuation and magnitude of the live load.

4.6 - Anchorage zone cracking in prestressed concrete Longitudinal cracks frequently occur in the anchorage zones of prestressed concrete members due to transverse tensile stresses set up by the concentrated forces.4.22T 4.23 Such cracks may lead to (or in certain cases are equivalent to) the failure of the member. Transverse reinforcement (stirrups) must be designed to restrict these cracks. Two types of cracks may develop: spalling cracks which begin at the end face (loaded surface) and propagate parallel to the prestressing force, and bursting cracks which develop along the line of the force or forces, but away from the end face. For many years stirrups were designed to take the entire calculated tensile force based on the analysis of the uncracked section. Classical and finite-element analyses show similar stress distributions for which the stirrups are to be provided. However, since experimental evidence shows that higher stresses can result.4.23 than indicated by these analyses, and the consequences of under-reinforcement

224R-20

ACI COMMITTEE REPORT

can be serious, it is advisable to provide more steel

than required by this type of analysis. More recently, designs have been based on cracked section analyses. A design procedure for post-tensioned members using a cracked section analysis4.24 has found acceptance with many designers. For pretensioned members, an empirical equation has proven to be quite usefu1.4.25 Spalling cracks form between anchorages and propagate parallel to the prestressing forces and may cause gradual failure, especially when the force acts near and parallel to a free edge. Since analyses show that the spalling stresses in an uncracked member are confined to near the end face, it is important to place the first stirrup near the end surface, and to distribute the stirrups over a distance equal to at least the depth of the member to fully account for both spalling and bursting stresses. Precast blocks with helical reinforcement may be used when the prestressing forces are large.

4.7 - Tension cracking The cracking behavior of reinforced concrete members in tension is similar to that of flexural members, except that the maximum crack width is larger than that predicted by the expressions for flexural members.4.26T 4.27 The lack of strain gradient, a n d resultant restraint imposed by the compression zone of flexural members, is probably the reason for the larger tensile crack width. Data are limited but it appears that the maximum tensile crack width may be expressed approximately in a form similar to that used for flexural crack width. w = O.lOf,&tA x 10-3

(4.10)

References 4.1. Leonhardt, Fritz, “Crack Control in Concrete Structures,” IABSE Surveys No. S4/77, International Association for Bridge and Structural Engineering, Zurich, 1977, 26 pp . 4.2. Yerlici, V. A., “Minimum Wall Thickness of Circular

Concrete Tanks,” Publication No. 35-11, International Association for Bridge & Structural Engineering, Zurich, 1975, p. 237. 4.3. ACI Committee 224, “Causes, Mechanism, and Control of Cracking in Concrete,” ACI Bibliography No. 9, American Concrete Institute, Detroit, 1971, 9.2 pp. 4.4. Nawy, Edward G., and Neuwerth, G. E., “Behavior of Concrete Slabs, Plates and Beams with Fiber Glass as Main Reinforcement,” Proceedings, ASCE, V. 103, ST2, Feb. 1977, pp. 421-440. 4.5. Clark, Arthur P., “Cracking in Reinforced Concrete Flexural Members,” ACI JOURNAL , Proceedings V. 52, No. 8, Apr. 1956, pp. 851-862. 4.6. Gergely, Peter, and Lutz, Leroy A., “Maximum Crack Width in Reinforced Concrete Flexural Members,”

Causes, Mechanism, and Control of Cracking in Concrete, SP-20, American Concrete Institute, Detroit, 1968, pp. 87-117. 4.7. Nawy, Edward G., and Blair, Kenneth W., “Further Studies on Flexural Crack Control in Structural Slab Systems,” Cracking, Deflection, and Ultimate Load of Concrete Slab Systems, SP-30, American Concrete Institute, Detroit, 1971, pp. 1-41. 4.8. Nawy, Edward G., “Crack Control Through Reinforcement Distribution in Two-Way Acting Slabs and Plates,” ACI JO U R N A L, Proceedings V. 69, No. 4, Apr. 1972, pp. 217-219. 4.9. Nawy, Edward G., “Crack Control in Reinforced Concrete Structures,” ACI JOURNAL , Proceedings V. 65, No. 10, Oct. 1968, pp. 825-836. 4.10. Nilson, Arthur H., Design of Prestressed Concrete, John Wiley and Sons, New York, 1978, 526 pp. 4.11. Abeles, Paul W., “Cracks in Prestressed Concrete Beams,” Proceedings, Fifth IABSE Congress (Lisbon, 1956), International Association for Bridge and Structural Engineering, Zurich, 1956, pp. 707-720. 4.12. Bennett, E. W., and Dave, N. J., “Test Performances and Design of Concrete Beams with Limited Prestress,” The Structural Engineer (London), V. 47, No. 12, Dec. 1969, pp. 487-496. 4.13. Holmberg, Ake, and Lindgren, Sten, “Crack Spacing and Crack Widths Due to Normal Force and Bending Moment,” Document D2:1970, National Swedish Council for Building Research, Stockholm, 1970, 57 pp. 4.14. Rao, A.S.P.; Gandotra, K.; and Ramaswamy, G. S., “Flexural Tests on Beams Prestressed to Different Degrees of Prestress,” Journal, Institution of Engineers (Calcutta), V. 56, May 1976. 4.15. Bate, Stephen C. C., “Relative Merits of Plain and Deformed Wires in Prestressed Concrete Beams Under Static and Repeated Loading,” Proceedings, Institution of Civil Engineers (London), V. 10, Aug. 1958, pp. 473-502. 4.16. Bennett, E. W., and Chandrasekhar, C. S., “Calculation of the Width of Cracks in Class 3 Prestressed Beams,” Proceedings, Institution of Civil Engineers (London), V. 49, July 1971, pp. 333-346. 4.17. Hutton, S. G., and Loov, R. E., “Flexural Behavior of Prestressed, Partially Prestressed, and Reinforced Concrete Beams,” ACI JO U R N A L , Proceedings, V. 63, No. 12, Dec. 1966, pp. 1401-1410. 4.18. Krishna, Raju N.; Basavarajuiah, B. S.; and Ahamed Kurty, U. C., “Flexural Behavior of Pretensioned Concrete Beams with Limited Prestress,” Building Science, V. 8, No. 2, June 1973, pp. 179-185. 4.19. Stevens, R. F., “Tests on Prestressed Reinforced Concrete Beams,” Concrete (London), V. 3, No. 11, Nov. 1969, pp. 457-462. 4.20. Nawy, E. G., and Huang, P. T., “Crack and Deflection Control of Pretensioned Prestressed Beams,” Journal, Prestressed Concrete Institute, V. 22, No. 3, May-June 1977, pp. 30-47. 4.21. Beeby, A. W., “Corrosion of Reinforcing Steel in Concrete and Its Relation to Cracking,” The Structural Engineer (London), V. 56A, No. 3, Mar. 1978, pp. 77-81. 4.22. Gergely, Peter, “Anchorage Systems in Prestressed Concrete Pressure Vessels; Anchorage Zone Problems,” ORNL-TM-2378, Oak Ridge National Laboratory, U.S. Atomic Energy Commission, Oak Ridge, Tenn., 1969, pp. l-49. 4.23. Zielinski, J. L., and Rowe, R. E., “An

CONTROL OF CRACKING

Investigation of the Stress Distribution in the Anchorage Zones of Post-Tensioned Concrete Members,” Technical Report No. 9, Cement and Concrete Association, London, Sept. 1960, 32 pp. 4.24. Gergely, P., and Sozen, M. A., “Design of Anchorage Zone Reinforcement in Prestressed Concrete Beams,” Journal, Prestressed Concrete Institute, V. 12, No. 2, Mar.-Apr. 1967, pp. 63-75. 4.25. Marshall, W. T., and Mattock, A. H., “Control of Horizontal Cracking in the Ends of Pretensioned Concrete Girders,” Journal, Prestressed Concrete Institute, V. 7, No. 5, Aug.-Oct. 1962, pp. 56-74. 4.26. Broms, Bengt B., “Crack Width and Crack Spacing in Reinforced Concrete Members,” ACI JOURNAL , Proceedings, V. 62, No. 10, Oct. 1965, pp. 1237-1256. 4.27. Broms, Bengt B., and Lutz, Leroy A., “Effects of Arrangement of Reinforcement on Crack Width and Spacing of Reinforced Concrete Members,” ACI JOURNAL, Proceedings V. 62, No. 11, Nov. 1965, pp. 1395-1410.

Chapter 5 - Long-term effects on cracking* 5.1 - Introduction

Cracking in concrete is affected by the long-term conditions to which the concrete element is subjected. In most cases, long-term exposure and longterm loading extend the magnitude of cracks in both reinforced and plain concrete. The discussion in this chapter summarizes the major long-term factors which affect the crack control performance of concrete. 5.2 - Effects of long-term loading

As discussed in Chapter 2, both sustained and cyclic loading increase the amount of microcracking in concrete. The total amount of microcracking appears to be a function of the total strain and is largely independent of the method by which the strain is induced. Microcracking due to long-term loading may well be an effect, rather than a major cause, of creep, and microcracks formed at service load levels do not seem to have a great affect on the strength or serviceability of concrete. The effect of sustained or repetitive loading on macroscopic cracking, however, may be an important consideration in the serviceability of reinforced concrete members, especially in terms of corrosion of reinforcing steel and appearance. The increase in crack width due to long-term or repetitive loading can vary between 10 percent and 1,000 percent over the span of several years. 5.1-5.8 While there is a large scatter in the data, information obtained from sustained loading tests of up to 2 .7,5.8 and fatigue tests with up to one million 5.4, 5.5,5.8,5.9 indicate that a doubling of crack cycles width with time can be expected. Under most conditions, the spacing of cracks does not change with time at constant levels of stress. 5.4,5.7,5.8 An excep*Principal authors: David Darwin and Ernest K. Schrader.

224-21

tion to this occurs at low loads or in beams with high percentages of reinforcement, in which case the total number and width of cracks increase substantially after the loading has begun.5.2,5.4,5.8 The largest percentage increase in crack width is then expected in flexural members subject to low levels of load, since the cracks take more time to develop. For both prestressed and reinforced concrete flexural members, long-term loading and repetitive loading seem to give about the same crack widths and spacing.5.9 The rate of crack development, however, is considerably faster under repetitive loading. 5.5,5.8-5.10 As discussed in Chapter 4, crack width is a function of cover. For short-term static and fatigue loading, surface crack width is approximately proportional to the steel s t r a i n 5 . 7 , 5 . 8 , 5 . 1 0 Cracks grow in width under sustained loading at a decreasing rate. However, the rate of growth is faster than the average observed surface strain at the level of the steel. For long term loading, crack width is proportional to the steel strain (including the effects of creep), plus the strain induced in the concrete due to shrinkage.5.7 Under initial loads, cracks adjacent to reinforcement are restricted by the bond between the steel and the concrete,5.7-5.11 and thus the width of surface cracks do not provide a good indication of the exposure of the reinforcing steel to corrosive conditions. Over a period of time, however, the adhesion bond between the steel and the concrete undergoes breakdown. After about 2 years, the crack width at the reinforcement is approximately equal to the crack width at the surface.5.7 At this stage, cracks in flexural members are triangular in shape increasing in width from the neutral axis to the soffit, and are approximately uniform across the width of the beam. Therefore, after a few years, the width of a surface crack provides a good estimate of the crack width at the level of the reinforcing steel. Many questions remain as to the importance of crack width on the serviceability of reinforced and prestressed concrete members. 5.12-5.14 Added cover is generally acknowledged as a method of improving the corrosion protection for reinforcing steel. Since additional cover also results in added surface crack width, and since this surface crack width appears to provide a good estimate of the crack width at the level of the steel, the entire question of the importance of crack width on corrosion protection remains open. It does seem clear that crack widths predicted on the basis of short term static tests do not provide a precise guide to crack widths in structures actually in service. 5.3 - Environmental effects

The long-term effects of an adverse environment in both producing and in enlarging concrete cracks 5.15,5.16 can be damaging to both concrete and

