Comparison between thermal, rheological and failure properties for the performance grading of asphalt cements

Construction and Building Materials 220 (2019) 196–205

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Comparison between thermal, rheological and failure properties for the performance grading of asphalt cements Ahmad Nawaz Khan, Simon A.M. Hesp ⇑ Department of Chemistry, Queen’s University, Kingston, Ontario K7L 3N6, Canada

h i g h l i g h t s
Thermal properties fail to correlate with asphalt cement rheological and failure properties.
Asphalt cement ductility correlates with approximate critical crack tip opening displacement (CTOD).
Phase angles are able to differentiate asphalt cements with a high degree of precision and sensitivity.

a r t i c l e

i n f o

Article history: Received 15 February 2019 Received in revised form 3 April 2019 Accepted 30 May 2019

Keywords: Asphalt cement Performance grading Glass transition Fragility CTOD Ductility Phase angle

a b s t r a c t Thermal, rheological and failure properties of nine extracted and recovered asphalt cements were investigated to predict cracking susceptibility. Glass transition temperatures, non-isothermal Ozawa crystallization constants and binder fragilities were determined but found to lack a strong correlation with established performance indicators. All binders exhibited thermorheologically simple behavior and master curves for complex modulus and phase angle were readily constructed by applying the timetemperature superposition (TTS) principle. Failure properties of the binders were determined using ductility and double-edge-notched tension (DENT) tests. It was found that the ductility and approximate critical crack tip opening displacement (CTOD) are correlated, although the latter provides more reasonable (lower) values of strain tolerance. The Williams-Landel-Ferry (WLF) equation described the shift factors in dynamic shear, and the WLF constants C1 and C2 correlate reasonably well with ductility and CTOD. It was found that the phase angle was sensitive to asphalt cement quality, with poor performing binders showing a limiting phase angle temperature, T(d = 45°), nearly 25 °C higher than those of superior performing materials. This investigation demonstrates that recovered binders can be prematurely hardened, likely due to overheating during production or the incorrect use of reclaimed asphalt. As this problem is not recognized when testing material from the asphalt cement supply tank, it is recommended to test extracted and recovered binder from the loose hot mix asphalt. This will assure that the material as placed actually meets performance requirements for the local climate. Ó 2019 Elsevier Ltd. All rights reserved.

1. Introduction The properties of asphalt cement play an important role in designing pavements for a long lifespan. Asphalt consists of hydrocarbons with small amounts of sulphur, nitrogen, oxygen and trace Abbreviations: AASHTO, American Association of State Highway and Transportation Officials; BBR, bending beam rheometer; CTOD, critical crack tip opening displacement; DCM, dichloromethylene; DENT, double-edge-notched tension; DSR, dynamic shear rheometer; EBBR, extended bending beam rheometer; MDSC, modulated differential scanning calorimetry; MTO, Ontario Ministry of Transportation; PAV, pressure aging vessel; SHRP, Strategic Highway Research Program; TTS, time-temperature superposition; WLF, Williams-Landel-Ferry. ⇑ Corresponding author. E-mail address: [email protected] (S.A.M. Hesp). https://doi.org/10.1016/j.conbuildmat.2019.05.187 0950-0618/Ó 2019 Elsevier Ltd. All rights reserved.

metals such as vanadium, nickel and iron [1,2]. The chemistry is complex, with a heterogeneous mixture of large numbers of similar yet different molecules such as saturates, resins, aromatics and asphaltenes [3]. Consequently, development of specifications for the performance grading of asphalt is largely based on the use of physical, thermorheological and failure measurements of the material rather than a chemical analysis due to its complex chemical nature. The U.S. Strategic Highway Research Program (SHRP) developed the SuperpaveTM specification system and associated acceptance criteria in the late 1980s [1]. Superpave was implemented as American Association of State Highway and Transportation Officials (AASHTO) standard M 320 by numerous user agencies around

