Creep of compacted recycled asphalt pavement

Chirayus Viyanant, Ellen M. Rathje, and Alan F. Rauch

Abstract: Recycled asphalt pavement (RAP) can be beneficially used as fill in the construction of earth structures suchas embankments and retaining structures. Experiments were conducted to evaluate the creep response of compactedRAP under sustained deviatoric stresses. Constant stress, consolidated-drained triaxial tests were performed on 100 mmdiameter, compacted RAP specimens at multiple confining stresses and deviator stress levels. The test data displayedclassic creep behavior, with clearly identifiable primary and secondary creep observed in all specimens. Tertiary creepand creep rupture were observed in specimens tested at larger deviator stress levels. The creep response of RAP wassignificant at confining pressures less than about 272 kPa, while the creep response was less severe at larger confiningpressures. Upper yield stress levels, representing the deviator stress levels below which creep rupture does not occur,were identified and shown to be confining stress dependent. In general, the creep potential of RAP is significant andshould be considered in design. The developed creep models can be used to predict the time-dependent deformation ofearth structures utilizing RAP backfill.

Key words: creep, recycled asphalt pavement, time-dependent behavior, rupture.

Résumé : Les pavages d’asphalte recyclé (PAR) peuvent être utilisés avec bénéfice comme remblai dans la construc-tion de structures en terre, telles les remblais et les structures de soutènement. On a conduit des expériences pour éva-luer la réaction en fluage du PAR sous des contraintes déviatoriques soutenues. On a réalisé des essais triaxiauxconsolidés drainés à contrainte connstante sur des spécimens de PAR consolidés de 100 mm de diamètre sous plusieurscontraintes de confinement et niveaux de contraintes déviatoriques. Les résultats des essais ont montré un comporte-ment classique de fluage, avec fluage primaire et secondaire clairement identifiable dans tous les spécimens. Le fluagetertiaire et la rupture en fluage ont été observés dans des spécimens testés à de plus hauts niveaux de contrainte dévia-torique. La réaction en fluage du PAR était significative à des pressions de confinement inférieures à environ 272 kPaalors que la réaction en fluage était moins sévère à des pressions de confinement plus élevées. Les niveaux supérieursde contrainte de limite élastique, représentant les niveaux de contrainte déviatorique en decà desquels la rupture enfluage ne se produit pas, ont été identifiés et se sont révélés dépendants de la contrainte de confinement. En général, lepotentiel de fluage du PAR est significatif et devrait être pris en compte dans la conception. Les modèles de fluage dé-veloppés peuvent être utilisés pour prédire la déformation en fonction du temps des structures en terre utilisant un rem-blai de PAR.

Mots-clés : fluage, pavage d’asphalte recyclé, comportement en fonction du temps, rupture.

[Traduit par la Rédaction] Viyanant et al. 697

Introduction

Recycled asphalt pavement (RAP) is removed and (or) re-processed pavement material containing bituminous asphaltcement and aggregate. More than 73 million tons of RAP areprocessed each year in the United States (Kelly 1998) withmuch of it re-used in pavement construction. RAP is an at-tractive alternative for backfill because pavement demolitionmaterials can be re-used on site with minimal disposal costs.Additionally, in areas where select backfill is scarce, usingrecycled materials such as RAP can eliminate the need totransport select fill from a significant distance. Thus, there

are economic and environmental incentives for using RAPas backfill in new construction. However, asphalt pavementis susceptible to creep under the large cyclic loads impartedon roadways, and RAP may also display creep under thesustained shear stresses found in embankments and retainedfill. This experimental study was undertaken to quantify thecreep response of RAP under deviatoric loading and to de-velop empirical models to predict creep deformations andcreep rupture for RAP. Constant axial stress, consolidated-drainedtriaxial tests were performed on 100 mm diameter, compactedRAP specimens. Tests were performed at multiple confiningstresses and at different deviator stress levels. Axial strains

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687

Received 8 August 2006. Accepted 14 February 2007. Published on the NRC Research Press Web site at cgj.nrc.ca on 20 July2007.

C. Viyanant. Bechtel Corporation, 3000 Post Oak Boulevard, P.O. Box 2166, Houston, TX 77056-6503, USA.E.M. Rathje.1 University of Texas, Austin, TX 78712-1076, USA.A.F. Rauch. Fuller, Mossbarger, Scott & May Engineers, Inc., Lexington, KY 40511, USA.

1Corresponding author (e-mail: e.rathje@mail.utexas.edu).

were recorded with time during the creep tests until creeprupture occurred or until 1 week had elapsed.

Material description

RAP is derived from demolished asphalt pavement and isgenerated by milling or full-depth removal of asphalt pave-ment. Milling involves the mechanical removal of up to50 mm of pavement in a single pass. Full-depth removal isusually achieved with a pneumatic pavement breaker or arhino horn on a bulldozer. The broken materials are trans-ferred to a central facility for a series of recycling processes,which include crushing, screening, conveying, and stacking.Asphalt pavement can also be pulverized in place and in-corporated into granular or stabilized base courses using aself-propelled pulverizing machine (FHWA 2000), whicheliminates the cost of transporting material to and from theprocessing facility. The processing practice generally yieldsRAP with a consistent gradation.

