Home > Guide for Modeling and Calculating Shrinkage and Creep in Har~ened Concrete

Guide for Modeling and Calculating Shrinkage and Creep in Har~ened Concrete

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Disclaimer: Bazant voted against the statistical model comparisons in this guide and believes them to be misleading. His name appears since this was mandatory for committee members.
ACI 209.2R-08
Guide for Modeling and Calculating Shrinkage
./
and Creep in Har~ened Concrete
Reported by ACI Committee 209
Akthem A. AI-Manaseer Zdenek P. Bazant
J eff~ey J. Brooks
Ronald G. Burg Mario Alberto Chiorino Carlos C. Videla'
Chair
MaI'\van A. Daye Walter H. Dilger Noel J. Gardner' Will Hansen Hesham Marzouk *Members of the subcommittee that prepared this guide.
This guide is intended for the prediction of shrinkage and creep in compression in hardened concrete. It may be assumed that predictions apply to concrete under tension and shear. It outlines the problems and limitations in developing prediction equations jor shrinkage and compressive creep of hardened concrete. It also presents and compares the prediction capabilities offour different numerical methods. The models presented are valid jar hardened concrete moist cured for at least 1 day and loaded after curing or later. The models are intended jar concretes with mean compressive cylindrical strengths at 28 days within a range oj at least 20 to 70 MPa (3000 to 10,000 psi). This document is addressed to designers who wish to predict shrinkage and creep in concrete without testing. For structures that are sensitive to shrinkage and creep, the accuraCj of an individual model's predictions can be improved and their applicable range expanded if the model is calibrated' with test data oj the actual concrete to be used in the project.
Keywords: creep; drying shrinkage; prediction models; statistical indicators.
ACI Committee Reports, Guides, Manuals, Standard Practices, and Commentaries are intended for guidance in planning, designing, executing, and inspecting construction. This document is intended for the use of individuals who are competent to evaluate the significance and limitations of its content and recommendations and who will accept responsibility for the application of the material it contains. The American Concrete Institute disclaims any and all responsibility for the stated principles. The Institute shall not be liable for any loss or damage arising therefrom. Reference to this document shall not be made in contract documents. If items found in this document are desired by the ArchitectiEngineer to be a part of the contract documents, they shall be restated in mandatory language for incorporation by the ArchitectlEngineer.
Domingo J. Carreira' Secretary David B. McDonald' Harald S. Mueller Ham H. A. Nassif Lawrence C. Novak Klaus Alexander Rieder Ian Robertson Kenji Sakata
K. Nam Shiu
W. Jason Weiss
CONTENTS Chapter 1-lntroduction and scope, p. 209.2R-2
1. I-Background
1.2-Scope 1.3-Basic assumptions for development of prediction models Chapter 2-Notation and definitions, p. 209.2R-3 2. I-Notation 2.2-Defmitions Chapter 3-Prediction models, p. 209.2R-5 3 . I-Data used for evaluation of models 3 .2-Statistical methods for comparing models 3 .3-Criteria for prediction models 3.4-Identification of strains 3 .5-Evaluation criteria for creep and shrinkage models Chapter 4-Model selection, p. 209.2R-7 4.l-ACI 209R-92 model 4.2-Bazant-Baweja B3 model 4.3-CEB MC90-99 model 4.4-GL2000 model 4.5-Statistical comparisons 4.6-Notes about models
ACI 209.2R-08 was adopted and published May 2008. Copyright © 2008, 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 electronic or mechanical device, printed, 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.
209.2R·1

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209.2R-2 ACI COMMITIEE REPORT
Chapter 5-References, p. 209.2R-13 5.1-Referenced standards and reports 5.2-Cited references Appendix A-Models, p. 209.2R-16 A.1-ACI 209R-92 model A.2-Bazant-Baweja B3 model A.3-CEB MC90-99 model A.4--GL2000 model Appendix B-Statistical indicators, p. 209.2R-28 B.I-BP coefficient of variation (til Bp%) method B.2-CEB statistical indicators B.3-The Gardner coefficient of variation (IDG) Appendix C-Numeric examples, p. 209.2R-30 C.l-ACI 209R-92 model solution C.2-Bazant-Baweja B3 model solution C.3-CEB MC90-99 model solution C.4--GL2000 model solution C.5-Graphical comparison of model predictions CHAPTER 1-INTRODUCTION AND SCOPE 1.1-Background To predict the strength and serviceability of reinforced and prestressed concrete structures, the structural engineer requires an appropriate description of the mechanical properties of the materials, including the prediction of the time-dependant strains of the hardened concrete. The prediction of shrinkage and creep is important to assess the risk of concrete cracking, and deflections due to stripping-reshoring. As discussed in ACI 209.lR, however, the mechanical properties of concrete are significantly affected by the temperature and availability of water during curing, the environmental humidity and temper- ature after curing, and the composition of the concrete, including the mechanical properties of the aggregates. Among the time-dependant properties of concrete that are of interest to the structural engineer are the shrinkage due to cement hydration (self-desiccation), loss of moisture to the environment, and the creep under sustained loads. Drying before loading significantly reduces creep, and is a major complication in the prediction of creep, stress relaxation, and strain recovery after unloading. While there is a lot of data on shrinkage and compressive creep, not much data are available for creep recovery, and very limited data are available for relaxation and tensile creep.
Cre~p under variable stresses and the stress responses
under constant or variable imposed strains are commonly determined adopting the principle of superposition. The limitations of this assumption are discussed in Section 1.3. Further, the experimental results of Gamble and Parrott (1978) indicate that both drying and basic creep are only partially, not fully, recoverable. In general, provided that water migration does not occur as in sealed concrete or the interior of large concrete elements, superposition can be used to calculate both recovery and relaxation. The use of the compressive creep to the tensile creep in calculation of beam's time-dependant deflections has been successfully applied in the work by Branson (1977), Bazant and Ho (1984), and Carreira and Chu (1986). The variability of shrinkage and creep test measurements prevents models from closely matching experimental data. The within-batch coefficient of variation for laboratory- measured shrinkage on a single mixture of concrete was approximately 8% (Bafant et al. 1987). Hence, it would be unrealistic to expect results from prediction models to be within plus or minus 20% of the test data for shrinkage. Even larger differences occur for creep predictions. For structures where shrinkage and creep are deemed critical, material testing should be undertaken and long-term behavior extrapolated from the resulting data. For a discussion of testing for shrinkage and creep, refer to Acker (1993), Acker et al. (1998), and Carreira and Burg (2000). 1.2-Scope This document was developed to address the issues related to the prediction of creep under compression and shrinkage- induced strains in hardened concrete. It may be assumed, however, that predictions apply to concrete under tension and shear. It outlines the problems and limitations in developing prediction equations, presents and compares the prediction capabilities of the ACI 209R-92 (ACI Committee 209 1992), Bazant-Baweja B3 (Bafant and Baweja 1995, 2000), CEB MC90-99 (Muller and Hillsdorf 1990; CEB 1991, 1993, 1999), and GL2000 (Gardner and Lockman 2001) models, and gives an extensive list of references. The models presented are valid for hardened concrete moist cured for at least 1 day and loaded at the end of 1 day of curing or later. The models apply to concretes with mean compressive cylindrical strengths at 28 days within a range of at least 20 to 70 MPa (3000 to 10,000 psi). The prediction models were calibrated with typical composition concretes, but not with concretes containing silica fume, fly ash contents larger than 30%, or natural pozzolans. Models should be calibrated by testing such concretes. This document does not provide information on the evaluation of the effects of creep and shrinkage on the structural performance of concrete structures. 1.3-Basic assumptions for development of prediction models Various testing conditions have been established to stan- dardize the measurements of shrinkage and creep. The following simplifying assumptions are normally adopted in the development of prediction models.
1.3.1 Shrinkage and creep are additive-Two nominally
identical sets of specimens are made and subjected to the same curing and environment conditions. One set is not loaded and is used to determine shrinkage, while the other is generally loaded from 20 to 40% of the concrete compressive strength. Load- induced strains are determined by subtracting the measured shrinkage strains on the nonloaded specimens from the strains measured on the loaded specimens. Therefore, it is assumed that the shrinkage and creep are independent of each oft.1er. Tests carried out on sealed specimens, with no moisture movement from or to the specimens, are used to determine autogenous shrinkage and basic creep.

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1.3.2 Linear aging model for creep-Experimental
research indicates that creep may be considered approxi- mately proportional to stress (L'Hermite et al. 1958; Keeton 1965), provided that the applied stress is less than 40% of the concrete compressive strength. The strain responses to stress increments applied at different times may be added using the superposition principle (McHenry 1943) for increasing and decreasing stresses, provided strain reversals are excluded (for example, as in relaxation) and temperature and moisture content are kept constant (Le Camus 1947; Hanson 1953; Davies 1957; Ross 1958; Neville and Dilger 1970; Neville 1973; BaZant 1975; Gamble and Parrot 1978; RlLEM Technical Committee TC-69 1988). Major deviations from the principle of superposition are caused by the neglect of the random scatter of the creep properties, by hygrothermal effects, including water diffusion and time evolution of the distributions of pore moisture content and temperature, and by material damage, including distributed cracking and fracture, and also frictional microslips. A comprehensive summary of the debate on the applicability of the principle of superposition when dealing with the evaluation of creep structural effects can be found in the references (BaZant 1975, 1999, 2000; CEB 1984; RILEM Technical Committee TC-1 07 1995; Al Manaseer et al. 1999; Jirasek and BaZant 2002; Gardner and Tsuruta 2004; Bazant 2007). 1.3.3 Separation of creep into basic creep and drying
creep-Basic creep is measured on specimens that are sealed
to prevent the ingress or egress of moisture from or to its environment. It is considered a material constitutive property and independent of the specimen size and shape. Drying creep is the strain remaining after subtracting shrinkage, elastic, and basic creep strains from the total measured strain on nominally identical specimens in a drying environment. The measured average creep of a cross section at drying is strongly size- dependant. Any effects of thermal strains have to be removed in all cases or are avoided by testing at constant temperature. In sealed concrete specimens, there is no moisture movement into or out of the specimens. Low-water-cement-ratio concretes self-desiccate, however, leading to autogenous shrinkage. Normal-strength concretes do not change volume at relative humidity in the range 95 to 99%, whereas samples stored in water swell (L'Hermite et al. 1958).
1.3.4 Differential shrinkage and creep or shrinkage and creep gradients are neglected-The shrinkage strains deter-
mined according to ASTM C157/C157M are measured along the longitudinal axis of prismatic specimens; however, the majority of reported creep and shrinkage data are based on surface measurements of cylindrical specimens (ASTM C512). Unless fmite element analysis (BaZant et al. 1975) or equivalent linear gradients (Carreira and Walser 1980) are used, it is generally assumed that shrinkage and creep strains in a specimen occur uniformly through the specimen cross section. Kristek et al. (2006) concluded that for box girder bridges, the classical creep analysis that assumes the shrinkage and creep properties to be uniform throughout the cross section is inadequate. As concrete ages, differences in strain gradients reduce (Carreira and Walser 1980; Aguilar 2005).
1.3.5 Stresses induced during curing phase are negligible-
Most test programs consider the measurement of strains from the start of drying. It is assumed that the restrained stresses due to swelling and autogenous shrinkage are negligible because of the large creep strains and stress relaxation of the concrete at early ages. For restrained swelling, this assumption leads to an overestimation of the tensile stresses and, therefore, it may be an appropriate basis for design when predicting deflections or prestress losses. For predicting the effects of restrained autogenous shrinkage or relaxation, however, the opposite occurs. Limited testing information exists for tensile creep.
CHAPTER 2-NOTATION AND DEFINITIONS 2.1-Notation
a,b
=
a
=
Co(t,to) = Cl.t,to,te) = c
=
d=4V/S = E
=
Eem
=
Eem28
=
Eemt
=
Eemto
=
e=2V1S = fem
=
fern28
=
fernt
=
fernte
=
femto
=
constants used to describe the strength gain development of the concrete, ACI 209R-92 and GL2000 models agfregate content of concrete, kg/m
3
or lb/ yd ,B3 model compliance function for basic creep at concrete age t when loading starts at age to' B3 model compliance function for drying creep at concrete age t when loading and drying starts at ages to and te, respectively, B3 model cement content of concrete, kg/m3 or Ib/yd3, ACI 209R-92 and B3 models average thickness of a member, mm or in., ACI 209R-92 model modulus of elasticity, MPa or psi mean modulus of elasticity of concrete, MPa or psi mean modulus of elasticity of concrete at 28 days, MPa or psi mean modulus of elasticity of concrete at age
t, MPa or psi
mean modulus of elasticity of concrete when loading starts at age to' MPa or psi effective cross section thickness of member or notional size of member according to B3 or CEB MC90 and CEB MC90-99 models, respectively, in mm or in.; defined as the cross-section divided by the semi-perimeter of the member in contact with the atmo- sphere, which coincides with the actual thick- ness in the case of a slab concrete mean compressive cylinder strength, MPaorpsi concrete mean compressive cylinder strength at 28 days, MPa or psi concrete mean compressive cylinder strength at age t, MPa or psi
, lit
concrete mean compressive cylinder strength when drying starts at age te, MPa or psi concrete mean compressive cylinder strength when loading starts at age to' MPa or psi

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209.2R-4
fd
=
H(t)
=
h
=
J(t,to)
=
J(to,to)
=
kh,13RJlh)
or 13(h)
=
ks
=
ql
=
S(t - te), 13s<t - te)
or 13(t - te)=
s
=
T
= =
t- te
=
te
=
to
=
VIS
=
w
=
Cl
=
Cli or k
=
Cl2
=
Clas' Cldsl
and Clds2 =
13as(t!
=
13e(t - to) =
13ds(t - te) =
13e
13RH,T
=
ACI COMMITTEE REPORT
concrete specified cylinder strength at 28 days, MPa or psi spatial average of pore relative humidity at concrete age t, B3 model relative humidity expressed as a decimal compliance at concrete age t when loading starts at age to' IlMPa or lIpsi elastic compliance at concrete age to when loading starts at age to' IIMPa or l/psi correction term for effect of humidity on shrinkage according to B3, CEB MC90 and CEB MC90-99, or GL2000 models, respec- tively cross-section shape factor, B3 model inverse of asymptotic elastic modulus, IIMPa or lIpsi, B3 model correction term for effect of time on shrinkage according to B3, CEB MC90, or GL2000 models, respectively slump, mm or in., ACI 209R-92 model. Also, strength development parameter, CEB MC90, CEB MC90-99, and GL2000 models temperature, DC, OF, or OK age of concrete, days duration of drying, days age of concrete when drying starts at end of moist curing, days age of concrete at loading, days volume-surface ratio, mm or in. water content of concrete, kg/m3 or Ib/yd3, B3 model air content expressed as percentage, ACI 209R-92 model shrinkage constant as function of cement type, according to B3 or GL2000 models, respectively shrinkage constant related to curing conditions, B3 model correction coefficients for effect of cement type on autogenous and drying shrinkage, CEB MC90-99 model function describing time development of autogenous shrinkage, CEB MC90-99 model correction term for effect of time on creep coefficient according to CEB MC90 and CEB MC90-99 models function describing time development of drying shrinkage, CEB MC90-99 model factor relating strength development to cement type, GL2000 correction coefficient to account for effect of temperature on notional shrinkage, CEB MC90modei
~sc = correction coefficient that depends on type of
cement, CEB MC90 model correction coefficient to account for effect of temperature on time development of shrinkage, CEB MC90 model autogenous shrinkage strain at concrete age t,
mm1mm or in.lin., CEB MC90-99
drying shrinkage strain at concrete age t since the start of drying at age te, mm1mm or in.lin., CEB MC90-99 model
Eeso
= notional shrinkage coefficient, mm1mm or
in.lin., CEB MC90 model
Eeasoifem2S) = notional autogenous shrinkage coefficient, mm1mm or in.lin., CEB MC90-99 model Eedsoifem2S)= notional drying shrinkage coefficient, mmI
mm or in.lin., CEB MC90-99 model
Esh(t,te) = shrinkage strain at concrete age t since the
start of drying at age te, mm1mm or in.lin.
Eshu or Eshoo= notional ultimate shrinkage strain, mm1mm IP(t, to) IP2S(t, to) IPo IPRJlh) 'tsh
= = = =
= =
=
or in.lin., ACI 209R-92 and GL2000 models and B3 model, respectively creep coefficient (dimensionless) 28-day creep coefficient (dimensionless), CEB MC90, CEB MC90-99, and GL2000 models notional creep coefficient (dimensionless), CEB MC90 and CEB MC90-99 models correction term for effect of relati ve humidity on notional creep coefficient, CEB MC90 and CEB M90-99 models correction term for effect of drying before loading when drying starts at age te, GL2000 model ultimate (in time) creep coefficient, ACI 209R-92 model unit weight of concrete, kg/m3 or Ib/ft
3
shrinkage and creep correction factor, respec- tively; also used as product of all applicable corrections factors, ACI 209R-92 model shrinkage half-time, days, ACI 209R-92 and B3 models
= ratio of fine aggregate to total aggregate by
weight expressed as percentage, ACI 209R-92 model 2.2-0efinitions autogenous shrinkage-the shrinkage occurring in the absence of moisture exchange (as in a sealed concrete specimen) due to the hydration reactions taking place in the cement matrix. Less commonly, it is termed basic shrinkage or chemical shrinkage. basic creep--the time-dependent increase in strain under sustained constant load of a concrete specimen in which moisture losses or gains are prevented (sealtid sp~~men). compliance J(t,to)-the total load induced strain (elastic strain plus creep strain) at age t per unit stress caused by a unit uniaxial sustained load applied since loading age to'

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creep coefficient-the ratio of the creep strain to the initial strain or, identically, the ratio of the creep compliance to the compliance obtained at early ages, such as after 2 minutes. 28-day creep coefficient-the ratio of the creep strain to the elastic strain due to the load applied at the age of 28 days
(cP2s(t,to) = cI>(t,to) . Ecm2SIEcmto)·
creep strain-the time-dependent increase in strain under constant load taking place after the initial strain at loading. drying creep--the additional creep to the basic creep in a loaded specimen exposed to a drying environment and allowed to dry. drying shrinkage-shrinkage occurring in a specimen that is allowed to dry. elastic compliance or the nominal elastic strain per unit stress J(to,to)-the initial strain at loading age to per unit stress applied. It is the inverse of the mean modulus of elasticity of concrete when loading starts at age to' initial strain at loading or nominal elastic strain-the short-term strain at the moment of loading and is frequently considered as a nominal elastic strain as it contains creep that occurs during the time taken to measure the strain. load-induced strain-the time-dependent strain due to a constant sustained load applied at age to' shrinkage-the strain measured on a load-free concrete specimen. specific creep--the creep strain per unit stress. total strain-the total change in length per unit length measured on a concrete specimen under a sustained constant load at uniform temperature.
CHAPTER 3-PREDICTION MODELS 3.1-Data used for evaluation of models
In 1978, BaZant and Panula started collecting shrinkage and creep data from around the world and created a comput- erized databank, which was extended by Muller and Panula as part of collaboration between the ACI and the CEB established after the ACI-CEB Hubert Rusch workshop on concrete creep (Hillsdorf and Carreira 1980). The databank, now known as the RILEM databank, has been extended and refined under the sponsorship of RILEM TC 107-CSP, Subcommittee 5 (Kuttner 1997; Muller et al. 1999). Problems encountered in the development of the databank have been discussed by Muller (1993) and others (Al-Mana- seer and Lakshmikantan 1999; Gardner 2000). One problem involves which data sets should be included. For example, some investigators do not include the low-modulus sandstone concrete data of Hansen and Mattock (1966), but do include the Elgin gravel concrete data from the same researchers. A further problem is the data of some researchers are not inter- nally consistent. For example, the results from the 150 mm. (6 in.) diameter specimens of Hansen and Mattock are not consistent with the results from the 100 and 200 mm (4 and 8 in.) diameter specimens. Finally, it is necessary to define the relative humidity for sealed and immersed concrete specimens. A major problem for all models is the description of the concrete. Most models are sensitive to the type of cement and the related strength development characteristics of the material. Simple descriptions, such as ASTM C150 Type I, used in the databank are becoming increasingly difficult to interpret. For example, many cements meet the requirements of Types I, II, and III simultaneously; also, the multiple additions to the clinker allowed in ASTM C595 or in other standards are unknown to the researcher and designer. N ominaIly identical concretes stored in different environments, such as those tested by Keeton (1965), have different strength development rates. If this information exists, it should be taken into account in model development.
In addition, cement descriptions differ from country to
country. The data obtained from European cement concretes may not be directly compared with that of United States cement concretes. Some researchers have suggested that correlation should only be done with recent and relevant data and that different shrinkage and creep curves should be developed for European, Japanese, North American, and South Pacific concretes (McDonald 1990; McDonald and Roper 1993; Sakata 1993; Sakata et al. 2001; Videla et al. 2004; Videla and Aguilar 2005a). While shrinkage and creep may vary with local conditions, research has shown that short-term shrinkage and creep measurements improve the predictions regardless oflocation (Bazant 1987; Bazant and Baweja 2000; Aguilar 2005). For this reason, the committee recommends short-term testing to determine the shrinkage, creep, and elastic modulus of the concrete to improve the predictions of the long-term deformations of the concrete. Other issues include: • The databank does not include sufficient data to validate modeling that includes drying before loading or loading before drying, which are common occurrences in practice; • Many of the data sets in the databank were measured over relatively short durations, which reduces the usefulness of the data to predict long-term effects; and • Most of the experiments were performed using small specimens compared with structural elements. It is debatable if the curing environment and consequent mechanical properties of concrete in the interior of large elements are well represented by small specimen experiments (BaZant et al. 1975; Kristek et al. 2006). Despite these limitations, it is imperative that databanks such as the RILEM databank are maintained and updated as they provide an indispensable source of data in addition to a basis for comparing prediction models.
3.2-5tatistical methods for comparing models
Several methods have been used for the evaluation of the accuracy of mOdels to predict experimental data. Just as a single set of data may be described by its mean, mode, median, standard deviation, and maximum and minimum, a model for shrinkage or creep data may have several methods to describe its deviation from the data. The committee could not agree on a single method for comparison of test data with predictions from models for shrinkage and creep. Reducing the comparison between a large num~r of experimental
IJ·
results and a prediction method to a single number is fraught with uncertainty. Therefore, the committee strongly recom- mends designers to perform sensitivity analysis of the response of the structure using the models in this report and