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North America in the mid to late 1990s [4]. The Ontario Ministry of Transportation (MTO) adopted the AASHTO M 320 acceptance criteria but soon realized that premature and excessive cracking remained problematic [5–9]. In 2001, the MTO initiated research projects to investigate the premature cracking which eventually resulted in an improved asphalt cement specification by incorporating the double-edgenotched tension (DENT) test [5,6,10] and the extended bending beam rheometer (EBBR) test [9,11,12]. Recently, researchers at MTO investigated eight 2011 paving contracts using these enhanced methods and found that both the DENT and EBBR test methods were once again able to provide significantly improved performance grading compared to the regular AASHTO M 320 specification [13,14]. A consistent finding has been that poor performing binders in service lack strain tolerance in the ductile state and lose their low temperature grade during cold storage [7,9,11,14]. Another important finding is that many pavements provide extracted and recovered asphalt cement properties that fall well short of the design criteria, suggesting there are problems with the sampling and/or laboratory aging protocols [9,15,16]. Estimates for the average error in the low temperature grade range from 4 °C to 6 °C for Ontario paving contracts, reducing life cycles by anywhere from 5 to 10 years or more. The DENT test provides an approximate critical crack tip opening displacement (CTOD) value for the strain tolerance of the binders in their ductile state [5,6]. As such it provides a measure by how much the pavement can stretch before the asphalt cement undergoes tensile damage within the mixture. In the EBBR test the low temperature grades for the asphalt binder are derived from the stiffness and relaxation ability by determining the creep stiffness (S-value) and creep rate (m-value). Binders are conditioned for 1, 24 and 72 h at 10 °C and 20 °C above the pavement design temperature. The warmest limiting temperature determines the limiting low temperature grade equal to the temperature at which the creep stiffness at 60 s loading reaches 300 MPa or the m-value reaches 0.300. The three day grade loss is determined by the difference between the limiting low temperature grades as determined in the EBBR from those obtained after just one hour of conditioning in the regular BBR protocol [7,9,11]. While the DENT test is highly correlated with fatigue cracking performance, it does require a relatively large amount of aged asphalt cement. Similarly, the EBBR is able to identify inferior asphalt binders and predict poor cracking performance in service with high accuracy but the method is time consuming and also requires large quantities of recovered binder. Hence, there is a need to find simpler test methods that can be completed in less time with less material to replace the DENT and EBBR protocols for future specifications. Modulated differential scanning calorimetry (MDSC) is a powerful technique that allows one to measure glass transition temperatures and crystallization kinetic parameters in an automated fashion on milligram quantities of material. Using different cooling rates in the MDSC allows for the determination of non-isothermal crystallization constants which have been reported in one of our previous studies to correlate strongly with low temperature EBBR results [17]. Hence, this paper investigates the use of MDSC further with the aim of potentially simplifying the DENT and EBBR based specification protocols. The dynamic shear rheometer (DSR) is an effective tool to characterize the viscoelastic behavior of asphalt binders. The complex modulus and phase angle as measured in the DSR on gram quantities of binder provide measures of stiffness and relaxation ability, the latter of which has been found to be very sensitive to changes in performance [18–22]. Recent research has reported that limiting phase angle temperatures are more than twice as sensitive to thermal cracking compared to the Superpave loss modulus (G*sind)

197

parameter and about 43% more sensitive than the EBBR grading [23]. As DSR testing of the asphalt binder is much more convenient and somewhat more accurate it deserves further evaluation. In the research described hereinafter, the performance behavior of nine recovered asphalt binders is classified based on their thermal, rheological and failure properties. For thermal characterization, glass transition temperatures (Tg) are determined as well as the non-isothermal Ozawa crystallization constants. The rheological properties are measured and master curves of complex modulus and phase angle are constructed by applying the timetemperature superposition (TTS) principle. The glass-forming ability of all binders was assessed based on their fragility index. Failure properties were determined using ductility and DENT tests which are known to separate the good from poor performing binders [9,11,24]. Finally, the limiting phase angle temperature is evaluated as it correlates with EBBR properties [23]. It is hoped that the findings of this research will help user and producers alike to select more practical test methods and acceptance criteria for asphalt cement used in road construction. 2. Materials and experimental methods 2.1. Materials The asphalt core samples were obtained from different contract locations in Ontario, Canada. Prior studies on properties of some of these and other materials and how such properties relate to field performance has been discussed [13,14,16]. Nine asphalt binders were recovered from cores provided by the user agency. All pavements ranged in age from four to five years. The binder codes, performance grades and other pertinent characteristics of the recovered binders are listed in Table 1 [13,14]. The asphalt cement was carefully extracted using dichloromethylene (DCM) solvent by soaking overnight. The aggregate was washed several times with fresh solvent after which the asphalt solution was left to sediment overnight. Binders were recovered under a nitrogen gas atmosphere and high vacuum in a rotary evaporator. Most of the solvent was evaporated between 70 °C and 90 °C, followed by treatment at 160 °C for an additional 60 ± 10 min to assure that all the solvent had been removed. None of the recovered binders were aged in the pressure aging vessel (PAV) as they had already been exposed to several years of service. After recovery the samples were analyzed according to the methods described next. Samples in Table 1 are ranked according to their DENT test performance with three showing high, three intermediate and three low CTOD values. 2.2. Experimental methods 2.2.1. Modulated differential scanning calorimetry The thermal behavior of asphalt binders was determined using a model Q2000 modulated differential scanning calorimeter (MDSC) made by TA Instruments in the United States of America. The instrument was calibrated with Indium and Zinc standards. All samples were sealed in Tzero aluminium hermetic pans and the weight of sample was kept between 10 mg and 15 mg. For the glass transition temperature, all the binders were kept at 140 °C for 10 min and subsequently quenched to 90 °C at various rates of cooling. The heating scans were recorded at 0.5 °C/min (30 °C/h) with a nitrogen gas flow rate of 50 mL/min. In the case of non-isothermal kinetic analyses, all binders were kept at 140 °C for 10 min and then cooled down to 90 °C at rates of 3, 10, 30 and 100 °C/h. Afterwards, binders were heated from 90 °C to 140 °C at a rate of 10 °C/min (600 °C/h) and heating curves were used for the analysis of the Ozawa crystallization constants.