The parent material for RAP is obviously the original as-phalt pavement. Asphalt pavement is a blend of aggregateand bituminous asphalt cement binder, with typical mixproportions ranging from 3% to 7% asphalt cement (Rob-erts et al. 1996). The processed RAP material contains ag-gregate particles that are coated with asphalt cement, suchthat asphalt cement is found at most of the grain contacts.Thus, while the properties of RAP are affected by tradi-tional geotechnical parameters such as in-place density,material gradation, and particle shape, the properties arealso affected by the presence and character of the asphaltcement binder. The asphalt cement binder is a hydrocarbonderived from the distillation of crude oil, and its properties(e.g., viscosity, ductility) are controlled by the type of vir-gin crude oil and the distillation process (Roberts et al.1996). AASHTO MP1a-04 (AASHTO 2004) provides stan-dard specifications for asphalt cement binders that arebased on achieving specific properties at specific tempera-tures, such that the appropriate asphalt cement performancegrade can be selected for the range of temperatures ex-pected in a region.

For this study, a bulk sample of RAP was obtained from aTexas Department of Transportation stockpile within theCorpus Christi District. The asphalt cement content of theRAP was estimated using a nuclear gauge, which measuresthe hydrogen content of a material (ASTM method 4125-05)(ASTM 2005a). After correcting for the water content of theRAP, the asphalt cement content was estimated as 3.5%. Itwas not possible to determine the type of asphalt cementused in the parent hot-mix asphalt for the RAP used in thisstudy, but performance grades 70–22 and 76–22 are usedmost often in Texas.

Figure 1 displays the grain size distribution of RAP sam-ples taken from four different locations in the RAP stock-pile. The grain size distribution of the RAP was veryconsistent in the stockpile. Less than 5% of the materialwas larger than 40 mm, and no particles larger than 75 mmwere observed. Only 2% of the material passes the No. 40sieve (0.425 mm), and there were no fines passing theNo. 200 sieve (0.075 mm). The Unified Soil ClassificationSystem (USCS) classification of this material is well-gradedgravel (GW). The gradation of the RAP across the stock-

pile was consistent with gradations generated by commer-cial producers of RAP (Rathje et al. 2006) and generallymeets gradation specifications for earth structures such asretaining walls.

Atterberg limit testing indicated that the RAP wasnonplastic, as the plastic limit could not be determined. Thespecific gravity (Gs) of RAP was determined by a weightedaverage of the measured values for particles larger than theNo. 4 sieve (ASTM method C127) (ASTM 2005b) and forparticles smaller than the No. 4 sieve (ASTM method D854)(ASTM 2005c). The weighted specific gravity was equal to2.33, based on measured values of 2.36 (ASTM 2005b) and2.28 (ASTM 2005c). This specific gravity is smaller thanthat for typical soil because it represents an effective specificgravity for the aggregate, asphalt cement coating, and voidsencapsulated by the asphalt cement.

Creep behavior

Deviatoric creep represents time-dependent shear defor-mations that occur under sustained shear stress (Mitchell1993). Figure 2a is a plot of axial strain versus time under aconstant deviator stress (σd = σl – σ3) for a soil experienc-ing creep. The curve displays three distinct regions of creepbehavior, primary, secondary, and tertiary creep, followed bycreep rupture (Mitchell 1993). Primary creep occurs imme-diately after application of the shear stress, and during thisstage the strain rate (slope of the strain–time curve) de-creases with time. During secondary creep, the strain ratereaches a minimum value (�εmin) and remains essentially con-stant over an extended period of time before the strain ratestarts to accelerate. This point of accelerating deformationrepresents the initiation of tertiary creep and leads to com-plete creep rupture at the end of the tertiary creep stage.

Generally, creep is a significant concern in clay soils dueto the viscous nature of clay minerals (Mitchell 1993).

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Fig. 1. Gradation of RAP used in this study.

Nonetheless, creep has also been observed in sands, al-though the resulting creep deformations tend to be smallerthan those for clay (Augustesen et al. 2004). The mechanismof creep in RAP is related to the presence of the viscous as-phalt cement at the grain contacts. Because shear loads insoil are carried by friction at the grain contacts, the presenceof the viscous asphalt cement material found at the graincontacts will affect the shear stress and deformation re-sponse of RAP. Thus, although RAP is coarse grained, itexhibits creep deformations that are on the order of thoseobserved for clays.

Various models have been developed to describe the creepbehavior of soils. These models range from simple empiricalmodels to general elasto-visco-plastic models, as discussedby Liingaard et al. (2004). Empirical models are simple inthat they are fit to experimental data in an effort to mathe-matically describe the soil behavior, but they are limited tothe loading conditions applied in the laboratory experiments.Elasto-visco-plastic models are general stress–strain–timemodels that can be applied to loading conditions beyond

those used to develop the experimental data, but they aresignificantly more complex than the empirical models. Thisstudy will focus on applying simplified empirical models tothe RAP creep data, in an effort to develop a general under-standing of the creep response of RAP as compared withother soils.

Singh and Mitchell (1968) developed a three parameterempirical model to describe the time-dependent deformationbehavior of soil during the primary and secondary stages ofcreep. The basic relationship developed by Singh and Mitch-ell (1968) to predict the time-dependent strain rate for agiven shear load is

[1] �( )ε αt Att

Dm

=

e 1

where �( )ε t is the strain rate as a function of time, D is thestress level (D = σ σd df/ , where σd is the deviator stress andσdf is the deviator stress at failure), t is time, t1 is an arbi-trary reference time, and A, α, and m are model parameters.Singh and Mitchell (1968) stated that eq. [1] is most suitablefor stress levels (D) from about 0.30 to 0.90, which arewithin the range of engineering interest. While eq. [1] hasmostly been used to describe the undrained creep-time re-sponse of clays, Singh and Mitchell (1968) indicated thatsimilar trends in creep behavior are observed in drained con-ditions and that the same functional model can be applied.

