Construction and Building Materials 25 (2011) 4202–4209

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Construction and Building Materials

journal homepage: www.elsevier .com/locate /conbui ldmat

Fatigue and rutting performance of lime-modified hot-mix asphalt mixtures

Sangyum Lee a, Sungho Mun b,⇑, Y. Richard Kim c

a Road Management Division, Seoul Metropolitan Government, Deoksugung-gil 15, Jung-gu, Seoul 100-110, South Koreab School of Civil Engineering, Seoul National University of Science & Technology, 172 Gongreung-2 Dong, Nowon-Gu, Seoul 139-743, South Koreac Department of Civil, Construction and Environmental Engineering, Campus Box 7908, North Carolina State University, Raleigh, NC 27695-7908, USA

a r t i c l e i n f o a b s t r a c t

Article history:Received 31 December 2010Received in revised form 10 March 2011Accepted 22 April 2011Available online 25 May 2011

Keywords:Lime-modified asphaltViscoelasticityContinuum damageDynamic modulusHot-mix asphalt

0950-0618/$ - see front matter � 2011 Elsevier Ltd. Adoi:10.1016/j.conbuildmat.2011.04.059

⇑ Corresponding author.E-mail addresses: smundyna@gmail.com, smun@s

This paper describes the effect of lime on composite hot-mix asphalt (HMA) mixtures of aggregate andasphalt binder evaluated using viscoelastic continuum damage analysis, which is based on predictingthe effective stress vs. strain equations and microcrack growth. The performance characteristics evalu-ated in this study included fatigue cracking and rutting resistance in both moisture-damaged and undam-aged states. The test methods used in this evaluation were the dynamic modulus test for stiffnesscharacterization, the direct tension test for fatigue cracking characterization, and the triaxial repeated-load permanent deformation test for rutting characterization. The main contribution of this paper isthe demonstration of advanced test methods and models that can be used to evaluate the performanceof various mixtures with respect to the fatigue damage and rutting performance.

� 2011 Elsevier Ltd. All rights reserved.

1. Introduction

The use of mineral filler for improving the performance ofhot-mix asphalt (HMA) mixtures has been demonstrated [1,2].Hydrated lime, in particular, has proven to be a superior additiveor filler for HMA mixtures because of its ability to improve the fa-tigue performance and moisture damage resistance. Hydrated limereacts with asphalt mixtures in a physico-chemical manner. As in-ert filler, hydrated lime physically reduces the volumetric opti-mum asphalt content by filling voids. It also improves thestability and fatigue resistance via dispersion of tiny particlesthroughout the mixture to stop cracks. In addition, it can have anactive filler effect that reduces age hardening and moisture suscep-tibility by a chemical reaction with the asphalt mixture. Moisturedamage can occur when the bond between the asphalt cementand the aggregate breaks down due to the action of moisture onthe interface of the asphalt–aggregate system and when the binderis separated from the aggregate surface.

Even though the use of hydrated lime in asphalt concrete mix-ture has been repeatedly shown to be beneficial, the benefit of limein a volumetric mix design, which may or may not change the opti-mum asphalt binder content, has still not been demonstrated.

The primary objective of this study was to use advanced testsand analysis methods to investigate the effect of lime on the fun-damental behaviour of asphalt concrete in such areas as fatiguecracking, stripping resistance, and rutting performance for various

ll rights reserved.

eoultech.ac.kr (S. Mun).

volumetric optimum asphalt content levels. This paper summa-rises findings based on the performance evaluation of differentmixtures.

The outline of the paper is as follows. Section 2 contains the de-tails of the materials and specimen fabrication. Section 3 describesthe test setup and methods used to determine the material param-eters. Section 4 presents the results of the testing and performanceevaluation. Section 5 contains the concluding remarks andsummary.

2. Materials and specimen fabrication

Table 1 shows the target gradation for the selected mixture. The mixing temper-ature was 150 �C, and the compaction temperature was 143 �C. Hydrated lime wasadded to the HMA mixture using one of the dry introduction techniques [2]. In thismethod, hydrated lime (1% by aggregate weight) was substituted for a portion ofthe baghouse fines. Mix designs for the unmodified HMA mixture (referred to asthe control mixture) and the lime-modified HMA mixture (referred to as the Lsubmixture) were developed following the standard Superpave volumetric mix designprocedure [3,4].

