Laboratory evaluation of fatigue characteristics of recycled asphalt mixture

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Construction and Building Materials 22 (2008) 1323–1330

and Building

MATERIALS

Laboratory evaluation of fatigue characteristics of recycledasphalt mixture

Xiang Shu, Baoshan Huang *, Dragon Vukosavljevic

Department of Civil and Environmental Engineering, The University of Tennessee, Knoxville, TN 37996, USA

Received 6 February 2007; received in revised form 12 April 2007; accepted 12 April 2007Available online 19 June 2007

Abstract

This paper presents the results of a laboratory study of evaluating the fatigue characteristics of hot-mix asphalt (HMA) mixtures usingdifferent testing methods. In this study, the fatigue performance of HMA mixtures was evaluated with the Superpave indirect tension(IDT) tests and beam fatigue test. The HMA mixtures containing 0%, 10%, 20%, and 30% of recycled asphalt pavement (RAP) wereplant prepared with one source of aggregate, limestone, and one type of binder, PG 64–22. The fatigue properties tested included indirecttensile strength (ITS), failure strain, toughness index (TI), resilient modulus, DCSEf, energy ratio, plateau value, and load cycles to fail-ure. The results from this study indicated that both Superpave IDT and beam fatigue tests agreed with each other in ranking the fatigueresistance of mixtures when proper procedures were followed.� 2007 Elsevier Ltd. All rights reserved.

Keywords: Hot-mix asphalt; Fatigue; Dissipated energy; Indirect tensile strength

1. Introduction

Fatigue cracking is one of the three major distresses(fatigue cracking, low temperature cracking, and rutting)of flexible pavements. Fatigue cracking is mainly causedby repeated traffic loading and it can lead to significantreduction in the serviceability of flexible pavements. Thecracking resistance of hot-mix asphalt (HMA) mixtures isdirectly related to the fatigue performance of flexible pave-ments. Therefore, the laboratory characterization of thefatigue behavior of HMA mixtures has been a topic ofintensive study for many years.

Many laboratory testing methods are available to char-acterize the fatigue behavior of HMA mixtures. Probablythe one that possesses the most similar stress condition toHMA field mixtures under traffic loading is the repeatedflexural test (also called beam fatigue test) [1]. This testwas developed under SHRP-A-003A to evaluate the fatigue

0950-0618/$ - see front matter � 2007 Elsevier Ltd. All rights reserved.

doi:10.1016/j.conbuildmat.2007.04.019

* Corresponding author. Tel.: +1 865 974 7713.E-mail address: [email protected] (B. Huang).

response of HMA mixtures and to summarize what isknown about the factors that influence pavement life usinga third point loading. The flexural beam fatigue test waslater modified in SHRP-A-404 to improve its simplicityand reliability.

This test uses a digitally controlled, pneumatic beamfatigue equipment, which subjects a beam specimen underrepeated stress or strain controlled loading, which isapplied at the center of the beam until failure occurs. Thefailure of the flexural fatigue test can be defined as a 50%reduction in initial stiffness, which is measured from thecenter point of the beam after 50th load cycle [1].

Recently, a new way to determine the failure of the flex-ural fatigue test was proposed by Carpenter et al. based onthe dissipated energy [2–4]. In this new method, the ratio ofdissipated energy change (RDEC) is defined as a ratio ofthe change in dissipated energy between two neighboringcycles divided by the dissipated energy of the first cycle.A plateau value (PV), or the nearly constant value ofRDEC, can be determined and it represents a period wherethere is a constant percent of input energy being turned

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Table 1Asphalt binder properties

Binder status Binder test Testresults

Specification

Original binder Rotational viscosity at135 �C, Pa Æ s

0.52 3 Pa Æ s max

DSR, G*/sind,kPa

70 �C 0.78 1.00 kPa min64 �C 1.63

RTFO agedbinder

DSR, G*/sind,kPa

70 �C 1.66 2.20 kPa min64 �C 3.54

PAV aged binder DSR, G*sind MPa, 25 �C 3725 5000 kPa maxBBR creep stiffness S,MPa

