2.7 EFFECT OF MODIFIED ASPHALTS ON THE RUTTING BEHAVIOR OF MIXTURES Rutting Behavior of Mixtures

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1 46 for using the shift factors. The derived model proves that using the master master curve concept can be an effective tool to calculate specification parameters at selected traffic speeds, selected temperatures, and selected strain levels in order to study the effect of traffic, pavement temperature, and pavement structure on pavement performance. 2.7 EFFECT OF MODIFIED ASPHALTS ON THE RUTTING BEHAVIOR OF MIXTURES The RSCH test as defined in AASHTO TP7 was used to measure the effect of modified asphalts on the rutting behavior of asphalt mixtures. The initial testing indicated that at the selected testing temperature (58 C), the level of shear stress used (68 kpa) could not result in a significant permanent strain. In addition, using the fitted models was considered questionable because it requires significant extrapolations that create a high level of uncertainty. To achieve the objective of inducing a higher level of permanent strain that simulates typical rutting in the field within a reasonable number of cycles, the testing conditions were changed to include a stress level of 207 kpa, which is three times the recommended stress level, and the testing temperature was changed to the high PG-grade temperature less 12 C. For PG 82 grade binders, the mixture testing was conducted at 70 C; for the PG 58 grade binders, it was conducted at 46 C. To measure the damage behavior of modified binders a new test, the repeated creep binder test, was developed. This test is conducted using the DSR and provides a direct, effective tool to describe the resistance of binders to accumulation of permanent strain under repeated cycles of constant load. It is in many ways similar to the mixture RSCH test protocol, but is conducted in the shear mode on binders. The following sections include the mixture and binder results, as well as the correlation found between mixture and binder rutting measurements Rutting Behavior of Mixtures The data collected according to the modified protocol significantly reduced the need for extrapolation of RSCH data. Figure 2.25 depicts an example of the data collected with the RSCH test at 70 C for mixtures produced with different gradations and aggregate sources. The results show that there are significant effects of the binders and aggregate properties. Each of these tests was done in two replicates with only minor variation. In order to evaluate the effect of binder modification on the results of the RSCH test, certain mixture behavior indicators Figure 2.25 Effect of binder type on permanent strain as a function of number of loading cycles measured in RSCH tests for limestone coarse aggregate mixture at 70 C.

2 47 are needed. These indicators are commonly derived from the typical power-law model recommended for representing rutting. The model, defined in the following equation, includes an initial strain factor, p1, and a slope factor, S: where p is the total accumulated permanent strain and N is the number of cycles. From the data collected for a large number of samples, it appeared that the first 10 cycles have a significant influence that results in the lack of fit of the model. To get a better evaluation of the effect of the controlled variables, curve fitting was conducted without including the first 10 cycles. The logarithmic slope, S, was the only parameter that could be considered reliable for studying the role of the binders in permanent deformation of the mixtures. It is believed that the initial permanent strain can be affected by many mixture factors that are not related to the binders and that this effect influences the total strain commonly used in comparing mixtures. Figure 2.26 shows an example of the fitted curves with the first 10 cycles excluded. The values of the logarithmic rate of strain S are significantly different than the initial values estimated with the inclusion of the first 10 cycles, and the ranking of the mixtures based on these values follows a more reasonable trend. The results indicate that the elaslog p = log p1 + S log N ( 10) tomeric binder is offering a slightly better resistance to rutting than is the plastomeric binder and that both of these binders offer significantly better resistance than does the oxidized binder. These differences are significant and indicate that the binder can have an important effect on rutting performance. Figure 2.27 shows the permanent strain rate, S, estimated as a function of aggregate type and binder type. It is observed that the modification type is very important and varies within the grade of the binder and among the grades. It is also observed that the effect of modification is dependent on the aggregate source and gradation. For the gravel aggregate, the binder modification effect is less important than the effect of gradation. For the limestone aggregate, the effect of modification is higher than the effect of gradation. These results indicate high interactive effects between binders and aggregates. Similar trends can be observed for the PG 76 and the PG 58 binders for both the initial strain and the rate of accumulation parameters. There does not appear to be a consistent trend showing one type of modification being superior to the others. In summary, the two main findings regarding the effect of modification that can be derived from the mixture RSCH data and analysis are 1. The rutting behavior of mixtures, as measured in the laboratory, is not consistent for binders of the same PG grade. The grade of the binder as determined by Permanent Strain (mm/mm) 1.0E E E E E-04 LS Coarse LS Fine GRVL Coarse GRVL Fine LS Coarse Fit LS Fine Fit GRVL Coarse Fit GRVL Fine Fit log p=[log p1+slogn] Binder Aggregate Stress (kpa) Temp (C) log p1= S= N 5 = R 2 = SBSrL S Coarse SBSrL S Fine SBSrG RVL Coarse SBSrG RVL Fine E E E E E E+05 Number of Cycles Figure 2.26 Curve fitting with the exclusion of the first 10 cycles (data collected using the RSCH for one mixture produced with three PG 82 binders).