Optimum asphalt contents of 5.6% and 5.3% for the control and Lsub mixtures,respectively, were determined from the mix design based on Table 2. The Super-pave volumetric criteria were met for both mixtures. Table 2 refers to the voidsin mineral aggregate (VMA) and the volume of voids filled with asphalt (VFA).

Performance testing was conducted for lime-modified HMA mixtures with theoptimum asphalt content. All test specimens were compacted by a Superpave gyra-tory compactor to a height of 178 mm and a diameter of 150 mm. To obtain spec-imens of uniform quality and air void distribution, they were cut and cored tocylindrical specimens with dimensions of 100 mm in diameter and 150 mm inheight for the dynamic modulus test and TRLPD test, and 75 mm in diameter and150 mm or 140 mm in height for the constant crosshead test. The details of this pro-cess can be found elsewhere [5,6].

Table 1Gradation used for control mixture.

Sieve size % Passing

3/400 (19 mm) 100.01/200 (12.5 mm) 99.23/800 (9.5 mm) 95.3#4 (4.75 mm) 77.5#8 (2.36 mm) 59.2#16 (1.18 mm) 45.1#30 (0.600 mm) 32.5#50 (0.300 mm) 19.9#100 (0.150 mm) 12.1#200 (0.075 mm) 6.4Pan 0.0

Table 2Volumetric mix design results.

Property Control Lsub Criteria

Optimum % AC 5.6 5.3 n/a% Air voids 3.9 3.9 n/aVMA 16.2% 15.7% 15% minVFA 75.7% 75.1% 65–80%Passing #200 6.4 6.3 4.0–8.0Dust portion 1.28 1.30 0.6–1.4

S. Lee et al. / Construction and Building Materials 25 (2011) 4202–4209 4203

After obtaining specimens of the appropriate dimensions, air void measure-ments were conducted using the CoreLok method [6], and specimens were storedfor a maximum of two weeks until testing took place. During storage, specimenswere sealed in bags and placed in an unlit cabinet to reduce aging effects.

3. Test setup and methods

Both tensile and compressive loading protocols were used witha closed-loop servo-hydraulic testing machine composed of a uni-versal testing machine (UTM-25; IPC Global, Boronia, Australia)and a loading frame (MTS 810) equipped with either a 25-kN or8.9-kN load cell, depending on the nature of the test. Measure-ments of axial deformation, load, and crosshead movement forall tests were obtained via a data acquisition system consisting ofa National Instruments 16-bit data acquisition card and Labviewsoftware. In all the tests conducted during this study, axial mea-surements were taken at 90� intervals over the middle 100 mmof the specimen with loose-core linear variable differential trans-formers (IPC Global).

The purpose of the moisture conditioning process was to intro-duce a certain amount of moisture damage in the specimen prior tofatigue testing. Three parameters were controlled in the pre-condi-tioning process (moisture saturation before conditioning): mois-ture content (saturation level), conditioning temperature, andconditioning duration. Of these three parameters, the conditioningtemperature had the most significant effect on the moisture resis-tance of the asphalt mixes, followed by the conditioning durationand the moisture content.

Multiple-freeze–thaw (F–T)-cycle conditioning was conductedin accordance with AASHTO T-283 [7]. Because the fatigue testmethod used in this study was different from that in AASHTO T-283, the conditioning procedure was adjusted for appropriate sat-uration levels and post-conditioning temperatures.

A greater number of air voids in asphalt pavement generally in-creases the permeability and permits more water to penetrate. Inother words, the permeability of asphalt mixtures is proportionalto air void content when the properties and the structure of theaggregate are otherwise similar. In this study, all the mixtureshad the same aggregate and structure with a 4% air voids; this isless than the 7% air voids used for the tensile strength ratio testing.

Due to the small air void percentage, the test required three timeslonger than the maximum 10 min specified in AASHTO T-283.

The preconditioned (saturated) specimens were subjected tothree multiple-F–T cycles that consisted of freezing at �18 �C for24 h followed by thawing at 60 �C for 24 h. The conditioned spec-imens were stabilised in a water bath at 25 �C for 2 h. This proce-dure was not part of the AASHTO T-283 test but used to minimizethe weight creep of the specimens. All conditioned specimens weredried to a certain level of saturation using a core dryer to minimizethe effects of different saturation levels. Because setting up theconstant crosshead strain rate test took 18 h, the specimens werewrapped in plastic film (Parafilm) to avoid the evaporation of themoisture until the test setup was complete. The test saturation le-vel (post-conditioning) was determined based on the testing setuptime. Table 3 shows the saturation levels for pre-conditioning andpost-conditioning.