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209.2R-6 ACI COMMITTEE REPORT
to carry out short-tenn testing to calibrate the models to improve their predictions. The summary of the statistical indicators given in Section 4 provides the user with basis for comparison without endorsing any method. One of the problems with the comparison of shrinkage and creep data with a model's prediction is the increasing divergence and spread of data with time, as shown in the figures of Section 4. Thus, when techniques such as linear regression are used, the weighting of the later data is greater than that of the earlier data (Bazant 1987; BaZant et al. 1987). On the contrary, comparison of the percent deviation of the model from the data tends to weight early-age data more than later-age data. The divergence and spread are a measure of the limitation of the model's capabilities and variability in the experimental data. Commonly used methods for detennining the deviation of a model from the data include: • Comparison of individual prediction curves to individual sets of test data, which requires a case-by-case evaluation; Comparison of the test data and calculated values using linear regression; • Evaluation of the residuals (measured-predicted value) (McDonald 1990; McDonald and Roper 1993; Al- Manaseer and Lakshmikantan 1999). This method does not represent least-square regression and, if there is a trend in the data, it may be biased; and • Calculation of a coefficient of variation or standard error of regression nonnalized by the data centroid.
In the committee's opinion, the statistical indicators available
are not adequate to uniquely distinguish between models. 3.3-Criteria for prediction models Over the past 30 years, several models have been proposed for the prediction of drying shrinkage, creep, and total strains under load. These models are compromises between accuracy and convenience. The committee concludes that one of the primary needs is a model or models accessible to engineers with little specialized knowledge of shrinkage and creep. Major issues include, but are not restricted to: • How simple or complex a model would be appropriate, and what input infonnation should be required; • What data should be used for evaluation of the model; • How closely the model should represent physical phenomena/behavior; • What statistical methods are appropriate for evaluating a model.
Th~re is no agreement upon which infonnation should be
required to calculate the time-dependent properties of concrete; whether the mechanical properties of the concrete specified at the time of design should be sufficient or if the mixture proportions are also required. At a minimum, the committee believes that shrinkage and creep models should include the following infonnation: • Description of the concrete either as mixture propor- tions or mechanical properties such as strength or modulus of elasticity; • Ambient relative humidity; • Age at loading; Duration of drying; • Duration of loading; and • Specimen size. . Models should also: • Allow for the substitution of test values of concrete strength and modulus of elasticity; • Allow the extrapolation of measured shrinkage and creep compliance results to get long-tenn values; and • Contain mathematical expressions that are not highly sensitive to small changes in input parameters and are easy to use. As described in ACI 209.1R, it has long been recognized that the stiffness of the aggregate significantly affects the shrinkage and creep of concrete. Some models account for the effect of aggregate type by assuming that the effects of aggregate are related to its density or the concrete elastic modulus. Models that use concrete strength can be adjusted
to use a measured modulus of elasticity to account for aggregate
properties. Models that do not use the mechanical charac- teristics of the concrete and rely on mixture proportion information alone may not account for variations in behavior due to aggregate properties. Until recently, autogenous shrinkage was not considered significant because, in most cases, it did not exceed 150 microstrains. For concretes with water-cement ratios (w/c) less than 0.4, mean compressive strengths greater than 60 MPa (8700 psi), or both, however, autogenous shrinkage may be a major component of the shrinkage strain. Some models consider that basic creep and drying creep are independent and thus additive, while other models have shrinkage and creep as dependent, and thus use multiplicative factors. The physical phenomenon occurring in the concrete may be neither. 3.4-ldentification of strains Equations (3-1) and (3-2) describe the additive simplification discussed in 1.3.1 total strain = shrinkage strain + compliance x stress (3-1)
r
(elastic strain + basic creep + drying creep) (3-2) comp lance = stress
The total and shrinkage strains are measured in a creep and shrinkage test program from which the compliance is deter- mined. Errors in the measured data result in errors in the compliance. The elastic strain is detennined from early-age measurement, but as discussed previously, it is difficult to separate early-age creep from the elastic strain. Thus, the assumed elastic strain is dependent on the time at which the strain measurement is made and, therefore, on the ignored early creep. Basic creep and drying creep are detennined from the compliance by subtracting the elastic strain, which may have implicit errors, from the strains measured on dfNjng and nondrying specimens. Errors in the measured elastic strain used to determine the modulus of elasticity (ASTM C469), in the measured total strain, or in the measured shrinkage

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209.2R-7
strain, are all reflected in the calculated creep strain, the compliance, and creep coefficient. For sealed specimens, the equations for compliance and total strain simplify significantly if autogenous shrinkage is ignored as in Eq. (3-3) and (3-4) total strain = compliance x stress (3-3) compliance = (elastic strain + basic creep) (3-4) stress 3.5-Evaluatlon criteria for creep and shrinkage models In 1995, RILEM Committee TC 107 published a list of criteria for the evaluation of shrinkage and creep models (RILEM 1995; BaZant 2000). In November 1999, ACI Committee 209, which has a number of members in common with RILEM TC 107, discussed the RILEM guidelines and agreed on the following: • Drying shrinkage and drying creep should be bounded. That is, they do not increase indefinitely with time; • Shrinkage and creep equations should be capable of extrapolation in both time and size; • Shrinkage and creep models should be compared with the data in the databank limited by the conditions of applicability of the model(s). That is, some experi- mental data, such as those with high water-cement ratios or low-modulus concrete, may not be appropriate to evaluate a model; • Equations should be easy to use and not highly sensitive to changes in input parameters; • The shape of the individual shrinkage and creep curves over a broad range of time (minutes to years) should agree with individual test results; • Creep values should be compared as compliance or specific creep rather than as the creep coefficient. The immediate strain/unit stress and the modulus of elasticity are dependent on the rate of loading; however, for developing the creep equations to determine long-term deformations, this effect should not playa major role; • Creep expressions should accommodate drying before loading. Results by Abiar reported by Acker (1993) show that predried concrete experiences very little creep. Similarly, the very late-age loaded (2500 to 3000 days) results of Wesche et al. (1978) show reduced creep compared with similar concrete loaded at early ages. The effect of predrying may also be significantly influenced by the size of the specimen; • Shrinkage and creep expressions should be able to accommodate concretes containing fly ash, slag (Videla and Gaedicke 2004), natural pozzolans (Videla et al. 2004; Videla and Aguilar 2005a), silica fume and chemical admixtures (Videla and Aguilar 2005b); • The models should allow for the effect of specimen size; and
• The models should allow for changes in relative humidity.
Success in achieving the following guidelines is consequent to the method of calculation; that is, if the principle of super- position is valid and if the model includes drying before loading, and how they are considered under unloading: • Recovery of creep strains under complete unloading should not exceed the creep strain from le>ading, and should asymptotically approach a constant value; and • Stress relaxation should not exceed the initially applied stress. Yue and Taerwe (1992, 1993) published two related papers on creep recovery. Yue and Taerwe (1992) commented, "It is well known that the application of the principle of superposition in the service stress range yields an inaccurate prediction of concrete creep when unloading takes place." In their proposed two-function method, Yue and Taerwe (1993) used a linear creep function to model the time-dependent deformations due to increased stress on concrete, and a separate nonlinear creep recovery function to represent concrete behavior under decreasing stress. CHAPTER 4-MODEL SELECTION There are two practical considerations in the models for prediction of shrinkage and creep, namely: • Mathematical form of their time dependency; and • Fitting of the parameters and the resulting expressions.
If the mathematical form of the model does not accurately
describe the phenomena, extrapolations of shrinkage and creep results will deviate from reality. After the mathematical form has been justified, the fit of the prediction to measured results should be compared for individual data sets. The models selected for comparison are the ACI 209R-92 (ACI Committee 209 1992), the Bazant-Baweja B3 devel- oped by BaZant and Baweja (1995,2000), the CEB Model Code 1990-99 (CEB MC90-99) (Muller and Hillsdorf 1990; CEB 1991, 1993, 1999), and the GL2000 developed by Gardner and Lockman (2001). Table 4.1 lists the individual model's applicable range for different input variables (adapted from AI-Manaseer and Lam 2005). Comparison of models with experimental data is complicated by the lack of agreement on selection of appropriate data and on the methods used to compare the correlation. Descriptions of the ACI 209R-92, BaZant-Baweja B3, CEB MC90-99, and GL2000 models are given in Appendix A. Kristek et al. (200 1) and Sassone and Chiorino (2005) developed design aids for determination of shrinkage, compliance, and relax- ation for ACI 209R-92, BaZant-Baweja B3, CEB MC90-99, and GL2000 models. Figures 4.1 through 4.8 (Gardner 2004) compare the predicted values for two sets of input information for RILEM data sets extending longer than 500 days, concrete 28-day mean cylinder strengthsfcm28 between 16 and 82 MPa (2320 and 11,890 psi), water-cement ratios between 0.4 and 0.6, duration of moist curing longer than 1 day (possibly biased against ACI 209R-92 because this model was developed for standard conditions considering 7 days of moist curing and 7 days of age at loading), age of loading greater than the duration of moist curing, and volume- surface ratios VIS greater than 19 mm (3/4 in.). The humidity range for compliance was 20 to 100%, and below 80% for

Page 8
209.2R-8 ACI COMMITTEE REPORT
Table 4.1-Parameter ranges of each model
Input variables ACI209R-92 Bazant-Baweja B3
fcm28 , MPa (psi)
-
I7 to 70 (2500 to 10,000) ale
-
2.5 to 13.5 Cement content, 279 to 446 160 to 720 kg/m3 (lb/yd3) (470 to 752) (270 to 1215) w/c
-
0.35 to 0.85 Relative humidity, % 40 to 100 40 to 100 Type of cement, RorRS R,SL, RS European (U.S.)
(I or III)
(I, II, III)
tc (moist cured)
~ I day ~ I day
tc (stearn cured)
1 to 3 days -
to
~ 7 days
to ~ tc
shrinkage. Consequently, swelling was not included even if some specimens were initially moist cured. Two sets of comparisons are shown in each figure. One set, identified as "fem only," assumes that only the measured 28-day strength fem is known. The second set, identified as "all data," uses the fem calculated as the average of the measured fem' and that back -calculated from the measured Eem using the elastic modulus formula of the method and mixture proportions if required by the model. Calculated compliance is the calculated specific creep plus calculated elastic compliance for the fem graphs and the calculated specific creep plus measured elastic compliance for the all data graphs. The reported mixture composition was used for ACI 209R-92 and Bazant-Baweja B3. It was assumed that if mixture data were available, the strength development data and elastic modulus would also be available. Cement type was determined by comparison of measured strength gain data with the GL2000 strength gain equations. The same cement type was used for predictions in all methods. For CEB MC90-99, ASTM C 150 Type I was taken as CEB Type N cement, Type III as CEB Type R, and Type II as CEB Type SL. It should be noted that each model should use an appropriate value of elastic modulus for which the model was calibrated. Therefore, for CEB, the elastic modulus was taken as Eem = 9500(fem)1I3 in MPa (262,250[fem]1I3 in psi). For Bazant- Baweja B3, using the shape factor ks = 1.00 in 's (the shrinkage time function) improved the results of the statistical
analy~is, and all concretes were assumed moist cured; that is, u2
= 1.20 for calculations using the BaZant-Baweja B3 model.
To calculate a coefficient of variation (Gardner 2004), the durations after drying or application of load were divided into seven half-log decade intervals: 3 to 9.9 days, 10 to 31 days, 32 to 99 days, 100 to 315 days, 316 to 999 days, 1000 to 3159 days, and greater than 3160 days. That is, each duration is 3.16 times the previous half-log decade; these are similar to the CEB ranges. The root mean square (RMS) (calculated- observed) was calculated for all comparisons in each half-log decade. The coefficient of variation was the average RMSI average experimental value for the same half-log decade.
Model CEBMC90 CEB MC90-99 GL2000 20 to 90 15 to 120 16 to 82 (2900 to 13,000) (2175 to 17,400) (2320 to 11,900) .
-
- - - -
-
-
- 0.40 to 0.60 40 to 100 40 to 100 20 to 100 R, SL,RS R,SL,RS R,SL,RS
(I, II, III) (I, II, III)
(I, II, III)
< 14 days < 14 days
~ 1 day
- - - > 1 day > 1 day
to ~ tc ~ I day
4.1-ACI 209R-92 model The model recommended by ACI Committee 209 (1971) was developed by Branson and Christiason (1971), with minor modifications introduced in ACI 209R-82 (ACI Committee 2091982). ACI Committee 209 incorporated the developed model in ACI 209R-92 (ACI Committee 209 1992). Since then, it has not been revised or updated to the RlLEM databank, and it is compared with very recent models. This model, initially developed for the precast- prestressing industry (Branson and Ozell 1961; Branson 1963, 1964, 1968; Branson et al. 1970; Meyers et al. 1970; Branson and Kripanayanan 1971; Branson and Chen 1972), has been used in the design of structures for many years. Advantages of this model include:
• It is simple to use with minimal background knowledge;
and
• •
It is relatively easy to adjust to match short-term test data simply by modifying ultimate shrinkage or creep to produce the best fit to the data. Its disadvantages include: It is limited in its accuracy, particularly in the method of accommodating member size when its simplest form is used. This disadvantage, however, can be overriden if the methods provided for accommodating the shape and size effect on the time-ratio are applied; and It is empirically based, thus it does not model shrinkage or creep phenomena. At its most basic level, the ACI 209R-92 method only requires: Age of concrete when drying starts, usually taken as the age at the end of moist curing; • Age of concrete at loading; • Curing method; • Relative humidity expressed as a decimal; • Volume-surface ratio or average thickness; and • Cement type. This model calculates the creep coefficient rather than the compliance, which may introduce problems due to the assumed value of elastic modulus. Figures 4.1 and 4.2 show

Page 9
MODELING AND CALCULATING SHRINKAGE AND CREEP IN HARDENED CONCRETE
209.2R-9
the calculated and measured shrinkages and compliances, respectively. The comparison of shrinkage data in Fig. 4.1 clearly shows that the ACI 209R-92 model overestimates measured shrinkage at low shrinkage values (equivalent to short drying times) and underestimates at high shrinkage values (typical of long drying times). This result indicates the limitation of the model's equation used to predict shrinkage. The ACI 209R-92 compliance comparison is rather insensitive to using all of the available data, including mixture proportions, compared with just using the measured concrete strength. 4.2-Bazant-Baweja B3 model The Bazant-Baweja B3 model (Bazant and Baweja 1995, 20(0) is the culmination of work started in the 1970s (Bazant et al. 1976, 1991; Bazant and Panula 1978, 1984; Jirasek and Bazant 2(02), and is based on a mathematical description of over 10 physical phenomena affecting creep and shrinkage (Bazant 2(00), including known fundamental asymptotic properties that ought to be satisfied by a creep and shrinkage model (Bazant and Baweja 2000, RILEM Technical Committee TC 107 1995). This model has been found to be useful for those dealing with simple as well as complex structures. The Bazant-Baweja B3 model uses the compli- ance function. The compliance function reduces the risk of errors due to inaccurate values of the elastic modulus. The model clearly separates basic and drying creep. The factors considered include: • Age of concrete when drying starts, usually taken as the age at the end of moist curing; Age of concrete at loading; Aggregate content in concrete; Cement content in concrete; • Cement type; Concrete mean compressive strength at 28 days; Curing method; Relative humidity; • Shape of specimen; • Volume-surface ratio; and • Water content in concrete. Both Bazant-Baweja B3 shrinkage and creep models may require input data that are not generally available at time of design, such as the specific concrete proportions and concrete mean compressive strength. Default values of the input parameters can be automatically considered if the user lacks information on some of them. The authors suggest when only /cm28 is known, the water-cement ratio can be determined using Eq. (4-1), and typical values of cement content and aggregate cement ratio should be assumed wlc = ((fcm28!22.8) + O.535r
l
in SI units
-I
wlc = [(fcm2s133OO)+0.535] inin.-Ibunits
(4-1)
Equation (4-1) represents the best-fit linear regression equation to the values reported in Tables A1.5.3.4(a) and A6.3.4(a) of ACI 211.1-91 (ACI Committee 211 1991) for non-air-entrained concretes made with Type 1 portland cement; for air-entrained concretes, similar equations can be
.,
Q ~
.. .,
.:!
'tI
S ..
'S
u
'ii
(J
1500 ..---....... --.,..---.,..---.,..---.,-----,.
t. fem only CV = 34%
• All data CV = 41%
1250 I---+----+---+---+---if-----t 1000 750 500 250 0 0 250 500 750 1000 1250 1500
Measured Soh X 10";
Fig. 4.I-ACI 209R-92 versus RILEM shrinkage databank (Gardner 2004).
.. CI.
~
Of
Q
-..
i ...
:::;
'tI
S ..
'S
u
'ii
(J
300 r---------~-__,--__,--...."
A fem only CV = 30%
• All data CV = 30%
250 ~--t_--t__--t__-_Ir--_7I"------f 200 150 100 50 AC1209R·92 0 0 50 100 150 200 250 300
Measured J(t,to) x 10"; (1/MPa)
Fig. 4.2-ACI 209R-92 versus RILEM compliance databank (Gardner 2004). derived by regression analysis of the reported values on ACI 211.1-91. For other cement types and cementitious materials, ACI 211.1-91 suggests that the relationship between water- cement or water-cementitious material ratio and compressive strength of concrete be developed for the materials actually
to be used.
Figures 4.3 and 4.4 show the comparison between the calcu- lated and measured shrinkages and compliances, respectively. The shrinkage equation is sensitive to the water content. The model allows for extrapolation from short-term test data using short-term test data and a test of short-term moisture- content loss. 4.3-CEB MC90-99 model In 1990, CEB presented a model for the prediction of shrinkage and creep in concrete developed by Muller and

Page 10
209.2R-10
ACI COMMITTEE REPORT
" co
..
><
.c
J
"'CI
.s to
:;
u
iii
CJ
1500 r----------,...---,...---,...--~
I:> fom only CV= 31%
• All data CV = 20%
1250 I---r----r----+--,,---+--~>l"--~ 1000 750 500 250 0 0 250 500
I:> I:>
750 1000
Measured Esh x 10"
B3
1250 1500
Fig. 4.3-BaZant-Baweja B3 versus RILEM shrinkage databank (Gardner 2004).
ii'
Do
~
" co
.. Ie
S
::!
-.
"'CI
.s
.!!
:::I U
iii
CJ
300 r---------,...---,...---,...--~
I:> fom only CV = 29%
• All data CV = 27%
250 I---r----r---~+---+--_.>l"--~ 200 150 100 50
B3
0 0 50 100 150 200 250 300
Measured J(t,to) x 10" (lIMPa)
Fig. 4.4-BaZant-Baweja B3 versus RILEM compliance databank (Gardner 2004). Hilsdorf (1990). The model was revised in 1999 (CEB 1999) to include normal- and high-strength concretes and to separate the total shrinkage into its autogenous and drying shrinkage components, and it is called CEB MC90-99. While the revised models for the drying shrinkage component and for the compliance are closely related to the approach in CEB MC90 (MUller and Hilsdorf 1990, CEB 1993), for autogenous shrinkage, new relations were derived, and some adjustments were included for both normal- and high-strength concrete. For these reasons, the CEB 1990 and the revised CEB 1999 models are described in Appendix A. Some engineers working on creep and shrinkage-sensitive structures have accepted this model as preferable to the ACI 209R-92 model (based on the 1971 Branson and Christiason model). The CEB models do not require any information regarding the duration
" co
..
>< "Ii
'" "'CI
.s to
:;
u
iii
CJ
1500 .-----,----,----.,---..,---.,..--"""?I
I:> fom only CV = 32%
• All data CV = 25%
1250 1---+----t----f------i----,.t"-------1 1000
...
750 500 250 0-=--"------1.----'----'---...... ----'
o
250 500 750 1000 1250 1500
Measured Esh x 10"
Fig. 4.5-CEB MC90-99 versus RlLEM shrinkage databank (Gardner 2004).
300 r---------.,..----z--.-..,...---.,.
I:> fom only CV = 37%
••
• All data CV = 29%
250 I---+---+---+---.~+----.-if-----i ii'
Do
~ 200
1:>1:>
CEB MC90-99
o ~--~----~----~--~----~--~
o
50 100 150 200 250 300
Measured J(t,to) x 10" (lIMPa)
Fig. 4.6-CEB MC90-99 versus RlLEM compliance databank (Gardner 2004). of curing or curing condition. The duration of drying might have a direct impact on the shrinkage and creep of concrete, and should not be ignored when predicting the shrinkage and compliance. The correction term used for relative humidity in the creep equation is extremely se!lsitive to any variation in relative humidity. Figures 4.5 and 4.6 compare the calculated and measured shrinkages and compliances, respectively. The method requires: • Age of concrete when drying starts, usually taken as the age at the end of moist curing; • Age of concrete at loading; • Concrete mean compressive strength at ~8 da)!s; Relative humidity expressed as a decimal; • Volume-surface ratio; and • Cement type.