224R-22

ACI COMMll-i=EE REPORT

reinforcement. If concrete is not resistant to freezing and thawing when critically saturated, it will develop cracks when frozen. The lack of such resistance may be due to either the use of non-frost-resistant coarse aggregate or the failure to produce a satisfactory air-void system or failure to protect the concrete from freezing prior to the reduction of the freezable water content by maturity to a tolerable range. The achievement of critical saturation in nonfrost-resistant concrete may be facilitated by the presence of preexisting cracks which allow entry of water more readily than would be the case otherwise. The initiation of D-cracking near joints or other cracks in pavements is a good example. In more extreme cases, it is not uncommon for cracks in the roadway deck of dams and navigation locks (caused either by thermal stress or shrinkage of the richer topping mix) to spall due to water which freezes in the cracks themselves (independent of the frost resistance of the concrete). On the otherhand, preexisting cracks may also function to allow concrete to dry below critical saturation before freezing, when this might not occur in the absence of such cracks. Hence, the role of cracks as they effect the deficiencies in frost resistance will vary with the environmental conditions (e.g., typical time of drying after wetting before freezing), crack width, ability of cracks to drain, etc. If the aggregate used in the concrete is durable under freeze-thaw conditions and the, strength of the concrete is high, the concrete durability will better. (AC1 201.2R). Field exposure tests of reinforced concrete beams5*17 (subjected to freezing and thawing and an ocean side environment) indicate that the use of air-entrained concrete made the beams more resistant to weathering than the use of nonair-entrained concrete. Beams with modern deformed bars were found to be more durable than those using old-style deformations. Maximum crack widths did not increase with time when the steel stress was less than 30 ksi, (210 MPa) but did increase substantially (50 to 100 percent) over a 9 year period when the steel was 30 ksi (210 MPa) or more. 5.4 - Aggregate and other effects Concrete may crack as the result of expansive reactions between aggregate and alkalis derived from cement hydration, admixtures or external sources (e.g., curing water, ground water, alkaline solutions stored or used in the finished structure). Possible solutions to these problems include limitations on reactive constituents in the aggregate, limitations on the alkali content of cement, or addition of a satisfactory pozzolanic material. The potential for some expansive reactions, e.g., alkali-carbonate, is not reduced by pozzolanic admixtures. AC1 201.2R and Reference 5.18 give details on identification and evaluation of aggregate reactivity. Based on reports of AC1 Committees 201 and

212 9 5.15.5.16 the possible hazard of using calcium chlo-

ride in a water-soluble salt environment warrants a recommendation against its use under such circumstances. Also, the use of calcium chloride in reinforced structures exposed to unusually moist environments is to be avoided regardless of the presence or absence of water-soluble salts in adjacent waters and soils. Detrimental conditions may also result from the application of deicing salts to the surface of hardened concrete. When such applications are necessary, calcium chloride or sodium chloride should be used and only within recommended application rates. Concrete subjected to water soluble salts should be air entrained [6.5 to 7.5 percent for normal 3L4 in. (19 mm) MSA concrete and 4.5 to 5.5 percent for F/2 in. (38 mm) MSA concrete], should have adequate cover (about 2 in.), and should be made with a high-quality mix yielding low permeability. 5.5 - Use of polymers in improving cracking c h a r acterisitics Extensive work is available on the use of polymers in modifying the characteristics of concrete.5*1gy 5.20p 5.21 Polymer-portland cement concretes have a large deformation capacity, high tensile and compressive strengths and negligible permeability. The tensile splitting strength can be as high as 1550 psi (10.7 MPa).5-22 Polymer impregnation is another method of introducing beneficial polymer systems into concrete. This procedure creates a ‘layer’ of high quality material to the depth that has been impregnated. These materials are discussed in greater detail in Chapter 6. Because of these desirable characteristics, it is expected that structural elements made with polymer modified concrete will exhibit superior serviceability in cracking, deflection, creep, shrinkage, and permeability. Referenees 5.1. Bate, Stephen C. C., “A Comparison Between Prestressed Concrete and Reinforced Concrete Beams Under Repeated Loading,” Proceedings Institution of Civil Engineers (London), V. 24, Mar. 1963, pp. 331-358. 5.2. Brendel, G., and Ruhle, H., “Tests on Reinforced Concrete Beams Under Long-Term Loads (Dauerstandversuche mit Stahlbetonbalken),” Proceedings, Seventh IABSE Congress (Rio de Janeiro, 1964), International Association of Bridge and Structural Engineering, Zurich, 1964, pp. 916-922. 5.3. Lutz, LeRoy A.; Sharma, Nand K.; and Gergely, Peter, “Increase in Crack Width in Reinforced Concrete Beams Under Sustained Loading,” ACI JOURNAL , Proceedings, V. 64, No. 9, Sept. 1968, pp. 538-546. 5.4. Abeles, Paul W.: Brown, Earl L. II; and Morrow, Joe W., “Development and Distribution of Cracks in Rectangular Prestressed Beams During Static and Fatigue Loading,” Journal, Prestressed Concrete Institute, V. 13, No. 5, Oct. 1968, pp. 36-51.

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5.5. Bennett, E. W., and Dave, N. J., “Test Performances and Design of Concrete Beams with Limited Prestress,” The Structural Engineer (London), V. 47 No. 12,

Dec. 1969, pp. 487-496.

.

5.6. Holmberg, A., and Lindgren, S., “Crack Spacing and Crack Width Due to Normal Force or Bending Moment,” Document D2, National Swedish Council for Building Research, Stockholm, 1970, 57 pp. 5.7. Illston, J. M., and Stevens, R. F., “Long-term Cracking in Reinforced Concrete Beams,” Proceedings, Institution of Civil Engineers (London), Part 2, V. 53, Dec. 1972, pp. 445-459. 5.8. Holmberg, Ake, “Crack Width Prediction and Minimum Reinforcement for Crack Control,” Dansk Selskab for Byaningsstatik (Copenhagen), V. 44, No. 2, June 1973, pp. 41-50. 5.9. Rehm, Gallus, and Eligehausen, Rolf, “Lapped Splices of Deformed Bars Under Repeated Loadings (Ubergreifungsstosse von Rippenstahlen unter nicht ruhender Belastung),” Beton und Stahlbetonbau (Berlin), No. 7, 1977, pp. 170-174. 5.10. Stevens, R. F., “Tests on Prestressed Reinforced Concrete Beams,” Concrete (London), V. 3, No. 11, Nov. 1969, pp. 457-462. 5.11. Broms, Bengt B., “Technique for Investigation of Internal Cracks in Reinforced Concrete Members,” ACI JOURNAL, Proceedings, V. 62, No. 1, Jan. 1965, pp. 35-44. 5.12. Atimtay, Ergin, and Ferguson, Phil M., “Early Chloride Corrosion of Reinforced Concrete - A Test Report,” ACI JOURNAL , Proceedings V. 70, No. 9, Sept. 1973, pp. 606-611. 5.13. Beeby, A. W., “Concrete in the Oceans - Cracking and Corrosion,” Technical Report No. 1, Cement and Concrete Association (London), 1978. 5.14. Beeby, A. W., “Corrosion of Reinforcing Steel in Concrete and Its Relation to Cracking,” The Structural Engineer (London), V. 56A, No. 3, Mar. 1978, pp. 77-81. 5.15. Mather, Bryant, “Cracking Induced by Environmental Effects,” Causes, Mechanism, and Control of Cracking in Concrete, SP-20, American Concrete Institute, Detroit, 1968, pp. 67-72. 5.16. Mather, Bryant, “Factors Affecting Durability of Concrete in Coastal Structures,” Technical Memorandum No. 96, Beach Erosion Board, Washington, D.C., June 1957. 5.17. Roshore, Edwin C., “Field Exposure Tests of Reinforced Concrete Beams,” ACI JOURNAL , Proceedings V. 64, No. 5, May 1967, pp. 253-257. 5.18. Woods, Hubert, Durability of Concrete Construction, Monograph No. 4, American Concrete Institute/Iowa State University, Detroit, 1968, 187 pp. 5.19. Brookhaven National Laboratory, “Concrete Polymer Materials,” BNL Report 50134 (T-5091, 1968. 5.20. Polymers in Concrete, SP-40, American Concrete Institute, Detroit, 1973, 362 pp. 5.21. Polymers in Concrete, SP-58, American Concrete Institute, Detroit, 1978, 420 pp. 5.22. Nawy, Edward G.; Ukadike, Maurice M.; and Sauer, John A., “High Strength Field Polymer Modified Concretes,” Proceedings, ASCE, V. 103, ST12, Dec. 1977, pp. 2307-2322.

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Chapter 6 - Control of cracking in concrete layered systems* 6.1 - Introduction A “layered” concrete system can be created by a mortar or concrete overlay (topping) placed on an existing concrete surface. The use of “layered” concrete systems has been increasing during the last 10 years in the renovation of deteriorating bridge decks, strengthening and/or renovation of concrete pavements, warehouse floors, walkways, etc., and in new two-course construction of decks and pavements. The overlay can be portland cement low slump dense concrete (LSDC), polymer-portland cement concrete (PPCC), more commonly referred to as latex modified concrete (LMC), fiber reinforced concrete (FRC), or internally sealed concrete. A “layered” system can also be created by impregnating the upper portion [l/z to 3 in. (10 to 80 mm1 ] of existing concrete with a monomer system that requires polymerization after soaking. The major sources and types of cracking in these layered concrete systems are: 1. Differential shrinkage cracking 2. Reflective cracking (stress cracking) 3. Differential temperature cracking 4. Edge curling and delamination 5. Incorrect construction practices Long term observations 6.1-6.3 of many “layered” concrete systems have shown that differential shrinkage cracks are by far the most common and most likely to increase and widen with time. 6.2 - Fiber reinforced concrete (FRC) overlays When properly proportioned, mixed, and placed, a crack resistant topping layer of FRC can be the solution to certain field problems. Fibrous concrete overlays of highways, airfields, warehouse floors, walkways, etc., have been used since the early 1970s. Fibers are usually steel with lengths between 10 and 60 mm (l/2 to 2l/2 in.). The effects of fibrous concrete on cracking in a “layered” system depend largely on the field conditions of each situation. Some typical observations for similar field or laboratory conditions are discussed below.6*2-6*7 6.2.1 Bond to underlying concrete - During early fibrous concrete overlay work, it was thought that a “partially bonded” layer was the ideal system. The term “partially bonded” means that no deliberate attempt is made to bond or to debond the topping layer to the underlying material through agents, fasteners, polyethylene sheet, etc. The surface to be overlaid is cleaned of all loose material, usually by hosing, and generally left in damp condition. After the evaluation of partially bonded projects, this procedure has become the least desirable technique to *Principal authors: Alfred G. Bishara and Ernest K. Schrader.

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use. Over a period of several years many partially bonded FRC overlays have shown noticeable amounts of reflective cracking and edge curling. The curled edges are typical in thin overlays [less than about 3 in. (76 mm)] and can result in cracks if subjected to long-term dynamic loading. If the base slab is relatively crack free, or if the overlay is of sufficient thickness and strength to resist the extension of cracks in the original slab, a bonded layer with matched joints is generally the best approach. If the FRC layer is of sufficient thickness, a totally unbonded overlay is generally best where severe cracking is present or may develop in the base slab. Essentially unbonded systems have been constructed satisfactorily where FRC is placed over an asphalt layer. The asphalt itself will act as a debonding layer if it has a reasonably smooth surface without potholes. This type of construction lends itself particularly well to deteriorated airfield slabs which have been resurfaced with asphaltic concrete but require additional rigid pavement to take increased loads imposed by heavy aircraft. Another technique, which has been used when the base material to be overlaid is reasonably smooth, consists of placing the FRC over a layer of polyethylene sheet. On irregular, spalled, or potholed surfaces a thin leveling and debonding layer of sand or asphalt is desirable. 6.2.2 Fiber size and volume - The crack arresting mechanism on which the basic theory of FRC is founded depends on fiber spacing.6.8 Although fiber size and volume have little effect on the formation of the first crack they are major factors influencing subsequent crack development. As fiber diameter increases for any given volume percentage, the number of fibers decreases and the spacing between fibers increases. Also, as the volume percentage decreases, the spacing increases. If the fiber spacing becomes relatively large [more than about 5 mm (0.2 in.)], the crack arresting mechanism is limited. Regardless of the reason, as the fiber spacing increases, the number of small cracks decreases, but the number and width of larger cracks increase. For concrete with 20 mm t3/4 in.) aggregate, about 0.9 percent fibers by total volume will provide substantial crack resistance. For concrete with 10 mm (3/8 in.) aggregate about 1.2 percent is normal, and for mortar, 1.4 to 1.8 percent is adequate. If fiber contents much greater than these are used, or if aggregate gradations are not suitable, high cement and water requirements result and the FRC layer is susceptible to shrinkage cracks. 6.2.3 Fiber type and shape - Because of their increased resistance to pullout, deformed steel fibers have an advantage over smooth ones with regard to both pre- and post-cracking behavior. However, the advantage is not always worth the additional expense.