In eq. [1], the parameters A and α predict the strain rate att = t1 as a function of the stress level (D), while the parame-ter m models the change in strain rate over time. Equa-tion [1] can be integrated with respect to time to deriveexpressions for strain as a function of time. Figure 2b showsschematically the variation of strain with log(time) for m <1.0, m = 1.0, and m > 1.0. Distinctly different behaviors areobserved for these three ranges of m. For m > 1.0 the straineventually stabilizes and approaches an asymptotic value.For m = 1.0 the strain increases linearly with log(time),while for m < 1.0 the strain continuously increases with thesoil potentially reaching creep rupture. Thus, m is an impor-tant parameter for characterizing the creep potential of soils,with smaller values of m indicating high creep potential(Singh and Mitchell 1969). For clays, most m values reportedin the literature range from 0.7 to 1.2 (Augustesen et al.2004), although Singh and Mitchell (1969) indicate that val-ues as low as 0.4 are possible. For sands, less creep data isavailable, but most m values reported in the literature (e.g.,Murayama et al. 1984; Mejia et al. 1988) are close to 1.0. Iceand frozen soil are highly susceptible to creep, and m valuesof about 0.5 have been reported for these materials (Lacasseand Berre 2005). Although m is an important creep parameter,Lacasse and Berre (2005) emphasize that the actual value ofthe strain rate at a given stress level and time is also importantin understanding a material’s creep susceptibility. Because Aand α control strain rate values in the Singh and Mitchell(1968) model, these parameters should also be consideredwhen assessing the creep susceptibility of a material.

Creep rupture is defined as the failure of soil at the end oftertiary creep, where the axial strain becomes unbounded(Fig. 2a), and the time to rupture (tr) is the total elapsed timefrom the initiation of creep until final rupture. Traditionally,creep rupture has been treated separately from the Singh and

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Fig. 2. (a) Time-dependent creep deformation under a constantstress level. (b) Effect of creep parameter m on predicted creepstrains (modified from Mitchell 1993).

Mitchell (1968) creep model. Saito and Uezawa (1961) de-veloped a relationship between log(tr) and log(�εmin), where�εmin is the minimum strain rate at the end of secondarycreep. This expression can be written using a power law

[2] log(tr) = log(b) – a log(�εmin)

where a and b are experimentally determined coefficients.Based on this relationship, the time to rupture (tr) is in-versely proportional to the minimum strain rate (�εmin), suchthat soils with smaller strain rates take longer to rupture.This log-linear relationship between tr and �εmin has also beenobserved in metals (Monkman and Grant 1956), suggestingthat this relationship is applicable to a range of materialsthat are susceptible to creep. Researchers have also relatedthe logarithm of tr to the stress level, D, (e.g., Campanellaand Vaid 1974) and have shown the relationship to be in-versely proportional.

Experimental program

Constant-stress, consolidated-drained triaxial creep testswere performed on 100 mm diameter compacted specimensof RAP. During testing the axial load was adjusted to main-tain a constant axial stress as the cross-sectional area of thespecimen changed. Drained tests were performed because adrained response is expected in the field due to the largehydraulic conductivity of RAP (k ~ 0.005 cm/s, Viyanant2006). The RAP sample was scalped at 16 mm such that themaximum particle size of the RAP specimens was smallerthan one sixth of the specimen diameter. Specimens werecompacted at approximately 3% moisture content to a dryunit weight of 18.4 kN/m3, resulting in a void ratio of 0.24.The specimens were tested at room temperature at effectiveconfining pressures ( ′σ3) between 34 kPa (5 psi) and 340 kPa(50 psi), and at varying stress levels (D). Each test lasted atleast 1 week (~10 000 min) or until creep rupture occurred.A constant deviator stress was achieved throughout the1 week testing period by increasing the axial load appropri-ately to account for the increase in specimen diameter.Consolidated-drained, strain-controlled, monotonic triaxialtests were also performed at each confining stress to deter-mine the deviator stress at failure for each confining pres-sure (Fig. 3). These tests were performed at a strain rate of0.5%/min until an axial strain of about 12%–14% wasreached. At smaller confining pressures, the RAP displayeda slight peak in deviator stress that was accompanied bydilatant volumetric strain behavior. At larger confining pres-sures, the volumetric strain response was compressive andno peak was observed in the stress–strain curve. The secantfriction angles for these tests ranged from 54° at 34 kPa to40° at 340 kPa, indicating a curved failure envelope over thisrange in confining pressure.

Three series of creep tests were performed. The first testseries was performed at an effective confining pressure of136 kPa (20 psi), with ten specimens tested at stress levels(D) between 0.40 and 0.88. The second test series includedten specimens tested at an effective confining pressure of272 kPa (40 psi) and at stress levels between 0.50 and 0.90.The final test series included six specimens tested at stresslevels of 0.8 and at confining pressures of 34, 68, 136, 204,272, and 340 kPa (5, 10, 20, 30, 40, and 50 psi). Together,

these test series allow for an assessment of the creep-timeresponse of RAP as a function of stress level and confiningpressure, as well as an evaluation of the creep rupture re-sponse of RAP.