Page 11
MODELING AND CALCULATING SHRINKAGE AND CREEP IN HARDENED CONCRETE
209.2R-11 GL2000
a ~--~----~----~----~----~--~
a
250 500 750 1000 1250 1500
Measured &"h x 10'"
Fig. 4.7-GL2000 versus RILEM shrinkage databank (Gardner 2004). Using only the data with reported concrete strength, the model generally underestimates the shrinkage of North American concretes, and substantially underestimates the shrinkage of concretes containing basalt aggregates found in Hawaii, Australia, and New Zealand (McDonald 1990; McDonald and Roper 1993; Robertson 2000). The main reason is that primarily European concretes (lower cement content and other types of cement) were considered when optimizing the model. The shrinkage model does not respond well to early-age extrapolation using the simple linear regression method suggested by Bazant (1987); however, the creep model does (Robertson 2000). 4.4-GL2000 model The GL2000 model was developed by Gardner and Lockman (2001), with minor modifications introduced by Gardner (2004). The model is a modification of the GZ Atlanta 97 model (Gardner 2000) made to conform to the ACI 209 model guidelines given in Section 3.5. Except for the concrete compressive strength, the model only requires input data that are available to engineer at time of design. Figure 4.7 and 4.8 compare the calculated and measured shrinkages and compliances, respectively. The method requires: Age of concrete when drying starts, usually taken as the age at the end of moist curing; Age of concrete at loading; Relative humidity expressed as a decimal; Volume-surface ratio; • Cement type; and • Concrete mean compressive strength at 28 days. 4.5-Statistical comparisons As stated previously, there is no agreement as to which statistical indicator(s) should be used, which data sets should be used, or what input data should be considered. To avoid revising any investigator's results, the statistical comparisons of
'i'
~
)(
300 ,.---r---r---.,.----,----,---."
" fem only CV = 26%
• All data CV = 22%
250 1----,.---,.---+---+------.,----.,jL---~ 200 1----+---+-------:.L-----',-O-7i"-''-----,rt''-+---~
! 150 I----+---~
::;
"\:J
S
~
100 I----+_,
::I U
ii
t)
50
o ~____ ~__ ~____ ~____ ~____ ~____ _J o
50 100 150 200 250 300
Measured J(t,to) x 10'" (l/MPa)
Fig. 4.8-GL2000 versus RILEM compliance databank (Gardner 2004). BaZant and Baweja (2000), AI-Manaseer and Lam (2005), and Gardner (2004) are summarized in Table 4.2 for shrinkage and in Table 4.3 for compliance. As the statistical indicators represent different quantities and the investigators used different experimental results, comparisons can only be made across a row, but cannot be made between lines in the tables. Descriptions of the statistical indicators are given in Appendix B. AI-Manaseer and Lam (2005) noted that careful selection and interpretation of concrete data and the statistical methods can influence the conclusions on the performance of model prediction on creep and shrinkage. Brooks (2005) also reported the accuracy of five prediction models, including ACI 209R-92, Bazant-Baweja B3, CEB MC90, and GL2000 models, in estimating 30-year deformation, concluding that most methods fail to recognize the influence of strength of concrete and type of aggregate on creep coefficient, which ranged from 1.2 to 9.2. Brooks (2005) also reported that shrinkage ranged from 280 to 1460 x 1O---{i, and swelling varied from 25 to 35% of shrinkage after 30 years. 4.6-Notes about models The prediction capabilities of the four shrinkage and compliance models were evaluated by comparing calculated results with the RlLEM databank. For shrinkage strain prediction, Bazant-Baweja B3 and GL2000 provide the best results. The CEB MC90-99 underestimates the shrinkage. For compliance, GL2000, CEB MC90-99, and Bazant- Baweja B3 give acceptable predictions. The ACI 209R-92 method underestimates compliance for the most of the RlLEM databank. It should be noted that for shrinkage predictions, Bazant-Baweja B3 using Eq. (4-1) instead of experimental values for water, cement, and aggregate masses provides less accurate, but still acceptable, results. Except for ACI 209R-92, using more inf~rmation improved the prediction for all other methods. The predictions from the CEB, GL2000, and Bazant-Baweja B3 models were signifi- cantly improved by using measured strength development

Page 12
209.2R-12 ACI COMMITTEE REPORT
Table 4.2-Statlstical indicators for shrinkage
Indi- ACI Investigator cator 209R-92 BaZant and Baweja

TJJBP (2000) VCEB

AI- FCEB

Manaseer and Lam MCEBt (2005)

TJJBP Gardner (2004),

fcm only
roo
Gardner (2004),

roo
all data 'Perfect correlation = 0%. tPerfect correlation = 1.00. 55% 46% 83% 1.22 102% 34% 41% Model BaZant- CEB CEB BawejaB3 MC90 MC90-99 GL2000 34% 46% - - 41% 52% 37% 37% 84% 60% 65% 84% 1.07 0.75 0.99 1.26 55% 90% 48% 46% 31%
-
32% 25% 20%
-
25% 19%
Table 4.3-5tatistical Indicators for compliance
Indi- ACI Investigator cator 209R-92 BaZantand Baweja
· 58%
(2000), TJJBP basic creep BaZantand Baweja

(2000), TJJBP 45% drying creep VCEB · 48% AI- FCEB · 32% Manaseer and Lam MCEBt 0.86 (2005)
· 87%
TJJBP Gardner (2004),
• 30%
roo
fcmonly Gardner (2004),
• 30%
all data
roo
'Perfect correlation = 0%. tPerfect correlation = 1.00. Model BaZant- CEB CEB BawejaB3 MC90 MC90-99 GL2000 24% 35%
- -
23% 32%
-
-
36% 36% 38% 35% 35% 31% 32% 34% 0.93 0.92 0.89 0.92 61% 75% 80% 47% 29%
-
37% 26% 27%
-
29% 22%
and measured elastic modulus of the concrete to modify the concrete strength used in creep and shrinkage equations. It should be noted that the accuracy of the models is
limi~ by the many variables outlined previously and
measurement variability. For design purposes, the accuracy of the prediction of shrinkage calculated using GL2000 and Bazant-Baweja B3 models may be within ��20%, and the prediction of compliance ��30%. Parametric studies should be made by the designer to ensure that expected production variations in concrete composition, strength, or the environ- ment do not cause significant changes in structural response. The coefficients of variation for shrinkage measured by BaZant et al. (1987) in a statistically significant investigation were 10% at 7 days and 7% at 1100 days, and can be used as a benchmark for variations between batches. A model that could predict the shrinkage within 15% would be excellent, and 20% would be adequate. For compliance, the range of expected agreement would be wider because, experimen- tally, compliance is determined by subtracting two measured quantities of similar magnitude. . There is not an accepted sign convention for stress and strain. In this document, shortening strains and compressive stresses are positive. For all models, it is necessary to estimate the environmental humidity. The Precast/Prestressed Concrete Institute's PCI Design Handbook (2005) gives values of the annual average ambient relative humidity throughout the United States and Canada that may be used as a guide. Care should be taken when considering structures, such as swimming pools or structures near water. Although the models are not sensitive to minor changes in input values, the effect of air conditioning in moist climates and exposure to enclosed pool in dry climates can be significant. Therefore, the effects of air conditioning and heating on the local envi- ronment around the concrete element should be considered. Relaxation, the gradual reduction of stress with time under sustained strain, calculated using ACI 209R-92, BaZant- Baweja B3, CBB MC90-99, and GL2000, agreed with Rostasy et al.' s (1972) experimental results indicating that the principle of superposition can be used to calculate relaxation provided that calculations are done keeping any drying before loading term constant at the initial value (Lockman 2000). Lockman (2000) did a parametric comparison of models based upon the work of Chiorino and Lacidogna (1998a,b); see also Chiorino (2005). CBB MC90 and ACI 209R-92 underestimate the compliance compared with the GL2000 and BaZant-Baweja B3 models using the same input param- eters. Relaxations calculated by BaZant-Baweja B3 are significantly different than those calculated for the three other models. The elastic strains, calculated at 30 seconds after loading, for the BaZant-Baweja B3 model are very different from those calculated by the other three models. The method of calculating the elastic strain is unique to this model, and the initial stresses of relaxation differ radically from other models . For all ages of loading, especially in a drying environment, BaZant-Baweja B3 predicts more relaxation than the other models. Unlike the other models, BaZant-Baweja B3 uses an asymptotic elastic modulus (fast rate of loading), and not the conventional elastic modulus, which typically includes a significant early-age creep portion. The use of a larger asymptotic elastic modulus explains the comments about relaxation curves obtained from the BaZant-Baweja B3 model. For early ages of loading, the relaxations calculated using CBB MC90-99 and ACI 209R-92 are nearly 100% of the initial stress, with residual stresses close to zero. For creep recovery, GL2000 and Bazant-Baweja B3 are the only models that predict realistic recoveries by super- position. For partial creep recovery, that is, superposition not assumed, with complete removal of the load, no model provides realistic results. Calculating recovery by superpoltition is subject to more problems than calculating relaxation by superposition. If recovery is to be calculated by superposition, both basic and drying creep compliance functions have to be

Page 13
MODELING AND CALCULATING SHRINKAGE AND CREEP IN HARDENED CONCRETE
209.2R-13
parallel in time to give a constant compliance after unloading. As drying before loading reduces both basic and drying creep, it is not yet possible to determine a formulation that permits calculating recovery by superposition in a drying environment. Experimental evidence (Neville 1960) is inconclusive on whether either drying creep or basic creep is completely recoverable. High-strength concretes with water-cement ratios less than 0.40 and mean concrete strengths greater than 80 MPa (11,600 psi) experience significant autogenous shrinkage. The magnitude of the autogenous shrinkage also depends on the availability of moisture during early-age curing. Concretes containing silica fume appear to behave differently from conventional concretes. Few data on such concretes are held in the databank and hence, caution should be exercised using equations justified by the databank for such concretes. The models, however, can be used in such circumstances if they are calibrated with test data. CHAPTER 5-REFERENCES 5.1-Referenced standards and reports The latest editions of the standards and reports listed below were used when this document was prepared. Because these documents are revised frequently, the reader is advised to review the latest editions for any changes. American Concrete Institute 116R Cement and Concrete Terminology 209.1R Report on Factors Affecting Shrinkage and Creep of Hardened Concrete ASTM International C150 Specification for Portland Cement C595 Specification for Blended Hydraulic Cements C157 Test Method for Length Change of Hardened Hydraulic Cement, Mortar, and Concrete C512 Test Method for Creep of Concrete in Compression C469 Test Method for Static Modulus of Elasticity and Poisson's Ratio of Concrete in Compression 5.2-Clted references ACI Committee 209, 1971, "Prediction of Creep, Shrinkage and Temperature Effects in Concrete Structures," Designing for the Effects of Creep, Shrinkage and Temperature, SP-27, American Concrete Institute, Farmington Hills, MI, pp. 51-93. ACI Committee 209, 1982, "Prediction of Creep, , Shrinkage and Temperature Effects in Concrete Structures," Designing for Creep and Shrinkage in Concrete Structures, A Tribute to Adrian Pauw, SP-76, American Concrete Insti- tute, Farmington Hills, MI, pp. 193-300. ACI Committee 209, 1992, "Prediction of Creep, Shrinkage, and Temperature Effects in Concrete Structures (ACI 209R-92)," American Concrete Institute, Farmington Hills, MI, 47 pp. ACI Committee 211, 1991, "Standard Practice for Selecting Proportions for Normal, Heavyweight, and Mass Concrete (ACI 211.1-91) (Reapproved 2002)," American Concrete Institute, Farmington Hills, MI, 38 pp. ACI Committee 318, 2005, "Building Code Requirements for Structural Concrete (ACI 318-05) and Commentary (318R-05)," American Concrete Institute, Farmington Hills, MI,430pp. ACI Committee 363, 1992, "Report on High Strength Concrete "(ACI 363R-92)," American Concrete Institute, Farmington Hills, MI, 55 pp. Acker, P., 1993, "Creep Tests of Concrete: Why and How?" Creep and Shrinkage of Concrete, Proceedings of the Fifth International RILEM Symposium, E&FN Span, London, (ij(,pp.3-14. Acker, P.; BaZant, Z. P.; Chern, 1. c.; Huet, C.; and Wittman, F. H., 1998, RILEM Recommendation on "Measurement of Time-Dependent Strains of Concrete," Materials and Structures, V. 31, No. 212, pp. 507-512. Aguilar, C., 2005, "Study of the Behavior and Develop- ment of a Prediction Methodology for Drying Shrinkage of Concretes," PhD thesis, School of Engineering, Vniversidad Cat6lica de Chile, Santiago, Chile. AI-Manaseer, A.; Espion, B.; and Vim, F. J., 1999, "Conclusions: ACI Paris Chapter Workshop on Creep and Shrinkage in Concrete Structures," Revue Fran~aise de Genie Civil, V. 3, No. 3-4, pp. 15-19. AI-Manasser, A., and Lakshmikantan, S., 1999, "Comparison between Currents and Future Design Codes Models for Creep and Shrinkage," Revue Fran~aise de Genie Civil, special issue: Creep and Shrinkage of Concrete, pp.35-59. AI-Manaseer, A., and Lam, J. P., 2005, "Statistical Evalu- ation of Shrinkage and Creep Models," ACI Materials Journal, V. 102, No.3, May-June, pp. 170-176. BaZant, Z. P., 1975, ''Theory of Creep and Shrinkage in Concrete Structures: a Precis of Recent Developments," Mechanics Today, V. 2, Pergamon Press, 1975, pp. 1-93. BaZant, Z. P., 1987, "Statistical Extrapolation of Shrinkage Data-Part I: Regression," ACI Materials Journal, V. 84, No. I, Jan.-Feb., pp. 20-34. BaZant, Z. P., 1999, "Criteria for Rational Prediction of Creep and Shrinkage of Concrete," Revue Fran~aise de Genie Civil, V. 3, No. 3-4, pp. 61-89. BaZant, Z. P., 2000, "Criteria for Rational Prediction of Creep and Shrinkage of Concrete," The Adam Neville Symposium: Creep and Shrinkage-Structural Design Effects, SP-194, A. AI-Manaseer, ed., American Concrete Institute, Farmington Hills, MI, pp. 237-260. BaZant, Z. P., 2007, "Critical Appraisal of Methods of Creep and Shrinkage Analysis of Concrete Structures," Internal Report, Infrastructure Technology Institute of Northwestern University, also presented to ACI Committee 209,11 pp. BaZant, Z. P., and Baweja, S., 1995, "Creep and Shrinkage Prediction Model for Analysis and Design of Concrete Structures-Model B3," Materials and Structures, V. 28, pp. 357-365, 415-430, 488-495. BaZant, Z. P., and Baweja, S., 2000, '\Creep"and Shrinkage Prediction Model for Analysis and Design of Concrete Structures: Model B3," The Adam Neville Symposium: Creep and Shrinkage-Structural Design Effects, SP-194, A.

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209.2R-14 ACI COMMITIEE REPORT
AI-Manaseer, ed., American Concrete Institute, Farmington Hills, MI, pp. 1-83. Bazant, Z. P.; Carreira, D. J.; and Walser, A., 1975, "Creep and Shrinkage in Reactor Containment Shells," Journal of
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pp.2117-2131. Bazant, Z. P., and Ho, B. H., 1984, "Deformation of Progressively Cracking Reinforced Concrete Beams," ACI JOURNAL, Proceedings v. 81, No.3, May-June, pp. 268-278. Bazant, Z. P.; Kim, J.-K.; Panula, L.; and Xi, Y., 1991, "Improved Prediction Model for Time-Dependent Defor- mations of Concrete: Parts 1-6," Materials and Structures,
v. 24, No. 143, pp. 327-345; V. 24, No. 144, pp. 409-421;
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Bazant, Z. P., and Panula, L., 1978, "Practical Prediction of Time Dependent Deformations of Concrete, Parts I-IV,"
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Branson, D. E., 1963, "Instantaneous and Time-Dependent Deflections of Simple and Continuous Reinforced Concrete Beams," Report No.7, Part I, Alabama Highway Research Department, Bureau of Public Roads, Aug., pp. 1-78. Branson, D. E., 1964, "Time-Dependent Effects in Composite Concrete Beams," ACI JOURNAL, Proceedings V. 61, No.2, Feb., pp. 213-230. Branson, D. E., 1968, "Design Procedures for Computing Deflections," ACI JOURNAL, Proceedings V. 65, No.9, Sept., pp. 730-742. Branson, D. E., 1977, Deformation of Concrete Structures, McGraw Hill Book Co., New York. Branson, D. E., and Chen, C. I., 1972, "Design Procedures for Predicting and Evaluating the Time-Dependent Deforma- tion of Reinforced, Partially Prestressed and Fully Prestressed Structures of Different Weight Concrete," Research Report, Civil Engineering Department, University of Iowa, Iowa City, lA, Aug. Branson, D. E., and Christiason, M. L., 1971, "Time Dependent Concrete Properties Related to Design-Strength and Elastic Properties, Creep and Shrinkage," Creep,
Shrinkage and Temperature Effects, SP-27, American
Concrete Institute, Farmington Hills, MI, pp. 257-277. Branson, D. E., and Kripanarayanan, K. M., 1971, "Loss of Prestress, Camber and Deflection of Noncomposite and Composite Prestressed Concrete Structures," PCI Journal, V. 16, No.5, Sept.-Oct., pp. 22-52. Branson, D. E.; Meyers, B. L.; and Kripanarayanan, K. M., 1970, "Loss of Prestress, Camber, and Deflection of Noncomposite and Composite Structures Using Different Weight Concretes," Final Report No. 70-6, Iowa Highway Commission, Aug., pp. 1-229. Branson, D. E., and Ozell, A. M., 1961, "Camber in Prestressed Concrete Beams," ACI JOURNAL, Proceedings V. 57, No. 12, June, pp. 1549-1574. British Standards Institution, 1985, "BS 8110: Part 2: Structural Use of Concrete: Code of Practice for Special Circumstances," BSI, Milton Keynes. Brooks, J. J., 2005, "30-year Creep and Shrinkage of Concrete," Magazine of Concrete Research, V. 57, No.9, Nov., pp. 545-556. Carreira, D. J., and Burg, R. G., 2000, "Testing for Concrete Creep and Shrinkage," The Adam Neville Symposium: Creep
and Shrinkage-Structural Design Effects, SP-194, A. Al-
Manaseer, ed., American Concrete Institute, Farmington Hills, MI, pp. 381-422. Carreira, D. J., and Chu, K. H., 1986, "Time Dependent Cyclic Deflections in RIC Beams," Journal of Structural
Engineering, ASCE, V. 112. No.5, pp. 943-959.
Carreira, D. J., and Walser, A., 1980, "Analysis of Concrete Containments for Nonlinear Strain Gradients,"
Paper 1317, Fifth International Conference on Structural
Mechanics in Reactor Technology, Nov., pp. 77-83. CEB, 1984, "CEB Design Manual on Structural Effects of Time-Dependent Behaviour of Concrete," M. A. Chiorino, P. Napoli, F. Mola, and M. Koprna, eds., CEB Bulletin
d'lnformation No. 142/142 bis, Georgi Publishing Co.,
Saint-Saphorin, Switzerland, 391 pp. (See also: Final Draft,
CEB Bulletin No. 136, 1980).
CEB, 1991, "Evaluation of the Time Dependent Properties of Concrete," Bulletin d'lnformation No. 199, Comite Euro- pean du BetonlFederation Internationale de la Precontrainte, Lausanne, Switzerland, 201 pp. CEB, 1993. "CEB-FIP Model Code 1990," CEB Bulletin
d'lnformation No. 2131214, Comite Euro-International du
Beton, Lausanne, Switzerland, pp. 33-41. CEB, 1999, "Structural Concrete-Textbook on Behaviour, Design and Performance. Updated Knowledge of the CEBI FIP Model Code 1990," fib Bulletin 2, V. 2, Federation Inter- nationale du Beton, Lausanne, Switzerland, pp. 37-52. Chiorino, M. A., 2005, "A Rational Approach to the Analysis of Creep Structural Effects," Shrinkage and Creep of
Concrete, SP-227, N. J. Gardner and J. Weiss, eds., American
Concrete Institute, Farmington Hills, MI, pp. 107-141. Chiorino, M. A., and Lacidogna, G. 1998a, "General Unified Approach for Analysis of Concrete Structures: Design Aids for Different Code-Type Models," Revue
Fran<;aise de Genie Civil, V. 3, No. 3-4, pp. 173-217.
Chiorino, M. A., and Lacidogna, G., 1998b, "General Unified Approach for Creep Code-Type Models," Depart- ment of Structural Engineering, Politecnico di Torino, Turin, Italy, 41 pp. Davies, R. D., 1957, "Some Experiments on the Appli- cability of the Principle of Superposition to the Strain of Concrete Subjected to Changes of Stress, with Particular Reference to Prestressed Concrete," Magazine of Concrete
Research, V. 9, pp. 161-172.