The basic crack theory is applicable to both glass and metallic fibers, but the two types do exhibit some difference in physical crack behavior. Test+* have shown that glass FRC has less ability to store energy after its failure in flexure than steel FRC. Also, microcracking in the general vicinity of a major crack is typically more prominent with steel than glass. The failure (crack) zone for glass is more localized. 6.2.4 Fibers in open cracks - There has been considerable discussion about the condition and effectiveness of steel fibers that bridge over or through a crack. At the time of cracking, the fibers lose their bond to the concrete but continue to provide a “mechanical resistance to pullout.” This post-cracking strength is one of the most important characteristics of FRC. The “obvious” problem is that after cracking, steel fibers will oxidize and provide no long-term benefit. However, the majority of investigations 6.3,6.5,6.6 have shown, that if the cracks are tight (0.001 - 0.003 in. (0.03-0.08 mm)], the fibers will not oxidize, even after several years of exposure. Longterm evaluations are currently underway.6.3 6.2.5 Mix proportion conditions-ACI 544.3R provides detailed information on suitable mixture proportions for steel fiber reinforced concrete. The water requirement for fibrous concretes is higher than that of normal concrete due to the high surface area of the fibers. The high water content provides the basic ingredient for shrinkage cracks. Through the use of water reducing admixtures, the mix water can be held to reasonable levels.6-gp ‘JO If possible, these admixtures should be used to adjust the mix proportioning for a bonded overlay so that the water/cement ratio and cement factor approach the same values as used in the underlying material, If possible, the overlay should have aggregates of similar physical properties unless the original aggregates are unsuitable. 6.2.6 Joint overlays - Different methods of joint overlaying have been tried; most have been unsuccessful.6.7 As with conventional concrete overlays, if joints in a base slab are overlayed with FRC without taking special design precautions to prevent reflective cracking, the overlay will crack at joint locations. 6.3 - Latex modified concrete (LMC) overlays Latex modified mortar and concrete bonded overlays [3/4 to 1 l/z in. (20 to 40 mm)] have been used in

the renovation of deteriorated bridge decks and in new two-course construction to effectively resist the penetration of chloride ions from deicing salts and prevent the subsequent corrosion of the reinforcing steel and the spalling of the concrete deck.6.11,6.12 Some of these decks have been in use for over 10 years. Inspections of a large number of bridge decks overlaid with LMC6.1 have indicated that there is a

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high incidence of fine, random, shrinkage cracks in a large portion of the renovation jobs. This type of cracking is not as extensive in new two-course construction. Transverse cracks, spaced 3 to 4 ft (0.9 to 1.2 m) apart, also appear on many of the bridges inspected. However, there may be a relationship between the degree of transverse cracking and the intensity of heavy truck traffic during reconstruction. To keep the bridges in service, traffic is normally diverted to one lane, while renovation and application of the overlay proceed on an adjacent traffic lane. The quality of the overlay may be affected by the movement of the deck, although extensive data do not exist linking the effect of traffic-induced vibrations during reconstruction to deterioration or cracking in bridge decks. If traffic must be maintained, consideration should be given to placing overlays when traffic is low and/or when vehicle speed is restricted. To reduce the incidence of cracking and sub sequent loss of latex modified concrete overlays it is recommended 6 .1 that: 1. The surface of the underlying concrete should be cleaned by sand blasting to assure adequate bonding with the overlay. To reduce air pollution, particularly in urban areas, high pressure water jet cleaning [5000 to 6000 psi (35-40 MPa) at the nozzle] may be used just prior to placement of the overlay, in lieu of sand blasting; 2. The slump of latex modified concrete mixtures should be between 3 to 4 in. (75 to 100 mm) to reduce differential shrinkage and the high incidence of random cracking; 3. The finishing equipment should have been proven to be effective for adequately placing the concrete to the required density; 4. A thin coating of the overlay mixture should be thoroughly scrubbed into the surface of the underlying clean concrete immediately before placing the overlay mix to increase the bonding between the layers; coarser particles of the mixture which cannot be scrubbed into immediate contact with the surface of the underlying concrete, should be removed; 5. In new two-course construction, the overlay should be placed after removing the forms from the base concrete, so that stresses caused by the weight of the overlay are born by the underlying concrete. If placed before the forms are removed, the overlay will have to carry a portion of its own weight and may crack in negative moment regions; 6. Overlays should be placed only when the ambient weather conditions are favorable, as defined in ACI 308 on curing, or when appropriate actions are taken for cold-weather concreting (ACI 306R) or hotweather concreting (ACI 305R).

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6.4 - Polymer impregnated concrete (PIC) systems Surface impregnation and polymerization of concrete in place is a relatively new process but has been used successfully in a number of field projects.6.13,6.15 There has been considerable discussion about this procedure due to observations of cracks during or immediately after the drying step of these p r o j e c t s . I n t h e c a s e s t h a t h a v e b e e n evaluated, 6.14,6.15 the cracks were determined to either have been in the concrete prior to the impregnation or they were caused by improperly controlled drying during initial stages of the impregnation procedure. Temperatures during drying are usually in the range of 120 C (240 F) to 150 C (310 F) for about 4 to 12 hr. To some extent, thermal expansion will offset drying shrinkage until the concrete cools. Ideally, during the soak period and after cooling, the monomer will fill any cracks that have been created in the top surface of the concrete due to drying. The cracks will be mended when the monomer is polymerized. If a crack is open and can drain (as is the case with vertical surfaces and cracks through the full depth of a slab), the monomer can run out of the crack before it is polymerized, and no mending will occur. If a more viscous monomer is used, so that it does not drain from the crack, the depth of penetration into the concrete will be adversely affected. If there is a water source behind the material to be polymerized it is possible for moisture to re-enter the crack, after drying has been completed, but before the monomer soak starts. In this case, the presence of moisture prevents the monomer from entering the concrete adjacent to the crack, and the crack will not mend. The engineer should thoroughly evaluate all effects of the drying cycle in a PIC project and plan the drying temperatures and duration, the cooling cycle, and the monomer system to prevent the occurrence of unmended cracks. The strain capacity, thermal expansion, and specific heat of the material should be considered. Restraints, preventing movement at the perimeter of the concrete to be polymerized, should be avoided. The long-term influence of polymer impregnation on the behavior of cracking in concrete is not known at this time but will be established by the evaluation of currently completed field projects. References 6.1. Bishara, A. G., “Latex Modified Concrete Bridge Deck Overlays - Field Performance Analysis,” Report No. FHWA/OH/79/004, Federal Highway Administration, Washington, D.C., Oct. 1979, 97 pp. 6.2. Gray, B. H., “Fiber Reinforced Concrete - A General Discussion of Field Problems and Applications,” Technical Manuscript M-12, U.S. Army Construction Engineering Research Laboratory, Champaign, Apr. 1972. 6.3. Schrader, Ernest K., and Munch, Anthony V. “Deck

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Slab Repaired by Fibrous Concrete Overlay,” Proceedings, ASCE, V. 102, C01, Mar. 1976, pp. 179-196. 6.4. Gray, B. H.; Williamson, G. R.; and Batson, G. B., “Fibrous Concrete - Construction Material for the Seventies,” Conference Proceedings M-28, U.S. Army Construction Engineering Research Laboratory, Champaign, May 1972, 238 pp. 6.5. Hefner, S., “Fibrous Concrete McCarran International Airport,” Las Vegas, Nevada, Dec. 1974. 6.6. Rice, John L., “Fibrous Concrete Pavement Design Summary,” Technical Report No. M-134, U.S. Army Construction Engineering Research Laboratory, Champaign, June 1975, 13 pp. 6.7. Gray, B. H., and Rice, John L., “Fibrous Concrete for Pavement Applications,” Report No. M-13, U.S. Army Construction Engineering Research Laboratory, Champaign, Apr. 1972, 9 pp. 6.8. Shah, S. P., and Naaman, A. E., “Mechanical Properties of Glass and Steel Fiber Reinforced Mortar,” Department of Materials Engineering, University of Illinois, Chicago, Aug. 1975. 6.9. “Utilization of ‘Wirand’ Concrete in Bridge Decks,” Report by General Analytics, Monroeville, Pa., for Battelle Memorial Institute, May 1971. 6.10. Walker, A. J., and Lankard, D. R., “Bridge Deck Rehabilitation with Steel Fibrous Concrete,” Presented at the Third International Exposition on Concrete Construction (New Orleans, Jan. 1977), Battelle Columbus Laboratories, 1977. 6.11. Bishara, A. G., and Tantayanondkul, P., “Use of Latex in Concrete Bridges Decks,” Report No. EES 435 (ODOT-12-74) Ohio Department of Transportation, The Ohio State University, 1974. 6.12. Clear, K. C., “Time to Corrosion of Reinforcing Steel in Concrete Slabs,” Transportation Research Record, No. 500, Transportation Research Board, 1974, pp. 16-24. 6.13. Schrader, Ernest K.; Fowler, David W.; Kaden, Richard A., and Stebbins, Rodney J., “Polymer Impregnation Used in Concrete Repairs on Cavitation/Erosion Damage,” Polymers in Concrete, SP-58, American Concrete Institute, Detroit, 1978, pp. 225-248. 6.14. Depuy, G. W., “Recent Developments in ConcretePolymer Materials,” Second International Symposium on Concrete Technology (Monterrey, Mexico, Mar. 19751, U.S. Bureau of Reclamation, Denver, 1975. 6.15. Smoak, W. G., “Polymer Impregnation of New Concrete Bridge Deck Surfaces,” Interim Report No. FHWA-RD-75-72, U.S. Bureau of Reclamation, Denver, Prepared for Federal Highway Administration, Washington, D.C., June 1975.

Chapter 7 - Control of cracking in mass concrete* 7.1 - Introduction Temperature induced cracking in a large mass of

concrete can be prevented if proper measures are taken to reduce the amount and rate of temperature change. Measures commonly used include precooling, post-cooling or a combination of the two, and more

*Principal authors: Donald L. Houghton and Roy W. Carlson.

recently, thermal insulation has been used to protect exposed surfaces. The degree of temperature control necessary to prevent cracking varies greatly with such factors as the location, the height and thickness of the structure, the character of the aggregate, the properties of the concrete and the external restraints. Although a large amount of the data for this chapter has been obtained by experience gained from the use of mass concrete in dams, it applies equally well in mass concrete used in other structures such as steam power plants, powerhouses, bridge and building foundations, navigation locks, etc. Tremie concrete, a specialized type of mass concrete, has been amply covered in Chapter 8 of ACI 304 and will not be discussed in this report. The location of the structure affects the degree of temperature control which will be required. Generally at high altitudes the daily variations in temperature are greater than at low altitudes. Often at high altitudes, the ambient temperature variation alone may be sufficient to cause cracks to form at exposed surfaces. These surface cracks continue inward with only approximately half the stress which is necessary to cause internal cracking. A similar condition is likely to be found when a structure is located at a high latitude; only in this case the temperature variations are seasonal, rather than daily. In the case of a dam, the height affects the need for crack control. If the dam is very high, the design stresses will be high and more cement must be used to provide the stipulated factor of safety. This makes for more heat generation and a consequent tendency toward higher internal temperatures. Also, the higher dam will have greater horizontal dimensions which cause greater restraint and the need for still closer temperature control. The properties of the concrete affect the problem of crack control. Concretes differ widely in the amount of tensile strain they can withstand before cracking. For strain which is applied rapidly, the two factors which govern the strain capacity are the modulus of elasticity and the tensile strength. For strain which is applied slowly, the creep (or relaxation) of the concrete is important. The factors affecting strain capacity and creep rate are discussed more fully in Section 7.2. Another important property of concrete is the coefficient of thermal expansion. The amount of strain which a temperature change will produce is directly proportional to the coefficient of thermal expansion of the concrete. The average coefficient of thermal expansion of mass concrete is about 9 millionths per deg C (5 millionths/F), but with some aggregates, the coefficient may be as high as 15 millionths or as low as 7 millionths (4 to 8 millionths/F). Thus, in the extreme case, where a concrete has a low tensile strength, a high modulus o f elasticity, a high coefficient of thermal expansion, and is fully restrained, it may crack when there is a quick drop in

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temperature of only 3 C (6 F). On the other hand, some concretes can withstand a quick drop in temperature of as much as 10 C (20 F), even when fully restrained. More data on the thermal expansion of concrete may be found in the reports of ACI Committee 207 (ACI 207.1R and ACI 207.2R). From these considerations, it is apparent that the degree of crack control necessary for the safe elimination of joints may vary from nothing at all, for a dam near the equator with favorable aggregates, to very costly measures, in a location where temperature variations are great and where the only economical aggregates have high elastic moduli and high thermal expansion. In the latter case, present practice calls for both precooling and post-cooling, and for the application of thermal insulation to exposed surfaces during cold weather. The insulation is left in place long enough to permit the concrete temperature at the surfaces to slowly approach the ambient, or until additional concrete is placed on or against the surface being protected. Additional research into the most effective use of thermal insulation is needed particularly for regions having severe or sub-arctic climates. There are two measures which can be taken to provide safety against cracking. The first is to modify the materials and mix proportions to produce concrete having the best cracking resistance, or the greatest tensile strain capacity. This may require careful aggregate selection, using the minimum cement content for interior concrete, restricting the maximum aggregate size, or using other specialized procedures. The second measure to prevent cracking is to control the factors which produce tensile strain. This may mean precooling, post-cooling, insulating (and possibly heating) the exposed surfaces of the concrete during cold weather and designing to minimize strains around galleries and other openings. 7.2 - Crack resistance The tensile strain which concrete can withstand varies greatly with the composition of the concrete and the strain rate. When strain is applied slowly, the strain capacity is far greater than when the action is rapid. Thus, concrete in the interior of a large mass which must cool slowly, can undergo a large strain before failure. If concrete contains rough textured aggregate of small maximum size, the strain capacity will be high. However, there is an optimum with respect to the aggregate size. Smaller aggregate requires more cement for a given strength which results in more heat, a higher maximum temperature, and greater subsequent strain due to cooling. Thus, the gain through greater strain capacity of the richer concrete with smaller aggregate may be more than offset by the greater strain that must be withstood, if the size is reduced too much.