Creep-time response

Test results at ′�3 = 136 kPaFigure 4a is a plot of the measured axial creep strains (ε)

versus log(time) for the ten tests performed at an effectiveconfining pressure of 136 kPa. Tests with D ≥ 0.64 reachedthe tertiary creep stage and creep rupture within 1 week afterstress application. The creep tests performed at D < 0.64may have eventually reached creep rupture if the tests hadcontinued, but based on the axial strain rates observed at theend of 1 week, it would likely have taken more than a monthfor these specimens to reach complete creep rupture. At anygiven time, the magnitudes of the axial creep strains inFig. 4a are generally larger at the larger stress levels. Thevolumetric strains from these tests indicated compressive be-havior during the primary and secondary stages of creep, butthe specimens began to dilate as creep rupture was ap-

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Fig. 3. Results from consolidated-drained, monotonic triaxialtests on compacted RAP.

proached. Creep rupture will be discussed further in the nextsection of this paper.

The axial strain data in Fig. 4a can be differentiated withrespect to time to obtain the axial strain rate (�ε) as a functionof time (Fig. 4b). The initial strain rates are in the range of0.1%/min to 1.0%/min and larger strain rates are observedfor larger D values at any given time. The linear relationshipbetween log(�ε) and log(time) is clearly noticeable, particu-larly for D ≤ 0.70. At D ≥ 0.64, this linear behavior is followedby an increase in strain rate as creep rupture is approached.

The slopes of the axial strain rate curves in Fig. 4b repre-sent the creep parameter m in the Singh and Mitchell (1968)creep model (eq. [1]). Although eq. [1] assumes that m is in-dependent of stress level, best-fit m values were evaluatedseparately for each stress level by fitting a line through therange of strain rate data that appeared to be most linear(Viyanant 2006). In general, the m values at each stress levelare in good agreement with each other and vary between

0.67 and 0.79. The exceptions are the m values for D ≥ 0.8,which are affected by the rapid onset of tertiary creep. Thedata for D ≥ 0.8 do not follow a linear trend for a significantamount of time, and it is difficult to define accurate m val-ues for these tests. Based on the data from D < 0.8, the aver-age m value was about 0.70, a value which representssignificant creep potential (Singh and Mitchell 1969).

The creep strain data can also be used to assess the A and αcreep parameters (eq. [1]), which together predict the strainrate at the reference time (t1) as a function of D. These pa-rameters are evaluated from plots of log(�ε) versus D at the se-lected reference time, t1. Using a reference time of t1 = 5 min,the resulting A and α values were 2.8 × 10–3%/min and 6.12,respectively. Kuhn and Mitchell (1993) report A and α val-ues for soil ranging from 2 × 10–8%/min to 4 × 10–3%/minand 1.0–7.0, respectively, for t1 = 1 min. Despite the factthat a larger value of t1 was used to model RAP (5 min ver-sus 1 min), the A and α values determined for RAP are closeto the upper end of the ranges reported by Kuhn and Mitch-ell (1993), again suggesting that RAP exhibits high creepsusceptibility.

Test results at ′�3 = 272 kPaAnother test series was performed at an effective confin-

ing pressure of 272 kPa, and the axial strain rate data fromthese tests are shown in Fig. 5. The initial strain rates ob-served at ′σ3 = 272 kPa are similar to those observed at ′σ3 =136 kPa (i.e., about 0.1%/min to 1.0%/min). However, theslopes of the axial strain rate curves in Fig. 5 are larger thanthose in Fig. 4b. An average m value of 0.87 was obtainedfor the complete test series at 272 kPa, and less scatter wasobserved among the individual best-fit m values as comparedwith the ′σ3 = 136 kPa test series (Viyanant 2006). The largeraverage m value at ′σ3 = 272 kPa signifies that the strain ratedecreases more quickly with time, resulting in less creepstrain at any given time. Although m is less than 1.0 andcreep rupture is still a concern (Fig. 2b), the larger m valueindicates less creep susceptibility at ′σ3 = 272 kPa than at ′σ3 =136 kPa. Finally, the primary creep stage, during which theaxial strain rate decreases log-linearly with time, extendedfor a longer period of time for the tests performed at ′σ3 =272 kPa than for the tests at ′σ3 = 136 kPa (Fig. 5).

Again, A and α values were determined from the plot oflog(�ε) versus D at the selected reference time of t1 = 5 min.The resulting A and α values are 2.1 × 10–3%/min and 6.46,respectively. These values differ only slightly from the A andα values determined earlier for the tests at ′σ3 = 136 kPa (i.e.,2.8 × 10–3%/min and 6.12). When the ′σ3 = 136 kPa and272 kPa data are plotted together (Fig. 6), the resulting A and αvalues are 2.7 × 10–3%/min and 6.15, respectively, and thebest-fit line has an R2 of 0.95. Thus, the strain rates after theinitial application of D, which are controlled by the A and αparameters, are similar at the two confining pressures, and thedifferences in creep behavior observed with time at the two con-fining pressures are represented by the different values of m.

Test results at constant stress levelTo further investigate the effect of confining pressure on

the creep response of RAP, a series of tests was performed ata constant value of D = 0.8 at six confining pressures rang-ing from 34 to 340 kPa. A large value of D was chosen to

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Fig. 4. (a) Axial strains and (b) axial strain rates measured intests on RAP specimens at ′σ3 = 136 kPa.

ensure that most of the specimens experienced creep rupturewithin the 1 week testing period. The measured axial strainsand axial strain rates are plotted in Fig. 7. The creep behav-ior model (eq. [1]) predicts that tests performed at the sameD but at different confining pressures will produce the samestrain rate behavior. However, the data in Fig. 7 reveal dif-ferent curves for each confining stress level, indicating thatconfining pressure significantly affects the creep behavior ofRAP. However, the effect of confining pressure is not sys-tematic. The creep strains increase as confining pressure in-creases from 34 to 136 kPa, but then decrease at largerconfining pressures up to 340 kPa. Also, the tests at 34 and68 kPa initially experience small creep strains but quicklyaccelerate to creep rupture, while the tests at larger confin-ing pressures initially experience large creep strains but takemore time to reach creep rupture.