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MODELING AND CALCULATING SHRINKAGE AND CREEP IN HARDENED CONCRETE
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Gamble, B. R., and Parrott, L. J.,1978, "Creep of Concrete in Compression During Drying and Wetting," Magazine of
Concrete Research, V. 30, No. 104, pp. 129-138.
Gardner, N. J., 2000, "Design Provisions for Shrinkage and Creep of Concrete," The Adam Neville Symposium:
Creep and Shrinkage-Structural Design Effects, SP-194, A.
AI-Manaseer, ed., American Concrete Institute, Farmington Hills, MI, pp. 101-134. Gardner, N. J., 2004, "Comparison of Prediction Provisions for Drying Shrinkage and Creep of Normal Strength Concretes," Canadian Journalfor Civil Engineering, V. 31, No.5, Sept.-Oct., pp. 767-775. Gardner, N. J., and Lockman, M. J., 2001, "Design Provisions for Drying Shrinkage and Creep of Normal Strength Concrete," ACI Materials Journal, V. 98, No.2, Mar.-Apr., pp. 159-167. Gardner, N. J., and Tsuruta, H., 2004, "Is Superposition of Creep Strains Valid for Concretes Subjected to Drying Creep?" ACI Materials Journal, V. WI, No.5, Sept.-Oct., pp.409-415. Hansen, T. C., and Mattock, A. H., 1966, "Influence of Size and Shape on the Shrinkage and Creep of Concrete," ACI JOURNAL, Proceedings V. 63, No.2, Feb., pp. 267-290. Hanson, J. A., 1953, "A 10-Year Study of Creep Proper- ties of Concrete," (checked and reviewed by V. Jones and D. McHenry), Concrete lAboratory Repon Sp-38, U.S. Depart- ment of Interior, Bureau of Reclamation, Denver, CO, 14 pp. Hillsdorf, H. K., and Carreira, D. J., 1980, "ACI-CEB Conclusions of the Hubert Rusch Workshop on Creep of Concrete," Concrete International, V. 2, No. II, Nov., p. 77. Jirasek, M., and BaZant, Z. P., 2002, Inelastic Analysis of
Structures, J. Wiley & Sons, London and New York,
Chapters 27 and 28. Keeton, J. R., 1965, "Study on Creep in Concrete,"
Technical Repon No. R333-1, R333-2, R333-3, U.S. Navy
Civil Engineering Laboratory. Kristek, V.; BaZan!, Z. P.; Zich, M.; and Kohoutkova, A., 2006, "Box Girders Box Deflections," Concrete Interna-
tional, V. 23, No. I, Jan., pp. 55-63.
Kristek, V.; Petrik, V.; and Pilhofer, H.-W., 2001, "Creep and Shrinkage Prediction on the Web," Concrete Interna-
tional, V. 28, No. I, Jan., pp. 38-39.
Kuttner, C. H., 1997, "Creep and Shrinkage for Windows:
the Program for the RILEM Databank," Karlsruhe University,
Version 1.0, Weimar, Berlin and Karlsruhe, Germany. Le Camus, B., 1947, "Recherches experimentales sur la ,deformation du heton et du heton arme," Part n, Annales de
l'Institut du Batiment et des Travaux Publics. (in French)
L'Hermite, R.; Mamillan, M.; and Lefevre, C., 1958, "Noveaux Resultats de Recherche sur la Deformation et la Rupture du Beton," Supplement aux Annales de Institut
Technique du Batiment et des Travaux Publics No. 207/208,
p.325. Lockman, M. J., 2000, "Compliance, Relaxation and Creep Recovery of Normal Strength Concrete," MASc thesis, University of Ottawa, ON, Canada, 170 pp. McDonald, D. B., 1990, "Selected Topics on Drying Shrinkage, Wetting Expansion, and Creep of Concrete," PhD thesis, School of Civil and Mining Engineering, Sydney University, Australia. McDonald, D. B., and Roper, H., 1993, "Accuracy of Prediction Models for Shrinkage of Concrete," ACI Materials
Journal, V. 90, No.3, May-June, pp. 265-271.
McHenry, D., 1943. "A New Aspect of Creep in Concrete and its Application to Design," Proceedings, ASTM, V. 43, pp. 1069-1084. Meyers, B. L.; Branson, D. E.; Schumann, C. G., and Chris- tiason, M. L.. 1970, ''The Prediction of Creep and Shrinkage Properties of Concrete," Final Repon No. 70-5, Iowa Highway Commission, Aug., pp. 1-140. Muller, H. S., 1993, "Considerations on the Development of a Database on Creep and Shrinkage Tests," Creep and
Shrinkage of Concrete, Z. P. BaZant and I. Carol, eds.,
Barcelona, Spain, pp. 3-14. Muller, H. S.; BaZan!, Z. P.; and Kuttner, C. H., 1999, "Data Base on Creep and Shrinkage Tests," Rilem Subcom-
mittee 5 Repon RILEM TC 107-CSP, RILEM, Paris, 81 pp.
Muller, H. S., and Hilsdorf, H. K., 1990, "General Task Group 9," CEB Comire Euro-International du Beton, Paris, France, 201 pp. Neville, A. M., 1960, "Recovery of Creep and Observations on the Mechanism of Creep of Concrete." Applied Scientific
Research, V. 9, pp. 71-84.
Neville, A. M., 1973, Properties of Concrete, second edition, Wiley, New York; Third Edition-1981, Pitman, London and Marshfield, 779 pp.; 4th Edition-1995, Longman Group, 844 pp. Neville, A. M., and Dilger, 1970, Creep of Plain and
Structural Concrete, North-Holland, Amsterdam; new
edition: Neville, A. M.; Dilger, W. H.; and Brooks, J. J.,
1983, Creep of Plain and Structural Concrete, Construction
Press, London and New York, 361 pp. PrecastlPrestressed Concrete Institute, 2005, PCI Design
Handbook, sixth edition.
RILEM Technical Committee TC 69, 1988, "Material Models for Structural Creep Analysis," (principal author Z. P. BaZant) Chapter 2 in Mathematical Modeling of Creep
and Shrinkage of Concrete, Z. P. BaZant, ed., J. Wiley,
Chichester & New York, pp. 99-215. RILEM Technical Committee TC 107, 1995, "Guidelines for Characterising Concrete Creep and Shrinkage in Structural Design Codes or Recommendations," Materials and
Structures, V. 28, pp. 52-55.
Robertson .. I. N., 2000, "Correlation of Creep and Shrinkage Models with Field Observations," The Adam
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Institute, Farmington Hills, MI, pp. 261-282. Ross, A. D., 1958, "Creep of Concrete under Variable Stress," ACI JOURNAL, Proceedings V. 54, pp. 739-758. Rostasy, F. S.; Teichen, K. T.; and Engelke, H., 1972,
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ACI COMMITTEE REPORT
Sakata, K., 1993, "Prediction of Concrete Creep and Shrinkage," Proceedings of 5th International RILEM Symposium (Concreep5), Barcelona, Spain, pp. 649-654. Sakata, K.; Tsubaki, T.; Inoue, S.; and Ayano, T., 2001, "Prediction Equations of Creep and Drying Shrinkage for Wide-Ranged Strength Concrete," Proceedings of 6th Inter- national Conference CONCREEP-6@MIT, pp. 753-758. Sassone, M., and Chiorino, M. A., 2005, "Design Aids for the Evaluation of Creep Induced Structural Effects", Shrinkage and Creep of Concrete, SP-227, D. J. Gardner and
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APPENDIX A-MODELS A.1-ACI209R·92 model
This is an empirical model developed by Branson and Christiason (1971), with minor modifications introduced in ACI 209R-82 (ACI Committee 2091982). ACI Committee 209 incorporated the developed model in ACI 209R-92 (ACI Committee 209 1992). The models for predicting creep and shrinkage strains as a function of time have the same principle: a hyperbolic curve that tends to an asymptotic value called the ultimate value. The form of these equations is thought to be convenient for design purposes, in which the concept of the ultimate (in time) value is modified by the time-ratio (time-dependent development) to yield the desired result. The shape of the curve and ultimate value depend on several factors, such as curing conditions, age at application of load, mixture propor- tioning, al.l1bient temperature, and humidity. The design approach presented for predicting creep and shrinkage refers to standard conditions and correction factors for other-than-standard conditions. The correction factors are applied to ultimate values. Because creep and shrinkage equations for any period are linear functions of the ultimate values, however, the correction factors in this procedure may be applied to short-term creep and shrinkage as well. The recommended equations for predicting a creep coefficient and an unrestrained shrinkage strain at any time, including ultimate values, apply to normal weight, sand lightweight, and all lightweight concrete (using both moist and steam curing, and Types I and III cement) under the standard conditions summarized in Table A.l. Required parameters: Age of concrete when drying starts, usually taken as the age at the end of moist curing (days); Age of concrete at loading (days); Curing method; Ambient relative humidity expressed as a decimal; Volume-surface ratio or average thickness (mm or in.); Concrete slump (rum or in.); Fine aggregate percentage (%); Cement content (kg/m3 or Ib/yd3); Air content of the concrete expressed in percent (%); and ... Cement type
A.I.1 Shrinkage-The shrinkage strain ssh(t,tc) at age of
concrete t (days), measured from the start of drying at tc (days), is calculated by Eq. (A-I) (A-I) where f (in days) and a are considered constants for a given member shape and size that define the time-ratio part, sshu is the ultimate shrinkage strain, and (t - tc) is the time from the end of the initial curing. For the standard conditions, in the absence of specific shrinkage data for local aggregates and conditions and at ambient relative humidity of 40%, the average value suggested for the ultimate shrinkage strain £shw is £shu = 780 x 10-
6
mmfmm (in.!in.) (A-2) For the time-ratio in Eq. (A-I), ACI 209R-92 recommends an average value for f of 35 and 55 for 7 days of moist curing and 1 to 3 days of steam cUling, respectively, while an average value of 1.0 is suggested for a (flatter hyperbolic form). It should be noted that the time-ratio does not differentiate between drying, autogenous, and carbonation shrinkage. Also, it is independent of member shape and size, because f and a are considered as constant. The shape and size effect can be totally considered on the time-ratio by replacing a = 1.0, andfas given by Eq. (A-3), in Eq. (A-I), where VIS is the volume-surface ratio in mm or in.

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MODELING AND CALCULATING SHRINKAGE AND CREEP IN HARDENED CONCRETE 209.2R-17
Table A.1-Factors affecting concrete creep and shrinkage and variables considered in recommended prediction method
Factors Variables considered Standard conditions Cement paste content Type of cement Type I and III Water-cement ratio Slump 70 mm (2.7 in.) Air content :5:6% Concrete composition Mixture proportions Fine aggregate percentage 50% Aggregate characteristics 279 to 446 kg/m3 Cement content Concrete Degrees of compaction (470 to 752 Ib/yd3
)
(creep and shrinkage) Moist cured 7 days Length of initial curing Steam cured 1 to 3 days Initial curing Moist cured 23.2 ��2 ��C Curing temperature (73.4�� 4 oF) Steam cured :5: 100 ��C (:5: 212 oF) Curing humidity Relative humidity
~95%
Concrete temperature Concrete temperature 23.2 �� 2 ��C Environment (73.4�� 4 oF) Member geometry and Concrete water content Ambient relative humidity 40% environment (creep and shrinkage) Volume-surface ratio
VIS = 38 mm (1.5 in.)
Geometry Size and shape or minimum thickness 150 mm (6 in.) Moist cured 7 days Concrete age at load application Steam cured I to 3 days Loading history During of loading period Sustained load Sustained load Loading (creep only) Duration of unloading period
-
-
Number of load cycles
-
-
Type of stress and distribution Compressive stress Axial compression Stress conditions across the section Stress/strength ratio Stress/strength ratio :5:0.50
-2
f = 26.0e{1.42 x 10 (VIS)}
in SI units (A-3) The ambient relative humidity coefficient Ysh,RH is
f = 26.0e{O.36(VIS)}
in in.-Ib units (A-7) For conditions other than the standard conditions, the average value of the ultimate shrinkage Eshu (Eq. (A-2)) needs to be modified by correction factors. As shown in Eq. (A-4) and (A-5), ACI 209R-92 (ACI Committee 209 1992) suggests multiplying Eshu by seven factors, depending on particular conditions _ {1.40 - 1.02h for 0.40 S h S 0.80
Ysh, RH -
3.00 _ 3.0h for 0.80 S h S 1 where the relative humidity h is in decimals. For lower than 40% ambient relative humidity, values higher than 1.0 should be used for shrinkage Y sh,RH' Because
Ysh,RH= 0 when h = 100%, the ACI method does not predict
swelling.
Eshu = 780ysh x 10-6
mmlmm (in.lin.)
(A-4) with
Ysh = Ysh,teYsh,RHYsh,vsYsh,sYsh,IjIYsh,eYsh,a
(A-5) where Y sh represents the cumulative product of the applicable correction factors as defined as follows. The initial moist curing coefficient Y sh,te for curing times different from 7 days for moist-cured concrete, is given in Table A.2 or Eq. (A-6); for steam curing with a period of 1 to 3 days, Ysh,te = 1. The Ysh,ep correction factors shown in Table A.2 for the initial moist curing duration variable can be obtained by linear regression analysis as given in Eq. (A-6)
Ysh,te = 1.202 - 0.233710g(te)
R2 = 0.9987
(A-6) Coefficient Y sh, vs allows for the size of the member in terms of the volume-surface ratio, for members with volume-surface ratio other than 38 mm (1.5 in.), or average thickness other than 150 mm (6 in.). The average thickness d of a member is defmed as four times the volume-surface ratio; that is d = 4V/S, which coincides with twice the actual thickness in the case of a slab - 1 2 {-O.00472(VIS)}
Ysh, vs - . e
- 1 2 {-O.12(VIS)}
Ysh, vs - . e
in SI units (A-8) in in.-Ib units where V is the specimen volume in mm
3
or in.3, and S the specimen surface area in mm
2
or in2. Alternatively, the method also allows the use of the average-thickness method to account for the effect of member size on Eshu' The average-thickness method tends to compute

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ACI COMMITTEE REPORT
Table A.2-Shrinkage correction factors for initial moist curing, 'Y sh,te' for use in Eq. (A-5), ACI 209R-92 model
Moist curing duration tc' days
Y5h,Ie
1 1.2 3 1.1 7 1.0 14 0.93 28 0.86 90 0.75
correction factor values that are higher, as compared with the volume-surface ratio method. For average thickness of member less than 150 mm (6 in.) or volume-surface ratio less than 37.5 mm (1.5 in.), use the factors given in Table A.3. For average thickness of members greater than 150 mm
(6 in.) and up to about 300 to 380 mm (12 to 15 in.), use
Eq. (A-9) and (A-IO). During the first year drying, (t - tc) $ 1 year
Ysh,d = 1.23 - 0.0015d Ysh,d = 1.23-0.006(VIS)
Y sh, d = 1.23 - 0.038d
Ysh,d = 1.23 - 0.152(V IS)
For ultimate values, (t - tc) > 1 year
Ysh,d = 1.17-0.00114d
Y sh, d = 1.17 - 0.00456( VIS) Y sh, d = 1.17 - 0.029d
Ysh,d = 1.23-0.116(VIS)
in SI units in in.-Ib units in SI units in in.-Ib units (A-9) (A-IO) where d = 4V/S is the average thickness (in mm or in.) of the part of the member under consideration. For either method, however, Ysh should not be taken less than 02. Also, use YshEshu ~ 100 x 10--6 mmlmm (in.lin.) if concrete is under seasonal wetting and drying cycles and
YshEshu ~ 150 x 10--6 mmlmm (in.lin.) if concrete is under
sustained drying conditions. The correction factors that allow for the composition of the concrete are:
• Slump factor Ysh,s' where s is the slump of fresh
concrete (mm or in.)
Ysh,s = 0.89 + 0.00161s in SI units Ysh,s = 0.89 + 0.041s in in.-Ib units
(A-ll)
Table A.3-Shrinkage correction factors for average thickness of members, 'Ysh,d, for use in Eq. (A-5), ACI 209R-92 model



Average thickness of Volume/surface ratio VIS, Shrinkage factor Y slr,d member d, mm (in.) mm (in.) 51 (2) 12.5 (0.50) 1.35 76 (3) 19 (0.75) 1.25 102 (4) 25 (1.00) 1.17 127 (5) 31 (1.25) 1.08 152 (6) 37.5 (1.50) 1.00
Fine aggregate factor Ysh,IjI' where", is the ratio of fme aggregate to total aggregate by weight expressed as percentage
Ysh,\jJ = 0.30+0.014", for",$50% Ysh,\jJ = 0.90+0.002", for",>50%
(A-12) Cement content factor ~h,c> where c is the cement content in kg/m
3
or lb/yd
Ysh,e = 0.75 + 0.OOO61c in SI units
Y sh, e = 0.75 + 0.OOO36c in in.-Ib units (A-13)
Air content factor Ysh,a' where a is the air content in percent
Ysh,a = 0.95 + 0.OO8a ~ 1
(A-14) These correction factors for concrete composition should be used only in connection with the average values suggested for Eshu = 780 x 10--6 mmlmm (in.lin.). This average value for Eshu should be used only in the absence of specific shrinkage data for local aggregates and conditions determined in accordance with ASTM C512.
A.l.2 Compliance-The compliance function J(t,to) that
represents the total stress-dependent strain by unit stress is given by (A-15) where Ecmto is the modulus of elasticity at the time of loading to (MPa or psi), and eIl(t,to) is the creep coefficient as the ratio of the creep strain to the elastic strain at the start of loading at the age to (days). a) Modulus of elasticity-The secant modulus of elasticity of concrete Ecmto at any time to of loading is given by
Emero = 0.043y!·5 Jfemto (MPa) in SI units Emeto = 33y;.5 Jfemto (psi) in in.-Ib units
(A-16)
".
where Y c is the unit weight of concrete (kg/m
3
or Ib/ft3), and
fcmto is the mean concrete compressive strength at the time of
loading (MPa or psi).