3 48 Figure 2.27 Effect of binder modification on the permanent strain in the first cycle and permanent strain rate measured with the RSCH test. the performance system (AASHTO MP1) does not offer a reasonable indicator of the differences between the rutting behaviors of mixtures measured with the RSCH test. 2. The effect of binder modification at a given performance grade of binder is comparable with the effect of aggregate gradation. This is particularly true for the angular limestone angular aggregates. In fact, it is more important than changing the gradation in the case of angular aggregates. These main findings led the research team to focus on the binder parameter G*/sinδ, used in the current specification. The focus was on understanding the parameter and explaining

4 49 its value in rating the contribution of binders to mixture performance. This subject is discussed in the next subsections Rutting Behavior of Binders As indicated in the previous section, there was poor correlation between the binder performance grade and the laboratorymeasured mixture rutting performance. To evaluate the direct correlation between mixture rutting properties and G*/sinδ, the binders were aged in the RTFO and tested at the same temperatures at which the mixture RSCH tests were conducted. Figure 2.28 shows the correlation between the mixture rate of accumulated strain, S, and the parameter G*/sinδ measured at 10 rad/s. The correlation is poor. In addition to confirming the earlier findings, this result made it necessary to search for a better indicator of the binder s contribution to mixture rutting performance. Several protocols were tried in order to select a test procedure and a rheological parameter that could be used as a more-effective indicator of the role of binders in mixture rutting than was the parameter G*/sinδ. The selection process was based on two main hypotheses: 1. The strain in the binder domains is significantly larger than the strain at which the binders are tested in the DSR. Because the majority of modified binders exhibit nonlinear behavior, there is a possibility that by conducting strain sweeps, or stress sweeps, the main differences in binder contribution to mixture behavior can be detected and appropriately rated. 2. Cyclic loading with complete reversal in strain or stress is not appropriate for rating the contribution of binders to the mixture s resistance to rutting caused by cyclic irreversible loading (also called non-steady-state cyclic deformation or, more simply, repeated creep ). The energy dissipation under repeated creep conditions per cycle is not equivalent and cannot be directly related to the energy dissipated under fully reversible cyclic loading, such as the testing used currently in the specification. The first hypothesis is derived from the data collected in the earlier phases of the project, which indicated that modified binders vary significantly in their strain dependency. The second hypothesis is derived from the concept of the RSCH and the review of literature related to the concept of energy dissipation. It is also based on the recommendations from the internal advisory group of the project. To test these hypotheses, different binder testing protocols were used to find a relationship with mixture rutting performance. The protocols used to test the first hypothesis included strain sweeps, stress sweeps, time sweeps at constant Figure 2.28 Correlation between average mixture rate of accumulation of strain in the RSCH test and G*/sinδ.

5 50 strain and time sweeps at constant stress. To test the second hypothesis, a repeated creep test was developed and conducted to measure the permanent strain behavior of binders. The results of the strain sweep testing indicated that at temperatures in the range of high pavement design temperatures, the binders are sensitive to strain only at very high strain levels that exceed the reasonable range of approximately 50 percent. The results of the time-sweep testing indicated that asphalts are not very sensitive to repeated cyclic loading at these HTs. The collective analysis of the strain and time sweeps, conducted as part of Task 4 of the project, led to the conclusion that the strain sensitivity could not be used to explain the ranking observed in the mixture rutting tests. The focus was therefore turned to the second hypothesis. The testing of the first hypothesis led to the realization that the cyclic reversible tests do not offer a good indicator of resistance to rutting. It also led to the recognition that there is a fundamental problem with estimating energy dissipation during repeated creep from cyclic reversible loading. It was, therefore, decided that a new approach that utilizes repeated creep testing is required. Evaluation of the DSR software and capabilities indicated that a repeated creep test could be conducted on binders using the present same geometry and temperature range. Figure 2.29 depicts the results of the repeated creep testing of three binders of the PG 82 grade at conditions of 1 s load- ing and 9 s unloading. The results show a clear distinction between the accumulated permanent strains of the binders that could not be detected using the G*/sinδ parameter. It can be observed that the elastomeric binder SBSr is offering significantly higher recovery during the creep testing and thus exhibits less accumulated deformation. The plastomeric binder (PE) also offers higher recovery than does the third binder, which is an oxidized binder (Oxd.). These results are logical and could be related to the molecular nature of these materials. In addition, the ranking from the binder creep test matches well with the ranking from the mixture test results in terms of the rate of accumulation of permanent strain. Detailed analysis of these results was carried out so that the accumulated strain behavior could be understood. Figure 2.30 shows the creep and recovery of the same binders at the first cycle of testing. Figure 2.31 shows the same data for the 100th cycle. Some important observations can be made: The recovery of the elastomeric binder (PG82 SBSr) significantly increases with accumulated strain. This is not the case for the other binders, particularly for the oxidized binder, which hardly shows any change in recovery. The strain at the end of each loading cycle is not changing with accumulated strain for any of the binders. This indicates that the resistance of the binders to deformation Figure 2.29 Results of the accumulated strain under repeated creep testing for three PG 82 binders at 1 s loading and 9 s recovery (70 C, 300 Pa).