The moisture-conditioned specimens were then subjected tothe same performance tests used for the dry specimens, includingthe dynamic modulus test, triaxial repeated-load permanent defor-mation (TRLPD) test, and constant crosshead-rate direct tensiontest.

3.1. Dynamic modulus tests

The dynamic modulus test was performed in load-controlledmode under axial compression (zero maximum stress) generallyfollowing the protocol given in AASHTO TP-62 [8], as shown inFig. 1a. Tests were completed for all mixtures at �10�, 10�, 35�,and 54.4 �C and at frequencies of 25, 10, 5, 1, 0.5, 0.1, 0.05, and0.01 Hz. Load levels were determined by a trial-and-error processto ensure that the resulting strain amplitudes remained between50 and 70 microstrains. Based on the work of other researchers,this criterion was judged appropriate to ensure accurate linear vis-coelastic characterization [9,10].

3.2. TRLPD tests

TRLPD or flow number tests (Fig. 1b) were performed to as-sess the rutting potential of the asphalt concrete mixtures fol-lowing the protocol presented in NCHRP Report 465 [3]. Usingthis procedure, the specimen was subjected to a deviatoric hav-ersine stress pulse for 0.1 s and then allowed to rest for 0.9 s. Forthe mixtures presented here, only a single confining stress of138 kPa, temperature of 54.4 �C, and deviatoric stress level of827 kPa were examined. This test protocol is one of the proposedsimple performance tests used to ensure the satisfactory perfor-mance of Superpave mixtures in rutting. Note that due to limitedavailability of material, the TRLPD test was conducted on thesame specimens that were used for the dynamic modulus (|E⁄|)test. That is, the TRLPD test was performed after allowing thespecimen subjected to the |E⁄| test to restabilise. This was possi-ble because the |E⁄| test is nondestructive.

3.3. Constant crosshead-rate direct tension tests

Constant crosshead-rate tests were performed by applying aconstant rate of deformation over the complete loading train asshown in Fig. 1c. Because each component in the loading train(e.g., machine ram and load cell) deformed slightly, the on-specimen displacement rate or strain rate was not constant [11].The tests were conducted at 5 �C and at multiple rates (0.000055,0.000030, and 0.000021 s�1) for the purpose of viscoelastic contin-uum damage (VECD) characterization. These rates were fastenough not to cause any significant plastic or viscoplastic deforma-tions during loading. Constant crosshead rate monotonic directtension tests were used to generate the damage characteristic

Table 3Level of saturation for moisture-conditioned procedure.

Test type Saturation level (%)

Mix Control Lsub

Stage Pre-conditioned Post-conditioned Pre-conditioned Post-conditioned

Dynamic modulus test 1 56.0 33.3 56.7 38.42 60.2 34.8 57.6 38.03 58.3 37.3 54.2 37.9Avg. 58.2 35.1 56.2 38.1

Direct tension test 1 59.1 33.7 59.5 37.12 58.3 35.7 60.1 36.63 58.6 35.6 60.3 38.0Avg. 58.7 35.0 60.0 37.2

LVDT LVDT

LVDT

(b)(a)

(c)

Fig. 1. Testing setups: (a) dynamic modulus test, (b) triaxial repeated-load permanent deformation test, and (c) direct tension test.

4204 S. Lee et al. / Construction and Building Materials 25 (2011) 4202–4209

curve, which represents the fundamental resistance of a mixture todamage, based on the VECD model and Schapery’s work potentialtheory [12].

3.4. Determination of pseudo stiffness C(S)

The VECD model and the constant crosshead rate monotonic di-rect tension test method were used to evaluate the fatigue crackingperformance of HMA mixtures. The VECD model requires the dy-namic modulus and damage characteristic curve to predict theHMA cracking performance.

On the simplest level, the VECD model considers a damagedbody with some stiffness as an undamaged body with reducedstiffness. The VECD model thus quantifies the amount of damagethat has occurred in the sample and determines the impact of thisdamage on the material behaviour. An internal state variable Sbased on Schapery’s work potential theory [12] was used to quan-

tify the damage in the material. The relationship between S and thematerial integrity is referred to as the damage characteristic curve,or the C vs. S curve, which is discussed below.