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As stated above, the two factors governing the tensile strain which a concrete can withstand are the tensile strength and the modulus of elasticity. Many tests on very lean concretes, such as are used for the interior of large dams, have shown that tensile failure occurs without much “plastic” strain when loading is applied rapidly. For such concrete, the tensile strain which the concrete can withstand is approximately equal to the tensile strength divided by the modulus of elasticity of the concrete. For many purposes, then, it is sufficiently accurate to assume that the tensile strain capacity is inversely proportional to the modulus of elasticity of the concrete. It follows that the modulus of elasticity of the aggregate is important because of its large effect on the deformability of the concrete. Tensile strength is also important, and for this reason, crushed aggregates are apt to be superior to natural aggregates for crack prevention. Strain capacity can be measured directly on cylindrical specimens loaded in tension, or it can be determined on concrete beams located at the third points.7.1 A high creep rate of concrete is helpful in preventing cracking when the tensile strain is applied gradually. Since the tensile strength of concrete is nearly independent of prior loading, creep tends to increase the strain capacity. In the case of Dworshak Dam, for example, the strain to failure was almost three times as great for strain applied over 2 months as for quickly applied strain.7.1 The creep of concrete under sustained stress is affected by the stiffness of the aggregate. When the modulus is high, the creep is low and vice versa. The importance of aggregate rigidity on creep of concrete may be illustrated by two examples. First, assume that the aggregate and the cement paste have the same modulus of elasticity. When compressive stress is applied, the stress and the corresponding strain will be the same in the aggregate as in the cement paste. The aggregate does not creep under moderate stress but the paste does, and the paste which is between aggregate particles relaxes and loses stress. The lost stress must be shifted to the aggregate to maintain equilibrium. This imposes an elastic strain on the aggregate which accounts for a large part of the creep of the concrete. The amount of this elastic strain is directly related to the modulus of elasticity of the aggregate; the more rigid the aggregate, the lower the creep. Next, assume that the aggregate has a much higher modulus than the cement paste. When compressive stress is applied, the average stress in the aggregate will be higher than that in the cement paste and the paste will creep less than it did when the moduli were equal. The elastic strain in the aggregate due to the creep of the paste will then be less than it was when the moduli were equal. Thus, an increase in the rigidity

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ACI COMMITTEE REPORT

of the aggregate acts in two ways to reduce the creep of the concrete. 7.3 - Determination of temperatures and tensile strains Tensile strain in mass concrete results mainly from the restraint of thermal contraction, and to a lesser degree from autogenous shrinkage. Drying shrinkage is important only because it may cause shallow cracks to occur at surfaces. Thus, temperature change is the main contributor to tensile strain in mass concrete. The prediction of probable strain requires the prediction of the temperature to be expected. This prediction can be made quite reliably if the adiabatic temperature curve for the concrete is known, as well as the thermal diffusivity, boundary temperatures and dimensions. The finite element method can be used for the prediction of temperature distribution.7.3a 7.4 The main problem is that of choosing the correct boundary temperatures, which often depend upon the ambient temperatures. It is often satisfactory to use air temperatures found in weather reports as the surface temperatures to be used in the computations. For information on other methods of predicting temperatures in mass concrete, see the report ACI 207.lR. After the predicted temperature history is known, the determination of probable tensile strain is the next step. This can be accomplished using finite element computer programs.7.5a 7.6 Even with the finite element method, a thorough analysis is laborious because of the time-dependent variables. The analysis must include many steps of time to properly account for the creep (or relaxation) and the different and changing properties of every lift of concrete. On the other hand, strains near a boundary due to brief thermal shocks can be computed quite readily because in such cases the concrete can be assumed to be fully restrained. In this case, the strain is simply the temperature drop multiplied by the coefficient of expansion. This is important, because in many cases, the control of boundary strain is sufficient to prevent cracking. Internal strains usually develop slowly enough to be tolerable, even if large. Descriptions of test methods suitable for measuring the physical properties necessary for the prediction of temperatures and strains are given in Section 7.5. 7.4 - Control of cracking Given the probable temperatures and strains, the designer must determine what measures are most practicable to provide ample safety against cracking. The preventative measures will vary from nothing where weather and materials are favorable, to very expensive measures, where conditions are unfavorable. Some of the conditions which facilitate crack prevention are:

1. Concrete with large tensile strain capacity. 2. Small daily and seasonal temperature variations. 3. Low cement content (permitted by low design stresses). 4. Cement of low heat generation. 5. Short blocks. 6. Slow rate of construction when no cooling is used. 7. Low degree of restraint, as with yielding foundation, or in portions of structure well removed from restraining foundation. 8. High yearly average temperature. 9. Absence of stress raisers, such as galleries. 10. Low casting temperature. This list suggests many of the measures which can be taken to prevent cracking. First, an attempt should be made to produce a concrete with large tensile strain capacity. This may mean limiting the maximum aggregate size to a value somewhat below that which might be the most economical otherwise. Where several sources of aggregate are available economically, preference should be given to that which yields best crack resistance; usually this will be a crushed material of low thermal expansion and low modulus of elasticity. The heat producing characteristics of cement play an important role in the amount of temperature rise. ASTM Type II (moderate heat) cement should be used for mass concrete construction (Note: Type IV, low heat cement is, also, recommended, but is not readily available). Pozzolans can be used to replace a portion of the cement to reduce the peak temperature due to the heat of hydration (207.2R). In some cases, up to 35 percent or more of the cement can be replaced by an equal volume of a suitable pozzolan and still produce the same strength at 90 days or 1 year. Some of the more common pozzolans used in mass concrete include calcined clays, diatomaceous earth, volcanic tuffs and pumicites and fly ash. The actual type of pozzolan to be used and its appropriate replacement percentage are normally determined by test, cost, and availability. The lowest practical cement content permitted by the strength and durability requirements should be used to reduce the heat of hydration and the consequent thermal stresses and strains. More than the necessary amount of cement is a detriment rather than an advantage. In general, a reduction in the water content of concrete permits a corresponding reduction in the cement content. The concrete with less water and cement is superior in two important ways: it undergoes less temperature change and less drying shrinkage. Minimum water content can be achieved by such measures as specifying powerful vibrators which permit low slump, by using a water-reducing

CONTROL OF CRACKING

agent, and by placing the concrete at a low temperature. Precooling the concrete during its production and post-cooling it with embedded pipe systems after it is placed are especially effective measures. Details on pipe cooling are given in Section 7.6. One measure which offers promise is that of placing crack resistant concrete at boundaries (sides and top of lifts). Even though the more crack resistant concrete may be too costly to be used throughout the structure, it can be used to this limited extent without serious effect on economy. But thin layers of concrete next to the forms cannot be placed easily with present-day construction methods, which make use of very large buckets. Therefore, it appears more promising to use precast concrete panels for forms and to leave these panels as a permanent part of the structure. These panels should be of good quality for durability, and preferably lightweight so as to provide good thermal insulation. Since most cracks originate at boundaries, this partial measure may make the whole structure crack free. More information on the use of precast panels for protection of mass concrete can be found in ACI 347.1R. Thermal insulation on exposed surfaces during cold weather can protect concrete from cracking, if enough insulation is used and it is left in place long enough. If the insulation is sufficient to allow slow cooling, the tensile strain need never exceed the dangerpoint. The concrete can relax as rapidly as the tensile stress tends to develop, until finally, stable temperatures are reached. However, if the concrete has a very slow relaxation rate (or creep rate) the amount of insulation and the long protection time required may make this measure impractical. In extreme environments, where large amounts of insulation will be required during severely cold months, it may be necessary to remove the insulation in stages as the warmer months approach, Temperatures within the concrete just below the insulation should be allowed to slowly approach the environmental temperature. This is to prevent the occurrence of thermal shock which could induce cracking at the surface with possible, subsequent, deeper propagation into the mass. Precautions must be taken against using too much insulation or leaving it in place too long, which could result in stopping the desired cooling of the interior mass, and, in some cases, cause the interior temperature to begin to increase again. Insulation, as currently used for concrete, can be obtained in a variety of forms and materials having practical installed conductances ranging from 3.6 to 0.5 kg cal/m’/hr/C (0.75 to 0.10 BTU/hr/sq ft/F). It can be obtained in semirigid board type panels, rollon flexible rubber type material, and foamed sprayon material which becomes semirigid in place. The semirigid panels are usually installed on the inside

224R-29

face of the forms. Temporary anchors embedded in the newly placed lift of concrete retain the insulation on the concrete surface when the forms are lifted. The insulation is easily removed from the surface when desired. Roll-on insulation is particularly applicable for use on horizontal lift joints. It is easy to install and remove and can be reused many times. Spray-on insulation can be used on either horizontal or vertical surfaces. This type of insulation is particularly useful for increasing the thickness and effectiveness of insulation already in place and for insulating forms. Experience has shown that insulation which permits transmission of light rays should not be used because a temperature rise occurs between the insulation and the concrete when the insulation is subjected to direct sunlight. Spray-on insulation of timed longevity for frost protection of agricultural plants and trees, also, appears to have potential for the insulation of concrete lift joints during the active construction season. This insulation can be formulated to disintegrate at a given time after application. Thus, it can be timed to remain effective on the lift joints for approximately the period of time between successive placements and be easily removed by a final washing prior to placement of the new lift. Precast panels made of low conductance lightweight concrete or regular weight concrete cast with laminated or sandwich layers of low conductance cellular concrete also are acceptable as a means of insulating the interior concrete. The panels would then serve as both forms and face concrete.

7.5 - Testing methods and typical data 7.5.1 Adiabatic temperature rise - The temperature rise which would occur if there were no heat loss is defined as adiabatic temperature rise. The reader is referred to ACI 207.1R for methods of test. That report gives data on adiabatic temperature rise of concretes having a single cement content but having different types of portland cement. Fig. 7.1 gives typical adiabatic curves for Type II cement and various quantities of cement and pozzolan. Curves A and B in Fig. 7.1 represent data from mixes containing equal volumes of cementitious materials (cement plus pozzolan) thereby showing the effect of pozzolan replacement of cement on temperature reduction. 7.5.2 Thermal properties of concrete - Thermal diffusivity and thermal expansion are important in the control of cracking due to temperature change, and their determination is detailed in References ACI 207.1R and 7.8 through 7.10. The approximate range of thermal properties is shown in Table 7.1. 7.5.3 Creep of concrete - Creep may be defined as the continued deformation of concrete under sustained stress. A standard test for creep of concrete in compression is detailed in ASTM C 512-76.’ l5 Creep of concrete in tension is difficult to measure;

ACI COMMITTEE REPORT

224R-30

50 -

A

40

c _,_ 20 1

0 D I

.