Page 19
MODELING AND CALCULATING SHRINKAGE AND CREEP IN HARDENED CONCRETE 209.2R-19
The general equation for predicting compressive strength at any time t is given by
fcmt = [a: btJfcm28
(A-17) where fcm28 is the concrete mean compressive strength at 28 days in MPa or psi, a (in days) and b are constants, and t is the age of the concrete. The ratio alb is the age of concrete in days at which one half of the ultimate (in time) compressive strength of concrete is reached. The constants a and b are functions of both the type of cement used and the type of curing employed. The ranges of
a and b for the normal weight, sand lightweight, and all light-
weight concretes (using both moist and steam curing, and Types I and III cement) are: a = 0.05 to 9.25, and b = 0.67 to 0.98. Typical recommended values are given in Table A.4. The concrete required mean compressive strength fcm28 should exceed the specified compressive strengthf~ as required in Section 5.3.2 of ACI 318 (ACI Committee 3182005). b) Creep coefficient-The creep model proposed by ACI 209R-92 has two components that determine the asymptotic value and the time development of creep. The predicted parameter is not creep strain, but creep coefficient <I>(t,to) (defined as the ratio of creep strain to initial strain). The latter allows for the calculation of a creep value independent from the applied load. Equation (A-18) presents the general model (A-I8) where <I>(t,to) is the creep coefficient at concrete age t due to a load applied at the age to; d (in days) and", are considered constants for a given member shape and size that define the time-ratio part; (t - to) is the time since application of load, and <l>u is the ultimate creep coefficient. For the standard conditions, in the absence of specific creep data for local aggregates and conditions, the average value proposed for the ultimate creep coefficient <l>u is
<l>u = 2.35
(A-I9) For the time-ratio in Eq. (A-18), ACI-209R-92 recom- mends an average value of 10 and 0.6 for d and", (steeper curve for larger values of (t - to»' respectively. The shape and size effect can be totally considered on the time-ratio by replacing", = 1.0 and d = f as given by Eq. (A- . 3), in Eq. (A-I8), where VIS is the volume-surface ratio in mm or in. For conditions other than the standard conditions, the value of the ultimate creep coefficient <l>u (Eq. (A-19» needs to be modified by correction factors. As shown in Eq. (A-20) and (A-21), ACI 209R-92 suggests multiplying <l>u by six factors, depending on particular conditions.
<l>u= 2.35yc
(A-20)
Yc = Yc.loYc,RHYc,vsYc,sYc,IjIYsh,a
(A-21) Table A.4-Values of the constant a and b for use in Eq. (A-17), ACI 209R-92 model
Type of Moist-cured concrete Steam-cured concrete cement
a
b
a
b
I
4.0 0.85 1.0 0.95 III 2.3 0.92 0.70 0.98
where Yc represent the cumulative product of the applicable correction factors as defined as follows. For ages at application of load greater than 7 days for moist- cured concrete or greater than I to 3 days for steam-cured concrete, the age of loading factor for creep Y c,lo is estimated from
Y = 1 25t -D. lIS for moist curing
C,lo .
0
(A-22)
Y
= 1 13t -0.094 for steam curing
c,lo .
0
(A-23) where to is the age of concrete at loading (days). The ambient relative humidity factor Yc,RH is
Yc,RH = 1.27 - 0.67h for h ~ 0.40
(A-24) where the relative humidity h is in decimals. For lower than 40% ambient relative humidity, values higher than 1.0 should be used for creep Yh' Coefficient Y c, vs allows for the size of the member in terms of the volume-surface ratio, for members with a volume- surface ratio other than 38 mm (1.5 in.), or an average thickness other than 150 mm (6 in.)
= ~(1 I 13 {-O.0213(VIS)})
Yc vs
+ .. e , 3
in SI units (A-25)
= ~(I 1 13 {-O.S4(V IS)})
Yc vs
+ .. e , 3
in in.-Ib units where V is the specimen volume in mm
3
or in
3 , and S the
specimen surface area in mm2 or in2
.
Alternatively, the method also allows the use of the average-thickness method to account for the effect of member size on <l>u' The average-thickness method tends to compute correction factor values that are higher, as compared with the volume-surface ratio method. For the average thickness of a member less than 150 mm (6 in.) or volume-surface ratio less than 37.5 mm (1.5 in.), use the factors given in Table A.5. For the average thickness of members greater than 150 mm (6 in.) and up to about 300 to 380 mm (12 to 15 in.), use Eq. (A-26) and (A-27). During the first year after loading, (t - to) :=; 1 year
Y c, d = 1.14 - 0.00092d Yc,d = 1.14 - 0.00363(V IS) Yc,d= 1.14-0.023d Yc,d = 1.14-0.092(V/S)
in SI units
".
in in.-Ib units (A-26)

Page 20
209.2R·20
ACI COMMITTEE REPORT
Table A.5-Creep correction factors for average thickness of members, Yc,d, for use in Eq. (A-21), ACI 209R-92 model
Average thickness of Volume/surface ratio VIS, member do mm (in.) mm (in.) 51 (2) 12.5 (0.50) 76 (3) 19 (0.75) 102(4) 25 (1.00) 127 (5) 31 (1.25) 152 (6) 37.5 (1.50)
For ultimate values, (t - to) > 1 year
Ycod :::: 1.10 - 0.00067 d Yc,d:::: 1.10-0.00268(VIS) Ycod:::: 1.10-0.017d Yc.d :::: 1.10 - 0.068(V IS)
Creep factor Y c.d 1.30 1.17 1.11 1.04 1.00
in SI units (A-27) in in.-Ib units where d:::: 4(VIS) is the average thickness in mm or inches of the part of the member under consideration. The correction factors to allow for the composition of the concrete are: • Slump factor Ye,s' where s is the slump of fresh concrete (mm orin.)

Y c, s = 0.82 + 0.00264s in SI units Y c. s :::: 0.82 + 0.067 s in in.-Ib units
(A-28) Fine aggregate factor Ye,,!,, where 1jI is the ratio of fine aggregate to total aggregate by weight expressed as percentage Ye.'!' = 0.88 + 0.00241j1 (A-29) Air content factor Ye,a' where a is the air content in percent Ye,a = 0.46 + 0.09a ;::: 1 (A-30) These correction factors for concrete composition should be used only in connection with the average values suggested for cJ>u:::: 2.35. This average value for cJ>u should be used only in the absence of specific creep data for local aggregates and conditions determined in accordance with ASTMC512. A.2-Baiant-Baweja B3 model The Bazant-Baweja (1995) B3 model is the latest variant in a number of shrinkage and creep prediction methods developed by Bazant and his coworkers at Northwestern University. According to Bazant and Baweja (2000), the B3 model is simpler and is better theoretically justified than the previous models. The effect of concrete composition and design strength on the model parameters is the main source of error of the model. The prediction of the material parameters of the B3 model from strength and composition is restricted to portland cement concrete with the following parameter ranges: 0.35 ~ w/e ~ 0.85;
• 2.55: ale '5. 13.5;
• 17 MPa ~ fem28 ~ 70 MPa (2500 psi ~ fem28 ~ 10,000 psi); and • 160 kglm
3 ~ e ~ 720 kglm 3 (270 Ib/yd3 ~ e ~ 1215 Ib/yd3 )
where fem28 is the 28-day standard cylinder compression strength of concrete (in MPa or psi), w/c is the water-cement ratio by weight, e is the cement content (in kg/m
3 or Ib/yd3),
and ale is the aggregate-cement ratio by weight. If only design strength is known, thenfem28 :::: f; + 8.3 MPa (fcm28 =
f; + 1200 psi).
The Bazant-Baweja B3 model is restricted to the service stress range (or up to about 0.45fem28)' The formulas are valid for concretes cured for at least 1 day. Required parameters: Age of concrete when drying starts, usually taken as the age at the end of moist curing, (days); • Age of concrete at loading (days); • Aggregate content in concrete (kg/m
3
or Ib/yd
3 );
Cement content in concrete (kglm
3 or Ib/yd3);
• Water content in concrete (kg/m3 or Ib/yd3
);
• Cement type; • Concrete mean compressive strength at 28 days (MPa or psi); • Modulus of elasticity of concrete at 28 days (MPa or psi); Curing condition; Relative humidity expressed as a decimal; Shape of specimen; and Volume-surface ratio or effective cross-section thickness
(mm orin.).
A.2.t Shrinkage-The mean shrinkage strain Esh(t,te) in
the cross section at age of concrete t (days), measured from the start of drying at tc (days), is calculated by Eq. (A-31) (A-31) where Eshoo is the ultimate shrinkage strain, kh is the humidity dependence factor (Table A.6), Set - tc) is the time curve, and
(t - te) is the time from the end of the initial curing.
The ultimate shrinkage Eshoo is given by Eq. (A-32) (A-32) where Esoo is a constant given by Eq. (A-33), and Eem607/
Ecm(tc+'tsh) is a factor to account for the time dependence of
ultimate shrinkage (Eq. (A-34»
S,OO = -a,a2[O.019w2Ifc~O~2S+ 270] x 10-6 S,OO = -a,a2[0.02565w2Ifc~o;~+ 270] x 10-6
in SI units in in. -lb units
(A-33)

Page 21
MODELING AND CALCULATING SHRINKAGE AND CREEP IN HARDENED CONCRETE 209.2R-21
and E -E (
t )
emt - em28 4 + 0.85t
(A-34) where w is the water content in kglm
3
or Ib/yd
3 ,fem28 is the
concrete mean compressive strength at 28 days in MPa or psi, and 0.1 and 0.2 are constants related to the cement type and curing condition. (Note: The negative sign is the model authors' convention.) The values of 0.1 and 0.2 are given in Tables A.7 and A.8, respectively. This means that &shoo = &soo for te = 7 days, and 'tsh = 600 days. The time function for shrinkage S(t - te) is given by Eq. (A-35) (A-35) where t and te are the age of concrete and the age drying commenced or end of moist curing in days, respectively, and 'tsh is the shrinkage half-time in days as given in Eq. (A-36). The size dependence of shrinkage is given by
T,h = 0.085tc-O,08/cm28 -O,2s[2k,(V IS)]2 in SI units T,h = 190.8tc-O·08/cm28 -O,2S[2k,(V IS)]2 in in.-Ib units
(A-36) where ks is the cross-section shape-correction factor (Table A.9), and VIS is the volume-surface ratio in mm or in. A.2.2 Compliance-The average compliance function
J(t,to) at concrete age t caused by a unit uniaxial constant
stress applied at age to' incorporating instantaneous defor- mation, basic and drying creep, is calculated from (A-37) where ql is the instantaneous strain due to unit stress (inverse of the asymptotic elastic modulus) that is, in theory, approached at a time of about 10-
9
second; Co(t,to) is the compliance function for basic creep; Cj..Mo,te) is the additional compliance function for drying creep; and t, te, and to are the age of concrete, the age drying began or end of moist curing, and the age of concrete loading in days, respectively. The instantaneous strain may be written ql = lIEo, where
Eo is the asymptotic elastic modulus. The use of Eo instead
of the conventional static modulus Ecm is convenient because concrete exhibits pronounced creep, even for very short loads duration. Eo should not be regarded as a real elastic modulus, but merely an empirical parameter that can be considered age independent. Therefore, the instantaneous strain due to unit stress is expressed in Eq. (A-38) ql = 0.61Ecm28 (A-38) where Table A.6-Humidity dependence k", 83 model
Relative humidity
h~0.98
h=1.00 -D.2 0.98 < h < 1.00 Linear interpolation: 12.74 -12.94h
Table A.7-a.1 as function of cement type, 83 model
Type of cement
Type I
1.00
Type II
0.85
Type III
1.10
Table A.~ as function of curing condition, 83 model
Curing method Steam cured 0.75 Cured in water or at 100% relative humidity 1.00 Sealed during curing or normal curing in air with initial protection against drying
Eem28 = 4734 Jfem28 in SI units
Eem28 = 57,OOOJfem28
in in.-Ib units
1.20
(A-39) According to this model, the basic creep is composed of three terms: an aging viscoelastic term, a nonaging viscoelastic term, and an aging flow term where q2Q(t,to) is the aging viscoelastic compliance term. The cement content c (in kglm3 or Ib/yd3
) and the concrete
mean compressive strength at 28 days fem28 (in MPa or psi) are required to calculate the parameter q2 in Eq. (A-41) 185 4 10
-6 0.51" -0.9
q2 =
. x
C J em28
8 4 0
-6 0.5 -0.9
q2 = 86. 1 x 1 c fem28
in SI units (A-41) in in.-Ib units
Q(t,to) is an approximate binomial integral that must be multi-
plied by the parameter q2 to obtain the aging viscoelastic term (A-42) Equations (A-43) to (A-45) can be used to approximate the binomial integral (A-43)
(A-44)
(A-45)

Page 22
209.2R-22
ACI COMMITIEE REPORT
Table A.9-ks as function of cross section shape, B3 model
Cross section shape Infinite slab 1.00 Infinite cylinder 1.15 Infinite square prism 1.25 Sphere 1.30 Cube 1.55
Note: The analyst needs to estimate which of these shapes best approximates the real shape of the member or structure. High accuracy in this respect is not needed, and k, " 1 can be used for simplified analysis.
where m and n are empirical parameters whose value can be taken the same for all normal concretes (m = 0.5 and n = 0.1). In Eq. (A-40), q3 is the non aging viscoelastic compliance parameter, and q4 is the aging flow compliance parameter. These parameters are a function of the concrete mean compressive strength at 28 days f cm28 (in MPa or psi), the
3
I~h
.
cement content c (in kg/m or Ib/yd' ), the water-cement ratIo wlc, and the aggregate-cement ratio ale (A-46)
-6( )-07
q4 = 20.3 x 10 alc .
in SI units (A-47)
-6 -07
q4 = 0.14 x 10 (alc)'
in in.-lb units The compliance function for drying creep is defined by
Eq. (A-48). This equation accounts for the drying before
loading. Note that drying before loading is considered only for drying creep In Eq. (A-48), q5 is the drying creep compliance parameter. This parameter is a function of the concrete mean compressive strength at 28 days fcm28 (in MPa or psi), and of Eshoo' the ultimate shrinkage strain as given in Eq. (A-32) (A-49) HU) and HUo) are spatial averages of pore relative humidity. Equations (A-50) to (A-53) and Eq. (A-36) are required to calculate H(t) and H(to).
H(t) ::: 1 - (1 - h)S(t - tc)
(A-50) (A-51) where S(t - tc) and S(to - tc) are the time function for shrinkage calculated at the age of concrete t and the age of concrete at loading to in days, respectively, and 'tsh is the shrinkage half-time (A-52) [( t - t) 1/2J
S(to - tc) = tanh ~
'tsh
(A-53) A.3-CEB MC90-99 model The CEB MC90 model (Muller and Hilsdorf 1990; CEB 1993) is intended to predict the time-dependent mean cross- section behavior of a concrete member. It has concept similar
to that of ACI209R-92 model in the sense that it gives a hyper-
bolic change with time for creep and shrinkage, and it also uses an ultimate value corrected according mixture propor- tioning and environment conditions. Unless special provi- sions are given, the models for shrinkage and creep predict the time-dependent behavior of ordinary-strength concrete (12 MPa 1I740 psi] ~f; ~ 80 MPa [11,600 psi)) moist cured at normal temperatures not longer than 14 days and exposed to a mean ambient relative humidity in the range of 40 to 100% at mean ambient temperatures from 5 to 30��C (41 to 86 oF). The models are valid for normaIweight plain structural concrete having an average compressive strength in the range of 20 MPa (2900 psi) ~fcm28 ~ 90 MPa (13,000 psi). The age at loading to should be at least 1 day, and the sustained stress should not exceed 40% of the mean concrete strength fcmto at the time of loading to' Special provisions are given for elevated or reduced temperatures and for high stress levels. The CEB MC90-99 model (CEB 1999) includes the latest improvements to the CEB MC90 model. The model has been developed for normal- and high-strength concrete, and considers the separation of the total shrinkage into autogenous and drying shrinkage components. The models for shrinkage and creep are intended to predict the time-dependent mean cross-section behavior of a concrete member moist cured at normal temperatures not longer than 14 days and exposed to a mean ambient relative humidity in the range of 40 to 100% at mean ambient temperatures from 10 to 30��C (50 to 86 OF).
It is valid for normal weight plain structural concrete having
an average compressive strength in the range of 15 MPa (2175 psi) ~fcm2'8 ~ 120 MPa (17,400 psi). The age at loading should be at least 1 day, and the creep-induced stress should not exceed 40% of the concrete strength at the time of loading. The CEB model does not require any information regarding the duration of curing or curing condition, but takes into account the average relative humidity and member size. Required parameters: • Age of concrete when drying starts, usually taken as the age at the end of moist curing (days); • Age of concrete at loading (days); • Concrete mean compressive strength at 28 days (MPa or psi)~ Relative humidity expressed as a decimal; Volume-surface ratio or effective cross-section thickness of the member (mm or in.); and • Cement type. A.3.1 Shrinkage CEB MC90-The total shrinkage strains of concrete Esh(t,tc) may be calculated from
\I'
(A-54)

Page 23
MODELING AND CALCULATING SHRINKAGE AND CREEP IN HARDENED CONCRETE
209.2R-23
where Eeso is the notional shrinkage coefficient. i3s(t - te) is the coefficient describing the development of shrinkage with time of drying, t is the age of concrete (days) at the moment considered, te is the age of concrete at the beginning of drying (days), and (t - te) is the duration of drying (days). The notional shrinkage coefficient may be obtained from with
-6
Eifem28) = [160 + lOi3se(9 - fem28IfemO)] x 10 i3RH(h) = -1.55[1-(~YJ for0.4~h<0.99 i3RH(h) = 0.25 for h ~ 0.99
(A-55) (A-56) (A-57) where fem28 is the mean compressive cylinder strength of concrete at the age of 28 days (MPa or psi), femo is equal to 10 MPa (1450 psi), i3se is a coefficient that depends on the type of cement (Table A. 10), h is the ambient relative humidity as a decimal, and ho is equal to 1. The development of shrinkage with time is given by where (t-te) is the duration of drying (days), tl is equal to 1 day,
VIS is the volume-surface ratio (mm or in.), and (VIS)o is
equal to 50 mm (2 in.). The method assumes that. for curing periods of concrete members not longer than 14 days at normal ambient temperature, the duration of moist curing does not significantly affect shrinkage. Hence, this parameter, as well as the effect of curing temperature, is not taken into account. Therefore, in Eq. (A-54) and (A-58), the actual duration of drying (t- t
e)
has to be used. When constant temperatures above 30��C (86 oF) are applied while the concrete is drying, CEB MC90 recom- mends using an elevated temperature correction for i3dh) and i3s(t - te), shown as follows. The effect of temperature on the notional shrinkage coefficient is taken into account by In SI units: (A-59) In in.-Ib units:
_ [ ( 0.08 )(18.778'TlTo-37.77~] ~RH,T - ~RH(h) I + 1.03 _ hlho 40 -)
Table A.1O-Coefficient i3scaccording to Eq. (A-56), CEe MC90 model
Type of cement according to Ee2 SL (slowly-hardening cements) Nand R (normal or rapid hardening cements) RS (rapid hardening high-strength cements) ~sc 4
5
8
The effect of temperature on the time development of shrinkage is taken into account by In SI units: (A-60) In in.-Ib units: where i3RH T is the relative humidity factor corrected by
temperatur~ that replaces i3RH in Eq. (A-55), i3s,T(t - te) is the
temperature-dependent coefficient replacing i3s(t - te) in Eq. (A-54), h is the relative humidity in decimals, ho is equal to 1, VIS is the volume-surface ratio (mm or in.); (VIS) is equal to 50 mm (2 in.), Tis the ambient temperature eC or oF), and To is equal to 1 ��c (33.8 OF). A.3.2 Shrinkage CEB MC90-99-With respect to the shrinkage characteristics of high-performance concrete, the new approach for shrinkage subdivides the total shrinkage into the components of autogenous shrinkage and drying shrinkage. While the model for the drying shrinkage component is closely related to the approach given in CEB MC90 (CEB 1993), for autogenous shrinkage, new relations had to be derived. Some adjustments, however, should also be carried out for the drying shrinkage component, as the new model should cover both the shrinkage of normal- and high-perfor- mance concrete; consequently, the autogenous shrinkage also needs to be modeled for normal-strength concrete. The total shrinkage of concrete Esh(t,te) can be calculated from Eq. (A-61) (A-61) where Esh(t,te) is the total shrinkage, Eeas(t) the autogenous shrinkage, and Eeds(t,te) is the drying shrinkage at concrete age t (days) after the beginning of drying at te (days). The autogenous shrinkage component Eeas(t) is calculated from Eq. (A-62) (A-62)

Page 24
209.2R-24 ACI COMMmEE REPORT
where Eeaso(fcm28) is the notional autogenous shrinkage coeffi- cient from Eq. (A-63), and f3as(t) is the function describing the time development of autogenous shrinkage from Eq. (A-64)
a ) _ ( fem28
1
femo )2.5 x 10-6 (A-63) EeasoVem28 - -aas 6 +/. If. em28 emo
(A-64) wherefcm28 is the mean compressive strength of concrete at an age of 28 days (MFa or psi),fcmo = 10 MPa (1450 psi), t is the concrete age (days), tl = 1 day, and aas is a coefficient that depends on the type of cement (fable A.II). The autogenous shrinkage component is independent of the ambient humidity and of the member size, and develops more rapidly than drying shrinkage. The drying shrinkage Eeds(t,te) is calculated from Eq. (A-65) (A-65) where Eedso(fcm28) is the notional drying shrinkage coefficient from Eq. (A-66), f3RH<h) is the coefficient that takes into account the effect of relative humidity on drying shrinkage from Eq. (A-67), and f3ds(t - te) is the function describing the time development of drying shrinkage from Eq. (A-68) (A-67)
f3R1l-h) = 0.25 for h ~ 0.99f3s1
(A-69) where adsl and ads2 are coefficients that depend on the type of cement (fable A.II), f3s1 is a coefficient that takes into account the self-desiccation in high-performance concrete, h is the ambient relative humidity as a decimal, ho = I, VIS is the volume-surface ratio (rom or in.), (VIS)o = 50 rom (2 in.),fcmo = 10 MPa (1450 psi), te is the concrete age at the beginning of drying (days), and (t - te) is the duration of drying (days). According to Eq. (A-67) for normal-strength concretes, swelling is to be expected if the concrete is exposed to an ambient relative humidity near 99%. For higher-strength grades, swelling will occur at lower relative humidities
Table A.11-Coefflcients according to Eq. (A-63) and (A-66), CEB MC90-99 model
Type of cement according to EC2
a.as a.ds! a.ds2
SL (slowly-hardening cements) 800 3 0.13 Nor R (normal or rapid hardening cements) 700 4 0.12 RS (rapid hardening high-strength cements) 600 6 0.12
because of the preceding reduction of the internal relative humidity due to self-desiccation of the concrete.
A.3.3 Compliance-The compliance function J(t,to) that
represents the total stress-dependent strain by unit stress is given by
J(t, to) = ~[l1(to) + «P28(t, to») = f-+ «P~(t, to)
em28 cmlo em28
(A-70) where TJ(to) = Eem281Eemto' Eem28 is the mean modulus of elasticity of concrete at 28 days (MPa or psi), Eemto is the modulus of elasticity at the time of loading to (MPa or psi), and the dimensionless 28-day creep coefficient (1)28(t,to) gives the ratio of the creep strain since the start of loading at
the age to to the elastic strain due to a constant stress applied
at a concrete age of 28 days. Hence, I/Eemto represents the initial strain per unit stress at loading. The CEB MC90-99 model is closely related to the CEB MC90 model; however, it has been adjusted to take into account the particular characteristics of high-strength concretes. a) Modulus of elasticity-For the prediction of the creep function, the initial strain is based on the tangent modulus of elasticity at the time ofloading as defined in Eq. (A-7I) and (A-72). The modulus of elasticity of concrete at a concrete age t different than 28 days may be estimated from (A-71) where Eem28 is the mean modulus of elasticity of concrete at 28 days from Eq. (A-72); the coefficient s depends on the type of cement and the compressive strength of concrete and may be taken from Table A.12; and tl = I day. The modulus of elasticity of concrete made of quartzitic aggregates at the age of 28 days Eem28 (MPa or psi) may be estimated from the mean compressive strength of concrete by Eq. (A-72) Eem28 = 2I,500~/.em28 in SI units
emo
Eem28 = 3'118,31O~em28 in in.-Ib units
femo
(A-72) where fcm28 is the mean compressive cylinder strength of concrete at 28 days (MPa or psi), and fema = 10 MPa (1450 psi).