Schapery [12] developed a theory using the method of thermo-dynamics of irreversible processes to describe the mechanicalbehaviour of elastic composite materials with growing damage.The work potential theory contains three fundamental elements:

Strain energy density function; W ¼Wðeij; SmÞ; ð1Þ

Stress—strain relationship; rij ¼@W@eij

and ð2Þ

Damage evolution law;� @W@Sm¼ @Ws

@Sm; ð3Þ

where rij and eij are stress and strain tensors, respectively; Sm rep-resents the internal state variables; and Ws = Ws(Sm) is the

Table 4Material parameters used in this study.

a (unitless) b (kPa) c (unitless)

Moisture-conditioned controlmixture

3.07692308 0.00071255 0.61141291

Control mixture 3.01204819 0.00172158 0.50582002Moisture-conditioned Lsub

mixture3.15258512 0.00111099 0.53989277

Lsub mixture 2.79485746 0.00073882 0.58084643

S. Lee et al. / Construction and Building Materials 25 (2011) 4202–4209 4205

dissipated energy due to structural changes. Using Schapery’s elas-tic–viscoelastic correspondence principle and the rate-type damageevolution law [12–15], the physical strains eij are replaced withpseudo strains eR

ij to include the effect of viscoelasticity. The corre-spondence principle proposes the extended elastic–viscoelastic cor-respondence principle, which is applicable to both linearviscoelastic (LVE) and non-LVE materials. Schapery [12,13] sug-gested that the constitutive equations for certain viscoelastic mediawere identical to those for the elastic cases, but stresses and strainswere not necessarily physical quantities in the viscoelastic body. In-stead, they were pseudo variables in the form of convolution inte-grals. According to Schapery [12,13], the pseudo strain in a uniaxialcase is defined as

eR ¼ 1ER

Z t

0Eðt � sÞ de

dsds; ð4Þ

where ER is the reference modulus and E(t) is the relaxation modu-lus. The use of pseudo strain, as defined in Eq. (4), accounts for allthe hereditary effects of the material through the convolution inte-gral. Thus, the strain energy density function W = W(e, S) becomesthe pseudo strain energy density function.

In its application to the uniaxial behaviour of HMA mixtures[14,15], the experimental stress–strain constitutive relationshipis incorporated into the one-dimensional pseudo strain energydensity function of the material in the following form,

WR ¼ 12

CðSÞðeRÞ2; ð5Þ

where the pseudo stiffness C(S) depends on a single damage param-eter S. Thus, the pseudo stiffness is a reduction factor, starting froman initial value of one when the damage parameter is zero.

The pseudo strain energy density function of the material is for-mulated using Eq. (2). The stress for the uniaxial case can be deter-mined using Eq. (5) to obtain

r ¼ @WR

@eR ¼ CðSÞeR: ð6Þ

The damage evolution law in Eq. (3) is reduced to the following sin-gle equation for S:

_S ¼ � @WR

@S

!a

; ð7Þ

where a is the material constant.The damage evolution law and experimental data are used to

characterize the function C(S) in Eq. (6). The C values can be deter-mined with Eqs. (4) and (6) using the measured stresses and calcu-lated pseudo strains. For uniaxial loading conditions, a singledamage variable S is used along with the associated power a. Usingthe experimental data, the following incremental relationship canbe obtained by combining Eqs. (5) and (7) to obtain

Siþ1 ¼ Si þ Dt �12eR

iþ1dCðSiÞ

dS

� �a

; ð8Þ

where the subscripts i and i + 1 denote the previous and currentstates, respectively.

The value of a can be found using the relationship a = 1 + 1/n, inwhich n = �log E(t)/log(t). Furthermore, this relationship is formedbetween C and S, based on a functional form,

CðSÞ ¼ expð�b � ScÞ; ð9Þ

where b and c are the parameters that are determined through thebest fit process between the functional form and the experimentaldata, as shown in Table 4. Thus, the damage characteristic curvecan be generated to explain the resistance of HMA mixtures todamage.