6 51 Figure 2.30 Creep and recovery of the first cycle for three PG 82 binders at 1 s loading and 9 s recovery (70 C, 300 Pa). Figure 2.31 Creep and recovery of the 100th cycle for three PG 82 binders at 1 s loading and 9 s recovery (70 C, 300 Pa).

7 52 (i.e., compliance) is similar and unaffected by the accumulated strain. The accumulated strain affects the recovery independent of the modulus under loading. This suggests that the creep stiffness (or creep compliance) may not be a good indicator of the creep recovery. These data raise some questions about the value of cyclic reversible testing used in the current Superpave protocols. Specifically, if G* and phase angle are measured under reversible conditions, their value in estimating the ability of the binder to recover after removal of a load application may be very limited. To compare the properties derived from the repeated creep testing with the G* and sinδ values, Table 2.10 was prepared. It presents the values of G*, δ sinδ, and G*/sinδ of these binders and compares them with the total permanent strain after one cycle and 100 cycles. The ratio of the strain under load to the permanent strain ( L / p ) is also shown. These comparisons indicate that G*/sinδ is not a good indicator of the accumulated strain. In fact, this parameter, which is used in the current specification, gives a ranking opposite to that of the permanent strain measured under repeated loading. It appears that the phase angle is the only property related to the permanent strain. However, the phase angle is not proportional to the total accumulated strain Selection of a Rutting Parameter Based on these promising results, the focus in this task shifted toward defining what are the main shortfalls of the G*/sinδ parameter and what is the best possible replacement for it. In addition to the finding that the G*/sinδ parameter did not show reasonable correlation with the mixture rutting performance measured using the RSCH procedure, two other fundamental problems with this parameter were identified: 1. It is derived from the linear viscoelastic response measured after a few cycles of testing, which does not allow measurements of the damage behavior of binders. 2. It is derived from cyclic reversible loading that does not allow direct measurement and is not a good scientific indicator of the accumulation of binder permanent strain during repeated creep loading. The use of the repeated creep test in the DSR could successfully solve both these problems. The repeated creep test can measure the damage behavior in the linear and nonlinear range. This test is similar to the test used in the mixture analysis protocols of the Superpave system, which has been successfully related to the performance of pavements in many studies. To identify a practical and reliable testing protocol, a plan was developed to answer the following three questions: 1. Can we predict the accumulated deformation from the first few cycles, or do we have to run the repeated cycles for a significant time? A prediction model is needed to simplify the use of the DSR and to derive a specification parameter. A four-element Burgers model (i.e., a Kelvin model and a Maxwell model in series) was used to fit the data of the first cycle. The initial analysis indicated that there is a need for more flexibility in the model, so a model with six parameters (i.e., two Kelvin models and a Maxwell model in series) was also used in the regression analysis. In addition, the Christensen Anderson Marasteanu (CAM) model recently introduced by the Pennsylvania State University (PSU) research group was also used in the analysis (53). Figure 2.32 depicts an example of the fitting process for the creep and recovery data using the Burgers model. Figure 2.33 depicts the Burgers model components and the division of the response into the components of the elastic, delayed-elastic, and viscous response for each cycle. For the analysis of the initial cycles, it appears that the CAM model introduced by the PSU group gives the best estimate. Checking the influence of the test data in the first five cycles on the final prediction indicates that the convergence of the four parameters for this model is very good. Generally, after three cycles, the values of the parameters all remain nearly constant. The main problem with the PSU model is that it does not give a direct measure of the viscosity, which is the TABLE 2.10 Comparison of the creep and recovery indicators with the values of G* and sin measured with the DSR at small strains

8 Figure 2.32 Using statistical modeling to predict accumulated strain of binders from repeated creep test results. Figure 2.33 Burgers model and its response.