Page 25
MODELING AND CALCULATING SHRINKAGE AND CREEP IN HARDENED CONCRETE
209.2R-25
For concrete made of basalt, dense limestone, limestone, or sandstone, CEB MC90 recommends calculating the modulus of elasticity of concrete by multiplying Ecm28 (MPa or psi) according to Eq. (A-72) with the coefficients (J.E from TableA.13. The mean compressive cylinder strength of concrete (MPa or psi) is given by Eq. (A-73)
fcm28 = f; + 8.0 in SI units fcm28 = f; + 1160 in in.-Ib units
(A-73) wheref; is the specified/characteristic compressive cylinder strength (MPa or psi) defined as that strength below which 5% of all possible strength measurements for the specified concrete may be expected to fall. b) Creep coefficient-Within the range of service stresses (not larger than 40% of the mean concrete strengthfcmto at the time of loading to)' the 28-day creep coefficient CP28(t,to) may be calculated from Eq. (A-74)
(A-74)
where CPo is the notional creep coefficient, !3c(t - to) is the coefficient that describes the development of creep with time after loading, t is the age of concrete (days) at the moment considered, and to is the age of concrete at loading (days), adjusted according to Eq. (A-81) and (A-87). The notional creep coefficient CPo may be determined from Eq. (A-75) to (A-81) with
_ [3.5f cmoJO.7
(J.I- - -
fcm28
_ [3.5f cmoJO.2
(J.2 - - - fcm28
(A-75) (A-76)
(A-77)
(A-78) (A-79) (A-80) where fcm28 is the mean compressive strength of concrete at the age of 28 days (MPa or psi),fcmo = 10 MPa (1450 psi), h
is the relative humidity of the ambient environment in decimals, ho = 1, VIS is the volume-surface ratio (mm or in.), (VIS)o =
Table A.12--coefflcient s according to Eq. (A-71), CEB MC90 and CEB MC90-99 models
Icm28
Type of cement RS (rapid hardening high-strength cement) :560 MPa (8700 psi) Nor R (normal or rapid hardening cements) SL (slowly-hardening cement) >60 MPa (8700 psi)' All types 'Case not considered in CBB MC90.
Table A.13-Effect of type of aggregate on modulus of elasticity, CEB MC90 model
Aggregate type Basalt, dense limestone aggregates 1.2 Quartzitic aggregates 1.0 Limestone aggregates 0.9 Sandstone aggregates 0.7
s
0.20 0.25 0.38 0.20
50 mm (2 in.), tl = 1 day, to is the age of concrete at loading (days) adjusted according to Eq. (A-81) and (A-87), and (J.l and (J.2 are coefficients that depend on the mean compressive strength of concrete «(J.l = (J.2 = 1 in CEB MC90). The effect of type of cement and curing temperature on the creep coefficient may be taken into account by modifying the age at loading to according to Eq. (A-81)
to = to, r[ 9
12+ 1 r ~ 0.5 days
(A-81)
2 + (to, Tltl, T) .
where to,Tis the age of concrete at loading (days) adjusted to the concrete temperature according to Eq. (A-87) (for T = 20��C [68 oF], to,Tcorresponds to to) and tl,T= 1 day. (J. is a power that depends on the type of cement; (J. = -1 for slowly hardening cement; (J. = 0 for normal or rapidly hardening cement; and (J. = 1 for rapid hardening high-strength cement. The value for to according to Eq. (A-81) has to be used in Eq.
(A-78).
The coefficient !3c(t - to) that describes the development of creep with time after loading may be determined from Eq. (A-82) to (A-84)
(A-82)
with
13H = 150[1 + (1.2 . hlho)18](VIS)/(VIS)0 + 2500.3 ~ 15000.3 (A-83)
_ [3.5f cmoJO.5
(J.3 - - - fcm28 (A-84)
where tl = 1 day, ho = 1, (VlS)o = 50 mm (~in.), and (J.3 is a coefficient that depends on the mean compressi ve strength of concrete «(J.3 = 1 in CEB MC90).

Page 26
209.2R-26 ACI COMMITTEE REPORT
The duration of loading (t - to) used in Eq. (A-82) is the actual time under load.
Temperature effects-The effect of elevated or reduced
temperatures at the time of testing on the modulus of elasticity of concrete, at an age of 28 days without exchange of moisture, for a temperature range 5 to 80��C (41 to 176 OF), may be estimated from (A-85)
Ecm28 (D = Ecm28 [1.06 - 0.OO3(18.778T - 600.883)ITo] in in.-Ib units
where T is the temperature eC or oF), and To = 1 ��c (33.8 oF). Equation (A-85) can also be used for a concrete age other than t = 28 days. The 28-day creep coefficient at an elevated temperature may be calculated as (A-86) where <Po is the notional creep coefficient according to
Eq. (A-75) and temperature adjusted according to Eq. (A-90),
/3c(t - to) is a coefficient that describes the development of creep with time after loading according to Eq. (A-82) and temperature adjusted according to Eq. (A-88) and (A-89), and Ll<PT,trans is the transient thermal creep coefficient that occurs at the time of the temperature increase, and may be estimated from Eq. (A-92). The effect of temperature to which concrete is exposed before loading may be taken into account by calculating an adjusted age at loading from Eq. (A-87)
toJ = ��t.tieXP[13.65 -
4000
1 in SI units
i= I
273 + T(Mi )
To
(A-87)
to.T = ��t.tieXP[13.65-
4000
1 inin.-Ibunits
i = I
273 + (l8.778T(M,) - 600.883)
To
where to,T is the temperature-adjusted age of concrete at loading, in days, from Eq. (A-81), T(Lltj) is the temperature (OC or OF) during the time period Lltj, Mj is the number of days where a temperature T prevails, n is the number of time intervals considered, and To = 1 ��c (33.8 oF). The' effect of temperature on the time development of creep is taken into consideration using PH,T (Eq. (A-88» (A-88) with
I3T = exp[ 1500 - 5.12J in SI units (273 + TlTo)
(A-89)
13 = exp[
1500 - 5.12J in in.-Ib units
T
[273 + (18.778T - 600.883)ITo]
where /3H,T is a temperature-dependent coefficient that replaces /3H in Eq. (A-82), /3H is a coefficient according to Eq. (A-83), T is the temperature (OC or OF), and To = 1 ��c (33.8 oF). The effect of temperature conditions on the magnitude of the creep coefficient <Po in Eq. (A-74) and (A-75), respec- tively, may be calculated using Eq. (A-90) (A-90) with
q,T = exp[0.015(T ITo - 20)] in SI units
(A-91)
q,T = exp[0.015[(18.778T - 6oo.883)/To- 20]] in in.-Ib units
where <PRH,T is a temperature-dependent coefficient that replaces <PRH<h) in Eq. (A-75), <PRH<h) is a coefficient according to Eq. (A-76), and To = 1 ��c (33.8 OF). Transient temperature conditions, that is, an increase of temperature while the structural member is under load, leads to additional creep Ll<PT,trans that may be calculated from Eq. (A-92)
t.q,T"wo, = O,OOO4(TITo- 20)2 in SI units
(A-92)
t.q,T, "00' = 0.0004[(l8,778T - 6oo,883)ITo - 20]2 in in,-Ib units
Effect afhigh stresses-When stresses in the range of 40
to 60% of the compressive strength are applied, CEB MC90- 99 (CEB 1993, 1999) recommends using a high stress correction to the notional creep <Po as shown in Eq. (A-93)
<Po,k = <poexp{ 1.5(kcr - 0.4)}
(A-93) where ¢lo,k is the notional creep coefficient that replaces <Po in Eq. (A-74), and kcr is the stress-strength ratio at the time of application of the load. A.4-GL2000 model The model presented herein corresponds to the last version of the GL2000 model (Gardner 2004), including minor modifications to some coefficients and to the strength development with time equation of the original model developed by Gardner and Lockman (2001). It is a modified Atlanta 97 model (Gardner and Zhao 1993), which itself was influenced by CEB MC90. It presents a design-office procedure for calculating the shrinkage and creep of normal-strength concretes, defined as concretes with mean compressive strengths less than 82 MPa (11,890 psi) that do not experience self-desiccation, using the information available at design, namely, the 28-day specified concrete strength, the concrete strength at loading, element size, and relative humidity. According to Gardner and Lockman (2001), the method can be used regardless of what chemical admixtures or mineral by-products are in the concrete, casting terpperature, or
II'
curing regime. The predicted values can be improved by simply measuring concrete strength development with time and modulus of elasticity. Aggregate stiffness is taken into

Page 27
MODELING AND CALCULATING SHRINKAGE AND CREEP IN HARDENED CONCRETE
209.2R-27
account by using the average of the measured cylinder strength and that back-calculated from the measured modulus of elasticity of the concrete. The compliance expression is based on the modulus of elasticity at 28 days instead of the modulus elasticity at the age of loading. This model includes a term for drying before loading, which applies to both basic and drying creep. Required parameters: • Age of concrete when drying starts, usually taken as the age at the end of moist curing (days); • Age of concrete at loading (days); • Concrete mean compressive strength at 28 days (MPa or psi); • Concrete mean compressive strength at loading (MPa or psi); • Modulus of elasticity of concrete at 28 days (MPa or psi); • Modulus of elasticity of concrete at loading (MPa or psi); • Relative humidity expressed as a decimal; and • Volume-surface ratio (mm or in.).
A.4.1 Relationship between specified and mean compressive
strength of concrete-If experimental values are not available, the relationship between the specified/characteristic compressive strengthf; and the mean compressive strength of concretefcm28 can be estimated from Eq. (A-94)
fcm28 = 1.1f; + 5.0 in SI units fcm28 = 1.1f; + 700 in in.-Ib units
(A-94) Equation (A-94) is a compromise between the recommended equations of ACI Committee 209 (1982) and ACI Committee 363 (1992). It can be noted thatEq. (A-94) does not include any effects for aggregate stiffness or concrete density. Instead of making an allowance for the density of the concrete, it is preferable to measure the modulus of elasticity. If experimental values are not available, the modulus of elasticity Ecmt and the strength development with time fern! can be calculated from the compressive strength using Eq. (A-95) and (A-96).
A.4.2 Modulus of elasticity
E cmt = 3500 + 4300 JJ:::.t in SI units
Ecmt = 500,000 + 52,OOOJI:: in in.-Ib units
(A-95)
A.4.3 Aggregate stiffness-Aggregate stiffness can be
accommodated by using the average of the measured cylinder strength and that back-calculated from the measured modulus of elasticity using Eq. (A-95) in the shrinkage and specific creep equations. Effectively, Eq. (A-95) is used as an indicator of the divergence of the measured stiffness from standard values.
A.4.4 Strength development with time
(A-96) where (A-97) where s is a CEB (1993) style strength-development parameter (Table A.l4), and /3e relates strength development to cement type. Equation (A-96) is a modification of the CEB strength- development relationship. A single measured value of s permits values of k in the shrinkage equation to be interpolated, where k is a correction term for the effect of cement type on shrinkage (Table A. 14). If experimental results are available, the cement type is determined from the strength development characteristic of the concrete, regardless of the nominal designation of the cement. This enables the model to accommodate concretes incorporating any chemical or mineral admixtures.
A.4.5 Shrinkage-Calculate the shrinkage strain Esh(t,tc)
from Eq. (A-98) (A-98) where Eshu is the ultimate shrinkage strain, /3(h) is a correction term for the effect of humidity, and /3(t - tc) is a correction term for the effect of time of drying. The ultimate shrinkage Eshu is given by
Eshu = 9OOk( 30-'1 112
x 10-
6
in SI units
fcm21
900k(435~ 112
10-6 . . lb .
Eshu = -
E- X
m m.- umts
:Jcm2
(A-99) where fcm28 is the concrete mean compressive strength at 28 days in MPa or psi, and k is a shrinkage constant that depends on the cement type (Table A.l4). If test results for strength development are available, the shrinkage term can be improved by interpolating k from Table A.14 using the experimentally determined cement type/characteristic. The correction term for effect of humidity /3(h) is given by (A-100) Note that for a relative humidity of 0.96, there is no shrinkage. At a higher relative humidity, swelling occurs. The time function for shrinkage /3(t - tc) is given by [ (t-t)
]112
/3(t - tc) =
c
in SI units (t-tc )+0.12(V/S)2
(A-Wi) [ (t-t)
]1/2
/3(t - tJ =
c
in in.-Ib units (t - tc) + 77(V / S)2 where t and tc are the age of concrete and the age drying starts or end of moist curing in days, respectively, and VIS is the volume-surface ratio in mm or in.
A.4.6 Compliance equations~The "'compliance is
composed of the elastic and the creep strains. The elastic strain is the reciprocal of the modulus of elasticity at the age

Page 28
209.2R-28 ACI COMMITTEE REPORT
Table A.14-Parameters sand k as function of cement type, GL2000 model
Cement type
s
k
TypeJ 0.335 1.0 Type II 0.4 0.75 Type III 0.13 1.15
of loading Eemto' and the creep strain is the 28-day creep coefficient <P2S(t,to) divided by the modulus of elasticity at 28 days Ecm28 as in Eq. (A-102). The creep coefficient
<P28(t,to) is the ratio of the creep strain to the elastic strain due
to the load applied at the age of 28 days (A-I02) The 28-day creep coefficient <P28(t,to) is calculated using Eq. (A-103) In SI units: (A-I03) In in.-Ib units: The creep coefficient includes three terms. The first two terms are required to calculate the basic creep, and the third term is for the drying creep. Similar to the shrinkage Eq. (A-I 00), at a relative humidity of 0.96, there is only basic creep (there is no drying creep). <D(te) is the correction term for the effect of drying before loading.
If to = te
(A-104)
<D(t ) = 1 _
0 C
[ ( ( t - t )
J 0.5] 0.5
c
(to-tc)+0.12(V/S)2
in SI units (A-lOS)
<D(t) = [1- ( (to - tc)
J0.5]0.5 in in.-Ib units
c
(to-tc) + 77(V/S)2
To calculate relaxation, <D(te) remains constant at the initial value throughout the relaxation period. For creep recovery calculations, <D(te) remains constant at the value at the age of loading. APPENDIX B-STATISTICAL INDICATORS B.1-BP coefficient of variation (tlJ8P%) method· Developed by Bazant and Panula (1978), a coefficient of variation IDBP is determined for each data set. Data points in each logarithmic decade, 0 to 9.9 days, 10 to 99.9 days, and so on, are considered as one group. Weight is assigned to each data point based on the decade in which it falls and number of data points in that particular decade. The overall coefficient of variation (IDB3) for all data sets is the root mean square (RMS) of the data set values
IDij =
where n
N
(B-1) (B-2) (B-3)
(B-4)
= number of data points in data set number j;
= sum of the weights of all data points in a data set;
= number of data points in the k-th decade;
= number of decades on the logarithmic scale
spanned by measured data in data setj;
= number of data sets;
Oij = measured value of the shrinkage strain or creep
compliance for the i-th data point in data set numberj;
Cij = predicted value of the shrinkage strain or creep
compliance for the i-th data point in data set numberj;
Cij - Oij= deviation of the predicted shrinkage strain or
creep compliance from the measured value for the i-th data point in data set number,j;
IDij
= weight assigned to the i-th data point in data set
numberj;
IDj
= coefficient of variation for data set number j; and
IDB3
= overall coefficient of variation.
B.2-CEB statistical indicators The CEB statistical indicators: coefficient of variation
V CEB, the mean square error F CEB' and the mfian de;v.iation
MCEB were suggested by Muller and Hilsdorf (1990). The
indicators are calculated in six time ranges: 0 to 10 days, 11 to 100 days, 101 to 365 days, 366 to 730 days, 731 to 1095

Page 29
MODELING AND CALCULATING SHRINKAGE AND CREEP IN HARDENED CONCRETE
209.2R-29
days, and above 1095 days. The final values are the RMS of the six interval values. B.2.1 CEB coefficient o/variation
n
- 1
O· = -~ (0·.)
I
n~
I)
(B-5)
j = 1 n
Vi = 1 _1_ ~ (c.- o.l O. n -1 ~ I)
I)
) j = 1
(B-6) (B-7)
where n
= number of data points considered; N = total number of data sets considered; Vj
= coefficient of variation in interval i; and
V CEB = RMS coefficient of variation.
B.2.2 CEB mean square error-The mean square error uses the difference between the calculated and observed values relative to the observed value
(B-8)
FJ?n
2
F. = -~J;
I
n-l~)
j=l
(B-9) (B-1O)
where
Ij
= percent difference between calculated and
observed data pointj; and
F CEB = mean square error, %.
B.2.3 CEB mean deviation-The CEB mean deviation
MCEB indicates systematic overestimation or underestimation
of a given model
! ~ Cu
Mi = ~-'4
nO·.
j = 1 I)
(B-ll) (B-12)
where
Mj
= ratio of calculated to experimental values in time
range i; mean deviation;
N =
number of values considered in time interval; and total number of data sets considered. B.3-The Gardner coefficient of variation (OlG) Developed by Gardner (2004), the mean observed value and RMS of the difference between calculated and observed values were calculated in half logarithmic time intervals: 3 to 9.9 days, 10 to 31.5 days, 31.6 to 99 days, 100 to 315 days, 316 to 999 days, 1000 to 3159 days, and above 3160 days. That is, the duration of each time interval is 3.16 times the previous value. To obtain a crite- rion of fit, the average values and RMSs were averaged without regard to the number of observations in each half- decade. A coefficient of variation is obtained by dividing the average RMS normalized by the average value. It is necessary to emphasize that this is not the conventional definition of the coefficient of variation
_ 1 n
O.=-~(O .. )
)
n~
I)
i= 1 n
1 L
2
- (C.-O··)
n - 1
I) I)
_ 1 n _
O=-~(O.)
N~ ) j= 1
n i = 1
- 1
RMS = - ~ (RMS.)
N~
J
j = 1
RMS
OlG = ~
o
(B-13) (B-14) (B-15) (B-16) (B-17)

Page 30
209.2R-30 ACI COMMITTEE REPORT
APPENDIX C-NUMERIC EXAMPLES
Find the creep coefficients and shrinkage strains of concrete at 14, 28, 60, 90, 180, and 365 days after casting, from the following information: specified concrete compressive strength of 25 MPa (3626 psi), 7 days of moist curing, age of loading to = 14 days, 70% ambient relative humidity, and volume-surface ratio of the member = 100 mm (4 in.).
Problem data
Concrete data:
Specified 28-day strength
f~ = iAmbient conditions:
Relative humidity h= Temperature T=
Specimen:
Volume-surface ratio VIS = Shape
Initial curing:
Curing time te = Curing condition
Concrete at loading:
Age at loading to = Applied stress range ks =
C.1-ACI209R-92 model solution
C.I.I Estimated concrete properties Mean 28-day strength
fem28 =
Mean 28-day elastic modulus
Eem28 =
C.I.2 Estimated concrete mixture Cement type Maximum aggregate size Cement content
c=
Water content
w=
Water-cement ratio wlc= Aggregate-cement ratio alc= Fine aggregate percentage
\jI=
Air content
a=
Slump s= Unit weight of concrete Ye=
"Table A1.5.3.7.l and 6.3.7.1 of ACI 211.1-91.
C.I.3 Shrinkage strains Gsh(t,tc) Nominal ultimate shrinkage strain Moist curing correction factor SI units in. -lb units 25MPa 3626 psi 0.7 20��C 68 OF 100mm 4 in. Infinite slab 7 days Moist cured 14 days 40% SI units in.-lb units 33.3 MPa 4830 psi Table 5.3.2.2 ACI 318-05 28,178 MPa 4,062,346 psi (A-16) SI units in.-lb units I I 20mm 3/4 in. 409 kglm
3
6901b/yd
3
205 kglm
3
3451b/yd3 Table 6.3.3 ACI 211.1-91 0.50 (4-1) 4.23 40% 2% Table 6.3.3 ACI 211.1-91 75mm 2.95 in. 2345 kglm
3
39531b/yd3 146* Ib/ft3 SI units
I
in.-lb units Eshu = 780 x 10-6 (A-2) Ysh,te = 1.202 - 0.23371og(te) = 1.005 (A-6)