238 300.0 MPamax

BBR creep slope, m value 0.310 0.300 minPG grading 64–22

Table 2Properties of aggregates

Sievesize

Limestone D-rock

No. 10screening

Naturalsand

Manufacturedsand

5/800 100% 100% 100% 100%1/200 97% 100% 100% 100%3/800 70% 100% 100% 100%#4 21% 92% 98% 99%#8 7% 61% 93% 82%#30 4% 29% 63% 28%#50 3% 21% 13% 17%#100 2.0% 20.0% 2.0% 9.0%#200 1.8% 16.0% 1.0% 5.0%

Gsb 2.524 2.424 2.501 2.476

Note: Gsb – bulk specific gravity of aggregate.

0

20

40

60

80

100

0 2 4Sieve Size (0.45 Power)

Perc

ent P

assi

ng (%

)

AggregateRAPDesign EnvelopeMaximum Density Line

31

Fig. 1. Aggregate and RAP gradations.

1324 X. Shu et al. / Construction and Building Materials 22 (2008) 1323–1330

into damage. This PV can be used to characterize the fati-gue life of HMA mixtures. For a strain-controlled test, thelower the PV, the longer the fatigue life for a specific HMAmixture [4].

Since the 1970s, fracture mechanics theory has beenused to analyze the fatigue behavior in HMA mixtures[5]. In recent years, comprehensive laboratory and fieldstudies were conducted by Roque et al. at the Universityof Florida to characterize the crack growth rate of HMAmixtures using the Superpave indirect tension (IDT) tests[6–8]. They used the three Superpave IDT tests (IDTstrength test, resilient modulus test, and creep test) anddeveloped a viscoelastic fracture mechanics-based crackgrowth law for HMA mixtures. In addition, they intro-duced two thresholds, the dissipated creep strain energy(DCSE) limit and the fracture energy (FE) limit, to accountfor the crack development and propagation in HMA mix-tures. When these two thresholds are not exceeded, onlyhealable micro-damage occurs. Non-healable macro-dam-age appears unless one of the thresholds is exceeded. Thissuggests that the higher the values of DCSE or FE, thelonger the fatigue life of HMA mixtures [7].

In addition to the traditional fatigue approach and frac-ture mechanics approach, damage mechanics is alsoapplied to HMA mixtures to characterize their fatiguebehavior. Kim et al. developed a fatigue model for HMAmixtures using the elastic–viscoelastic correspondence prin-ciple and continuum damage mechanics [9,10]. This modelhas been successfully used to predict the fatigue life ofHMA mixtures with multiple rest periods based on thematerial’s viscoelastic properties, loading conditions, anddamage and micro-damage healing characteristics.

The objective of this study was to evaluate and comparethe fatigue performance of HMA mixtures based on theresults of different laboratory fatigue testing. In this study,HMA mixtures were plant prepared with one source ofcoarse aggregate (limestone), four percentages of recycledasphalt pavement (RAP) (0%, 10%, 20%, and 30%), andone asphalt binder (PG 64–22). The fatigue properties ofHMA mixtures were evaluated using the Superpave IDTtests and beam fatigue test.

2. Laboratory experiments

2.1. Materials

One type of asphalt binder, PG 64–22, was chosen in thestudy. Its properties are presented in Table 1.

The coarse aggregates selected in this study werecrushed limestone with a nominal maximum size of12.5 mm. The fine aggregates consisted of No. 10 screen-ings, natural sand, and manufactured sand. Their grada-tions and other properties are presented in Table 2. Allthe aggregate properties meet the specification of the Ten-nessee Department of Transportation (TDOT) [11].

The RAP used in this study was screened through theNo. 4 sieve (4.75 mm) to acquire a consistent gradation

that was comparable to the fine aggregate. RAP gradationwas determined on the bare aggregate after the binder wasextracted from RAP, as shown in Fig. 1. The asphalt con-tent of RAP was 5.5% and its maximum specific gravity(Gmm) was 2.412.