The strain rate data for the tests at each confining pressureare shown in Fig. 7b. Recall that the test series performed at136 and 272 kPa indicated that the A and α creep parameterswere not stress dependent. Thus, for tests performed at the

same stress level (D), eq. [1] predicts the same axial strainrate at t1 for all confining stress levels. Using the A and α pa-rameters derived previously in Fig. 6 (A = 2.7 × 10–3%/min,α = 6.15), eq. [1] predicts an axial strain rate of 0.37%/minat t = t1 = 5 min for D= 0.8. Considering the results inFig. 7b, the axial strain rates for ′σ3 ≥ 136 kPa all revealstrain rates between 0.22%/min and 0.32%/min at t = 5 min,which are generally consistent with the value of 0.37%/min pre-dicted by eq. [1]. However, the axial strain rate data at ′σ3 = 34 and68 kPa display much smaller values (0.08%/min to 0.1%/min) att = 5 min. The strain rate data used to derive the A and αparameters at ′σ3 = 136 and 272 kPa are plotted in Fig. 6along with the strain rate values at t = 5 min for the addi-tional confining stresses. The strain rate data for ′σ3 = 34 and68 kPa fall far below the other data, indicating that the previ-ously determined A and α parameters are not applicable atthese smaller confining pressures.

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Fig. 5. Axial strain rates measured in tests on RAP specimens at′σ3 = 272 kPa.

Fig. 6. Fit of A and α creep model parameters for t1 = 5 min.

Fig. 7. (a) Axial strains and (b) axial strain rates measured intests on RAP specimens at D = 0.8 and varying ′σ3.

The axial strain rate data in Fig. 7b were used to computebest-fit m values for each confining stress. These best-fit mvalues represent the linear portion of the plotted data at t <1000 min. The observed m values were equal to about 0.3for the tests at ′σ3 = 34 and 68 kPa, 0.6 for the tests at ′σ3 =136 and 204 kPa, and 0.8 for the tests at ′σ3 = 272 and340 kPa. These values of m indicate more creep susceptibil-ity (smaller m values) at smaller confining pressures.

The axial strain and axial strain rate data for the test per-formed at ′σ3 = 340 kPa are unlike the data from the otherconfining pressures (Fig. 7). This specimen experienced thelargest strains over the first 20 min, but thereafter the curveflattened out and did not reach creep rupture within the1 week test (Fig. 7a). An additional specimen was tested at

′σ3 = 340 kPa to confirm these data and the result was simi-lar. The strain rate data in Fig. 7b reveal that the strain ratedecreases with time most quickly at ′σ3 = 340 kPa and that theslope (i.e., m value) of the log(�ε) – log(time) curve changeswith time at this pressure. The m value at t < 1000 min isabout 0.8, increases to about 1.4 between t = 1500 min and10 000 min, and increases dramatically to about 2.6 betweent = 10 000 min and 15 000 min. This observed behavior is incontradiction with the other tests performed in this study andcontradicts various creep models. The decelerating strainrate over time (i.e., increase in the m value) might resultfrom the asphalt cement coating becoming compressed atlarger confining pressures such that there is little bitumen atthe grain contacts of the aggregate particles. Because thesedirect particle contacts are less susceptible to creep, lesscreep would be observed at higher confining pressures.However, there is no direct observation to confirm this hy-pothesis.

Figure 7 reveals two trends with respect to the effect ofincreasing confining pressure on the creep behavior of RAP.First, m generally increases with increasing confining pres-sure, which may be related to the transition from dilative(strain softening) to purely compressive (strain hardening)stress–strain behavior as confining pressure increases(Fig. 3) or may be caused by the compression of the bitumencoating at larger confining pressures. Second, the initialstrain rates at ′σ3 = 34 and 68 kPa (Fig. 7b) were signifi-cantly smaller than those observed at the larger confiningpressures, with the larger confining pressures experiencingvery similar initial strain rates. The smaller initial strainrates at the smaller confining pressures may be indicativeof an overconsolidation effect resulting from residualcompaction-induced stresses generated by the impact en-ergy used to prepare specimens.

DiscussionThe data in Figs. 4–7 reveal that the creep-time response

of RAP is complex and affected by such variables as confin-ing pressure, shape of the stress–strain curve, and, perhaps,compaction-induced stresses. The creep parameter m hasbeen used in the past (e.g., Singh and Mitchell 1969) to as-sess the creep susceptibility of soils. However, Lacasse andBerre (2005) suggest that a more complete assessment ofcreep susceptibility can be gained by considering m, A, andα because the parameters A and α control the absolute val-ues of creep strain rate. Additionally, the strain rate after ini-

tial loading can be used as an indicator of creep potential(Lacasse and Berre 2005).

The test series performed on RAP at ′σ3 = 136 kPa and272 kPa displayed average m values of 0.7 and 0.87, respec-tively. The test series at D = 0.8 revealed m values as low as0.3 at ′σ3 ≤ 68 kPa. At ′σ3 = 340 kPa and D = 0.8, the m valueinitially was about 0.8 but exceeded 1.0 after about 1500 min.As larger m values are associated with less creep susceptibil-ity, these data indicate that RAP is very susceptible to creep atsmaller confining pressures, but less so at larger confiningpressures. Table 1 lists values of m reported in the literaturefor various clays, sand, frozen soil, and ice. Most of the claysdisplay values between 0.7 and 0.9, while the materials withthe highest creep susceptibility (i.e., ice, Atchafalaya leveeclay) display values as low as 0.4. The RAP m values extendto this lower range, indicating high creep potential.