Page 31
MODELING AND CALCULATING SHRINKAGE AND CREEP IN HARDENED CONCRETE
209.2R·31
Ysh,RH= lAO -1.02h if 004 :$; h:$; 0.8
(A·7) Ambient relative humidity factor
Ysh,RH = 3.00 - 3h if 0.8 < h :$; 1
(A·7)
Ysh,RH = 0.686
(A·7) Volume-to-surface ratio factor y - 1 2e[--{).00472(VlS)]
sh,vs - .
(A-8)
Y
= 1 2e[--{).12(VIS)]
sh,vs .
(A-8)
Ysh,vs = 0.749
(A-8)
Ysh,vs = 0.743
(A-8)
Ysh,s = 0.89 + 0.00161s
(A-ll)
Ysh,s = 0.89 + 0.041s
(A-ll) Slump of fresh concrete factor
Ysh,s = 1.011
(A-II)
Ysh,s = 1.011
(A-11)
Ysh,IV = 0.30 + 0.014", if ",:$; 50%
(A-12) Fine aggregate factor
Ysh,IV = 0.90 + 0.002", if '" > 50%
(A-12)
Ysh,IV = 0.860
(A-12)
Ysh,e = 0.75 + 0.00061c
(A-13)
Ysh,e = 0.75 + 0.00036c
(A-13) Cement content factor
Ysh,e = 0.999
(A-13)
Ysh,e = 0.998
(A-13)
Ysh,a = 0.95 + 0.008a ~ 1
(A-14) Air content factor
Ysh,a = 1.000
(A-14)
Ysh = Ysh,teYsh,RJl'fsh,vsYsh,sYsh,IVYsh,eYsh,a
(A-5) Cumulative correction factor
Ysh = 0.448
(A-5)
Ysh = 0.444
(A-5)
Eshu = 780ysh x 10--6
(A-4) Ultimate shrinkage strain
Eshu = 350 x 10--6 Eshu = 347 x 10-
6
(A-4) (A-4) Shrinkage time function
f(t,te) = [(t - te)a/(f + (t - te)a)]
Shrinkage strains
Esh(t,tc) = [(t - tc)a/(f + (t - tc)a)]Eshu
(A-I)
a=l t, days f(t- te) Esh(t,te), X 10--6 t, days f(t- te) Esh(t,te), X 10--6
7 0.000 0 7 0.000 0 14 0.167 58 14 0.167 58 28 0.375 131 28 0.375 130
f= 35 days
60 0.602 211 60 0.602 209 90 0.703 246 90 0.703 244 180 0.832 291 180 0.832 288 365 0.911 318 365 0.911 316 Note that the 365-day shrinkage strain reduces to 268 x 10--6 when the effect of the volume-surface ratio on the shrinkage time function is considered, that is, iff = 26eo.0142(VlS) = 108 days (f= 26e��.36(VlS) = 110 days).
C.l.4 Compliance J(1,1o)
a) Elastic compliance J(1o,to)
SI units in.-Ib units I Cement type
a=4
(Table A.4)
b=0.85
(Table A.4) Mean strength at age to
femto = [ti(a + bto)]fem28
(A-17)
femto = 29.3 MPa
(A-17)
fcmto = 4253 psi
(A-17) Mean elastic modulus at
Ecmto = 0.043Ye 15 fcmto
(A-16)
Eemto = 33y/
5
femto
(A-16) age to
Eemto = 26,441 MPa
(A-16)
Eemto = 3,811,908 psi
(A-16)
J(to,to) = lIEemto
(A-15) Elastic compliance
J(to,to) = 37.82 x 10--6 (l/MPa)
(A-15)
J(to,to) = 0.262 x 10-6 (l/psi)
(A-15)

Page 32
209.2R-32
ACI COMMITTEE REPORT
b) Creep coefficient ~(t,to)
SI units in.-Ib units Nominal ultimate creep coefficient
<l>u = 2.35
(A-19)
"f - 1 25t -{l.118
(A-22) Age application of load factor
e,to -. 0 "fe, to = 0.916
(A-22) Ambient relative humidity factor
"f e,RH = 1.27 - 0.67 h if h ~ 0.4
(A-24)
"fe,RH=0.801
(A-24) Volume-to-surface ratio factor
"fevs = 2/3[1 + 1.13e(-D.0213(V/S)] (A-25) "fe,vs = 2/3[1 + 1.13i--{)·54(VIS)]
(A-25)
"fe,vs = 0.756
(A-25)
"fe,vs = 0.754
(A-25)
"fe,s = 0.82 + 0.OO264s
(A-28)
"fe,s = 0.82 + 0.067 s
(A-28) Slump of fresh concrete factor
"fe,s = 1.018
(A-28)
"fe,s = 1.018
(A-28) Fine aggregate factor
"fe, IV = 0.88 + 0.0024",
(A-29)
"fe, IV = 0.976
(A-29) Air content factor
"f e,a. = 0.46 + 0.09<1 ~ 1
(A-30)
"fe,a. = 1.000
(A-30) Cumulative correction factor
"fe = "fe,to"fe,RJlYe,vs"fe,s"fe,IV"fsh,a.
(A-21)
"fe = 0.551
(A-21)
"fe = 0.549
(A-21)
<l>u = 2.35"fe
(A-20) Ultimate shrinkage strain
<l>u = 1.29
(A-20)
<l>u = 1.29
(A-20) Creep coefficient time function
f(t - to) = [(t - to)IVJ(d + (t - to)IV)]
Creep coefficients
<I>(t,to) = [(t - to)IVJ(d + (t - to)IV)]<I>u
(A-18)
'" = 0.6 t, days f(t- te) <I>(t,to) t, days f(t- te) <I>(t,to) d= 10 days
14 0.000 0.000 14 0.000 0.000 28 0.328 0.424 28 0.328 0.423 60 0.499 0.646 60 0.499 0.643 90 0.573 0.742 90 0.573 0.740 180 0.682 0.883 180 0.682 0.880 365 0.771 0.998 365 0.771 0.995
c) Compliance J(t,to)= l/Ecmto+ «t,to}/Ecmto
SI units in.-Ib units
t, days J(to,to)' x 10-6 <I>(t,to)JEemto' x 10-
6
J(t,to) (lIMPa), x 10-6 J(to,to)' x 10-
6
<I>(t,to)JEemto' x 10-6 J(t,to) (lJpsi), x 10-6
14 37.82 0 37.82 0.262 0 0.262 28 37.82 16.04 53.86 0.262 0.111 0.373 60 37.82 24.42 62.24 0.262 0.169 0.431 90, 37.82 28.08 65.90 0.262 0.195 0.457 180 37.82 33.41 71.24 0.262 0.231 0.493 365 37.82 37.75 75.58 0.262 0.261 0.523 Note that when the effect of the volume-surface ratio is considered in the time function of the creep coefficient as d =
26eo.0142(VIS) = 108 days (f= 26e��.36(VIS) = 11Odays) and", = 1, the creep coefficient and the compliance rate of development
are initially smaller than when the effect of the volume-surface ratio is not considered; however, after 365 days under load, they are similar.

Page 33
,
MODELING AND CALCULATING SHRINKAGE AND CREEP IN HARDENED CONCRETE
C.2-Bazant-Baweja B3 model solution C.2.1 Estimated concrete properties SI units in.-Ib units
209.2R-33
Mean 28-day strength fcm28 = 33.3 MPa 4830 psi Table 5.3.2.2 ACI 318-05 Mean 28-day elastic modulus ECm28 = 27,318 MPa 3,961,297 psi (A-39) C.2.2 Estimated concrete mixture SI units in.-Ib units Cement type I Maximum aggregate size 20mm 3/4 in. Cement content c= 409kglm
3
690lb/yd3 Water content w= 205 kglm3 345lb/yd
3
Table 6.3.3 ACI 211.1-91 Water-cement ratio w/c= 0.50 (4-1) Aggregate-cement ratio alc= 4.23 Fine aggregate percentage
'If =
40% Air content U= 2% Table 6.3.3 ACI 211.1-91 Slump s= 75mm 2.95 in. Unit weight of concrete Yc= 2345 kglm3 3953lb/yd3 1461b/ft3*
·Table A1.5.3.7.l and 6.3.7.1 of ACI 211.1-91.
C.2.3 Shrinkage strains &sh(t,tJ SI units in.-Ib units kh = -0.2 if h = 1 (Table A.6) Ambient relative kh = 12.74 - 12.94h if 0.98 < h < 1 (TableA.6) humidity factor kh = 1 - h
3 if h ~ 0.98
(Table A.6) kh = 0.657 (Table A.6) Cement type factor ul = 1.000 (Table A.7) Curing condition factor u2 = 1.000 (Table A.8)
Esoo =-al~[0.01~.Ifcm28"'()28 Esoo = -alU2[0.02565~·1cm28 "'().28
Nominal ultimate
+ 270] x 10-6
(A-33)
+ 270] x 10-6
(A-33) shrinkage
Esoo = -780 x 10-6
(A-33)
E
= -781 x 10-6 (A-33)
SOC)
Member shape factor ks = 1.000 (Table A.9)
't - 0 085t -O.08!. -0.25 [2k (VIS)]2(A-36) 't = 190 8t -O.08!. -0.25 [2k (V/S)]2
(A-36) Shrinkage half-time sh-· c cm28 s sh ·c cm28 s 'tsh = 1211.323 (A-36) 'tsh = 1253.630 (A-36) Ecmfl.)7IEcm(tc+'tsh) = 1.167421[(tc + 'tsh)/(4 + 0.85(tc + 'tsh))] (A-32) & (A-34) Time dependence factor Ecm607IEcm(tc+'tsh) = 0.996 (A-32) & (A-34) Ecm607IEcm(tc+'tsh) = 0.996 (A-32) & (A-34) Eshoo = -EsooEcm607IEcm(tc+Tsh) (A-32) Ultimate shrinkage strain Eshoo = -777 x 10-6 Eshoo = -778 x 10-6 (A-32) (A-32) Shrinkage time function S(t - tc) = tanh[(t - tc)/'tsh]0.5 (A-35) Shrinkage strains Esh(t,tc) = -Eshookhtanh[ (t - tc)/'t sh]0.5 (A-31)

Page 34
209.2R-34 ACI COMMITTEE REPORT
t, days S(t - tc) Esh(t,tc)' x 10-6 t, days
7 0.000 14 0.076 28 0.131 60 0.206 90 0.256 180 0.361 365 0.496
C.2.4 Compliance J(t,ta) = q] + Co(t,to) + Cd(t,t(Ytc) a) Instantaneous compliance q]= O.6/Ecm2B
SI units Instantaneous 0 7 -39 14 -67 28 -105 60 -131 90 -184 180 -253 365
I
qI = lIEo = 0.6/Ecm28 S(t - tc) Esh(t,tc)' x 10-
6
0.000 0 0.075 -38 0.129 -66 0.203 -104 0.252 -129 0.355 -182 0.489 -250 in.-Ib units (A-38) compliance
qI = 21.96 x 10-6(lIMPa)
I
qI = 0.152 x lO-6(l/psi)
b) Compliance function for basic creep CoCt,to) = q2Q(t,ta) + q3ln[I + (t-talJ + q4In(tlto)
Aging viscoelastic tenn q2Q(t,to) SI units in.-Ib units
q = 185 4 x 1O-6cO.5f,
-0.9
2 · cm28
(A-41)
q = 86 814 x lO-6c0.5f,
-D.9
2 · cm28 q2 = 159.9 x 10-6 (lIMPa)
(A-41)
q2 = 1.103 x 10-6 (l/psi) Qt(to) = [0.086(to)2/9 + 1.21(to)4/9rI Qt(to) = 0.246
m =0.5 n = 0.1
r(to) =1.7(to)O.l2 + 8 r(to) = 10.333
Aging viscoelastic tenn Aging viscoelastic tenn
Z(t,to) = (tormln[1 + (t - to)n] Q(t,to) = Qt(to)[1 + {Qt(to)/Z(t,to)}',lolrIlr(tol q2Q(t,to) (l/MPa), t, days Z(t,to) Q(t,to)
x 10-6
t, days Z(t,to) Q(t,to)
14 0.000 0.000 0 14 0.000 0.000 28 0.223 0.216 34.59 28 0.223 0.216 60 0.241 0.228 36.41 60 0.241 0.228 90 0.249 0.232 37.02 90 0.249 0.232 180 0.262 0.236 37.78 180 0.262 0.236 365 0.275 0.240 38.30 365 0.275 0.240 Nonaging viscoelastic tenn q3ln[1 + (t - to)n] SI units
I
in.-Ib units
q3 = 0.29(w/c)4q2 q3 = 2.924 x 10-6 (lIMPa)
(A-46)1
q3 = 0.020 x 10-6 (l/psi)
n = 0.1 (A-41) (A-41) (A-43) (A-43) (A-45) (A-45) (A-44) (A-42)
q2Q(t,to) (l/psi),
x 10-6 0 0.239 0.251 0.255 0.261 0.264 (A-46) (A-46)

Page 35
MODELING AND CALCULATING SHRINKAGE AND CREEP IN HARDENED CONCRETE
209.2R-35
Nonaging viscoelastic term Nonaging viscoelastic term
t, days
In[ 1 + (t - to)n] q3ln[1 + (t - to)n] (llMPa), x 10--6 t, days In[1 + (t - to)n] I q31n[1 + (t - to)n] (l/psi), x 10-6 14 0.000 0 14 0.000 0 28 0.834 2.44 28 0.834 0.017 60 0.903 2.64 60 0.903 0.018 90 0.933 2.73 90 0.933 0.019 180 0.981 2.87 180 0.981 0.020 365 1.029 3.01 365 1.029 0.021 SI units in.-lb units
q4 = 20.3 x 1O-6(alcr-O·
7
(A-47)
q4 = 0.14 x 10--6(alc)-{)·7
(A-47)
q4 = 7.396 x 10--6 (llMPa)
(A-47)
q4 = 5.106 x 10- 8
(l/psi) (A-47) Aging flow term Aging flow term
t, days In(t,to) q41n(tlto) (llMPa), x 10-6 t, days In(t,to)
I q41n(tlto) (l/psi), x 10-6
14 0.000 0 14 0.000 0 28 0.693 5.13 28 0.693 0.035 60 1.455 10.76
60
1.455 0.074 90 1.861 13.76 90 1.861 0.095 180 2.554 18.89 180 2.554 0.130 365 3.261 24.12 365 3.261 0.167 SI units in.-Ib units
Co(t,to) = q2Q(t,to) + q31n[1 + (t - to)n] + q4ln(tlto)
(A-40)
q4ln(tlto)' Co(t,to) q41n (tlto), Co(t,to) t, days q2Q(t,to) q3ln[1 + (t - to)n] x 10--6 (llMPa), x 10--6 t, days q2Q(t,to) q31n[1 + (t - to)n]
x 10--6 (l/psi), x 10--6 14 0 0 0 0 14 0 0 0 0 28 34.59 2.44 5.13 42.15 28 0.239 0.Q17 0.035 0.291
60
36.41 2.64 10.76 49.81 60 0.251 0.Q18 0.074 0.344 90 37.02 2.73 13.76 53.51 90 0.255 0.019 0.095 0.369 180 37.78 2.87 18.89 59.54 180 0.261 0.020 0.130 0.411 365 38.30 3.01 24.12 65.42 365 0.264 0.021 0.167 0.451
c) Compliance function for drying creep Cit, to. tc) = C!5[ exp{ ~H(t)} - exp(~H(to)} f-5
SI units
I
in.-lb units
q5 = 0.757fcm28-IIEshoo x 106~.6
(A-49)
q5 = 419.3 x 10-6 (llMPa)
(A-49)1
q5 = 2.889 x 10--6 (l/psi)
(A-49)
S(to - tc) = tanh[(to - tc)/T. sh]O.5
(A-53)
S(to - tc) = 7.587 x 10-2
(A-53)1
S(to - tc) = 7.459 x 10-2
(A-53)
H(to) = 1 - (1 - h)S(to - tc)
(A-51)
H(to) = 0.977
(A-51)]
H(to) = 0.978
(A-51)
S(t - tc) = tanh[(t - tc)hsh]o.5
(A-52)

Page 36
209.2R-36
ACI COMMITTEE REPORT
H(t) = 1 - (1 - h)S(t - tc) (A-50)
f(H) = [exp{ -8H(t)} - exp{ -8H(to)} ]0.5
Cjt,to,tc) = q5[exp{ -8H(t)} - exp{ -8H(to)} ]0.5 (A-48)
f(H),
Cjt,to,tc) (1lMPa),
f(H),
Cjt,to,tc) (1/psi), t, days Set -tc) H(t) x 10-2 x 10-6 t, days Set -tc) H(t) x 10-2 x 10-6 14 0.076 0.977 0 0 14 0.075 28 0.131 0.961 0.754 3.16 28 0.129 60 0.206 0.938 1.216 5.10 60 0.203 90 0.256 0.923 1.475 6.19 90 0.252 180 0.361 0.892 1.988 8.34 180 0.355 365 0.496 0.851 2.646 11.10 365 0.489 SI units J(t,to) = q1 + Co(t,to) + Cjt,to,tc)
q1'
Co(t,to)' Cjt,to,tc)' J(t,to) (1IMPa),
q1'
t, days x 10-
6
x 10-6 x 10-6 14 21.96 0 0 28 21.96 42.15 3.16 60 21.96 49.81 5.10 90 21.96 53.51 6.19 180 21.96 59.54 8.34 365 21.96 65.42 11.10
C.3-CEB MC9O-99 model solution
C.3.1 Estimated concrete properties Mean 28-day strength Strength constant Mean 28-day elastic modulus C.3.2 Estimated concrete mixture Cement type Maximum aggregate size Cement content Water content Water-cement ratio Aggregate-cement ratio Fine aggregate percentage Air content Slump Unit weight of concrete
'Table A1.5.3.7.l and 6.3.7.1 of ACI 211.1-91.
x 10-6 t, days x 10-
6
21.96 14 0.152 67.27 28 0.152 76.87 60 0.152 81.66 90 0.152 89.84 180 0.152 98.48 365 0.152 SI units fcm28 = 33.0 MPa fcmo = lOMPa Ecm28 = 32,009 MPa SI units
N
20mm
c=
406 kg/m3
w= 205 kg/m
3
wlc= 0.504 alc = 4.27
"'=
40%
a=
2% s= 75mm Yc=
2345 kg/m3
C.3.3 CEB MC90 shrinkage strains csh(t,tc) SI units
I
Cement type factor
~sc = 5
0.978 0 0 0.961 0.746 0.022 0.939 1.202 0.035 0.925 1.458 0.042 0.893 1.964 0.057 0.853 2.613 0.076 in.-Ib units (A-37) Co(t,to)' Cjt,to,tc)' J(t,to) (1/psi), x 10-6 x 10-6 x 10-6 0 0 0.152 0.291 0.022 0.464 0.344 0.035 0.530 0.369 0.042 0.563 0.411 0.057 0.619 0.451 0.076 0.678 in.-Ib units 4786 psi (A-73) 1450 psi (A-72) 4,642,862 psi (A-72) in.-Ib units
3/4 in.
6851b/yd
3
3451b/yd
3
Table 6.3.3 ACI 211.1-91 (4-1) Table 6.3.3 ACI 211.1-91 2.95 in. 39531b/yd
3 146* Ib/ft3
in.-lb units (Table A.I0)