9 54 main parameter needed. It is necessary to use the model to generate a creep-and-recovery response and to use the Burgers model to estimate the viscous component of the response. This was not considered favorable, and thus the four-element model was considered a better option. It was also observed that the initial cycles are not as consistent and that the delayed elasticity can play a major role in changing the response from one cycle to the next. After several trials varying the number of cycles, it was apparent that, after 50 cycles, a more consistent response is observed and that the delayed elasticity effects are reduced. Because the interest is in the viscous part of the behavior, the test should include a certain number of cycles for conditioning before the response should be modeled. The advantage of the Burgers model is that the response could be successfully divided to estimate the viscous component, which is believed to be the cause of permanent deformation, and to estimate its accumulation in paving mixtures. Cycles 50 and 51 were used for many binders and proved to give a very good prediction of the accumulated strain. Based on this analysis, the Burgers model was selected to fit the response at Cycles 50 and 51. The analysis confirmed that the stress level could be an important factor. 2. Can we predict creep and recovery from the cyclic loading? The promising results collected with the PSU model raised the question of whether we can estimate the accumulated strain from a frequency sweep test. A literature review was con- ducted to identify procedures used in the past to relate cyclic loading to creep response (58 60). The review indicated that there are only empirical relationships that are based on deriving the quasi-static response from cyclic loading. These empirical methods are not based on asphalt behavior, but are mostly derived for other polymeric materials. A spreadsheet program was developed using the frequencysweep data to estimate the accumulated strain. The estimated strain after 100 cycles was compared with the measured strain and the strain estimated from the four-element model for a number of binders. Table 2.11 shows the comparison for the nine (and their base) binders used in the mixture RSCH testing. The results were not encouraging: using the frequencysweep data can result in very high errors that could rank materials differently and thus be misleading. The approach was abandoned because not only are the results not promising, but there are theoretical concepts that indicate the approach is not plausible. 3. Is the behavior linear viscoelastic, and what are the effects of an accumulated permanent strain or stress level during each cycle? Because the current specification parameters in AASHTO MP1 are based on linear viscoelastic behavior, it is important to determine if changing the stress level results in nonlinear behavior. Testing was conducted at different stress levels, and the model-fitting procedure was used to fit the data. Table 2.12 TABLE 2.11 Comparison between the prediction by frequency-sweep data and creep-recovery data (loading time 3 s; unloading time 30 s)

10 55 TABLE 2.12 Typical results of modeling the creep-and-recovery response of modified binders at different temperatures and stress levels lists the results for five binders of the PG 76 and PG 82 grades. The testing was also conducted at several temperatures. Based on the initial analysis, it appears that there is some level of nonlinearity because the accumulated strain after 100 cycles, after normalization for stress level, varies significantly depending on stress level. In addition the model parameters k and t c change with changing levels of stress. It is also important to notice that the stress dependency is not the same for all asphalts and appears to be small for the binders without additives (such as the oxidized binders). These limited results indicate that although there are major differences between 50 Pa and 100 Pa, changing the stress above 100 Pa was not as influential. Testing also indicated that, if the 50-cycle conditioning is included, the stress level appears to be less important for all binders. This indicates that some of the stress level effects are important only in the first few cycles at which the delayed elastic response plays an important role. The results collectively indicate that, although there is no strong indication that stress level is important, specification testing should be done at a stress level that simulates actual pavement conditions. The other possible nonlinear effect is related to the total amount of accumulated strain. For the majority of the samples tested, it is found that after 40 to 50 cycles at the high-grade temperatures, the rate of strain accumulation is constant and not dependent on the level of accumulated strain. Other factors such as the loading time, the number of points needed for reliable prediction, and rest periods were evaluated by conducting testing at multiple combinations of conditions. The analysis of the results led to the conclusion that a test protocol of 100 cycles carried out in less than 60 min can give a reliable measurement of the contribution of a binder to rutting behavior Derivation of a Specification Parameter Based on the analyses of the repeated creep tests, it was determined that a creep-and-recovery test will significantly

11 56 improve the estimation of resistance to accumulation of permanent strain of binders and their contribution to resistance of mixture rutting. The test can be conducted using the existing DSR equipment, but requires special software to measure the required properties. It was also found that a cyclic loading test could not be used to derive a rheological model that can give accurate estimates of creep behavior. To derive a new parameter for rutting resistance, a basic approach based on viscoelasticity was followed. The binder repeated creep data have indicated that the rate of secondary creep of binders is a simple direct function of the number of load cycles: a = I + SN ( 11) where a = the accumulated permanent strain, I = the intercept, S = the rate of strain, and N = the number of cycles. According to this model, a specification parameter is defined either in terms of the rate of secondary creep or by using the viscous component of the response as estimated from the Burgers model. It is difficult to use the parameter of creep rate S as a specification parameter; it is an experimental parameter that is affected by testing attributes such as stress, loading time, and number of cycles. A better choice for a specification parameter is one derived from a proper rheological model that characterizes the fundamental behavior of the material. Although several models have been used to describe the behavior of asphalt binders, the four-element Burgers model, shown in Figure 2.33, offers a good representation of the binder behavior. This model is a combination of a Kelvin model and a Maxwell model in series (see Figure 2.33). The total shear strain versus time is expressed as follows: τ0 τ0 tg η τ 1/ 1 0 γ() t = + ( 1 e ) + t ( 12) G G η 0 1 where (t) = shear strain, τ 0 = constant shear strain, G 0 = spring constant of Maxwell model, tg 1 = spring constant of Kelvin model, 1 = dashpot constant of Kelvin model, t = time, and 0 = dashpot constant of Maxwell model. The following equation represents the creep compliance, J(t), in terms of its elastic component (J e ), its delayed-elastic component (J de ), and its viscous component (J v ): Jt () = J + J () t + J() t ( 13) e de v 0 The viscous component is inversely proportional to the viscosity, η 0, and directly proportional to stress and time of loading. Based on this separation of the creep response, the compliance could be used as an indicator of the contribution of binders to rutting resistance. Instead of using the compliance (J v ), which has a unit of 1/Pa, and to be compatible with the concept of stiffness introduced during SHRP, the inverse of the compliance, G v, could be used. G v is defined as the viscous component of the creep stiffness. The creep-and-recovery response measured with the DSR may be used to estimate the value of G v and the accumulated permanent strain for any selected combination of loading and unloading times. This implies that the accumulated permanent deformation is a function of viscosity, load, and loading time. By selecting the appropriate testing stress, τ, and the appropriate time of loading, t, the viscous component of the stiffness, G v, could be directly related to the rate of accumulation of permanent deformation, S, and thus used as a fundamental indicator of rutting resistance of asphalt binders. The recommended test protocol is to use a shear stress in the range of 30 Pa to 300 Pa for 100 cycles at the rate of a 1.0-s loading time followed by a 9.0-s unloading time. The value of the viscous part of the creep stiffness is then estimated at selected loading times related to the expected traffic speed. Longer loading time is used for slower traffic that requires higher levels of G v Correlation with Mixture Performance To validate this test, the nine binders used in the mixture evaluation experiments were aged using the RTFO and tested according to the protocol described in the previous section. The rate of accumulation of permanent strain of mixtures as measured in the RSCH test was correlated with the rate of accumulation of permanent strain in the binder. The rate in the case of the mixture testing was estimated using the logarithmic model of strain as a function of number of cycles. In the case of the binder test, the rate was estimated from the linear plot of strain as a function of the number of cycles. Figure 2.34 depicts the correlation between the binder permanent-strain performance (i.e., rutting parameter) and mixture permanentstrain performance (i.e., rutting parameter) averaged for the four mixture types. The correlation coefficient (r 2 ) is 0.68, which is significantly better than the correlation with G*/sinδ shown previously in Figure 2.9. The correlation with G*/sinδ was 0.23, which is less than one-half the correlation shown in this figure for this new binder rutting parameter. Although this result was very promising, it was necessary to examine the correlation with the individual aggregates used in the testing. Figure 2.35 is prepared to show the individual correlations. As can be seen in this figure, the correlation within aggregate types is not as good as the correlation with the average mixture behavior. This result was expected because it is known that the aggregate has a major effect on rutting behavior of mixtures. To