LEGEND Curve A - Portland Cement 306Ib/cu yd(l8l kg/~);Pozzolan-None l

Curve B - Portland Cement 214 Ib/cuyd (127 kg/m3);l%zzoIon-74lb/cuyd(44kg/m) Curve C - Portland Cement . I81 Ib/cu yd (107kg/m3);Pozzolan-63Ib/cuyd (37 kg/m) Curve D - Portland Cement 148 lb/w yd (88 kg/d); PDZroh-50 Ibhu yd(30 kg/m) l

IO

-5

l

Type II Cement -0

0 0

4

I2

8

I6

20

24

28

Age , Days Fig. 7.1 - Typical adiabatic temperature curves for mass concrete (Reference 7.7)

TABLE 7.1 - Illustrative range of thermal and elastic properties of mass concrete Thermal properties

__-___-____~_-__~_--__Coefficient of linear expansion, millionths

Per O C _ --__ -_

Per O F

Diffusivity Conductivity _____--_- -__-----

__

_----__ ft x hr x O F

m x hr x O C-_- -

1 day ---___-_--- -_- _- --_----_, psi kg/cm’ kg/cm2 -3

__~---_.----~-

-3

-6

46.4

1

2.00

141

-6

2.56

kg/cm’ psi -3 -6 x10 -3 x 10 x 10 ~-_.__--_.__---

kg/cm’

psi

x 10 x 10 x 10 x 10 x 10 -_----_L--_----_.~-___--_-- - - - - - _,__---_ 0.66

0.22 -

Elastic~_~_____~---_~-----_ properties Static modulus of elasticity (E) for age _of__--__~-~-_--_ test indicated _----____-----p , 28 days 90 days 3 days 7 days __YPW ___---

- ____- -_- -_------

-6

-~-~

0.040 to 0.067 I__~--~-- ____------- - l----

4 7.2 to to 8 14.5 -_. _~____ _ _. _ ~_----_ ----- --

psi

-

ft’ hr

Specific heat ppp BTU/lb O F or Cal/g O C

180

4.00

281

psi, x 10 - 6 5.00

kg/cm’ x 10 -3 352

Poisson’s Ratio 0.15 to 0.25

224R-31

CONTROL OF CRACKING

thus, creep as measured in compression is assumed

to apply to tension as well. Such an assumption can be considered as reasonable when the stress is low. When the stress exceeds about 60 percent of the ultimate and microcracking occurs, not only does the instantaneous deformation increase, but the rate of creep increases, also. However, since the measured strain in a beam which is gradually loaded from the age of 1 month, to failure-at about 3 months, is only about 10 percent more than that computed using creep data as obtained from similar concrete in compression, it appears permissible to apply compression creep data to concrete stressed in tension in cases where approximate results will suffice. Creep of concrete is measured on carefully sealed specimens stored at a constant temperature and loaded to a constant stress. The measurement is usually made by means of embedded strain meters, although any reliable method of measuring strain can be employed. Butyl rubber is satisfactory for sealing the specimens, but neoprene should be avoided because it allows some moisture to escape. Specimens should be loaded at the same ages as specified for the modulus of elasticity tests, but loading at the early age of 1 day is not always practical. Again, the specimens should be large enough to permit concrete very nearly like that to be used in the structure. Cylinders of 9 x 18 in. (28 x 56 cm) size and with 3 in. (76 mm) maximum sized aggregate or 6 x 16 in (15 x 40 cm) cylinders with 11/2 in. (38 mm) maximum aggregate are frequently used. The symposium on creep of concrete,7.1’ gives useful coefficients for converting creep of smaller aggregate concrete to creep for mass concrete. Fig. 7.2 shows typical creep data obtained from laboratory investigations.7.1” Table 7.2 illustrates important computations that can be made using the data in the Fig. 7.2. Shown in Table 7.2 are values for sustained modulus of elasticity E, which in turn are used to develop tensile stress coefficients per degree temperature drop for the condition of full restraint. For example, concrete 2 days of age loaded at age 1 day would have a sustained modulus of elasticity (E,) of l/1.5 = 0.66 psi x lo6 (46.4 kg/cm2 x lo31 (see Fig. 7.2 and Table 7.2A), and if fully restrained would be stressed 0.66 x 5.5 psi per F = 3.6 psi/F (0.46 kg/cm2/C) for each degree drop in temperature (see Table 7.2B).

7.5.4 Modulus of elasticity - This subject is treated in detail in ACI 304. Table 7.1 shows values of the modulus of elasticity of a particular concrete after

various ages of curing.

7.5.5 Autogenous volume change - Autogenous

volume change7.7, 7.13 is the expansion or contraction of the concrete due to causes other than changes in temperature, moisture or stress. Thus, it is a selfinduced expansion or contraction. Expansion can be helpful in preventing cracks, but a contraction increases in tendency to crack. Autogenous volume

2.8

1

T o t a l Strain E=l481+00547~ (I+11 ---Icq c =0.521+0 0 7 0 0 LOG. (1+1) -3Oayr E =0384+0.0579 L O G . (1+1) -7Dg E =0.231+0.0500 Log (1+1) b -28asDays E =0.209+0.0294 LOG. (1+1 ) -9OCOyS

2.4 2.0

30 1 Day

1’

“E Y 2 20 ;

1.2

0

I”

0.6 0.5 0.4

Time,(t+l) Days Specific Creep Only

Fig. 7.2 - Typical concrete creep curves for mass concrete. change is usually measured by strain meters embedded in concrete cylinders which are carefully sealed (to insure that there is no loss in moisture) and kept at constant temperature. Measurements are begun as soon as the specimens are hardened and sealed, and continued periodically for months. 7.5.6 Tensile strain capacity-- The tensile strain ca-

pacity tests are generally performed on unreinforced concrete beams under third-point flexural loading. Relatively large beams ranging from 12 x 12 in. (30 x 30 cm) to 24 x 24 in. (60 x 60 cm) in cross section and 64 to 130 in. (160 to 325 cm) long are generally used.7.2 Strain capacity is determined from these tests under rapid and slow loading to simulate both rapid and slow temperature changes in the concrete. The loading rates are generally 40 psi (0.28 MPa) fiber stress per minute and 25 psi (0.17 MPa) fiber stress per week for rapid and slow loading tests, respectively. The strain for rapid loading can be measured using either surface or embedded strain gages or meters.7.1, 7.7 For long-term tests, embedded meters are best. The strain can also be determined from deflection measurements. The concrete test beam used for determining the strain capacity should be protected during the test to prevent loss of moisture by wrapping it with an impermeable material. Testing should be conducted at a constant temperature for maximum accuracy in measurement. Detailed test procedures can be found in References 7.1 and 7.14 Fig. 7.3 shows the unit strain values

ACI COMMITTEE REPORT

224R-32

B E A MSTRESS

OF OUTER

and using small ice particles as a replacement of part of the mixing water. Post-cooling of concrete is accomplished by circulating cool liquids (usually water) through pipes embedded in the concrete. Studies made during the design stage will establish such items as lift height, pipe spacing, water temperature and rate of flow, acceptable rate of temperature drop (for both rapid and slow drops), and approximate duration of cooling. In general, the duration of cooling and the heat removed by the pipe cooling should be sufficient to insure that a secondary internal temperature rise in the mass does not exceed the primary rise. It is, however, important that steep cooling gradients, which can result in cracking the mass, be avoided. This is particularly true in smaller masses where circulation of cooling water should be stopped when the maximum temperature has been reached and just begins to drop. A vulnerable location in pipe cooling systems is centered at the cooling coils where sharp gradients and cracking can be induced if termination of cooling water circulation is not timely.

I

FIBERS

Fig. 7.3 - Unit tensile strain versus beam stress (References 7.1 and 7.7). versus beam stress at outer fibers for a typical laboratory investigation.7.1* 7.11

In the preliminary studies of temperature and construction control plans for mass concrete projects, approximate methods for estimating tensile strain capacity under rapid and slow loadings given in References 7.5 and 7.20 may be used.

Resistance thermometers should be used in sufficient numbers to permit adequate monitoring and control of the internal concrete temperatures. 7.6 - Artificial cooling by embedded pipe systems Construction drawings should show basic pipe layThe overall program for cooling concrete, includout and spacing including minimum spacing, and the ing important field control criteria, should be deterlayout at dam faces, transverse construction joints, mined during the design stage. Precooling concrete interior openings and in sloping, partial, and isolated prior to placement is accomplished by a variety of concrete lifts. A pipe layout for a typical concrete methods, including cooling all ingredients of the mix lift is shown in Fig. 7.4.

TABLE 7.2 - Illustration of computation of sustained modulus of elasticity (Es) and stress coefficients A. Sustained modulus Es at age of concrete at time of loading, days 1 day Time after loading days

3 days 2 kg/cm - 10-3

psi -6 x 10 0.68 0.66 0.64 0.63

psi -6 x 10

47.6 46.2 44.8 44.1

(1) Sustained modulus of elasticity IE Fig. 7.2

/

7 days 2 kg /cm x 10-3

1.92 1.76 1.62 1.35

134 123 113 95

psi x 10-6b

t

I t

2.61 2.46 2.15 1.98

I !1

x 1 0 -3

t I

183 172 151 139

I

x 10 -6

I t--

I ’

4.33 3.76 3.34 2.99

I-

x 10 -3 ------

303 263 234 210 ._~

values are based on data given in

E,z

______

______

~._______-___.___

unit elastic strain/psi + Vz specific creep for time of loading R. Tensile stress coefficients for condition of full restraint and decreasing temperature Age of concrete at time of loading 1 day

0 1 3 7 (2) Coefficient

3 days

- ---- -- lb/in.‘/F

kg/cm’/C

3.7 3.6 3.5 3.5

0.47 0.46 0.45 0.44

I

m TGkg,em’)C ’

11.0 9.7 8.9 7.4

’ , j

of lineal thermal expansion of concrete assumed to be 5.5 mil lionths/F (9.9 millionths/C,

I

7 days

1.33 1.22 1.12 0.94

I

Ib/in.‘/F

1

kglcmJ/C

i

14 14 12 11

; i I

1.81 1.70 1.50 1.38

’

/

lb/in.,,:” i ~~g!~~/~ 24 21 18 16

I I !

3.00 2.60 2.31 2.08

__

CONTROL OF CRACKING

8

s

224R-33

Multiply

3

4@2'-O"

w

W

PLAN ELEV. I I35 COIL

LIN FEET

1" = 30'-0"

Inches .

Feet

By

To Obtain

0 0254 0 3048

Meters Meters

+_FlDw

~_

8

W

47

r-- -&--- Elev. 1140 K-1 Elev 1135

Detail " B "

Section

A-A

Fig. 7.4 - Typical cooling coil layout (Reference 7.11).

Fig. 7.5 - Schematic of embedded pipe cooling embedment system in mass concrete. In most areas of the dam, a uniform spacing can be maintained for the cooling pipe, but isolated areas always exist in all dams which tend to result in a concentration of pipes. These concentrations tend to occur at the downstream face of the dam where inlets and outlets to cooling pipes are located, adjacent to openings in the dam, and at isolated and sloping lifts of concrete. Proper planning will alleviate many of the undesirable conditions that can result from these concentrations. For example, it must be determined to what extent the cost saving procedure of concentrating cooling pipe inlets and outlets near contraction joints can be permitted at the face of the dam. Also, it must be decided if cooling pipes to isolated areas in the foundation and at openings such as

galleries can extend from the downstream face of the dam or if a vertical riser must be used. For ease of installation, the pipe used for postcooling should be thin wall tubing. Aluminum tubing is lightweight and easy to handle. However, breakdown from corrosion inducing elements of the concrete is a potential problem for aluminum pipe if cooling activities must be carried on over a period of several months. In this case, steel tubing is preferred. Compression type couplings are used because thin wall tubing cannot be threaded satisfactorily. Surface connections to the cooling pipe should be removable to a depth of 4 to 6 in. (102 to 152 mm) so

224R-34

ACI COMMITTEE REPORT

that holes can be reamed and dry packed when connections are removed. Forms should be designed and constructed so that shutdown of cooling activities is not necessary when forms are raised. Wire tiedowns embedded at the top of the concrete lift at about 10 ft (3 m) spacing satisfactorily secure the pipe during concrete placing. Coils must be pressure tested for leaks at the maximum pressure they will receive from the cooling system prior to placing concrete. Pressure must also be maintained during concrete placement to prevent crushing and permit early detection of damage, should it occur. After cooling is completed and the pipe is no longer needed, it should be thoroughly flushed with water at a high enough pressure to remove foreign matter and grouted full with a grout mixture compensated for plastic shrinkage or settlement. The grout should remain under pressure until final set is attained. Fig. 7.5 shows the schematic layout of a typical pipe cooling system. Sight flow indicators should be installed at the end of each embedded pipe coil to permit ready observance of cooling water flow. In addition to regular observance of flows, water temperatures and pressures and concrete temperatures should be observed and recorded at least once daily while the lift is being cooled. The refrigeration plant for cooling water may be centrally located, or several smaller complete portable plants may be used to permit moving the refrigeration system as the dam progresses upward. Sufficient standby components, equal in capacity to the largest individual refrigeration units should be provided.