4. Testing and performance evaluation results

4.1. Dynamic modulus test results

Fig. 2 shows a comparison of dynamic moduli and phase anglesfor the optimum asphalt content in each control and Lsub mixturesample. Little difference in the dynamic modulus was observed be-tween the control and Lsub mixtures. The Lsub mixtures exhibiteda greater phase angle than did the control mixture for frequenciesless than 0.0001 Hz. In general, the change in phase angle slope oc-curred at lower frequencies as the asphalt content increased. Note,however, that even though the lime-modified mixtures had lowerasphalt content (5.3%) than did control mixtures (5.6%), theyexhibited a change in phase angle slope at the same frequency of0.0001 Hz. The lime thus reduced the dynamic modulus slightlyand had no significant effect on the phase angle. In general, theaddition of lime increases the stiffness of HMA mixtures, and themanner in which the lime is added is a significant factor in the de-gree of stiffness change. It appears that the lime substitution meth-od used in this study did not significantly affect the dynamicmodulus or the phase angle. Figs. 3 and 4 show a comparison ofthe average dynamic moduli and phase angles for optimum asphaltcontents of the control and Lsub mixtures with moisture damage.

The moisture conditioning in the control mixture significantlyreduced the overall frequencies of the dynamic modulus, whereasthe reduction of the dynamic modulus due to moisture condition-ing was much smaller in the Lsub mixture. The phase angles of thecontrol mixture remained relatively the same in both the mois-ture-conditioned and dry cases, whereas the moisture-conditionedLsub mixture exhibited a lower phase angle than did the dry mix-ture. This implies that the effect of lime on the moisture damagewas sufficiently positive to increase the resistance. In addition,the slope change in the phase angle of the moisture-conditionedcontrol mixture occurred at higher reduced frequencies than thatin the dry samples. In the Lsub mixtures, this occurred at similarreduced frequencies, indicating that the Lsub mixture was rela-tively less sensitive to moisture damage.

4.2. TRLPD test results

The TRLPD tests were conducted for 10,000 cycles without theoccurrence of tertiary flow. The TRLPD test results shown inFig. 5 indicate the benefits of lime modification: lower volumetricoptimum asphalt content (e.g., 5.3%) and more resistance to ruttingpotential. Fig. 6 shows that the increase in the permanent straindue to moisture conditioning was much greater in the control mix-ture than in the Lsub mixture, demonstrating the beneficial effectof lime on moisture damage in terms of permanent deformation.

Note that due to the lack of available materials, the TRLPDtests were conducted on the same specimens used for the |E⁄|tests. A certain level of densification occurred even though thespecimens were allowed to restabilise. This may have affectedthe initial stage of permanent deformation because of the sensi-tivity of this initial phase to the effects of air voids. In other

0

10000

20000

30000

40000

Reduced Frequency (Hz)

|E*|

(MPa

)

LsubControl

(a)

100

1000

10000

100000

Reduced Frequency (Hz)

|E*|

(MPa

)

LsubControl

(b)

0

10

20

30

40

1E-08 0.000001 0.0001 0.01 1 100 10000

1E-08 0.000001 0.0001 0.01 1 100 10000

1E-08 0.000001 0.0001 0.01 1 100 10000Reduced Frequency (Hz)

Phas

e an

gle

(deg

) LsubControl

(c)

Fig. 2. Effect of lime modification at different volumetric optimum asphaltcontents: (a) |E�| in semi-log space; (b) |E�| in log–log space; and (c) phase angle.

0

10000

20000

30000

40000

Reduced Frequency (Hz)

|E*|

(MPa

)

CCM

100

1000

10000

100000

Reduced Frequency (Hz)|E

*| (M

Pa)

CCM

0

10

20

30

40

1E-08 0.000001 0.0001 0.01 1 100 10000

1E-08 0.000001 0.0001 0.01 1 100 10000

1E-08 0.000001 0.0001 0.01 1 100 10000Reduced Frequency (Hz)

Phas

e A

ngle

(deg

) CCM

(b)

(c)

(a)

Fig. 3. Dynamic modulus test results for control (C) and moisture-conditionedcontrol (CM) mixtures: (a) dynamic modulus in semi-log scale; (b) dynamicmodulus in log–log scale; and (c) phase angle.

4206 S. Lee et al. / Construction and Building Materials 25 (2011) 4202–4209

words, this initial phase is the phase during which the rate ofaccumulated permanent deformation stabilises and becomesconstant. Such predensification aside, the effect of lime on therutting performance of the Lsub mixture as a function of mois-ture damage relative to the inferior performance of the controlmixture was quite obvious, as shown in Fig. 6. The initial perma-nent strain showed a much higher increase in the control mix-ture than in the Lsub mixture. In fact, for the Lsub mixture,the permanent strain growth in the moisture-conditioned speci-mens was basically the same as that in the dry specimens.