12 Figure 2.34 The correlation between binder rutting parameter and the average mixture rutting parameter. Figure 2.35 Correlation of binder rutting parameter with mixture rutting parameter for individual types of aggregates.

13 58 Figure 2.36 Relationship between binder and mixture permanent strains. deal with this issue, it was decided to further explore the relationship between accumulation of permanent strain in binder and mixture. Using the model for mixture rutting and the Burgers model for binder rutting, the rate of accumulation of strain of the mixtures was plotted as a function of the rate for the binders, as shown in Figure To plot this figure, the mixture and binder data were normalized for the loading time and the stress level. The normalization assumptions were based on an understanding of the relationships between binder and mixture stiffness. The assumptions should not affect the trend in the relationship with binder as much as they result in shifting the relationship along one of the axes. As seen in the figure, it appears that the relationship is bimodal. It is believed that these two regions represent the role of the two different components that constitute asphalt mixtures. In the initial region, the mixture rutting appears to be very sensitive to the binder properties so that a small change in binder strain results in a major effect on mixture rutting. In the second region, the binder properties do not appear to play a substantial role in the mixture rutting. It is believed that in the second region the aggregate interlock starts taking a major role and thus the binder properties play a minor role. It is plausible that after a certain range of accumulated strain of mixtures, binders would return to play a major role in mixture rutting behavior. This range would be the tertiary creep range in the domain of mixture performance. The data collected in this project are not adequate to prove the latter hypothesis about role of binder in the tertiary region of rutting. It is a point worth exploring in the future for a better understanding of the role of the binder in the rutting of mixtures. Using the concept of bimodal relationship between binder and mixture behavior, the slope of the relationship was separated into K 1 and K 2, as shown in Figure K 1 represents the first asymptote to the binder-mixture relationship, which should be highly related to the binder properties, and K 2 is the second asymptote, which should be highly dependent on the aggregate properties. This approach was used to determine K 1 and correlate it with the binder viscosity, as shown in Figure The correlation significantly improved for the gravel and limestone aggregates. The data in Figure 2.37 are, however, not sufficient to prove that the concept is robust. As can be observed, there is one data point showing a high K 1 and a high viscosity value, which is improving the correlation. There is also a concern about the interdependency of K 1 and viscosity. However, the relationship is reasonable because it shows that a higher binder viscosity results in a higher K 1 value, which means that the mixture rutting is more sensitive to binder properties. The results of the correlation of the mixture rutting with the binder repeated creep results are fairly good. They represent in certain respects a major improvement from the current procedure of using G*/sinδ. The approach is more directly related to the loading patterns known to cause rutting, and it appears to result in better relationships to the rutting performance of some mixtures Summary of Findings from the Rutting Study Based on the results and analysis of the rutting studies, the findings may be summarized as follows:

14 59 Figure 2.37 Correlation of initial mixture-binder rutting relationship (K 1 ) with the viscosity of binders estimated from the repeated creep test. 1. The effect of binder modification at a given performance grade of binders is comparable with the effect of aggregate gradation. This is particularly true for the angular limestone aggregates. In fact it is more important than changing the gradation in the case of angular aggregates. 2. The rutting behavior of mixtures, as measured in the laboratory, is not consistent for binders of the same performance grade. It is therefore apparent that the grade of the binder as determined by the performance-grading system does not offer a reasonable indicator of the differences between the rutting behavior of mixtures of the same aggregates measured with the RSCH test. 3. There are critical questions about the validity of the binder parameter G*/sinδ. The correlation between the mixture rutting indicators and G*/sinδ is very poor. The parameter is derived from testing that does not provide a good representation of traffic loading in the field. The parameter was not found to be useful in describing the accumulation of permanent flow, which is important in rutting evaluation. 4. Repeated creep testing of binders is introduced as a better method for estimating binder resistance to permanent strain accumulation. The viscous component of the creep stiffness (G v ) is found to be a good indicator of the rate of permanent strain accumulation for binders. It is proposed as a better specification parameter. 5. The correlation between the proposed parameter G v and mixture rutting is good for certain aggregates, but not for others. It appears that the effect of aggregates and the interaction of aggregate characteristics with binder properties are very complex. Isolation of the binder effects to evaluate the validity of G v or any other binderonly property is very difficult. 6. Compared with the current binder protocol, the repeated creep test protocol for measuring the accumulated permanent strain of the binder represents an improvement in the theoretical and practical concepts for better rating binder properties related to pavement rutting. 2.8 EFFECT OF MODIFIED ASPHALTS ON THE FATIGUE BEHAVIOR OF MIXTURES IT temperature testing included the FSCH test and the flexural beam fatigue test. These tests were developed during SHRP as part of the A-003A contract at the University of California-Berkeley. Test procedures are described in AASHTO TP7 (FSCH) and TP8 (beam fatigue). The FSCH

15 60 results were reported earlier in Section 2.6 with the HT data. This section covers the results of the beam fatigue testing. This section also covers the development of a new binder test proposed for better estimation of the binder s role in pavement fatigue damage. It also includes a new approach for analysis and interpretation of fatigue data. Based on the Superpave recommendations, the testing was mostly conducted at intermediate pavement temperatures as determined by the binder performance-grading system Beam Fatigue Test Results of Mixtures Beam fatigue specimens were prepared using kneading compaction to air voids of 7.0 ± 0.5 percent. All mixtures were subjected to short-term oven aging (4 h at 135 C). In the fatigue-testing protocol, beam specimens (380-mm length, 50-mm height, and 63-mm width) are tested using a 10-Hz loading frequency at an initial controlled strain of 250 to 750 microstrains. At the end of 50 cycles, the initial specimen stiffness is determined. The strain rate is then adjusted, and the test is carried out to a minimum of 10,000 cycles. The test is terminated when the stiffness decreases to 50 percent of the initial stiffness. The output of the test is a curve of flexural stiffness as a function of load repetitions. Triplicate specimens are commonly tested for each mixture. The fatigue testing was done at the intermediate grade temperature of the binder (i.e, the temperature at which G*sinδ=5000 kpa at 1.59 Hz). Table 2.13 lists the binders and the testing temperatures. It also includes the values of G*sinδ of the binders measured after PAV aging at 1.59 Hz, which is the binder specification requirement. The table also includes the value of G*sinδ at 10 Hz after RTFO aging, which represents the conditions at which the mixtures were tested in the beam fatigue test. An example of the data set is shown in Figure 2.38 for one of the combinations of binder and aggregate. The figure also shows the model used to represent the fatigue response of the mixture, which is fitted to the average response value. The model estimated the average initial stiffness and the number of load cycles at which the stiffness of the mixtures is reduced to 50 percent of the initial value, N 50. In the current Superpave protocols, a simple power-law model is recommended to fit the data from the beam fatigue test (61). This model did not show good fit for these data and led to the search for a better model to fit the fatigue curves. The model selected included a simple modification to include the parameter S 1 in the formula: S = S S N n 0 1 ( 14) In the equation, S is the flexural stiffness, S 0 is the initial stiffness, and N is the number of cycles. S 1 and n are two constants determined from the statistical fit of the data. Using this model, the effects of aggregate gradation, aggregate source, and binders were evaluated. Figure 2.38 shows an example of the results comparing the fatigue life of mixtures made with the same binder, but with different aggregates. Using this modeling approach, the mixture and binder combinations were compared using the initial stiffness value and the fatigue life estimated as the number of cycles to reach 50 percent of the initial stiffness. Figure 2.39 shows the average N 50 value of mixtures with the limestone aggregates and the gravel aggregates. The results shown indicate that fatigue life is sensitive to the aggregate type (i.e., source and gradation) and binder type (i.e., modification). The effects of the aggregate gradations and sources, however, do not appear to be as significant as the effect of the binder grade or modification. Given the case that these binders were tested at an approximately equivalent value of the binder specification parameter (G*sinδ) after PAV aging, the results raise questions about the validity of the G*sinδ parameter. The DSR was used to test the binders at 10 Hz at the mixture testing temperatures after RTFO aging. The values of G*sinδ versus TABLE 2.13 Binder type and the temperature at which testing was conducted

16 61 Figure 2.38 An example of average fatigue response for one binder mixed with four aggregates. Figure 2.39 Fatigue life estimated for the limestone and the gravel mixtures with all of the nine binders used in the study.