7.7 - Summary - Basic considerations for construction controls and specifications The construction controls and specifications for mass concrete must be such that the structures will be safe, economical, durable, and pleasing in appearance. Each of these requirements in turn affects the crack resistance. Safety will be assured if the concrete has sufficient strength and continuity (absence of cracks). Economy will depend upon such features as the best choice of aggregates, adequate but not excessive temperature control, low cement content, etc. Durability will depend upon the quality of the concrete, exposure conditions, and freedom from chemical reactions of a deteriorating nature. Pleasing appearance will come from good workmanship, freedom from cracks and stains, absence of leakage and leaching, etc. The importance of a comprehensive materials test program to establish nec-

essary control prior to preparation of construction controls and specifications cannot be overemphasized. 7.7.1 Safety 7.7.1.1 Safety against crushing-concrete strength. A strength should be specified which will provide an adequate factor of safety against crushing of the concrete. The “nominal” factor of safety is merely the compressive strength divided by the maximum stress to be expected in the structure. However, neither the strength nor the maximum stress can be accurately determined. The strength is usually derived from tests on cylindrical specimens which are not completely representative of the structure. The maximum stress is usually taken as the design stress which is based upon assumed concrete properties. For such reasons, it is considered good practice to use a safety factor as high as three or four, meaning that the strength should be three or four times the expected maximum stress. The 90-day strength is often used and is derived from tests of job cylinders. Since the cylinders are made from wet screened concrete, the measured strength is corrected to a massconcrete equivalent by applying a reduction factor of about 0.80 for typical conditions. For specific data on appropriate reduction factors, the reader should refer to the U.S. Burau of Reclamation, Concrete Manual, 8th Edition. 7. The “factor of safety,” as defined above, is subject to a number of additional factors which, more or less, balance one another. Since the average strength of the job cylinders is used, half of the tests will be weaker. The strength at 90 days is not the ultimate strength. There can be a large gain after 90 days depending upon the composition of the cement. However, even a “factor of safety” of three is far more than enough to cover any likely differences between plus and minus corrections. For interior concrete, the lowest practical strength should be specified so as to reduce the cement content. This, in turn, will reduce the heat of hydration and the consequent thermal stresses, thus increasing the crack resistance of the concrete. More than the necessary amount of cement is detrimental rather than advantageous. 7.7.1.2 Safety against sliding. Sound, uncracked concrete provides a very large factor of safety against sliding. However, hardened horizontal lift joints may impair the safety. Therefore, the specifications should require care in the preparation of lift surfaces and in the placement and compaction of concrete thereon. Also, the lift surfaces should slope slightly upward toward the downstream edge (in the case of a dam) such that the downstream edge is higher than the upstream edge. It is not necessary to use a mortar layer on lift surfaces prior to the placement of the next lift. 7 . 7 2 Economy - Many factors which affect the

CONTROL OF CRACKING

economy also affect crack resistance. For example, the least expensive aggregate may have bad thermal properties and thus require expensive temperature control to prevent cracking. The aggregate which makes concrete of highest tensile-strain capacity may increase the water requirement and, therefore, also the cement requirement, thus offsetting the benefits of high strain capacity. Some of the factors which affect economy are discussed below. 7.7.2.1 Selection of aggregate. Aggregate should be chosen that makes good concrete with the lowest overall cost. If natural aggregate near the site has unfavorable properties for crack prevention, crushing to increase crack resistance may be an economical expedient because of the consequent saving in temperature control. When crushing is either advantageous or necessary, rock which has the most favorable properties should be chosen. The rock should have a low coefficient of thermal expansion, a low modulus of elasticity, and it should produce particles of good shape and surface texture. All of these factors are important in increasing the resistance of the concrete to cracking. 7.7.2.2 Aggregate size. The largest maximum size of aggregate, up to approximately 6 in. (150 mm) in diameter, should be specified as can be placed properly in the structure, except for concrete which must resist high-velocity water flow. Larger aggregate permits the use of less water and cement per cubic yard, resulting in savings in both the amount of cement and the amount of temperature control necessary for required crack resistance. 7.7.2.3 Water content. A reduction in the water content of concrete permits a corresponding reduction in the cement content. The concrete with less water and cement is superior in many ways: it undergoes less temperature change, less drying shrinkage, and as a result is more durable and crack resistant. As indicated in Section 7.4, minimum water content can be achieved by specifying adequately powerful vibrators which permit the use of low slump concrete, by using a water-reducing agent when appropriate, and by producing and placing the concrete at low temperature. 7.7.2.4 Use of pozzolan In most locations, good pozzolans such as fly ash are available, and they can be used to replace a portion of the cement. This can result in a considerable saving in cost, and possibly more important, it can reduce the heat generation and improve the resistance against cracking. Another advantage of using pozzolan is that when used in adequate amounts, it reduces the expansion due to reactive aggregates when such are encountered. The appropriate amount of pozzolan for a reactive aggregate should be based upon test data obtained with the pozzolan and cement being used. 7.7.3 Durability - Durability of concrete is closely related to the exposure conditions. In tropical cli-

224R-35

mates, for example, there may be no deteriorating influences acting on the concrete except that which is subject to high-velocity water flow. For the main structure in such a case, any concrete which has the required strength can be expected to last indefinitely, and the cement content should be kept low to minimize heat generation and resultant potential cracking. Where the climate is severe, such that there is much freezing and thawing in winter, the water-cement ratio of surface concrete should be kept lower than that necessary for strength alone. Air entrainment should be mandatory. For any concrete which might be subject to both alternations of freezing and water pressure, the water-cement ratio should be less than 0.40 by weight. The effect of the rich boundary concrete on thermally induced cracking will be minimized by keeping the thickness of the boundary layer to a minimum, probably 2 ft (0.6 m) or less. 7.7.4 Control of cracking - A detailed discussion of the control of cracking in massive structures has been presented in this chapter. With proper planning and execution, the procedures presented will serve as useful tools in developing a crack control program for mass concrete structures. References

7.1. Houk, Ivan E., Jr.; Paxton, James A.; and Hough-

ton, Donald L., “Prediction of Thermal Stress and Strain Capacity of Concrete by Tests on Small Beams,” ACI J O U R N A L , Proceedings V. 67, No. 3, Mar. 1970, p p . 253-261. 7.2. Houghton, Donald L., “Determining Tensile Strain Capacity of Mass Concrete,” ACI JO U R N A L, Proceedings V . 73, No. 12, Dec. 1976, pp. 691-700. 7.3. Wilson, E. L., “The Determination of Temperatures within Mass Concrete Structures,” Report No. 68-17, Structural Engineering Laboratory, University of California, Berkeley, Dec. 1968. 7.4. Polivka, R. M., and Wilson, E. L., “Finite Element Analysis of Nonlinear Heat Transfer Problems,” Report No. UC SESM 76-2, Department of Civil Engineering, University of California, Berkeley, June 1976. 7.5. Sandhu, R. S.; Wilson, E. L.; and Raphael, J. M., “Two-Dimensional Stress Analysis with Incremental Construction and Creep,” Report No. 67-34, Structural Engineering Laboratory, University of California, Berkeley, Dec. 1967. 7.6 Liu, Tony C.; Campbell, R. L.; and Bombich, A. A., “Verification of Temperature and Thermal Stress Analysis Computer Programs for Mass Concrete Structures,” Miscellaneous Paper No. SL-79-7, U.S. Army Engineer Waterways Experiment Station, Vicksburg, Apr. 1979. 7.7. Houghton, Donald L., “Concrete Volume Change for Dworshak Dam,” Proceedings, ASCE, V. 95, P02, Oct. 1969, pp. 153-166. 7.8. “Method of Test for Thermal Diffusivity of Mass Concrete,” (CRD-C 37-73), Handbook for Concrete and Cement, U.S. Army Corps of Engineers, Vicksburg, Dec. 1973, 3 pp.

224R-36

ACI COMMITTEE REPORT

7.9. “Method of Test for Coefficient of Linear Thermal Expansion of Concrete,” (CRD-C 39-55), Handbook for Concrete and Cement, U.S. Army Corps of Engineers, Vicksburg, 1939, 2 pp. 7.10. “Method of Test for Coefficient of Linear Thermal Expansion of Coarse Aggregate, Strain Gage Method,” (CRD-C 125-63), Handbook for Concrete and Cement, U.S. Army Corps of Engineers, Vicksburg, June 1963, 5 pp. 7.11. Symposium on Creep of Concrete, SP-9, American Concrete Institute, Detroit, 1964, 160 pp. 7.12. McCoy, E. E., Jr.; Thorton, H. T.; and Allgood, J. K., “Concrete Laboratory Studies, Dworshak (Bruce’s Eddy) Dam, North Fork Clearwater River Near Orofino, Idaho: Creek Tests,” Miscellaneous Paper No. 6-613, Report 2, U.S. Army Engineer Waterways Experiment Station, Vicksburg, Dec. 1964. 7.13. Houk, Ivan E., Jr.; Borge, Orville E.; and Houghton, Donald, “Studies of Autogenous Volume Change in Concrete for Dworshak Dam,” ACI JOURNAL , Proceedings V. 66, No. 7, July 1969, pp. 560-568. 7.14. McDonald, J. E.; Bombich, A. A.; and Sullivan, B. R., “Ultimate Strain Capacity and Temperature Rise Studies, Trumbull Pond Dam,” Miscellaneous Paper C-72-20, U.S. Army Engineer Waterways Experiment Station, Vicksburg, Aug. 1972. 7.15. Liu, Tony C., and McDonald, James E., “Prediction of Tensile Strain Capacity of Mass Concrete,” ACI JOURNAL , Proceedings V. 75, No. 5, May 1978, pp. 192-197. 7.16. Concrete Manual, 8th Edition, U.S. Bureau of Reclamation, Denver, 1975, 627 pp.

Chapter 8 - Control of cracking by correct construction practices* 8.1 - Introduction

Construction practices, as used in this chapter, include designs, specifications, materials, and mix considerations, as well as on-the-job construction performance. Before discussing control of construction practices which affect cracking, it is worthwhile to mention the basic cause of cracking. It is restraint. If all parts of the concrete in a concrete structure are free to move as concrete expands or contracts, particularly the latter, there will be no cracking due to volume change. Obviously, however, all parts of concrete structures are not free, and inherently, cannot be free to respond to the same degree to volume changes. Consequently, differential strains develop and tensile stresses are induced. When these differential responses exceed the capability of the concrete to withstand them at that time, cracking occurs. This points to the importance of protecting new concrete for as long as practicable from the loss of moisture or a drop in temperature. These considerations may result in stresses capable of causing cracks at an *Principal author: Lewis H. Tuthill.

early age but which might be sustained at greater maturity. Preferably, concrete should have a high tensile strain-to-failure capacity. This is influenced greatly by the aggregate, and a low modulus of elasticity in tension is desirable. 8.2 - Restraint Restraint exists in many circumstances under which the structure and its concrete elements must perform. Typical examples will illustrate how restraint will cause cracking, if the concrete is not strong enough to withstand the tensile stresses developed. 8.21 - A wall or parapet anchored along its base to the foundation or to lower structural elements less subject or responsive to volume change, will be restrained from shrinking when its upper portions shorten due to drying or cooling. Cracking is usually inevitable unless contraction joints (or at least grooves of a depth not less than 10% of the wall thickness on both sides, in which the cracks will occur and be hidden) are provided at intervals ranging from one (for high walls) to three (for low walls) times the height of the wall. 8.2.2 - Exterior and interior concrete, particularly in heavier sections, will change temperature or moisture content at different rates and to different degrees. When this happens, the interior concrete restrains the exterior concrete from shrinking, and tensile strains develop which may cause the exterior to crack. This occurs when the surface cools, while the interior is still warm from the heat of hydration, or when the surface concrete dries faster than the interior concrete. As noted earlier, it is often feasible to protect the surface for a time at early ages so that such stress-inducing differentials cannot develop before the concrete is strong enough to withstand the strain without cracking. 8.2.3 - Acting similarly to the interior concrete in the foregoing example, temperature reinforcement can restrain the shrinkage of surface concrete, but more and narrower cracks may result. 8.2.4 - Restraint will occur at sharp changes in section, since the effect of temperature change or drying shrinkage will be different in the two sections. If feasible, a contraction joint can be used to relieve the restraint. 8.2.5 - Restraint of flat work results from anchorage of slab reinforcement in perimeter slabs or footings. When a slab is free to shrink from all sides toward its center, there is a minimum of cracking. Contraction joints and perimeter supports should be designed accordingly (see Section 3.5.3). 8.2.6 - Wall, slabs, and tunnel linings placed against the irregular surface of a rock excavation are restrained from moving when the surface expands or contracts in response to changes in temperature or