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Table 1 Pertinent Properties of Nine Extracted and Recovered Asphalt Cements. Sample

Supply Tank PG XX-YY °C

CTOD mm

EBBR Grade °C

Three Day EBBR Loss °C

Ductility cm

we J/m2

bwp MJ/m3

A B C D E F G H I

58–40 64–34 64–34 58–34 64–34 70–28 64–34 70–28 70–28

39.3 32.8 31.2 15.8 15.7 10.4 8.2 5.6 1.4

37.0 32.0 31.0 27.0 30.0 22.0 14.0 15.0 17.0

5.0 3.4 3.9 6.6 4.4 6.4 13.8 8.3 6.2

15.7 13.9 16.1 9.3 6.3 5.2 2.7 2.2 1.4

11.8 12.9 10.8 13 11.2 3.2 8.4 16.2 14

0.82 0.58 0.52 0.64 0.44 0.18 0.29 0.8 0.1

Note: PG = performance grade, CTOD = critical crack opening displacement, EBBR = extended BBR protocol, we = essential work of failure, bwp = plastic work of failure term.

2.2.2. Dynamic shear rheometer testing A dynamic shear rheometer (DSR), model AR2000ex from TA Instruments with 8 mm diameter parallel plate geometry was used to determine all rheological properties. The frequency sweep was carried out from 0.1 rad/s to 10 rad/s in the temperature range of 34 °C to 8.0 °C at 6 °C intervals. Samples were soaked for 15 min at each temperature. A constant gap of 2 mm was kept between the parallel plates during testing. The test was performed in strain-controlled mode at 0.1% strain to keep the results in the linear viscoelastic region. The time-temperature superposition was performed using the rheometer software at a reference temperature of 2°C. Williams-Landel-Ferry (WLF) fitting was done with the rheometer software to calculate C1 and C2 for the WLF equation. 2.2.3. Ductility testing Ductility testing was carried out in a tensile testing bath manufactured by Petrotest in Germany. All binders were kept at 160 °C in an oven for at least one hour and poured into standard brass ductility molds. Prior to testing, all samples were soaked for 60 ± 10 min in the water bath at 15 °C and tested at a rate of 10 mm/min in accordance with AASHTO standard T 51–09 [25]. The ductility was determined by testing three replicates for each binder and average values are reported. 2.2.4. Double-edge-notched tension testing The double-edge-notched tension (DENT) test was developed and implemented in Ontario as laboratory standard LS-299 [10] and adopted by AASHTO as provisional standard TP 113 [26]. The DENT test was conducted in the tensile testing bath manufactured by Petrotest in Germany and three samples with separate ligament lengths were tested at once. All binders were kept at 160 °C in the oven and poured into silicon molds with 5, 10 and 15 mm ligaments. Prior to testing the samples were soaked for 60 ± 10 min in the water bath at 15 °C. Samples were tested in duplicate for a total of six force-displacement curves at 15 °C and a rate of 50 mm/min. 3. Results and discussion 3.1. Glass transition temperatures Fig. 1 shows the glass transition temperatures (Tg) for the nine recovered binders at a heating rate of 0.5 °C/min (30 °C/hour). The data are separated in three sets that are based on their overall performance in the rheological and failure tests to be discussed later. Among the nine binders, the Tg of sample A is the lowest at 32.2 °C while the highest transition temperature of 11.8 °C is for binder I. As such the glass transition agrees with the CTOD for both the best and worst performers (Table 1). The increase in the value of Tg reflects a decrease of the chain mobility in sample

Fig. 1. Glass transition temperatures (Tg) at 0.5 °C/min for (a) good quality binders, (b) fair quality binders, and (c) poor quality binders. Note: Tg at half height is indicated by arrows.