Considering the values of A and α , tests performed at′σ3 = 136 kPa and 272 kPa revealed similar strain rates as a

function of D and thus similar values of A and α. Averagevalues of A and α from these two test series were deter-mined to be 2.7 × 10–3%/min and 6.15, respectively. Thesevalues of A and α are large when compared to values re-ported for soils, indicating large initial strain rates for RAP.For example, the A value measured in the laboratory forAtchafalaya levee clay was 2.0 × 10–3%/min (Lacasse andBerre 2005). This clay is considered highly creep susceptibleand the source of creep deformations for the Atchafalayalevees (Edgers et al. 1973).

It is also important to consider the initial strain rates mea-sured for a given stress level. For D = 0.8, the initial strainrates for ′σ3 ≥ 136 kPa were between 0.5%/min and0.8%/min for RAP. Lacasse and Berre (2005) report valuesof 0.5%/min for Atchafalaya levee clay for this stress level,while highly creep susceptible ice and frozen sand displaythese types of strain rates at much smaller stress levels D(= 0.4–0.5). Thus, it appears that the creep potential ofRAP is similar to that of highly creep susceptible clays, butits creep potential is not as high as frozen soil or ice.

The initial strain rates measured during the tests per-formed at ′σ3 ≤ 68 kPa were significantly smaller than thosemeasured at larger confining pressures (i.e., 0.15%/min ver-sus 0.5%/min to 0.8%/min, Fig. 7b). Generally, smaller ini-tial strain rates indicate less creep susceptibility; however,the m values measured during these tests were as small as0.3 (Fig. 7b), and small values of m indicate high creep sus-ceptibility. Thus, it appears that RAP remains highly creepsusceptible at smaller confining pressures.

Creep rupture

Seven tests at ′σ3 = 136 kPa and five tests at ′σ3 = 272 kPaexperienced creep rupture within the testing period. Addi-tionally, several of the tests performed at D = 0.8 with vari-able confining pressures (Fig. 7) experienced creep rupture.All of the specimens that experienced creep rupture initiallyexhibited a log-linear decrease in axial strain rate with time,followed by an acceleration in strain rate leading to creeprupture. Additionally, these specimens initially experiencedcompressive volumetric strains that were followed bydilative behavior as creep rupture was approached. The initi-ation of tertiary creep and imminent creep rupture can be

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identified as the point where the axial strain rate reaches aminimum (�εmin) and starts to increase (e.g., Fig. 4b). Creeprupture was defined by the asymptotic point in the axialstrain versus log(time) plots (e.g., Fig. 4a). These data wereused to investigate models for creep rupture.

Figure 8 displays the measured values of time to rupture(tr) for RAP versus stress level, D. Previous work (e.g.,Campanella and Vaid 1974) has shown that tr is longer forsmaller stress levels. The RAP data generally agree withthese trends, but a confining-pressure dependence is also ob-served. The data at ′σ3 = 136 kPa reveal a linear relationshipbetween log(tr) and D for stress levels greater than about 0.7(Fig. 8). In this range, tr increases by approximately one or-der of magnitude as D decreases by 0.1. At D < 0.7, the datadeviate from the linear relationship such that tr is larger thanpredicted by the linear trend. At ′σ3 = 272 kPa, the tr data forRAP are larger than the values at ′σ3 = 136 kPa, and the datadisplay a larger slope. Additionally, the ′σ3 = 272 kPa datado not deviate from a linear trend over the D values tested.The rupture data at other confining stress levels from the testseries at D = 0.8 (Fig. 7) fall between the curves for ′σ3 =136 kPa and 272 kPa.

For comparison, Fig. 8 also displays the rupture data forHaney clay reported by Vaid and Campanella (1977) for un-drained conditions. As these tests were performed under un-drained conditions, this makes the comparison with RAPunder drained conditions somewhat tenuous. However, thecomparison allows one to consider the time to rupture forRAP in relation to clay. The stress levels for the Vaid andCampanella (1977) data were derived from the reported un-drained shear strength obtained from isotropically consoli-dated triaxial tests performed at a strain rate of 1%/min.Generally, these data fall within the data measured for RAP.The Haney clay data follow a linear trend at stress levelsgreater than about 0.8, but deviate at smaller stress levels.Vaid and Campanella (1977) interpreted the nonlinear rela-tionship at smaller stress levels as representing an asymptoteto an upper yield strength, which is defined as the stresslevel below which creep rupture does not occur. Based on theobserved rupture data and the fact that the specimen tested atD = 0.76 did not rupture after 3 weeks of testing, Vaid andCampanella (1977) inferred that the upper yield strength ofHaney clay was between stress levels of 0.76 and 0.78.

Finn and Snead (1973) suggested that the upper yieldstrength of a material can be evaluated from the y interceptof the relationship between stress level and the cube root ofthe minimum strain rate ( �εmin

3 ). The relationship between Dand �εmin

3 for the RAP and Haney clay creep test data are

shown in Fig. 9. A straight line through the Haney clay datahas a y intercept at D ~ 0.74, which is similar to the upperyield strength determined by Vaid and Campanella (1977).The Finn and Snead (1973) method indicates upper yieldstrengths for RAP of D ~ 0.59 at ′σ3 = 136 kPa and D ~ 0.78at ′σ3 = 272 kPa. These upper yield strength values are rea-sonable considering the data in Figs. 4 and 5, where no initi-ation of tertiary creep was detected within 1 week for creeptests with D < 0.64 at ′σ3 = 136 kPa and D < 0.8 at ′σ3 =272 kPa. An upper yield strength for RAP is significant be-cause creep rupture will not occur in RAP if the deviatoricstresses are below the measured upper yield strength.