Page 37
MODELING AND CALCULATING SHRINKAGE AND CREEP IN HARDENED CONCRETE
209.2R·37
Concrete strength
Es(fcm2S) = [160 + I013sc(9 - fcm2Slfcmo)] x lO-6
(A·56) factor
Es(fcm2S) = 445 x lO-6
(A·56)
13RIIh) = -1.55[1- (hlho)3] for 0.4 ~ h < 0.99
(A-57) Ambient relative
13RIIh) = 0.25 for h ~ 0.99
(A-57) humidity factor
ho = 1 f3RIIh) = -1.018
(A-57) Notional shrinkage
Ecso = Es(fcm2s)13RIIh)
(A·55) coefficient
Ecso = -453 x lO-6
(A-55)
Ecso = -453 x lO-6
(A-55)
13it - tc) = [{ (t - tc)/td/{350([(VIS)/(VIS)o]2 +(t - tc)td]O.5
(A-58) Shrinkage
tl = 1 day
time function
(VIS)o = 50 mm (VIS)o = 2 in.
Shrinkage strains
Esh(t,tc) = Ecso13it - tc)
(A-54)
t, days 13it - tc) Esh(t,tc)' x 10-6 t, days 13it - tc) Esh(t,tc)' x 10-6
7 0.000 14 0.071 28 0.122 60 0.191 90 0.237 180 0.332 365 0.451 C.3.4 CEB MC90-99 shrinkage strains GSh(t,tc)
a) Autogenous shrinkage &cas(t)
Cement type factor 0 -32 -55 -87 -lO7 -150 -205 SI units 7 0.000 0 14 0.071 -32 28 0.122 -55 60 0.191 -87 90 0.237 -107 180 0.332 -150 365 0.451 -205 in.-Ih units
aas= 700
(Table A. 11) Notional
Ecaso(fcm2S) = -aas[(fcm2Slfcmo)/{ 6 + (fcm2S/fcmo)} ]2.5 x 10-6
(A-63) autogenous shrinkage
Ecaso(fcm2S) = -52.5 x 10-6
(A-63)
Ecaso(fcm2S) = -52.5 x 10-6
(A-63) Autogenous shrinkage time function
f3as(t) = 1 - exp[-O.2(tlti)O.5]
(A-64)
tl = 1 day
Autogenous shrinkage strains
Ecas(t) = Ecaso(fcm2S)13as(t)
(A-62)
t, days 13as(t) Ecas(t), X lO-6 t, days 13ait) Ecas(t), X lO-6
0 0.000 0 0 0.000 0 7 0.411 -22 7 0.411 -22 14 0.527 -28 14 0.527 -28 28 0.653 -34 28 0.653 -34 60 0.788 -41 60 0.788 -41 90 0.850 -45 90 0.850 -45 180 0.932 -49 180 0.932 -49 365 0.978 -51 365 0.978 -51
b) Drying shrinkage &cdS<t,tc)
SI units
I
in.-Ib units
adsl = 4
. fl·
(Table A. 11) Cement type factors
ads2 = 0.12
(Table A. 11)

Page 38
209.2R-38 ACI COMMITTEE REPORT
Notional drying shrinkage coefficient
Eedso(fem28) = [(220 + l1Oudsl)exp(-uds2fem28ifemo)] x 10-6
(A-66)
Eedso(fem28) = 444 x 10-
6
(A-66)
Eedso(fem28) = 444 X 10-6
(A-66)
ho= 1
Pool = [3.5femdfem28]O.l S; 1.0
(A-69)
Pool = 1.000
(A-69)
Pool = 1.000
(A-69) Ambient relative humidity factor
PRH(h) = -1.55[1 - (h/ho)3] for 0.4 S; h < 0.99Psl
(A-67)
PRH(h) = 0.25 for h ~ 0.99Psl
(A-67)
PRH(h) = -1.018
(A-67)
PRH(h) = -1.018
(A-67)
Pdit - te) = [{ (t - te)/t 1 }/ {350([ (V/S)/(V/S)o]2 + (t - te)/ti} ]0.5
(A-68) Drying shrinkage time function
tl = 1 day (V/S)o = 50 mm (V/S)o = 2 in.
Drying shrinkage strains
Eedit,te) = Eedso(fem28)PRH<h)Pds(t - te)
(A-65)
t, days Pds(t - te) Eeds(t,tc)' X 10-6 t, days Pds(t - tc) f.cdit,tc)' x 10-6
7 0.000 0 7 0.000 0 14 0.071 -32 14 0.071 -32 28 0.122 -55 28 0.122 -55 60 0.191 -86 60 0.191 -87 90 0.237 -107 90 0.237 -107 180 0.332 -150 180 0.332 -150 365 0.451 -204 365 0.451 -205
c) Total shrinkage strains csh(t,tc)
SI units in. -lb units
f.sh(t,tc) = Ecait) + Ecds(t,tc)
(A-61)
t, days f.cas(t), X 10-6 f.cdit,tc)' x 10-6 Esh(t,tc)' x 10-6 t, days Ecait), x 10-6 Ecds(t,tc)' x 10-6 Esh(t,tc)' x 10-6
0 0
-
0 0 0
-
0 7 -22 0 -22 7 -22 0 -22 14 -28 -32 -60 14 -28 -32 -60 28 -34 -55 -89 28 -34 -55 -89 60 -41 -86 -127 60 -41 -87 -128 90 -45 -107 -152 90 -45 -107 -152 180 -49 -150 -199 180 -49 -150 -199 365 -51 -204 -255 365 -51 -205 -256
C.3.S Compliance J(t,lo) a) Elastic compliance J(lo,lo)
SI units
I
in.-Ib units N Cement type
s = 0.25
(Table A.12)
Pe = exp[s/2{ 1 - (28/to)0.5}]
(A-97) Mean strength at age to
Pe = 0.950
(A-97)
fcmto = P/fcm28
(A-96)
fcmto = 29.8 MPa
(A-96)1
fcmto = 4315.1 psi
(A-96) Mean elastic modulus at age to
Eemto = Ecm28exP[s/2{ 1 - (28/to)0.5)]
(A-71)
Ecmto = 30,394 MPa
(A-71)1
Ecmto = 4,408,587 psi
(A-71)

Page 39
MODELING AND CALCULATING SHRINKAGE AND CREEP IN HARDENED CONCRETE
209.2R·39
J(to,to) = lIEcmto (A·70) Elastic compliance J(to,to) = 32.90 x 10-6 (l/MPa) (A·70) J(to,to) = 0.227 x 10-6 (llpsi) (A-70) Ecm2S(1) = Ecm2S(1.06 - 0.003TITo) (A-85) Ecm2S(1) = Ecm2S(1.06 - 0.003 . [18.778T - 6OO.883]1To) (A-85) Effect of temperature Ecm2S(1) = 32,009 MPa (A-85) Ecm2S(1) = 4,642,853 psi (A-85) on modulus of elasticity Ecmti1) = EcmtoO.06 - 0.003TITo) (A-85) Ecmto(1) = Ecmto(1.06 - 0.003 . [18.778T - 600.883]ITo) (A-85) Ecmti1) = 30,394 MPa (A-85) EcmtO<1) = 4,408,579 psi (A-85) J(to,to) = 11Ecmto (A-70) Elastic compliance temperature adjusted J(to,to) = 32.90 x 10--6 (l/MPa) (A-70) J(to,to) = 0.227 x 10-6 (llpsi) (A-70)
b) Creep coefficient ¢2S< t, to)
SI units in.-Ib units 0.1 = [3.5fcmJfcm2S]0.7 (A-79) Compressive strength factors 0.2 = [3.5fcmJfcm2S]0.2 (A-79) 0.1 = 1.042 (A-79) 0.1 = 1.042 (A-79) 0.2 = 1.012 (A-79) 0.2 = 1.012 (A-79) tPREih) = [1 + {(l - hlho)al/(O.l (VIS)/(VIS)0}]a2 (A·76) Ambient relative humidity and ho = 1 volume-surface ratio factor (VIS)o = 50 mm (VIS)o = 2 in. tPREih) = 1.553 (A-76) tPREih) = 1.553 (A-76) Concrete strength factor
!3(fcm2S) = '5.3/(fcm2S/fcmo)0.5
(A-77)
!3(fcm2S) = 2.918
(A-77)
!3(fcm2S) = 2.917
(A-77) to,T = ~At,exp[13.65 - 40001 to,T= ~At,exp[13.65 -40001 {273 + (T(AtITo))}] (A-87) {273 + (18.778T(Atj) - 6OO.883ITo)}] (A-87) To = 1��C To = 33.8 OF Temperature-adjusted to,T= 14.0 days (A-87) to,T= 14.0 days (A-87) age of loading to = to,rl9/{2 - (to,Tltl,T)1.2} + 1]1l ~ 0.5 days (A-81) 0.=0 tl,T= 1 day to = 14.0 days (A-81) Adjusted age of loading factor !3(to) = 1/[0.1 + (tJt 1)0.2] (A-78) !3(to) = 0.557 (A-78) !3(to) = 0.557 (A-78)
tPo = tPREih)!3(fcm2S)!3(to)
(A-75) Notional creep coefficient
tPo= 2.524
(A-75)
tPo= 2.524
(A-75) 0.3 = [3.5fcmJfcm2S]0.5 (A-84) 0.3 = 1.030 (A-84) 0.3 = 1.030 (A-84) Creep coefficient time function !3H = 150[1 + (1.2hlho)IS](VIS)/(VlS)0 + 2500.3:5; 1500a.3 (A-83) !3H= 570.470 (A-83) !3H = 570.445 (A-83) !3c(t- to) = [(t- to)/tl/{!3H+(t- to)/t!l]O.3 (A-82) Creep coefficients tP2S(t,to) = tPo!3c(t - to) (A-74)

Page 40
209.2R-40
ACI COMMITTEE REPORT
t, days
f3c(t - to)
$28(t,to) t, days
f3c(t - to)
$28(t,to) 14 0.000 0.000 14 0.000 0.000 28 0.326 0.824 28 0.326 0.824 60 0.459 1.159 60 0.459 1.159 ' 90 0.526 1.328 90 0.526 1.328 180 0.640 1.614 180 0.640 1.614 365 0.749 1.890 365 0.749 1.889 SI units in.-lb units $T= exp[0.015(T1To - 20)] (A-91) $T = exp[0.015 {(18.778T - 600.883)ITo - 20}] (A-91) $T= 1.000 (A-91) $T= 1.000 (A-91) Effect of temperature $RH,T= $T+ fiRH<h) -1]$/2 (A-90) conditions $RH,T = 1.553 (A-90) $RH,T = 1.553 (A-90) $0 = $RH,rf3(fcm28)f3(to) (A-75) $0 = 2.524 (A-75) $0 = 2.524 (A-75) Effect of high stresses $o,k = $oexp[1.5(ka - 0.4)] (A-93) $o,k = 2.524 (A-93) $o,k = 2.524 (A-93) Notional creep $0 = $ck coefficient temperature $0 = 2.524 $0= 2.524 and stress adjusted
~T= exp[1500/(273 + T1To) - 5.12]
(A-89) ~T= exp[1500/{273 + (18.778T - 6OO.883)ITo}) -5.12] (A-89)
f3T= 0.999 (A-89) f3T= 0.999 (A-89) f3H,T = f3Hf3T (A-88) Effect of temperature f3H,T= 570.159 (A-88) f3H,T= 570.128 (A-88) conditions on creep coefficient time function
~$T.trans = 0.0004(T1To - 20)2 (A-92) ~$T.trans = 0.0004[(18.778T - 600.883)ITo - 20]2(A-92)
~$T.trans = 0.000
(A-92)
~$T.trans = 0.000
(A-92) f3c(t - to) = [(t - to)lt11 {f3H + (t - to)lttJ ]0.3 (A-82) Creep coefficients temperature and stress $28(t,to,1) = <Pof3c(t - to) + ~$T.trans (A-86) adjusted t, days f3c(t - to) $28(t,to,1) t, days f3c(t - to) $28(t,to,1) 14 0.000 0.000 14 0.000 0.000 28 0.326 0.824 28 0.326 0.824 60 0.459 1.159 60 0.459 1.159 90 0.526 1.328 90 0.526 1.328 180 0.640 1.615 180 0.640 1.614 365 0.749 1.890 365 0.749 1.890
c) Compliance J(t,1o)= llEcmto+ ¢2B<'t,lo)IEcm28
'.
SI units in.-lb units J(t,to) = llEcmto + $28(t,to)IEcm28 (A-70) $28(t,to)IEcm28' J(t,to) (1IMPa), $28(t,to)IEcm28' J(t,to) (l/psi), t, days J(to,to)' x 10--6 x 10-6 x 10--6 t, days J(to,to)' x 10--6 x 10-6 x 10-6 14 32.90 0 32.90 14 0.227 0 0.227 28 32.90 25.74 58.65 28 0.227 0.178 0.404 60 32.90 36.20 69.10
60
0.227 0.250 . 0.47,6 90 32.90 41.49 74.39 90 0.227 0.286 0.513 180 32.90 50.44 83.34 180 0.227 0.348 0.575 365 32.90 59.04 91.94 365 0.227 0.407 0.634

Page 41
MODELING AND CALCULATING SHRINKAGE AND CREEP IN HARDENED CONCRETE 209.2R-41
Compliance temperature and stress adjusted SI units in.-Ib units J(t,to) = lIEemto(T) + ~28(t,to)/Eem28(T) (A-70) tp28(t,to)/Eem28, J(t,to) (llMPa),
~28(t,to)/Eem28' J(t,tJ (l/psi),
t, days J(to,to)' x 10-6
x 10-6
14 32.90 0 28 32.90 25.75 60 32.90 36.21 90 32.90 41.50 180 32.90 50.44 365 32.90 59.04 C.4-GL2000 model solution C.4.1 Estimated concrete properties Mean 28-day strength fcm28 = Mean 28-day elastic modulus Ecm28 = C.4.2 Estimated concrete mixture Cement type Maximum aggregate size Cement content
c=
Water content
W=
Water-cement ratio wlc= Aggregate-cement ratio ale = Fine aggregate percentage
'If =
Air content
0.=
Slump s= Unit weight of concrete
Ye=
'Table AI.5.3.7.1 and 6.3.7.1 of ACI 211.1-91.
C.4.3 Shrinkage strains Esh(t,tJ Cement type factor
x 10-6
t, days J(to,to)' x 10-6
x 10-6 x 10-6
32.90 14 0.227 0 0.227 58.65 28 0.227 0.178 0.404 69.11 60 0.227 0.250 0.476 74.40 90 0.227 0.286 0.513 83.34 180 0.227 0.348 0.575 91.94 365 0.227 0.407 0.634 SI units in.-Ib units 32.5 MPa 4689 psi (A-94) 28,014MPa 4,060,590 psi (A-95) SI units in.-Ib units I 20mm
3/4 in.
402kglm
3
676lb/yd
3
205 kglm
3
345lb/yd
3
Table 6.3.3 ACI 211.1-91 0.510 (4-1) 4.33 40% 2% Table 6.3.3 ACI211.1-91 75mm 2.95 in. 2345 kglm
3
3953lb/yd
3
1461b/ft3* SI units in.-Ib units
k=1.ooo
(Table A.14) Ultimate shrinkage strain &shu = 9OOk[30ifcm28]o.5 x 10-6 (A-99) &shu = 9OOk[4350ifem28]0.5 x 10-6 (A-99) &shu = 865 x 10-6 (A-99) &shu = 867 x 10-
6
(A-99) Ambient relative humidity factor /3(h) = (1 - 1.18h4) (A-1 00) /3(h) = 0.717 (A-1 00) Shrinkage time function
J3(t-te) = [(t-te)/{ t - te + O.12(Vlst n��.5(A-IOl) J3(t - te) = [(t - te)!{t - te + 77(VIS)2} ]0.5
(A-lOI)
Shrinkage strains &sh(t,te) = &shu/3(h)/3(t - te) (A-98) t, days /3(t - te) &sh(t,te), x 10-6 t, days /3(t - te) &sh(t,te), x 10-6 7 0.000 0 7 0.000 0 14 0.076 47 14 0.075 47 28 0.131 81 28 0.129 80 60 0.206 128 60 0.203
". 126
90 0.254 158 90 0.251 156 180 0.355 220 180 0.351 218 365 0.479 297 365 0.475 295

Page 42
209.2R-42
ACI COMMITTEE REPORT
C.4.4 Compliance J(t,to)
a) Elastic compliance J(ta,ta)
SI units in.-Ib units Cement type
s = 0.335
I (Table A.14)
f3e = exp[s/2{ 1 - (28/to)0.5}]
(A-97) Mean strength at age to
f3e = 0.933
(A-97)
fcmto = f3/fcm28
(A-96)
fcmto = 28.3 MPa
(A-96)
fcmto = 4081.1 psi
(A-96) Mean elastic modulus at
Ecmto (MPa) = 3500 + 4300(fcmto)0.5
(A-95) Ecmto (psi) = 500,000 + 52,000(fcmto)0.5 (A-95) age to
E cmto = 26,371 MPa
(A-95)
E cmto = 3,821,929 psi
(A-95)
J(to,to) = lIEcmto
(A-102) Elastic compliance
J(to,to) = 37.92 x 10-6 (l/MPa)
(A-102)
J(to,to) = 0.262 x 10-6 (l/psi)
(A-102)
b) Creep coefficient ¢2sCt,to)
SI units in.-Ib units
J(t,to) = lIEcmto + ~28(t,to)/Ecm28
(A-102) Effect of drying before loading factor Effect of drying before loading factor
<P(tc) = 0.961
(A-I04) & (A-105)
<P(tc) = 0.962
(A-I04) & (A-105) Basic creep coefficient 1st term
2[(t-to)0.3/{(t-to)0.3 + 14}]
2nd term [7 /to]0.5[ (t - to)/ {(t ~ to) + 7} ]0.5
t, days
1st term 2nd term Basic creep coefficient
t, days
1st term 2nd term Basic creep coefficient 14 0.000 0.000 0.000 14 0.000 0.000 0.000 28 0.272 0.577 0.850 28 0.272 0.577 0.850 60 0.368 0.659 1.026 60 0.368 0.659 1.026 90 0.415 0.677 1.092 90 0.415 0.677 1.092 180 0.497 0.693 1.190 180 0.497 0.693 1.190 365 0.586 0.700 1.286 365 0.586 0.700 1.286 Drying creep coefficient Ambient 2.5(1 - 1.086h2) relative humidity factor 1.170 Time function
fit, to) = [(t - to)/{ (t -to) + 0.12(V/S)2} ]0.5
Time function
fit, to) = [(t - to)/{ (t -to) + 77(V/S)2}]0.5 f(t,to)
Drying creep coefficient
f(t,to)
Drying creep coefficient
t, days
3rd term
t, days
3rd term 14 0.000 0.000 14 0.000 0.000 28 0.107 0.126 28 0.106 0.124 60 0.192 0.225 60 0.190 0.222 90 0.244 0.285 90 0.241 0.282 180 0.349 0.408 180 0.345 0.403 365 0.476 0.556 365 0.471 0.551 Creep coefficient
~28(t,to) = <P(tc) x [basic + drying creep]
(A-103)
t, days
Basic + drying creep
~28(t,to)
t, days
Basic + drying creep
~28(t,to)
14 0.000 0.000 14 0.000 0.000 28 0.975 0.937 28 0.974 0.936 60 1.251 1.203 60 1.248 1.201

Page 43
MODELING AND CALCULATING SHRINKAGE AND CREEP IN HARDENED CONCRETE 209.2R-43
90 1.377 1.324 90 1.374 1.321 180 1.598 1.536 180 1.593 1.532 365 1.843 1.771 365 1.837 1.767
SI units
in. -Ib units J(t,to) = 1!Ecmto+ 'i>2s(t,lo)/Ecm2S (A-102) J(t,to) (l/MPa), I, days J(lo,to)' x 10-6 'i>2S(I,to)/Ecm2S' x 10-6
x 10-6
J(I,to) (1!psi), I, days J(tooto)' x 10-6 'i>2s(t,to)/Ecm2S' x 10-6
x 10-6
14 37.92 0 37.92 14 0.262 0 0.262 28 37.92 33.46 71.38 28 0.262 0.231 0.492 60 37.92 42.93 80.85 60 0.262 0.296 0.557 90 37.92 47.25 85.17 90 0.262 0.325 0.587 180 37.92 54.82 92.74 180 0.262 0.377 0.639 365 37.92 63.22
101.1
365 0.262 0.435 0.697

Page 44
209.2R-44
ACI COMMITTEE REPORT
C.S-Graphical comparison of model predictions
C.S.1 Shrinkage strains Gsh(t,tc)
v~
o -.
-II- GL2000 Model
o
50 100 150 200 250 300 350
Concrete age t (days)
Fig. C.l-Shrinkage strain predictions. C.S.2 Compliance l(t,to)
110 Ii 100
IL
~ 90
~ 80
II)
e 70
u
��.
60
S
e 50
..,
CD
40
u
c
_EI- .-
.-
--..t:
-:::.. ~
----- ----
..
--~ ----
~ ---- --
~~
--~
r- ,,",
, .;-~
J~~ .....
'I
:ll!
30
Do
E
0
20
(J
10
o
14 42 70 98 126 154 182 210 238 266 294 322 350
Concrete age t (days) 1-
400
0.70
n
o 0.60 3 'g,
AI
0.50 S c.. 0.40 ~
"3
:::: i
j
0.10 .:: 0.00
---¢-AC1209 Model (1IMPa) ...•.. ACI209 Model (I/psi) _83 Model (lM'a) - _ - 83 Model (Ilpsi)
- -+- -CEB 1.C90-99Model (1IMPa) ____ GL2000 Modal (1IMPa)
-"* -CEB 1£90-99 Model (Ilpsi) - -<>- - GL2000 Modol (1/psl)
Fig. C.2-Compliance predictions.
".

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