I as compared to sample A owing to the severe hardening of the binder. The rest of the binders have values in a rather narrow range of 8 °C from 17 °C to 25 °C (Table 2). The glass transition values for individual saturates, aromatics, resins and asphaltenes have been reported to be in the neighborhood of 60 °C, 15 °C, +20 °C and + 70 °C, respectively [27]. The average Tg found in all nine binders presumably reflects the chain

A.N. Khan, S.A.M. Hesp / Construction and Building Materials 220 (2019) 196–205 Table 2 Modulated Differential Scanning Calorimetry Findings.

y

Sample

Tgy °C

DC p J/g/°C

Tend °C

A B C D E F G H I

32.2 23.5 23.5 19.9 23.1 17.4 25.4 19.6 11.8

0.363 0.443 0.423 0.474 0.448 0.461 0.440 0.448 0.468

15.2 3.4 5.3 0.5 5.0 4.1 4.7 3.3 9.8

The glass transition temperature is for a heating rate of 30 °C/h (0.5 °C/min).

mobility of the aromatics blended with the saturates in the asphalt composition. It should be noted that the glass transition in asphalt can cover a rather wide range, likely due to the multiphase nature of the material [27]. In addition, the temperatures for the end of the glass transition range, Tend, are also listed in Table 2. The good performing binders like C and A have lower values of the end of the glass transition (5.3 °C and 15.2 °C) as compared to poor performing binders such as I and H (9.8 °C and 3.3 °C), respectively. However, anomalies exist for the Tend values; for instance sample G, considered a poor performer, yet displaying a Tend of 4.7 °C, which is higher than the 3.4 °C for sample B that is deemed a good performer. The values of DCP are also listed in Table 2 for all the binders and are in the range of 0.36 to 0.47 J/g/°C. On the whole, it is

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difficult to distinguish the good from the not-so-good and poor performing binders on the basis of Tg, Tend and DCP values, as their corresponding ductility and CTOD values differ widely. 3.2. Non-Isothermal kinetic analysis according to the Ozawa theory Fig. 2 shows the non-isothermal behavior of sample A by varying the cooling rate from 3 °C/h to 100 °C/h and heating rate of 600 °C/h (10 °C/min). The total heat flow and non-reversible heat flow curves are plotted against temperature to reflect the binder behavior under varying cooling rates. The trend for other binders is similar and therefore not shown here. By increasing the cooling rate from 3 °C/h to 100 °C/h, the overall heat flow curves become shallower. Similarly, for the non-reversible heat flow curves, the peak height is decreased with increasing cooling rates owing to the fact that less time is available for crystallization and/or phase separation at faster rates. Fig. 2 also provides the method of determining the enthalpy values from heat flow and non-reversible heat flow curves. For the total heat flow curve, the baseline originated from the liquid line on the high temperature side, and the area under the baseline was considered for the calculation of the melting enthalpy of waxes and other structured materials in the binders. As seen clearly, with increasing cooling rate from 3 °C/h to 100 °C/h, the area under the baseline is reduced. For the non-reversible heat flow curves, the peak area was used to determine the melting enthalpy values. Notably, the temperature range over which the enthalpy values are determined in the total heat flow and non-reversible heat flow curves varied significantly. In the total heat flow curves, the area under the baseline

Fig. 2. MDSC data of binder A for (a) total heat flow curves, and (b) non-reversible heat flow curves at various cooling rates and a heating rate of 600 °C/h (10 °C/min). The plots on the right side show how baselines were applied to calculate enthalpy values.

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is in the range of about 0 °C to +50 °C while it varied from about 50 °C to +50 °C for the non-reversible heat flow curves. The crystallized fraction, C(T), is measured using the following expression [17]:

DHðT Þ DH 1

-1.7 log(-ln(1-C(T)))

C ðT Þ ¼

-1.5

ð1Þ

DH(T) is obtained experimentally from the area under the crystallization and/or melting peak and DH1 is the total area under the peak for the dissolution enthalpy of n-alkanes. The value of DH1 used in this work was 200 J/g as suggested earlier [17]. The C(T) values are listed in Table 3 and are rather low for all binders [17]. The Ozawa theory is well known for the analysis of the nonisothermal kinetics for crystallizing systems and uses the following expression [17,28]:

logð ln½1  C ðT ÞÞ ¼ log XðTÞ  n logðbÞ

ð2Þ

where C(T) is the crystallized fraction of the system, X(T) indicates the rate of crystallization, n is the Ozawa exponent depending on the crystal dimension as well as nucleation mode and b is the cooling rate. X(T) and n are obtained from the intercept and slope of the Ozawa plot, respectively. Fig. 3 shows the Ozawa plots for binder I using the values obtained from the total heat flow and non-reversible heat flow curves. The plot illustrates the linear fit to the calculated data of the binder using Eq. (2) for the total heat flow and non-reversible heat flow curves along with the correlation coefficients (R2) of 0.99 and 0.97, respectively. The values of slope, intercept and their correlation coefficient for all nine binders are listed in Table 3. The n values of all binders obtained from the total heat flow curves varied from 0.17 to 0.28 and for the non-reversible heat flow curves it ranged from 0.33 to 0.41. Interestingly, the n values are almost 2-fold higher using non-reversible heat flow curves compared to when using the total heat flow curves. The nonreversible signal is derived from the total heat flow signal to separate processes such as oxidation, phase transitions, crystallization, and melting that overlap and/or are suppressed in the conventional total heat flow signal. Possibly, phase separation due to crystallization overlaps with the glass transition in the total heat flow signal, therefore providing relatively lower values of n as compared to the non-reversing heat flow signal. Generally, smaller values of n indicated slower rates of hardening through crystallization and vice versa. The n value for sample A is 0.17 and 0.37, whereas for sample H they are 0.22 and 0.41, as obtained from the total and non-reversible heat flow curves, respectively. It clearly indicates that sample H hardened slightly faster than sample A and this might explain why it failed in a brittle manner on the basis of its ductility and CTOD values. However, such trend cannot be generalized for all the binders. For instance, binder I also failed in a brittle manner but its n value