Saito and Uezawa (1961), and others, have investigatedcreep rupture through the relationship between tr and �εmin.Figure 10 is a plot of log(tr) versus log(�εmin) for RAP. Incontrast to the relationships shown in Figs. 8 and 9, therupture data for RAP plotted versus �εmin do not show aconfining-pressure dependence. All of the rupture datafrom the four different confining pressures lie on a straightline with little scatter. Fitting eq. [2] to the data in Fig. 10with units of min for tr and min–1 for �εmin resulted in a =0.93 and b = 0.208. Table 2 lists values of a and b reportedby others for various soils, frozen soil, and ice. Generally,the a values in Table 2, which represent the slope of the re-lationship, are close to 1.0, but the b values differ by morethan an order of magnitude. For RAP, a falls within the

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694 Can. Geotech. J. Vol. 44, 2007

Material m Reference

Ice (–18 °C) 0.4–0.5 Ting (1983)Frozen sand (–18 °C) 0.75–0.85 Ting (1983)San Francisco Bay Mud 0.75 Singh and Mitchell (1969)Haney clay 0.6–0.8 Campanella and Vaid (1974)Atchafalaya levee clay 0.5–0.8 Lacasse and Berre (2005)Norwegian clays 0.8–0.9 Lacasse and Berre (2005)Sand 0.9–1.0 Augustesen et al. (2004)RAP 0.3–0.9 This study

Table 1. Creep parameter m reported for different soils.

Fig. 8. Time to creep rupture measured in tests on RAP andHaney clay.

usual range, but b is much larger than the other reportedvalues. Campanella and Vaid (1974) suggested that the bparameter is related to the soil type and consolidation his-tory, based on the different b values they measured for dif-ferent soils tested under different stress conditions.Figure 10 includes the proposed relationship from Saito andUezawa (1961) and their ±1 standard deviation curves; thesecurves generally encompass all of the relationships listed inTable 2. The RAP relationship falls above the upper boundof Saito and Uezawa (1961), indicative of RAP’s larger bvalue (Table 1). These results suggest that for a given mini-mum strain rate, it takes longer for RAP to reach creep rup-ture compared to other soils.

To explain the longer rupture life for RAP, one may con-sider the axial strain at the initiation of tertiary creep (i.e.,t �εmin

, the time at �εmin). Campanella and Vaid (1974) reportedthat the cumulative axial strain at t �εmin

for Haney clay wasthe same (about 3%) for different stress levels, which indi-cates that the time to rupture is controlled by the time ittakes to reach a critical strain value. For Haney clay, the crit-ical strain value was approximately equal to the strain at the

peak deviator stress measured in monotonic, undrained tests.Ting (1983) confirmed this observation with data fromfrozen sand and ice. For RAP, the cumulative axial strain atthe minimum strain rate is plotted in Fig. 11 as a function ofstress level. In contrast to the data from Campanella andVaid (1974) and Ting (1983), the axial strain at t �εmin

variesconsiderably with stress level and confining pressure. For agiven confining pressure, the smallest strains at t �εmin

wereobserved at larger stress levels where the tests start close tofailure. For a given stress level (D = 0.80), the axial strain att �εmin

increases significantly with increasing confining pres-sure, from about 3% at ′σ3 = 34 kPa to over 20% at ′σ3 =272 kPa. This dramatic increase in strain at the initiation oftertiary creep is related to the shape of the monotonicstress–strain curve for RAP and its dilative response at smallconfining pressures, as discussed below.

Figure 3 displayed the monotonic stress–strain curves forRAP measured at confining pressures from 34 kPa to340 kPa. The stress–strain curves at smaller confining pres-sures display a slight peak and then strain softening due tothe dilatant response, but the curves at 272 kPa and 340 kPado not exhibit a peak within the axial strains tested. Forcomparison, at larger strains Haney clay displayed a 15% re-duction from the peak deviator stress and a sensitivity be-tween 6 and 10 (Vaid and Campanella 1977). This post-peakreduction in strength lends itself to a single strain value initi-ating tertiary creep for a given confining pressure, becausecreep rupture would occur shortly after specimens reachedthe strain level associated with the peak strength. Without asignificant peak in its stress–strain curve, RAP does not fol-low the same trend. Additionally, the axial strain at whichthe maximum deviator stress occurs in RAP increases withincreasing confining pressure, an indication of the increasedductility at larger confining pressures. The shape of thestress–strain relationship, the increased ductility at largerconfining pressures, and the large values of strain required toinitiate tertiary creep (Fig. 11) may explain why the tr − �εminrelationship for RAP (Fig. 10) falls above the relationshipsreported for clays.

Summary and implications

Three series of constant stress, consolidated-drainedtriaxial creep tests were performed on compacted RAP to as-sess its creep behavior under sustained deviatoric loading.Generally, RAP displayed classical creep behavior with theaxial strain rate decreasing log linearly with time until aminimum strain rate was reached and creep rupture oc-curred. However, the experimental data suggest that thecreep-time response of RAP is complex and influenced byconfining pressure, dilative behavior, and compac-tion-induced stresses.