CONTROL OF CRACKING

moisture content. As discussed in Section 8.2.1, closely spaced contraction joints or deep grooves must be provided to prevent or hide the cracks which often disfigure such surfaces. In tunnel linings, the shrinkage in the first few weeks is primarily thermal, and the use of cold concrete (50 F or 10 C) has reduced cracking materially. By the time drying is significant, the concrete lining is much stronger and better able to resist shrinkage cracking. However, circumferential cracks in tunnel linings and other cast-in-place concrete conduits and pipe lines can be greatly reduced in number and width. As shown in the Bureau of Reclamation Concrete Manual,8.1 this can be done if a bulkhead is used to prevent air movement through the tunnel, and shallow ponds of water are placed in the invert as soon as possible after lining, and left until the tunnel goes into service. If the tunnel carries water, there will be no further drying shrinkage. If it does not, the concrete will have become much stronger in the humid environment and will be better able to resist shrinkage-induced tensile stresses. 8.2.7 - The typical examples presented above clearly indicate that many crack control procedures must be considered by the engineer during design. While proper construction performance can contribute a great deal (as will be discussed below), the contractor cannot be expected to utilize the best procedures, unless these procedures are included in the designs and specifications on which the bid price is based. 8.3 - Shrinkage The following sections discuss the major causes of shrinkage, which is a key contributor to the formation of cracks in concrete. 8.3.1 Effect of water content - The greater the water content of concrete, the more it will shrink on drying. Such a hypothesis is clearly indicated in Fig. 3.2, as well as in Reference 8.1. The use of the lowest practical slump is important. Of major importance is the selection of mix proportions that require the least amount of water per cubic yard for the desired concrete strength. This means avoiding oversanded mixes (the richer the concrete, the coarser the sand should be and the less there should be of it in the mix); using the largest maximum aggregate size practical; using aggregate with the most favorable shape and grading conducive to best workability; and using well-graded sand with a minimum of fines passing the l00-mesh and free of clay, such that its sand equivalent value is not less than 80 percent AASHTO T176. Contrary to common belief, increasing the cement content of concrete, per se, does not necessarily cause an increase in shrinkage. This is because the water requirement of concrete does not change much with a change in cement content. Drying shrinkage is proportional to water content (Fig. 3.2),

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not cement content. Moreover, the reduction of the amount of fine aggregate to compensate for the added cement, in accordance with correct principles of concrete proportioning, will offset any tendency to increase the water requirement. 8.3.2 Surface drying - Surface drying will ultimately occur except when the surface is submerged or backfilled. It will cause shrinkage strains of up to 600 millionths or more. The amount of shrinkage cracking depends on 1. how dry the surface concrete becomes, 2. how much mixing water was in the concrete, 3. the character and degree of restraint involved, and 4. the extensibility of the concrete. The extensibility represents how much the concrete can be strained (stretched), without exceeding its tensile strength and is the sum of creep plus elastic strain capacity. The latter is largely related to the composition of the aggregate and may vary widely. Typically, some concretes of highly quartzitic gravels have a low strain capacity and a high modulus of elasticity, while some concretes of granitic and gneissic aggregate have a high strain capacity and a low modulus of elasticity. Concretes having a low strain capacity are much more sensitive to shrinkage due to drying (and to drop in temperature) and will be subject to a greater amount of cracking. Accordingly, as mentioned in connection with tunnel linings and conduits, a prime objective of crack control procedures is to keep the concrete wet as long as feasible, so that it will have time to develop more strength to resist cracking forces. The importance of this will vary with the weather and the time of year. Cold concrete (below 50 F, 10 C) dries very slowly, provided the relative humidity is above 40 percent. At some depth, concrete loses moisture slowly, as shown in Fig. 3.5. Where surface drying may be rapid, more care must be devoted to uninterrupted curing to get good surface strength. Cracking stresses will be further reduced by creep, if the surface is prevented from drying quickly at the end of the curing period. To accomplish this, the wet curing cover can be allowed to remain several days without wetting after the specified curing period (preferably 7 to 10 days), until the cover and the concrete under it appear to be dry. If job conditions are likely to be such that these measures will be worthwhile, they should be required in the specifications for the work. 8.3.3 Plastic shrinkage - Plastic shrinkage cracks occur most commonly, and objectionably, in the surfaces of floors and slabs when the ambient job conditions are so arid that moisture is removed from the concrete surface faster than it is replaced by bleed water from below. These cracks occur prior to final finishing and commencement of the curing process. As the moisture is removed, the surface concrete contracts, resulting in tensile stresses in the essentially strengthless, stiffening plastic concrete, that cause short random cracks or openings in the sur-

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face. These cracks are usually rather wide at the surface but only a few inches in depth. The cracks generally range from a few inches to a few feet in length and are a few inches to two feet apart. Sometimes plastic shrinkage cracks appear early enough to be worked out in later floating or first trowelling operations. When this is successful, it is advisable to postpone these operations as long as possible to get their maximum benefit without the recurrence of cracking. In other cases, an earlier than normal floating may destroy the growing tension by reworking the surface mortar and prevent plastic cracking that would otherwise occur. At the first appearance of cracking while the concrete is still responsive, a vigorous effort should be made to close the cracks by tamping or beating with a float. If firmly closed, they will be monolithic and are unlikely to reappear. However, they may reappear if they are merely trowelled over. In any event, curing should be started at the earliest possible time. Conditions most likely to cause plastic shrinkage cracking are high temperatures and dry winds. Accordingly, specifications should stipulate that effective moisture control precautions should be taken to prevent a serious loss of surface moisture under such conditions. Principal among these precautions are the use of fog (not spray) nozzles to maintain a sheen of moisture on the surface between the finishing operations. Plastic sheeting can be rolled on and off before and after floating, preferably exposing only the area being worked on at that time. Least effective but helpful are certain sprayed monomolecular films which inhibit evaporation. Windbreaks are desirable, and as such, it is desirable to schedule flatwork after the walls are up (ACI 305R, ACI 302.1R). Other helpful practices that may augment the bleeding and counteract the excessive loss of surface moisture, are 1. using a well dampened sub-grade, 2. cooling the aggregates by dampening and shading them, and 3. using cold mixing water or chipped ice as mixing water to lower the temperature of the fresh concrete. 8.3.4 Surface cooling - Surface cooling will shrink the surface of average unrestrained concrete about 10 millionths for each deg C (5.5 millionths per deg F) the temperature goes down. This amounts to 9 mm in a 30 m length with a drop of 30 C (l/3 in. in 100 ft with a drop of 50 F). The amount of shrinkage is reduced by restraint and creep, but tensile stresses are induced. The earlier the age and the slower the rate at which cooling or drying occur, the lower the tensile stresses will be. This is due to the relaxing influence of creep, which imparts more extensibility to concrete at early ages.

In ordinary concrete work, the winter protection required for the development of adequate strength will prevent the most critical effects of cooling. The system of contraction joints and grooves previously discussed for control of shrinkage cracking will serve the same purpose against substantial later drops in surface temperature. In addition to Chapter 7 of this report, Chapters 4 and 5 of ACI 207.1R discuss temperature controls for mass concrete to minimize the early temperature differences between interior and exterior concrete. Primarily, these controls lower the interior temperature rise caused by the heat of hydration by using 1. no more cement than necessary, 2. pozzolans for a portion of the cement, 3. water reducing admixtures, 4. air-entrainment, 5. large aggregate, 6. low slump, and 7. last but by no means least, where at all practicable, chipped ice for mixing water to reduce the temperature of the fresh concrete as much as possible. See Fig. 3.4 and Fig. 3.1 of ACI 207.2R. At no time should forms be removed to expose warm surfaces to low temperatures. As mentioned in Section 8.3.2, the extensibility, or strain the concrete will withstand before tensile failure, is a function of the aggregate and should be evaluated, especially on larger projects. What applies to one will not necessarily apply to another. 8.4 - Settlement Settlement or subsidence cracks develop while concrete is in the plastic stage, after the initial vibration. They are not due to any of the causes discussed above, but are the natural result of heavy solids settling in a liquid medium. Settlement cracks occur opposite rigidly supported horizontal reinforcement, form bolts or other embedments. Sometimes concrete will tend to adhere to the forms. A check will appear at these locations, if the forms are hot at the top or are partially absorbent. Cracks often appear in horizontal construction joints and in bridge deck slabs over reinforcing or form bolts with only a few inches cover. The cracks in bridge decks can be reduced by increasing the concrete cover.8.2 Properly executed late revibration can be used to close settlement cracks and improve the quality and appearance of the concrete in the upper portion of such placements, even though settlement has taken place and slump has been lost. 8.5 - Construction A great deal can be done during construction to minimize cracking, or in many cases to eliminate it. But, as noted in Section 8.2.7, such actions must be required by the specifications and by the engineering forces which administer them. Such actions include the following: 8.5.1 Concrete aggregates - The aggregate should be one which makes concrete of high strain capacity, if reasonably available (see Section 7.2). Fine and

CONTROL OF CRACKING

coarse aggregates have to be clean and free of unnecessary fine material, particularly clays. The sand should have a sand equivalent value in excess of 80 percent, and this should be verified frequently (AASHTO T176). The sand should have sufficient time in storage for the moisture content to stabilize at a level of less than 7 percent on an oven-dry basis. 8.5.2 Expansive cement - Expansive cement can be used to delay shrinkage during the setting of concrete in restrained elements reinforced with the minimum shrinkage steel required by ACI 318. T h e principal property of these cements is that the expansion induced in the concrete while setting and hardening is designed to offset the normal drying shrinkage. With correct usage (particularly with early and ample water curing on which maximum expansion depends), the distance between joints can sometimes be tripled without increasing the level of shrinkage cracking. Details on the types and correct usage of shrinkage compensating cements are given in ACI 223-83. 8.5.3 “Non-shrink” grout, mortar, or concrete - Ordinarily, the solids in grout, mortar, and concrete mixtures will settle before hardening, and water will rise, some of it to the top surface. This settlement can be objectionable if a space is to be filled up tightly without leaving a void at the top, such as under machine bases. Measures taken to prevent such subsidence have produced what is known in the trade as “Non-shrink” grout, mortar, or concrete. Some of the materials merely prevent settlement; others in addition, provide a slight expansion as the mixture hardens. The most widely used materials contain unpolished aluminum powder. These should contain no stearates, palmitates, or fatty acids. In an alkaline solution, such as exists in portland cement mixtures, the aluminum reacts to form aluminum oxide and hydrogen. The hydrogen gas tends to expand the mixture and thus prevents subsidence and may even cause expansion. The amount of aluminum powder used varies widely with conditions, but is usually in the neighborhood of 0.005 to 0.01 percent by weight of the cement. It is not possible to specify an exact percentage because the amount to be used varies with such factors as temperature, alkali content of the cement, and the richness of the mix. Therefore, it is advisable to make trial mixes with various percentages of aluminum powder to find which percentage gives the desired (slight) expansion under the prevailing conditions. The amount of aluminum powder used is so small that it is advisable to dilute it by blending with 50 parts

of sand or fly ash. This diluted mixture will have enough bulk so that it can be easily measured and properly dispersed in the mix. Among the admixtures that merely prevent settlement, a number of different mechanisms are in operation. One commercial grout is so highly acceler-

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ated that it starts setting before settlement takes place. Another is composed of organic gelling compounds of soluble cellulose which increase in viscosity so that the solid particles remain in suspension. Still another contains a form of carbon with a very large surface area. In the dry form, it contains a large amount of adsorbed air, which is released gradually into the mix producing an expansion. Gas forming agents and air releasing agents produce the same net effect, although all grouts, mortars and concretes employing these agents have no expansive properties after hardening, and have a drying shrinkage at least equal to similar plain grouts, mortars and concretes not employing them. Grouts which expand (if unconfined) after hardening can function as nonshrink grouts, as opposed to grouts that expand only in the plastic state and later suffer drying shrinkage. Among the commercial admixtures, there is one containing a metallic aggregate which, in addition to opposing settlement during hardening, provides a modest expansion after hardening. This acts to hold the grout tightly up under base plates, etc., and also tends to offset the effect of drying shrinkage. Where feasible, the problem of settlement can be solved by the use of dry tamped mortar, instead of a fluid grout or mortar. Grout mixed in a colloid mill will not readily settle. It should be noted that prepackaged “Non-shrink” grouts, like any portland cement grouts and mortars, are subject to shrinkage if exposed to drying and may deteriorate and lose serviceability if exposed to an aggressive environment (weathering, salt spray, etc.). 8.5.4 Handling and batching - Should be done with all practical care to avoid contamination, overlap of sizes, segregation, and breakage, s o t h a t e x t r a amounts of fines are not needed in the mixes to account for variations in grading without a serious loss of workability. This is best done by finish screening and rinsing as a combination of coarse aggregate sizes goes to the batch plant bins. Every effort should be made to uniformly batch and mix the concrete so that there will be a minimum of troublesome variation in slump and workability. These, invariably, lead to demands for a greater margin of workability, with more sand and more water in the concrete. 8.5.5 Excessive workability - Whether it is achieved with unneeded higher slump, oversanding, small aggregate, or even higher air content (which may reduce strength), is always popular and in demand on the job. It must be discouraged if the best concrete for the work (having adequate workability with proper handling and vibration, and having minimum shrinkage factors) is to be obtained. 8.5.6 Cold concrete - Cold concrete, when com-