R² = 0.9797

-1.9 -2.1 -2.3

R² = 0.9956 -2.5 -2.7 0

0.5

1

1.5

2

2.5

log (β) Fig. 3. Ozawa plots of the recovered asphalt binder from sample I for total heat flow curves (circles), and non-reversible heat flow curves (triangles). The plots for all other binders showed similar goodness of fit.

obtained from the total heat flow is close to that of the C binder and even less than C using the non-reversible heat flow curves. Moreover, a value of n  1 infers that crystal growth is to be nearly one-dimensional in all the binders owing to the highly disordered and restricted growth of crystallized entities. Recently, it was reported that the Ozawa analysis for various binders from MTO pavement test sections correlated moderately well with the performance grading obtained from the EBBR and field performance [17]. However, two municipal binders showed nearly the same n value even though one of these was the worst performer which cracked prematurely and excessively within 4–5 years after construction while the other was a relatively good performer with little to no cracking over the same time period. Overall, it appears to be difficult to differentiate good performing binders from not-so-good performing binders on the basis of the Ozawa theory, even though it fits the crystallization kinetic data with a high degree of accuracy (Table 3). 3.3. Rheological properties and time-temperature superposition (TTS) principle Fig. 4 shows the Black space diagram of complex modulus, G*, and phase angle, d, for three representative binders (one each for good, intermediate and low performance), in the temperature and frequency range of +34 °C to 8°C and 0.1–10 rad/s, respectively. As cracking distress occurs more likely in the sub-ambient temperature range, the intermediate temperatures are investigated for these binders. All the binders illustrated thermorheologically simple behavior as indicated by the smooth overlap of the data in this temperature and frequency range. At lower temperatures, poor performers such as binders G, H and I contained higher

Table 3 Ozawa Crystallization Theory Findings. Sample

y

Total Heat Flow, W/g

Non-Reversible Heat Flow, W/g

Slope

Intercept

R2

C(T), %y

Slope

Intercept

R2

C(T), %y

A B C

0.17 0.2 0.2

2.23 2.15 2.13

0.98 0.98 0.98

0.33 0.37 0.35

0.37 0.37 0.39

1.64 1.58 1.59

0.95 0.98 0.99

0.73 0.73 0.71

D E F

0.2 0.18 0.23

2.14 2.07 2.01

1.00 0.98 0.99

0.36 0.45 0.46

0.34 0.33 0.38

1.68 1.68 1.60

0.95 1.00 0.99

0.73 0.65 0.65

G H I

0.28 0.22 0.21

1.86 1.97 2.12

0.98 1.00 0.97

0.51 0.51 0.39

0.34 0.41 0.34

1.65 1.65 1.71

0.97 0.98 0.97

0.77 0.50 0.65

The crystallized fractions C(T) are for a heating rate of 30 °C/h (0.5 °C/min).

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Fig. 4. Black space diagrams for the three recovered asphalt binders (one each for high, intermediate and low performance).

complex moduli along with phase angles below 20°, inferring higher elastic behavior as compared to the other materials. At high temperatures, good performers such as A, B and C exhibited divergence from linearity in comparison to the other binders, indicating the likely presence of more SBS polymer modifier that may have phase separated at higher temperatures. Generally, binders with lower phase angles and higher complex moduli will be prone to cracking as these binders are unable to relax thermal stresses at low temperatures. The frequency sweep data for the nine binders at intermediate temperatures was used to construct the master curve by applying the time-temperature superposition (TTS) principle. Fig. 5 depicts the master curves of the same three representative binders for complex modulus and phase angle using a reference temperature

of 2.0 °C. At lower frequencies or higher temperatures, the difference of complex modulus values among the binders is very obvious compared to their values at higher frequencies and/or lower temperatures. Notably, along the whole frequency range, the value of the complex modulus was lowest for good performing binders AC, and highest for poor performers G-I. It clearly indicates that poor performing binders are more age hardened than good performing binders. Moreover, over the entire frequency range the phase angle for the poor performers is lower than what it is for the fair and good performers. The phase angle difference is as high as 10° to 15° between poor and good performing binders at higher frequencies and/or lower temperatures. It is important to note that fatigue and thermal loading are more likely to generate cracking in the pavement at lower temperatures and higher frequencies. It has been reported by a number of research groups that binders with phase angles above 28° show little cracking while those with a phase angle below 28° crack prematurely and severely [20 (7.8 Hz), 22 (0.1 rad/s), 16 (10 rad/s)]. Others have correlated the limiting m-value of 0.3 in the BBR to a phase angle of 26.2° in the DSR [[29] (10 rad/s)]. However, our work has clearly shown that the BBR m-value and DSR phase angle are not well correlated [16,22]. Thus, binders G-I with phase angles below 25° and higher stiffness showed more brittle failure compared to the binders with higher phase angles and lower stiffness. Fig. 6 illustrates the temperature dependence of the shift factor (aT) for the three representative binders by using the WilliamsLandel-Ferry (WLF) equation [30,31]:

log aT ¼

C 1 ðT  T ref Þ C 2 þ ðT  T ref Þ

ð3Þ

where C1 and C2 are the WLF material constants as listed in Table 4. For good performing binders such as A-C, the C1 and C2 values are in

Fig. 6. Temperature dependence of the shift factor for the recovered asphalt binders.

Table 4 Williams-Landel-Ferry Constants.

Fig. 5. Master curves of recovered asphalt binders for (a) complex modulus, and (b) phase angle. (One each for high, intermediate and low performance).

Sample

C1

C2

A B C

24 22 22

158 145 140

D E F

25 27 29

163 171 178

G H I

35 40 35

219 251 212

Note: C1 and C2 were determined with the rheometer software.

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the range of 22–24 and 140–158, respectively. Whereas, for poor performing binders such as G-I, C1 and C2 values are around 35– 40 and 212–251, respectively, which is considerably higher. This observation implies that stress relaxation occurs faster in good performing binders compared to poor performing binders. Hence, good performers will be able to heal cracks generated by thermal stress more readily than poor performers. Additionally, Fig. 7 shows the correlation of C1 and C2 with ductility and CTOD for all nine binders. Overall, both C1 and C2 show a reasonably good correlation with ductility and CTOD. However, C1 exhibited a slightly higher correlation compared to C2. The dynamic fragility (mf) was also determined which is a measure of the rate of change in the liquid’s dynamic properties when the glass transition temperature is approached [32–34] and quantified as follows:

2 3 dlog s m ¼ 4
5 T d Tg

T¼T g

2 3 dlog a T ¼ 4
5 T d Tg

ð4Þ

T¼T g

Using the WLF parameters when Tref – Tg, mf can be expressed as follows [35]:

T g C1C2 mf ¼  2 C 2 þ T g  T ref

ð5Þ

Usually, a lower value of mf corresponds to a strong glass forming liquid, which follows a nearly Arrhenius temperature dependence of the relaxation properties, while higher values

(a) 45 Duclity/CTOD, cm/mm

40 35 30 R² = 0.677

25 20 15 10 R² = 0.8782

5 0 15

25

35

45

C1 (b) 45 Duclity/CTOD, cm/mm

40

correspond to a fragile liquid displaying Williams-Landel-Ferry (WLF) or Vogel-Fulcher-Tamman (VFT) behavior. Moreover, the fragility of polymers covers a wide range of values from about 40 to over 200 [36]. In general, polymers with stiff backbone structures have higher fragility values in comparison to polymers with flexible backbones. The dynamic fragility of nine binders was calculated using Eq. (5) and the values are listed in Table 5. The mf value for the nine asphalt binders varies over a rather narrow range from 46 to 55, hence all binders can be considered strong glass formers. These values are similar to the mf values reported by others for asphalt binders [37]. Moreover, polymers like polyisobutylene and polyethylene also have mf  46, providing relatively strong glass forming tendencies because of their flexible chain structure and weak intermolecular interactions [38]. In the case of asphalt, weak intermolecular interactions exist among the different combinations of flexible and rigid chain structures present in saturates, aromatics, resins and asphaltenes, thereby providing relatively low mf values. Activation energy (Ea) was calculated using the WLF approach as follows [31]:

Ea ¼ R

dlnaT  d 1T

"

!

¼ 2:303R 

C 1 C 2 T 2g C 2 þ T  T ref

# ð6Þ

2

where R is the universal gas constant. Ea may be considered as an indicator of the energy required to mobilize the polymer chains. The values of Ea are listed in Table 5 for all the binders and varied from 224 to 263 kJ/mol, which is similar to the Ea values of asphalt binders reported elsewhere [37]. Interestingly, good performers such as samples C, B and A displayed an almost 20–40 kJ/mol higher activation energy in comparison to poor performing binders such as G-I. Fig. 8 shows the limiting phase angle temperature (Td) for phase angles reaching 25°, 35° and 45° for the nine binders at 10 rad/s. It has been observed previously that limiting phase angle temperatures in the range of 27–29° provide a better separation among poor and good performers in terms of cracking [16,20,22]. Moreover, at d = 45°, the elastic storage modulus is equivalent to viscous loss modulus and the limiting temperature at this point may give a measure of gelation onset in the binders. Gelation is considered when tan d is independent of frequency. Clearly, the limiting temperature is much higher for the poor performers such as G–I compared to the good performers A–C. The T(d = 25°) ranges from 13 °C to 20 °C while the T(d = 45°) ranges from 15 °C to 25 °C higher in poor performers compared to good performers. It is clearly demonstrated that poor performing binders are severely hardened during production and/or early service and thus will be more prone to thermal distress. Notably, the results on the basis of the WLF constants and limiting phase angle temperatures support the findings

35 30 25

Table 5 Fragility Constants and Activation Energies.