4.3. Damage-based fatigue performance evaluation

The development of a simple performance test for assessing fa-tigue performance has been less successful than the developmentof such a protocol for permanent deformation performance [3].Constant crosshead rate monotonic direct tension tests are gener-ally used to generate the damage characteristic curve. The C vs. Scurve represents the material’s resistance to damage. Thus, themixture with a higher C value for a given S value indicates that itis more resistant to damage.

In general, one must be very careful in examining such damagecharacteristic curves to assess the resistance of various mixtures tofatigue damage [16] because the damage characteristic curves onlyshow a material’s resistance to damage. The fatigue performance of

HMA, however, is affected by a mixture’s ability to resist bothdamage and deformation. The resistance to deformation is mea-sured by the dynamic modulus test, and the resistance to damageis indicated by the C vs. S curve. Both of these characteristicsshould be evaluated together.

The consideration of the modulus, damage characteristics andtheir interaction in asphalt pavements is of critical importance tosuch efforts. To meet these criteria, the asphalt concrete specimensconsisting of unmodified and lime-modified mixtures were evalu-ated based on the material properties of dynamic modulus anddamage characteristic curve C(S).

Fig. 7 shows the damage characteristic curves for the mixturesexamined in this study. Note that these curves are the average ofthe results obtained from constant crosshead-rate tests performedat 5 �C and multiple crosshead rates.

Fig. 7 shows that the lime-modified Lsub mixture with 5.3% bin-der content had greater resistance to damage than did the opti-mum unmodified control mixture. This means that the damagecharacteristic curve indicates that the lime-modified mixtureswith the optimum volumetric asphalt content in the dry stateshowed better resistance to fatigue than did the unmodified mix-tures because the dynamic modulus test results were the sameregardless of mixture type.

Fig. 8 shows the effects of lime on the performance propertieswith regard to moisture damage compared to the results for the

0

10000

20000

30000

40000

Reduced Frequency (Hz)

|E*|

(MPa

)

LsubLM

100

1000

10000

100000

Reduced Frequency (Hz)

|E*|

(MPa

)

LsubLM

0

10

20

30

40

1E-08 0.000001 0.0001 0.01 1 100 10000

1E-08 0.000001 0.0001 0.01 1 100 10000

1E-08 0.000001 0.0001 0.01 1 100 10000Reduced Frequency (Hz)

Phas

e A

ngle

(deg

) Lsub

LM

(c)

(b)

(a)

Fig. 4. Dynamic modulus test results for Lsub (Lsub) and moisture-conditionedLsub (LM) mixtures: (a) dynamic modulus in semi-log scale; (b) dynamic modulusin log–log scale; and (c) phase angle.

0

0.005

0.01

0.015

0 2000 4000 6000 8000 10000 12000No. of Cycles

Perm

anen

t Str

ain

Control 1Control 2Control 3Lsub 1Lsub 2

Fig. 5. Effect of lime modification at volumetric optimum asphalt content in TRLPDtest results of the control and Lsub mixtures.

0

0.005

0.01

0.015

No. of Cycles

Perm

anen

t Str

ain

C-1C-2C-3CM-1CM-2

0

0.005

0.01

0.015

0 2000 4000 6000 8000 10000 12000

0 2000 4000 6000 8000 10000 12000No. of Cycles

Perm

anen

t Str

ain

Lsub-1Lsub-2LM-1LM-2LM-3

(b)

(a)

Fig. 6. Effect of moisture conditioning found in TRLPD test results: (a) controlmixture (e.g., CM denotes moisture-conditioned); (b) Lsub mixture (e.g., LMdenotes moisture-conditioned).

0

0.2

0.4

0.6

0.8

1

0 100000 200000 300000 400000Damage Parameter, S

Nor

mal

ized

Pse

udo

Stiff

ness

C

ControlLsub

Fig. 7. Comparison of damage characteristics between the control and Lsubmixtures.

S. Lee et al. / Construction and Building Materials 25 (2011) 4202–4209 4207

dry state. This figure also shows the C vs. S curves before and aftermoisture conditioning for both mixtures. The resistance to damagewas significantly lower in the control mixture, whereas it wasessentially the same in the Lsub mixture.