17 62 Figure 2.40 Correlation between G*sinδ and mixture fatigue life measured at 10 Hz at the intermediate grade temperature for the limestone and the gravel mixtures. mixture fatigue life values (N 50 ) were plotted for each of the aggregate types individually, as shown in Figure The maximum correlation coefficient for the different mixtures is 0.23, which indicates a lack of relationship between the binder parameter G*sinδ and the mixture fatigue life for the nine binders tested. This lack of correlation could not be attributed to the effect of the aggregates. As shown in Figure 2.39, the role of the binder is found to be very important for each aggregate type and is found relatively independent of the aggregate type Development of a New Binder Fatigue Test In the existing specification, the parameter G*sinδ is used to rate the binder contribution to fatigue damage resistance. This fatigue parameter was selected based on the dissipated energy concept as applied to linear viscoelastic behavior. It was, however, based on a critical assumption that fatigue is a strain-controlled phenomenon typical of conditions of weak pavement structure (14,62). Fatigue is considered one of the most complicated damage phenomena in asphalt pavements. Many researchers believe that it is a pavement structure problem, while many others believe that it is a mixture problem. It is, however, recognized that fatigue cracking starts and propagates in the binder. Therefore, a binder-associated damage must be controlled in the binder specification. The main reason for the lack of correlation between the parameter G*sinδ and mixture fatigue indicators, as shown in Figure 2.40, may be that it is measured in the linear viscoelastic range using small strains. This approach is unlikely to be useful in representing the effect of repeated cyclic loading and the changes in binder properties with accumulation of damage. The effort to develop a new test focused on simulating the fatigue phenomenon in a binder-only fatigue test such that damage behavior could be directly monitored. The DSR was used to conduct what is called a time-sweep test. The test provides a simple method of applying repeated cycling of stress or strain loading at selected temperatures and loading frequency. The initial data collected were very promising and showed that time sweeps are effective in measuring binder damage behavior under repeated loading in shear. Effects of frequency, temperature, stress, and strain conditions were measured. Details of the binder fatigue were published by the

18 63 Association of Asphalt Paving Technologists in 1999 (44). These efforts led to the advancing of the concept of a binderonly fatigue test using the DSR. It was hypothesized that this test could be a better measurement of the effect of binders on mixture fatigue response. To test this hypothesis, all nine binders used in the production of the mixtures were tested in the DSR at conditions that match the mixture beam fatigue conditions. The binders were aged in the RTFO to simulate the effect of mixing and compaction, and the testing was conducted at 10 Hz at temperatures as close as possible to the mixture beam fatigue temperatures. The binder testing was conducted in a strain-controlled mode, and to match the mixture strain level, an estimated strain of 3 percent was used for all binders. Figure 2.41 shows the results of the binder testing. The initial G* values are similar, but the binders show significantly different fatigue behavior. Three of the binders showed high resistance to fatigue after more than 100,000 cycles, while the other binders showed significant variation in fatigue life when it was defined as the number of cycles to the initial G* value by 50 percent. To see if the binder fatigue life measured in the strain-controlled binder test has any relationship to the mixture strain-controlled fatigue life, Figure 2.42 was prepared to show the relationship between the average mixture performance and the binder fatigue life as determined by the number of cycles to 50 percent G*. As shown in the figure, there is a high correlation (r 2 = 0.84) for the nine binders that showed binder fatigue failure. This result is very encouraging. It indicates that the binder fatigue test is promising and may result in a good indicator of the fatigue damage of mixtures. Figure 2.43 shows the good-to-excellent correlation for each type of mixture Development of a Binder Fatigue Parameter Different loading modes can be used in fatigue testing. However, a reliable indicator of fatigue failure should be independent of the loading mode. It should also provide a consistent indication of the level of damage and the progression of damage in the material in terms of changes in mechanical behavior under any loading conditions. The most commonly used definition of fatigue failure in asphalt mixtures is a decrease in the initial stiffness by 50 percent. This arbitrary definition, however, does not allow evaluation of the distinctly different mechanism by which a material would respond to the energy input during a loading history for the different loading modes. Recent advancements in fatigue research have indicated that a better indicator of fatigue is the rate of change of dissipated (distortion) energy per load cycle. A few studies in the United States and Europe (63 70) have shown that the rate of dissipation is a more appropriate approach to the study of fatigue in asphalt mixture. Figure 2.41 Binder fatigue results at 10 Hz, 3 percent strain, and the temperatures selected for the mixture beam fatigue testing (binders were RTFO aged).