2.2. Mixture design

The Marshall mix design procedure was employed todesign the control mixture. For the HMA mixtures, 50%limestone D-rock, 15% No. 10 screenings, 25% natural

X. Shu et al. / Construction and Building Materials 22 (2008) 1323–1330 1325

sand, and 10% manufactured sand were selected. Theaggregate gradation in the experiment is shown in Fig. 1.The optimum asphalt content was 5.0%, namely asphaltfrom RAP together with the virgin asphalt was 5.0% bythe weight of the mix. Since mixtures contained differentcontents of RAP, the asphalt contribution from RAP andvirgin asphalt was different for each mixture, as presentedin Table 3. Mixture volumetric properties are listed inTable 4.

2.3. Sample preparation

In the mixing drum, virgin aggregate was superheatedand used to heat RAP added in the middle of the mixingprocess before both were mixed with liquid asphalt. Theplant-prepared HMA mixtures were collected at the mixingplant and then taken to the University of Tennessee labo-ratory for specimen preparation and fatigue testing. Cylin-drical 150-mm samples were compacted with the Superpavegyratory compactor (SGC) and then sliced into about50 mm thick specimens for the Superpave IDT tests. Theair voids for the Superpave IDT tests were 4 ± 0.5%.

2.4. Mixture performance testing

2.4.1. Superpave IDT Tests

The Superpave IDT tests include the resilient modulus,creep, and indirect tensile strength tests and they were per-formed following the procedures developed by Roque andButtlar [12,13]. Fig. 2 shows the test setup of the SuperpaveIDT tests. The testing system and associated analysis pro-cedures are described in detail by Roque and Buttlar[12,13]. It should be pointed out that the gage length wasnot exactly one quarter of the specimen diameter as sug-gested by Roque and Buttlar [12,13]. However, it is stillreasonable to use the equations developed by them to com-pare the relative fatigue performance of HMA mixtures inthis study. The tests were performed at 25 �C compared tothe 10 �C temperature used by Roque et al. [14].

Table 3Asphalt contribution from RAP and virgin asphalt for the mixtures

Mixture Asphalt contentfrom RAP (%)

Asphalt content fromvirgin asphalt (%)

0 % RAP added 0 5.510 % RAP added 0.55 4.4520 % RAP added 1.11 3.8930 % RAP added 1.66 3.34

Table 4Volumetric properties of mixture

AC (%) Gmm Gmb Air voids VMA Stability (kN) Flow (mm)

5.0 2.456 2.356 4.0 16 11.6 2.77

Note: AC, asphalt cement content; VMA, voids in mineral aggregate;Gmm, maximum specific gravity of mixture; Gmb, bulk specific gravity ofcompacted mixture.

2.4.1.1. Resilient modulus test. This test was performed onthe cylindrical samples by applying a repeated peak-loadresulting in horizontal deformations within the range of200–300 microstrains. Each load cycle consists of 0.1-s loadapplication followed by a 0.9-s rest period. The load anddeformation were continuously recorded and the resilientmodulus can be calculated as follows [12,13]:

MR ¼P �GL

DH � t � D� Ccmpl

ð1Þ

where,

MR = resilient modulus;P = maximum load;GL = gage length;DH = horizontal deformation;t = thickness of specimen;D = diameter of specimen;Ccmpl = nondimensional creep compliance factor,Ccmpl = 0.6354(X/Y)�1 � 0.332;(X/Y) = ratio of horizontal to vertical deformation.

2.4.1.2. Creep test. The creep compliance test wasperformed on the same specimen used for the resilient mod-ulus test. After allowing the specimen to re-stabilize (5–10 min) the creep compliance test was performed. Duringthis test the specimen was loaded with a constant loadfor 1000 s. The constant load was chosen such that itproduced a horizontal deformation within the range of200–750 microstrains after 1000 s of loading. The creepcompliance is calculated as follows [12,13]:

DðtÞ ¼ DH � t � D� Ccmpl

P �GLð2Þ

where, D(t) = creep compliance at time t;P, GL, DH, t, D, and Ccmpl are the same as described

above.The creep compliance D(t) can be represented using the

following power function [7]:

DðtÞ ¼ D0 þ D1tm ð3Þwhere, D0, D1 and m = parameters obtained from the creeptest.