The creep-time response of RAP was characterized usingthe axial strain rate data during primary and secondary creepand the Singh and Mitchell (1968) creep model. At moderateconfining pressures ( ′σ3 = 136 kPa), the Singh and Mitchell(1968) creep parameters m, A, and α were determined to be0.7, 2.7 × 10–3%/min, and 6.15, respectively, for RAP.Compared with soils, these values indicate high creep sus-ceptibility, although the creep susceptibility of RAP does notappear to be as severe as for frozen soil and ice. Similar val-

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Viyanant et al. 695

Fig. 9. Evaluation of upper yield strength for RAP and Haneyclay using the Finn and Snead (1973) method.

Fig. 10. Relationship between time to rupture and minimumstrain rate for RAP and various clays.

ues of A and α were obtained for RAP at larger confiningpressures ( ′σ3 = 272 kPa), but the m value was larger (m ~0.87). The larger m value signifies a more rapid decrease instrain rate with time, and thus less creep susceptibility atlarger confining pressures. In fact, at ′σ3 = 340 kPa an mvalue greater than 1.0 was observed at t > 1500 min, whichindicates limited creep potential and no potential for creeprupture (Mitchell 1993). At smaller confining pressures ( ′σ3 ≤68 kPa), the strains and strain rates upon initial loading weresmall, but the m value was measured as 0.3. This smallvalue of m indicates high creep potential at low confiningpressures. The increased creep susceptibility at low confin-ing pressures may be due to the strain softening stress–strainbehavior of RAP at low confining pressures.

Generally, creep is more significant at large stress levels;therefore creep is a potential problem for earth structures,such as slopes and retaining walls, where the stress levels inthe soil are relatively large. These stress levels are large dueto design factors of safety that are closer to 1.0, as opposedto typical factors of safety for foundation design that arecloser to 3.0. The Singh and Mitchell (1968) creep-timemodel allows one to characterize the time dependent defor-mations of an earth structure. Using the creep model param-

eters derived in this study, a numerical analysis could beused to assess the potential for excessive creep deformationsin an earth structure constructed with RAP. However, theconfining pressure dependence of the creep-time responsewould need to be adequately represented in the analysis, per-haps by varying the creep parameters with confining stress.Nonetheless, RAP has been used in highway embankmentswith satisfactory long-term performance (TxDOT (TexasDepartment of Transportation) 1999).

Creep rupture was observed within the 1 week testing pe-riod in RAP specimens tested at larger stress levels. Ascreep rupture was approached, the axial strain rate increasedsignificantly and specimens started to dilate. The time tocreep rupture exhibited a confining-pressure dependence,with creep rupture occurring more quickly at smaller confin-ing pressures. Stress levels that represent an upper yieldstrength, below which creep rupture does not occur, wereidentified and were shown to be larger at larger confiningpressures. The axial strain values at the initiation of creeprupture were variable, with axial strains as large as 20% ob-served at the initiation of creep rupture at larger confiningpressures and smaller stress levels. Creep rupture initiated ataxial strain levels as small as 3% at low confining pressures.

Because of the large axial strains required to initiate creeprupture (Fig. 11), creep rupture does not appear to be a con-cern for RAP. Excessive deformations will most likely occurbefore creep rupture occurs, particularly at smaller stresslevels and larger confining pressures. However, at smallerconfining pressures, where less axial strain is required to ini-tiate tertiary creep, creep rupture may be a concern. In thesecases, the strength of RAP used in design calculations maybe reduced to 60% of the measured value, based on the up-per yield strengths determined from Fig. 9.

Finally, the creep properties of RAP, as well as its otherrelevant mechanical properties, will be influenced by theproperties of the parent hot mix asphalt. Most importantly,the asphalt content and asphalt cement performance gradewill affect the creep response of RAP, with ageing and aggre-gate type also having an impact. Therefore, material-specificcreep testing of RAP may be warranted for projects where po-tential creep behavior is a concern.

Conclusions

Understanding the creep response of RAP is important if itis to be used in earth structures. Based on this experimentalstudy, creep should be a concern when RAP is used ingeotechnical applications. The most critical situations will be

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696 Can. Geotech. J. Vol. 44, 2007

Material a b Source

High plasticity silt 0.92 4.63×10–2 Saito and Uezawa (1961)

Haney clay – isotropic triaxial 1.11 1.73×10–2 Campanella and Vaid (1974)Haney clay – Ko triaxial 1.09 4.04×10–3 Campanella and Vaid (1974)Ice 1.25 1.24×10–3 Ting (1983)Frozen sand (S = 40%) 0.83 4.65×10–2 Ting (1983)Frozen sand (S = 100%) 0.83 7.96×10–2 Ting (1983)RAP – isotropic triaxial 0.93 2.08×10–1 This study

Note: a and b derived for �εmin in min–1 and tr in min. saturation; S, saturation.

Table 2. Typical a and b values for eq. [2].

Fig. 11. Axial strain at the initiation of creep rupture for RAPand Haney clay.

applications where RAP experiences large stress levels thatinitiate large creep deformations and, potentially, creep rup-ture. The creep-time relationships developed here providesome guidance to design engineers in trying to predict poten-tial creep deformations of RAP. To alleviate problems relatedto creep rupture, the design engineer may consider reducingthe strength of RAP by 40% for design calculations.

Acknowledgments

Financial support was provided by the Texas Departmentof Transportation under Grant 0–4177. This support is grate-fully acknowledged. Discussions with Prof. Stephen Wrightfrom the University of Texas were particularly insightful,and comments from two anonymous reviewers improved themanuscript.

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