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bined with factors to reduce water and cement content to a practical minimum will reduce temperature differentials which cause cracking. Cold concrete is particularly useful for massive concretes. It requires less mixing water and thus reduces drying shrinkage. In warm weather it expedites the work by reducing slump loss, increasing pumpability, and by improving the response to vibration. It is obtained by substituting chipped ice for all or a part of the batched mixing water. In cold weather, concrete is naturally cold and every effort should be made to use it as cold as possible without inviting damage from freezing. It is pointless to expect to protect surfaces, edges, and corners by placing needlessly warm concrete in cold weather. These vulnerable parts must be protected with insulation or protective enclosures (ACI 306R). 8.5.7 Revibration - When done as late as the formed concrete will respond to the vibrator, will eliminate cracks and checks where something rigidly fixed in the placement prevents a part of the concrete from settling with the rest of it. Settlement cracks are most apparent in the upper part of wall and column placements where revibration can be readily used. Deep revibration corrects cracks caused by differential settlement around blockout and window forms, and where slabs and walls are placed monolithically. 8.5.8 Finishing - Flatwork finishing can make a great difference in the degree of freedom from all types of cracking (ACI 302.1R). Low-slump concrete should be used. More than a 3 in. (76 mm) slump is rarely necessary except perhaps in very hot weather in which both slump and moisture are lost quite rapidly. Finishing should not be done in the presence of surface water. Precautions (see Section 8.3.4) should be taken to prevent plastic shrinkage. Any required marking and grooving should be carefully cut to the f u l l depth specified. Curing should be prompt, of full duration, and the wet cover should be allowed to dry before it is removed. 8.5.9 Curing and protection - Newly placed concrete must be brought to a level of strength maturity and protected from low temperatures and drying conditions which would otherwise cause cracking. The curing and protection should not be discontinued abruptly. If the new concrete is given a few days to gradually dry or cool, creep will have an opportunity to reduce the possibility of cracking when the curing and protection are fully discontinued. 8.5.10 Miscellaneous - Some items normally covered in specifications (or certainly which should be covered where appropriate) require special attention during construction because of their potential effects on cracking. 1. Reinforcement and embedments must be properly positioned with the designated thickness of cover in order to prevent corrosion, expansion and cracking.

2. Concrete should not be placed against hot reinforcement or forms. 3. Formwork support should be strong enough to be free of early failures and distortion causing cracking. 4. Subgrade and other supports must not settle unevenly, to prevent cracks due to overstress in the structure. 5. Contact between aluminum and steel embedded in the concrete must be eliminated, particularly if use of calcium chloride is permitted. If it is used, calcium chloride must be limited to the absolute minimum (see Section 3.4.4). 6. Special care is needed in handling precast units to prevent overstress due to handling. 7. Unvented salamanders in cold weather (ACI 306R) or gasoline operated equipment must be avoided where adequate ventilation is not furnished, because of the danger of carbonation shrinkage surface cracking. 8. Control joints, discussed in Sections 3.5.3 and 8.2.6, must not be omitted and grooves must be of the specified depth and well within the maximum permitted spacing. 9. In addition to cleanliness of aggregate, stipulated in Section 8.3.1, any reactive elements of aggregate should be neutralized through the use of low alkali cement or a suitable pozzolan, or preferably both. Certain cherts and other expansive aggregates and lignite can cause cracks at popouts. Job specifications should cover these aggregate properties and constructors should ensure observance of these requirements. 10. Correct amounts of entrained air should be specified and used to prevent cracking due to freezing and thawing and exposure to calcium or sodium chloride. 8.6 - Specifications to minimize drying shrinkage Actions during construction to obtain the lowest possible drying shrinkage must be supported by the specifications. Unless bids are taken on this basis, the contractor cannot be expected to provide other than ordinary materials, mixes, and procedures. The following items should be carefully spelled out in the specifications. 8.6.1 Concrete materials - They can have an important influence on drying shrinkage. 1. Cement should be Types I, II, V, or IS, preferably not Type III. 2. Aggregates favorable to low mixing water content are (a) well graded, (b) well shaped (not elongated, flat, or splintery), and (c) free of clay, dirt, and excess fines. 3. Aggregate should consist of rock types which will produce low-shrinkage concrete (see Section 3.4.2). 4. Calcium chloride should be prohibited.

CONTROL OF CRACKING

8.6.2 Concrete mixes - For least shrinkage, the mix proportioning should incorporate those factors that contribute to the lowest water content. This means: 1. The largest practical maximum size of aggregate (MSA). 2. The lowest practical sand content. 3. The lowest practical slump. 4. The lowest practical temperature. 5. Less than half the smooth grading curve amount of small coarse aggregate, No. 4 to 3/8 or 3/4 in. (4.75 mm to 9.5 or 19 mm), especially if it is crushed material. 8.6.3 Concrete handling and placing - Equipment (chutes, belts, conveyors, pumps, hoppers, and bucket openings) should be capable of working effectively with lower slump, larger MSA concrete wherever it is appropriate and feasible to use. (It is cautioned that too often, in order to expedite pumping, the actions taken are those which increase drying shrinkage and resultant cracking: more sand, more fines, more water, more slump, smaller aggregate. When pumping is to be permitted and freedom from shrinkage cracking is important, special emphasis must be placed on obtaining effective locations and an adequate number of contraction joints. Moreover, the use of pumping equipment capable of handling mixes favorable to least cracking should be required.) Vibrators should be the largest and most powerful that can be operated in the placement. Upper lifts of formed concrete should be revibrated as late as the running vibrator will penetrate under its own weight. 8.6.4 Finishing - Finishing should follow the

recommendations of ACI 302.1R to minimize or

avoid all forms of surface cracking. It is particularly important that flatwork joint grooves have a depth of at least l/5 of slab thickness, but not less than 1 in. (25.4 mm) deep.

8.6.5 Forms - Forms should have ample strength to sustain strong vibration of low slump concretes. Exposure of warm concrete surfaces to fast drying conditions or to low temperatures prior to curing, should be avoided during form removal, if drying and thermal shrinkage cracking is to be prevented.

8.6.6 Contraction joints - Plans should include an adequate system of contraction joints to provide for shrinkage. Formed grooves should be constructed in both sides of parapet, retaining, and other walls at the depth and spacing indicated in Sec. 8.2.1. 8.6.7 Curing and protection - These procedures should insure the presence of adequate moisture to sustain hydration and strength development in the surface concrete. Rapid drying of the surfaces at the conclusion of the specified curing period should be

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avoided. Providing time for adjustment and gradual, slow elongation will minimize cracking. Water curing should use a wet cover in contact with the concrete surfaces. At the end of the wet curing period, preferably at least 7 days, the cover should be left in place until it and the concrete surface appear to be dry, especially in arid weather. In less arid areas and for interiors, the forms will provide adequate curing if exposed surfaces are protected from drying and provided they can be left in contact with the concrete for at least 7 days. Thereafter, the forms should be left on with loosened bolts long enough to allow the concrete surfaces to dry gradually. Ponding is not a desirable method of curing in an arid climate because of the quick drying that occurs when it is discontinued. Because drying is slow and prolonged, a properly applied sealing compound provides good curing for flatwork placed on a well-wetted subgrade and provides adequate curing for massive sections. In an arid climate, sealing compounds are not adequate for thinner structural sections. When used on formed surfaces, they should be applied when the thoroughly wetted surface is still damp but no longer wet. 8.7 - Conclusion As noted early in this chapter, it is the responsibility of the engineer to develop effective designs and clear and specific specifications. To assure both the owner’s and the engineer’s satisfaction with the results, the engineer should have the owner arrange for inspection by either the owner’s personnel, the engineer, or a reliable professional inspection service who will insure that the construction is performed on the same basis as it was bid. Without the full and firm intent to confirm the specified character and degree of performance, there is a serious chance that undesirable results will be obtained. Without firm inspection and controls, and a clear understanding of the job requirements by the contractor, it is likely that concrete will contain more water than it should, finishing operations will be expedited with the water brush (or hose), and curing will be interrupted or abbreviated (not to mention other less obvious items which influence the later appearance of unsightly cracks). When properly applied, the procedures discussed in this chapter can be used to produce a high quality concrete with the least probable amount of cracking. References

8.1. Concrete M a n u a l 8th Edition, U.S. Bureau of Reclamation, Denver, 1975, 627 pp. 8.2. Dakhil, Fadh H.; Cady, Philip D.; and Carrier, Roger, E., “Cracking in Fresh Concrete as Related to

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Reinforcement,” ACI JOURNAL, Proceedings V. 72, No. 8, Aug. 1975, pp. 421-428.

517.2R 544.3R

Concrete

Chapter 9 - References 9.1- Recommended references The documents of the various standards producing organizations referred to in this document are listed below with their serial designation. American Association of State Highway and Transportation Officials Plastic Fines in Graded Aggregate and T176 Soils By Use of the Sand Equivalent Test American Concrete Institute Guide to Durable Concrete 201.2R Mass Concrete 207.1R Effect of Restraint, Volume Change, and 207.2R Reinforcement on Cracking of Massive Concrete Standard Practice for Selecting Propor211.1 tions for Normal, Heavyweight, and Mass Concrete Admixtures for Concrete 212.1R/ and Guide for Use of Admixtures in Con212.2R crete Standard Practice for the Use of Shrink223 age-compensating Concrete Guide for Concrete Floor and Slab Con302.1R struction Guide for Measuring, Mixing, Transpor304R tating, and Placing Concrete 305R Hot Weather Concreting Cold Weather Concreting 306R Standard Practice for Curing Concrete 308 Recommended Practice for Design and 313 Construction of Concrete Bins, Silos, and Bunkers for Storing Granular Materials Building Code Requirements for Rein318 forced Concrete 340.lR Design Handbook in Accordance with the Strength Design Method of ACI 318-83, Volume 1 - Beams, Slabs, Brackets, Footings, and Pile Caps (SP-17) Precast Concrete Units Used as Forms 347.1R for Cast-in-Place Concrete Guide to Joint Sealants for Concrete 504R Structures

Accelerated Curing of Concrete at Atmospheric Pressure - State of the Art Guide for Specifying, Mixing, Placing and Finishing Steel Fiber Reinforced

ASTM C 512 E 399

Test Method for Creep of Concrete in Compression Test Method for Plane-Strain Fracture Toughness of Metallic Materials

Cornit Euro-International du B&ton and F&i&&m Internationale de la Prkcontrainte CEB-FIP Model Code for Concrete Structures The above publications may be obtained from the following organizations: American Association of State Highway and Transportation Officials 444 North Capital St., N.W. Suite 225 Washington, DC 20001 American Concrete Institute P.O. Box 19150 Detroit, MI 48219 ASTM 1916 Race Street Philadelphia, PA 19103 Cornit Euro-International du B&on and Federation Internationale de la Precontrainte - English edition available from: British Cement Association Wexham Springs Slough SL# 6PL ENGLAND

9.2 - Cited references Cited references are provided at the end of each chapter.

This report was submitted to letter ballot of the committee which consists of 24 members; 21 were affirmative, 2 were not returned, and 1 abstained. It has been processed in accordance with the Institute procedure and is approved for publication and discussion.

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ACI Committee 224 Cracking Bernard L. Meyers Past Chairman

David Darwin Chairman Donald L. Houghton Paul H. Kaar Tony C. Liu J. P. Lloyd LeRoy Lutz V. M. Malhotra Dan Naus Edward G. Nawy

R. S. Barneyback, Jr. Eduardo Santos Basilio Alfred G. Bishara Roy W. Carlson Noel J. Everard J. Ferry-Borges Peter Gergely

Robert E. Philleo Milos Polivka Julius G. Potyondy Robert E. Price Ernest K. Schrader Lewis H. Tuthill Robert L. Yuan

The committee voting on the 1990 revisions was as follows: Randall W. Poston Secretary

Grant T. Halvorsen* Chairman Florian G. Barth Alfred G. Bishara Howard L. Boggs Merle E. Brander David Darwin* Fouad H. Fouad* Peter Gergely *Members contributing to these revisions.

Will Hansen Tony C. Liu Edward G. Nawy John D. Nicholas Harry Palmbaum Arnfinn Rusten Andrew Scanlon

Ernest K. Schrader Wimal Suaris Lewis H. Tuthill* Thomas D. Verti Zenon Zielinski