With these two parameters, D1 and m, Roque et al. pro-posed to use the term, DCSEmin, the minimum dissipatedcreep strain energy to characterize the cracking perfor-mance of HMA mixtures. DCSEmin is expressed as follows:

DCSEmin ¼m2:98 � D1

Að4Þ

The parameter A is a function of tensile strength and tensilestress in the asphalt pavement as follows:

A ¼ 0:0299r�3:10t ð6:36� StÞ þ 2:46� 10�8 ð5Þ

where, rt = applied tensile stress of asphalt layer, St = indi-rect tensile strength.

Fig. 2. The Superpave IDT test setup.

0

0.2

0.4

0.6

0.8

1

1.2

Strain

ITS

Nor

mal

ized

p

Ap A

ε ε

ε

Fig. 3. An example of normalized IDT curve for TI calculation.


2.4.1.3. Indirect tensile strength test. The IDT strength testwas used to determine tensile strength and strain of themixture specimens compacted to 4 ± 1% air voids. Cylin-drical specimens with 152.4 mm in diameter and 50.8 mmin thickness were monotonically loaded to failure alongthe vertical diametric axis at the constant rate of76.2 mm/min. The indirect tensile strength can be calcu-lated as follows:

St ¼2� P � Csx

p� t � Dð6Þ

where,

St = indirect tensile strength;P = failure load;Csx = horizontal stress correction factor;Csx = 0.948 � 0.01114 · (t/D) � 0.2693 · m + 1.436 · (t/D) · m;m = Poisson’s ratio, m = �0.1 + 1.480 · (X/Y)2 � 0.778 ·(t/D)2 · (X/Y)2; andt, D, and (X/Y) are the same as described above.

Toughness index (TI), a parameter describing the tough-ening characteristics in the post-peak region, was also cal-culated from the indirect tensile test results. Fig. 3 presentsan example of normalized indirect tensile stress and straincurve. A dimensionless indirect tensile toughness index, TIis defined as follows:

TI ¼ Ae � Ap

e� ep

ð7Þ

where,

TI = toughness index;Ae = area under the normalized stress–strain curve up tostrain e;

Ap = area under the normalized stress–strain curve upto strain ep;e = strain at the point of interest; andep = strain corresponding to the peak stress.

This toughness index compares the performance of aspecimen with that of an elastic perfectly plastic referencematerial, for which the TI remains a constant of 1. Foran ideal brittle material with no post-peak load carryingcapacity, the value of TI equals zero. In this study, the val-ues of indirect tensile toughness index were calculated up totensile strain of 10%.

With the stress–strain response from the IDT strengthtest, the dissipated creep strain energy threshold (DCSEf)was determined by Roque et al. as follows (Fig. 4) [7]:

DCSEf ¼ FE� EE ð8Þwhere, FE = fracture energy, it is defined as the area underthe stress–strain curve to the failure strain ef, andEE = elastic energy.

MR

0 f

DCSEf EE

St

ε ε ε

σ

Fig. 4. Determination of creep strain energy threshold (DCSEf).


FE ¼Z ef

0

SðeÞde ð9Þ

EE ¼ 1

2Stðef � e0Þ ð10Þ

where, e0 can be found in Fig. 4.With DCSEf and DCSEmin, energy ratio (ER) was

defined as follows [13]:

Fig. 5. Beam fat

ER ¼ DCSEf

DCSEmin

ð11Þ

2.4.2. Beam fatigue test

The flexural beam fatigue test was a strain-controlledtest to determine the fatigue life of 38.1 cm long by5.08 cm thick and by 6.35 cm wide beam specimens sawedfrom laboratory compacted samples subjected to repeatedflexural bending until failure (AASHTO T321-03).

Beam specimens were compacted using the vibratorycompactor to 7 ± 1% air voids and tested at 25 �C accord-ing to AASHTO T321-03. Specimens were placed in abeam fatigue fixture (Fig. 5) that would allow four-pointbending with free rotation and horizontal translation atall load and reaction points using an MTS closed loopcomputer controlled data acquisition system.