R² = 0.6023

20 15 10 5

R² = 0.8208

0 120

170

220

270

C2 Fig. 7. Correlation of (a) C1 with ductility and CTOD, and (b) C2 with ductility and CTOD. Note: Open symbols are for ductility while filled symbols are for CTOD correlations.

Sample

mf

Ea kJ/mol

A B C

55 53 55

251 252 263

D E

50 51

242 242

F G H I

51 49 46 47

247 234 224 234

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203

Double-edge-notched tension (DENT) testing was carried out to correlate the findings with ductility results. This test is conducted using 5 mm, 10 mm and 15 mm ligaments by varying the notch depths, allowing for the extrapolation to zero ligament length. The extrapolation determines the specific work of failure from which an essential work of failure (we) is calculated by dividing the ligament cross sectional area. The essential work is then used to calculate the approximate critical crack opening displacement (CTOD) [5,6,10,41]. The CTOD is generally considered to be related to the amount by which a thin fibre of asphalt binder can be stretched under severe constraint in the ductile state until failure. Fig. 9 shows the force-displacement curves at 5 mm and 15 mm ligaments for the nine binders. Binders A, B and C fail in a ductile manner and display the highest displacements among all the binders at all ligaments along with a lower maximum force. In contrast, binders G-I exhibited the lowest displacement at very high peak loads, suggesting that these could fail in a brittle manner when incorporated in the pavement. The data obtained from the force-displacement curves is used to calculate the total work of failure as follows:

W Total ¼ W essential þ W plastic

ð7Þ

It was Cotterell and Reddel [42] who first proposed in a rather simple manner that the essential work scales with the crosssectional area of the ligament (L  B) and the plastic work with the volume of the ligament (b  L2  B), where b is a shape factor for the plastic zone around the fracture surfaces. Mathematically, the relationship can be written as follows:

W Total ¼ wt  LB ¼ we  LB þ wP  bL2 B

ð8Þ

Dividing Eq. (8) by the cross-sectional area (LB) on either side gives the essential work of failure expression in the following manner:

wt ¼ we þ bwP  L

Fig. 8. Limiting phase angle temperature for the nine recovered asphalt binders.

obtained from ductility and DENT testing as discussed next by distinguishing the poor quality binders from the good quality binders.

3.4. Comparison of ductility and double-edge-notched tension test results The ductility of the nine binders was determined at 15 °C using a strain rate of 10 mm/min. The findings for this test are listed in Table 1. Binders A–C displayed higher ductility values of 161, 139 and 157 mm, respectively. While binders D, E and F gave values of 93, 63 and 52 mm, respectively. The lowest ductilities were obtained for G, I and H binders in the range of 27, 14 and 22 mm, respectively. It has been reported that pavements remain in relatively good condition if the ductility measures above 100 mm [39]. Others have reported that when ductility is reduced further the pavement starts to show loss of fines, raveling, cracking, and extensive cracking that needs resurfacing [40]. On the basis of the ductility values for these nine binders, it appears that A–C are the best performers and G–I are the worst performers.

ð9Þ

where wt is the specific total work of failure (J/m2), we is the specific essential work of failure (J/m2), and bwP is the specific plastic work of failure (MJ/m3). The essential work term (we) is typically divided by the average net-section peak stress for the 5 mm ligament specimen to obtain an approximate value for the critical crack tip opening displacement (CTOD) in the following manner:

CTOD ¼

we

rnet;5mm

ð10Þ

where, CTOD (mm) is a measure of strain tolerance in the ductile state, similar to ductility. The CTOD was determined for all the nine binders and their values are listed in Table 1. Fig. 10 shows the correlation of CTOD with ductility in which the binders exhibiting higher ductility also have higher CTOD values and vice versa. The G-I samples failed to achieve the passing criteria in terms of CTOD values, also have the lowest ductility values among all the binders. Moreover, the values of we and bwp are also listed in Table 1. For identical CTOD values like in D and E, both we and bwp differ a great deal, indicating the variation of chemical composition in these binders. Additionally, sample I shows an anomaly in its values which could be due to its high brittle failure as reflected in the forcedisplacement curve. In general, the binders with the lowest values for CTOD and ductility would reveal lower pavement performance at cold temperatures due to the higher thermal distresses as compared to the binders with higher CTOD and ductility values.