4.4. Reduced time vs. damage function calculation based on constantcrosshead strain rates

The damage function C(S) for unmodified control and modifiedLsub mixtures was calculated and investigated under the samecondition of constant crosshead strain rates (e.g., 0.001 and0.002 s�1) to evaluate the fatigue resistance. The reduced time asa function of damage function was calculated to evaluate the fati-gue cracking performance of both types of mixtures. The followingequation was created by combining Eqs. (4) and (6) to calculate thestress history, damage parameter, and damage function outputs gi-ven the constant crosshead strain rates:

rðtÞ ¼ CðSÞER

Z t

0Eðt � sÞdeðsÞ

dsds: ð10Þ

The numerical computation based on a state variable approachcan be found in [17]. As shown in Figs. 9 and 10, the Lsub mixtureswere less sensitive to moisture under constant crosshead strainrates of 0.001 and 0.002 s�1, resulting in no significant difference

0

0.2

0.4

0.6

0.8

1

Damage Parameter, S

Pseu

do S

tiffn

ess,

C CCM

0

0.2

0.4

0.6

0.8

1

0 100000 200000 300000 400000 500000 600000

0 100000 200000 300000 400000 500000 600000Damage Parameter, S

Pseu

do S

tiffn

ess,

C LsubLM

(a)

(b)

Fig. 8. Effect of moisture conditioning on the damage characteristic curve: (a)control mixture (e.g., CM denotes moisture-conditioned); (b) Lsub mixture (e.g., LMdenotes moisture-conditioned).

0

0.2

0.4

0.6

0.8

1

1.2

Reduced Time (sec)

Dam

age

Fun

ctio

n, C

(S) Moisture-conditioned Control Mix

Moisture-conditioned Lsub MixControl MixLsub Mix

0

1000

2000

3000

4000

5000

6000

0

0 0.5 1 1.5 2

0.0005 0.001 0.0015 0.002

Strain

Stre

ss (

kPa)

Moisture-conditioned Control Mix

Moisture-conditioned Lsub Mix

Control Mix

Lsub Mix

(a)

(b)

Fig. 9. 0.001 s�1 constant crosshead strain rate: (a) reduced time vs. damagefunction, and (b) strain vs. stress history.

0

0.2

0.4

0.6

0.8

1

1.2

Reduced Time (sec)

Dam

age

Fun

ctio

n, C

(S) Moisture-conditioned Control Mix

Moisture-conditioned Lsub Mix

Control Mix

Lsub Mix

0

1000

2000

3000

4000

5000

6000

7000

0 0.5 1 1.5 2

0 0.0005 0.001 0.0015 0.002 0.0025 0.003 0.0035 0.004Strain

Stre

ss (

kPa)

Moisture-conditioned Control MixMoisture-conditioned Lsub MixControl MixLsub Mix

(a)

(b)

Fig. 10. 0.002 s�1 constant crosshead strain rate: (a) reduced time vs. damagefunction, and (b) strain vs. stress history.

4208 S. Lee et al. / Construction and Building Materials 25 (2011) 4202–4209

in terms of reduced time vs. damage function and strain vs. stress.A large difference between the intact and moisture-conditionedcontrol mixtures is reflected in the strain vs. stress plots shownin Figs. 9b and 10b, resulting in a significant softening of the mois-ture-conditioned control mixtures.

5. Conclusions

The impacts of hydrated lime on the volumetric optimum andthe fundamental behaviour and performance of HMA mixtureswere assessed. A laboratory study was conducted to evaluate theeffect of lime on the resistance to moisture damage by comparingthe performance properties of both dry (unconditioned) and mois-ture-conditioned HMA mixtures. This evaluation included |E⁄|tests, TRLPD tests, and characterization tests for the VECD modelthat combined multiple F–T conditioning cycles. The followingconclusions were drawn based on the laboratory experimentsand analyses:

� Lime made the mixture less susceptible to moisture damage asthe temperature increased because of its effect on the |E⁄|.� The TRLPD tests showed that moisture damage had less impact

on the rutting performance of the Lsub mixture than on that ofthe control mixture due to less permanent strain accumulation.� According to the VECD characterization results, lime made the

mixtures less susceptible to moisture damage because of itseffect on the fatigue performance.

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