A strain level of approximately 600 microstrains and aloading frequency of 10 Hz were used such that the speci-men will undergo a minimum of 10,000 load cycles. Duringeach load cycle beam deflections were measured at the cen-ter of the beam to calculate maximum tensile stress, maxi-mum tensile strain, phase angle, stiffness, dissipated energy,

igue fixture.

0.000E+00

1.000E+05

2.000E+05

3.000E+05

4.000E+05

5.000E+05

100 50,100 100,100 150,100 200,100 250,100Loading Cycles

Flex

ural

Stif

fnes

s (lb

f/in^

2)

Fig. 6. Flexural stiffness versus loading cycles.

RD

EC

Load cycles

Plateau Value (PV)

I II III

Fig. 7. Typical RDEC plot with three behavior zones (after [2]).

0

0.2

0.4

0.6

0.8

1

1.2

1.4

1.6

0%

RAP content

IDT

stre

ngth

(MPa

)

0.0

0.5

1.0

1.5

2.0

2.5

RAP content

IDT

failu

re s

trai

n (%

)

0.8

1.0

30%20%10%

0% 30%20%10%

a

b

c


and cumulative dissipated energy. Fig. 6 represents a typi-cal stiffness versus load cycle plot using an automated fati-gue software.

For the beam fatigue test, fatigue life is traditionallydefined as the number of cycles corresponding to a 50%reduction in initial stiffness and initial stiffness was mea-sured at the 50th load cycle (AASHTO T321-03). Recently,Carpenter et al. proposed to use RDEC to determine thefatigue life [2–4]. Fig. 7 presents a typical RDEC plot. Asseen from Fig. 7, the curve can be divided into three differ-ent zones. RDEC value decreases with the load cycle inzone 1. RDEC value is approximately constant in zone 2,representing a period where there is a constant percent ofinput energy turned into damage. In zone 3, RDEC valueincreases with the load cycle, indicating that more andmore input energy are turned into damage and ultimatelythe mixture loses the load carrying capability.

0.0

0.2

0.4

0.6

RAP content

IDT

TI

0% 30%20%10%

Fig. 8. Results from IDT strength test. (a) IDT strength, (b) IDT failurestrain, and (c) IDT TI.

3. Discussion of test results

3.1. Superpave IDT test results

Fig. 8 presents the strengths, failure strains, and tough-ness index (TI) values from the IDT strength test. It can beseen that mixtures containing higher percentages of RAPexhibited higher indirect tensile strength (ITS), lower strainat peak-load, and lower toughness index than the control

0.00

0.05

0.10

0.15

0.20

0.25

0.30

0%

RAP content

Ener

gy R

atio

30%20%10%

Fig. 11. Energy ratios from Superpave IDT tests.


mixture (0% RAP mixture). These phenomena can beattributed to the aged, stiffened, and embrittled asphaltbinder in RAP due to the aging process. The test resultsshowed that incorporation of RAP increased the strengthof HMA mixtures. However, due to the increase in the brit-tleness (decreased failure strain) and the decrease in TI, thefatigue life of HMA mixtures may still be compromised.These test results are consistent with those from labora-tory-prepared HMA mixtures [15].

Fig. 9 presents the resilient modulus test results. As seenfrom Fig. 9, the resilient modulus of HMA mixturesincreased with the increase in the percentage of RAP, whichmeans that incorporation of RAP increased the elastic com-ponent of the viscoelastic material – HMA mixtures. On theother hand, HMA mixtures containing RAP became morebrittle (Fig. 8b). The overall effect of RAP was disadvanta-geous to the fatigue performance of HMA mixtures.

Fig. 10 presents the test results of the dissipated creepstrain energy threshold (DCSEf). It can be seen that withthe incorporation of RAP, DCSEf values became lower,which means that the energy required to fracture HMAmixtures decreased as RAP percentage increased. Thisclearly indicated that the fatigue behavior of HMA mixturewas compromised by the incorporation of HMA mixtures.