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Fig. 9. Force-displacement curves at (a) 5-mm ligaments and (b) 15-mm ligaments.

4. Summary and conclusions Nine asphalt binders were recovered from 4 to 5 year old pavements at various locations in Ontario, Canada. Their performance was differentiated using MDSC, DSR, ductility and DENT tests. The following conclusions are provided:

Fig. 10. Correlation of the values obtained from ductility and CTOD.


Thermal analysis provided accurate measures for the glass transition and the non-isothermal Ozawa crystallization parameters but neither of these showed a high correlation with established performance indicators such as CTOD and ductility.
The rheological phase angle, d, showed a reasonable correlation with failure properties such as CTOD and ductility ranking them similarly.
Ductility and CTOD are highly correlated with each other but the latter provides more reasonable (lower) values. Testing at

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lower temperatures and with higher constraint (deeper notches) will likely provide better performance ranking.
A significant portion of the binders investigated would fail current acceptance criteria and thus there are problems with the specification acceptance criteria as implemented. Considering the practicality of all the test methods it is recommended to investigate the phase angle and complex modulus further for performance ranking of asphalt binders. These properties are readily measured with a high degree of precision on a small amount of material and this would benefit both users and producers of asphalt cement. Given the pervasiveness of premature and excessive cracking in asphalt pavements in cold climates it is up to the user agencies to implement improved specification test methods and acceptance protocols. Testing extracted and recovered materials would likely reduce the use of materials that are overheated and/or contaminated with reclaimed asphalt pavement (RAP) during production, as this is a likely cause for the poor performance of a significant number of binders. Declaration of Competing Interest None. Acknowledgements Special appreciation goes out to the Ontario Ministry of Transportation, Imperial Oil of Canada and the Natural Sciences and Engineering Research Council of Canada for the financial support towards this research. Staff of the ministry is thanked for their help with the collection of samples. References [1] Asphalt Institute, Superpave performance graded asphalt binder specification and testing Superpave series No 1 (SP-1), 3rd Ed., Asphalt Institute, 2003. [2] P. Redelius, H. Soenen, Relation between bitumen chemistry and performance, Fuel 140 (2015) 34–43. [3] D. Lesueur, The colloidal structure of bitumen: consequences on the rheology and on the mechanisms of bitumen modification, Adv. Colloid Interface Sci. 145 (2009) 42–82. [4] D.A. Anderson, D.W. Christensen, H.U. Bahia, R. Dongre, M.G. Sharma, C.E. Antle, J. Button, Binder characterization and evaluation (SHRP A-369 Physical Characterization), National Research Council, Washington D.C., 1994. [5] A. Andriescu, S.A.M. Hesp, J.S. Youtcheff, Essential and plastic works of ductile fracture in asphalt binders, Transp. Res. Rec. J. Transp. Res. Board 1875 (2004) 1–7. [6] A. Andriescu, S. Iliuta, S.A.M. Hesp, J.S. Youtcheff, Essential and plastic works of ductile fracture in asphalt binders and mixtures, Proc. Can. Tech. Asphalt Assoc. 49 (2004) 93–121. [7] S. Iliuta, A. Andriescu, S. Hesp, K. Tam, Improved approach to low-temperature and fatigue fracture performance grading of asphalt, Proc. Can. Tech. Asphalt Assoc. 49 (2004) 123–158. [8] P. Yee, B. Aida, S. Hesp, P. Marks, K. Tam, Analysis of premature low temperature cracking in three Ontario, Canada, pavements, Transp. Res. Rec. J. Transp. Res. Board No. 2006 (1962) 44–51. [9] S.A.M. Hesp, A. Soleimani, S. Subramani, T. Phillips, D. Smith, P. Marks, K.K. Tam, Asphalt pavement cracking: analysis of extraordinary life cycle variability in eastern and northeastern Ontario, Int. J. Pavement Eng. 10 (3) (2009) 209– 227. [10] Ministry of Transportation of Ontario, ‘‘Determination of Asphalt Cement’s Resistance to Ductile Failure Using DENT,” LS-299, Rev. 29, MTO Laboratory Testing Manual, 2012. [11] M.O. Zhao, S.A.M. Hesp, Performance grading of the Lamont, Alberta C-SHRP pavement trial binders, Int. J. Pav. Eng. 7 (3) (2006) 199–211. [12] Ministry of Transportation of Ontario, ‘‘Determination of Performance Grade of Physically Aged Asphalt Using EBBR Method,” LS-308, Rev. 29, MTO Laboratory Testing Manual, 2015. [13] S. Tabib, O. Khuskivadze, P. Marks, E. Nicol, H. Ding, S.A.M. Hesp, Pavement performance compared with asphalt properties for five contracts in Ontario, Constr. Build. Mater. 171 (2018) 719–725.

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