Fig. 11 presents the energy ratios calculated from theSuperpave IDT tests. The energy ratio concept is more rea-

0

1

2

3

4

0%

RAP content

MR

(GPa

)

30%20%10%

Fig. 9. Resilient modulus results.

0.0

5.0

10.0

15.0

20.0

0%

RAP content

DC

SEf (

kJ/m

3 )

30%20%10%

Fig. 10. DCSEf results.

sonable to characterize the cracking resistance of HMAmixtures than DCSEf because it takes into account boththe energy required to fracture HMA mixtures and the dis-sipated energy accumulation in HMA mixtures under cer-tain loading condition. The energy ratios from this studywere usually lower compared to typical values obtainedby Roque et al. [13]. The reason is that a higher tempera-

0.00

0.50

1.00

1.50

2.00

2.50

3.00

0%RAP content

0

20000

40000

60000

80000

100000

120000

140000

RAP content

Load

cyc

les

to fa

ilure

30%20%10%

0% 30%20%10%

a

b

Plat

eau

Valu

e (

10-5

)

Fig. 12. Results from beam fatigue test. (a) Plateau value and (b) loadcycles to failure.


ture (25 �C) was used in this study compared to the 10 �Ctemperature used by Roque et al. [13]. The energy ratio val-ues from this study are presented in Fig. 11. It can be seenthat the energy ratio of HMA mixtures decreased with theincorporation of more RAP, which means that HMA mix-tures containing higher content of RAP were more likely tofracture than control mixture.

3.2. Beam fatigue test results

Fig. 12 presents the plateau values and load cycles tofailure based on the 50% reduction in stiffness. It can beseen that HMA mixtures incorporating more RAP exhib-ited higher plateau values (Fig. 12a), which indicated thatthey experienced more damage and would result in shorterfatigue life. However, from Fig. 12b mixtures containinghigher percentages of RAP appeared to experience longerfatigue life. Although the load cycle results were consistentwith the authors’ pervious study, it is still doubtful thatincorporation of RAP will increase the fatigue life ofHMA mixtures due to the controversy about the testmethod and the failure criterion. The plateau value methodseems more reasonable in evaluating the fatigue life ofHMA mixtures.

4. Summary and conclusions

A laboratory study was conducted to evaluate the fati-gue characteristics of HMA mixtures containing RAPusing different testing methods. HMA mixtures were plantprepared with one source of aggregate – limestone and onetype of binder – PG 64–22, containing 0%, 10%, 20%, and30% of RAP. The fatigue properties tested included indi-rect tensile strength (ITS), failure strain, toughness index(TI), resilient modulus, DCSEf, energy ratio, plateau value,and load cycles to failure. Based on the results from thestudy, the following conclusions can be summarized:

� The inclusions of RAP into HMA mixtures in this studygenerally increased the tensile strength and reduced thepost-failure tenacity in indirect tensile strength test.� The inclusions of RAP also generally decreased the dis-

sipated creep strain energy threshold and energy ratio,which may result in the short fatigue life of HMAmixtures.� Based on the failure criterion of 50% reduction in stiff-

ness, incorporation of RAP increased the fatigue lifeof HMA mixtures. Whereas using the plateau valuesfrom the beam fatigue test, inclusion of RAP wouldmake more input energy turn into damage, which mayresult in the shorter fatigue life.� The plateau value failure criterion appeared more rea-

sonable in evaluating the fatigue performance of HMAmixtures in this study.� The energy ratio concept seemed effective in the evalua-

tion of fatigue cracking behavior of HMA mixtures inthis study.

� The results presented in this paper were only the preli-minary findings of a more complete study. Further stud-ies would be needed before the relevant testing methodscan be recommended to evaluate the fatigue perfor-mance of HMA mixtures.

Acknowledgements

The authors would like to acknowledge the financialsupport from the Tennessee Department of Transporta-tion. Thanks are also due to Mr. William Gibbons andMr. Chun-Yip Chan who helped prepare testing specimens.

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Documents

Laboratory evaluation of fatigue characteristics of recycled asphalt mixture