19 64 Figure 2.42 Correlation between binder fatigue life and average mixtures fatigue life measured at the same temperature and frequency. For a viscoelastic material, the dissipated energy per cycle per unit volume (W i ) is the area covered by the stress strain curve: d () t Wi = σ() t d () t = σ() t dt dt where W i = dissipated energy per cycle per unit volume, = stress, = strain, and t = time. Under cyclic loading conditions: σ() t = σ sinωt ( 16) and ( t) = sin( ωt δ) ( 17) i i ( 15) where σ i = the stress amplitude at cycle i, i = the strain amplitude at cycle i, ω=2πf ( f is the frequency), and δ=the phase angle between the stress and strain signals. From the above equations, the energy dissipated per cycle is calculated as follows: W i =πσ i i W c The accumulated dissipated energy after n cycles is = n i = 1 sin δ ( 18) W i ( 19) where W c = accumulated dissipated energy after n cycles, n = number of cycles, and W i = dissipated energy per cycle per unit volume. As indicated earlier, several researchers have taken a different approach to evaluating fatigue that is based on the concept of change in dissipated energy (71). Typically, the rate of change in dissipated energy under strain-controlled and stress-controlled tests undergoes a sudden change after a certain number of loading cycles. This change is highly material specific and independent of loading conditions. This number of cycles is considered as a better indicator of materials resistance to fatigue. The premise behind this concept is related to the development of fatigue cracking within a material. Two important stages can be distinguished in the fatigue process: crack initiation (Stage 1) and crack propagation (Stage 2).

20 65 Figure 2.43 Correlation of binder fatigue life with mixture fatigue life as measured by the number of cycles at which modulus is 50 percent of initial value. When the fatigue damage progresses from Stage 1 to Stage 2, there is a significant change in the amount of energy dissipated into fatigue damage per cycle. The fatigue life to crack propagation (N p ) is used in this approach to mark the transition from crack initiation to crack propagation on the fatigue-life scale. This transition is a function of energy dissipated per cycle and is independent of the mode of loading. There are several approaches to present the criterion of fatigue based on the rate of change in the dissipated energy. The most promising approaches are the rate of change of dissipated energy, presented by Carpenter and Jansen (68), and the dissipated energy ratio (DER), presented by Pronk and Hopman (66,67). Applying the rate-of-change concept to binder data indicated that the approach is useful for stress-controlled testing, but is not useful for strain-controlled testing. The main reason is that it is almost impossible to clearly define an inflection point in the trend, which would allow a clear definition of fatigue life. Moreover, the rate of energy dissipation will stabilize when the damage starts accumulating because the material will soften, resulting in a reduction in the stress required to cause the same strain. The test can therefore take a long time to achieve a transition from the crack initiation to the crackpropagation stage. The cumulative DER as defined by Pronk and Hopman (66,67) is based on the following calculations: n Wi i DER = = 1 ( 20) W n where W i = the dissipated energy at cycle i, and W n = the dissipated energy at cycle n. Pronk has published several examples of the determination of the fatigue life of an asphalt mixture under constant stress and constant strain modes using this definition. Figures 2.44 and 2.45 are examples of the application of this concept to binder data. The results shown indicate that binders can be evaluated effectively using this method. The curves of the binder data are similar to the mixture data published earlier by other researchers, which is very promising. This also implies that the main factor in the fatigue behavior of mixtures could well be related to fatigue damage in the binder. It is also observed that the slope of the relationship between the energy ratio and the number of cycles to failure is equal to 1.0 when the material is not undergo-

21 66 Figure 2.44 An example of using the concept of dissipated energy ratio to analyze fatigue data of asphalt binders. Figure 2.45 Typical results of a stress-controlled time sweep of asphalt binders.

22 67 ing fatigue damage as indicated by the constancy of the dissipated energy per cycle. Assume that W i is constant and is equal to W n : n Wi = nw i = 1 ( 21) Substituting Equation 2-21 into Equation 2-20 leads to the dissipated energy ratio being equal to the number of load cycles. The change in rate of dissipated energy with number of cycles is therefore equal to 1.0. Based on this derivation, an understanding of the fatigue curve may be introduced. It is assumed that the first portion of the curve represents the stage during which the energy per cycle is dissipated in viscoelastic damping with negligible damage. In the next stage, cracking initiation consumes an additional amount of energy beyond the viscoelastic damping. In the third stage, crack propagation begins, and a noticeable increase in dissipated energy per cycle is observed. This is assumed to be the most critical stage during which the damage per cycle is so high that healing and recovery occur with difficulty. The hypothesis introduced here is that an effective fatigue parameter can be derived if it is assumed that, when a binder reaches the third stage under any combination of energy input, load rate, and temperature, it undergoes irrecoverable fatigue damage. A protocol for binder testing was developed to identify the critical value of N p. In this context, N p is defined as the inter- section between the viscoelastic damping asymptote and the irrecoverable fatigue asymptote. This N p value is considered for the new fatigue specification parameter. The Internal Advisory Group of the project raised several issues about the reliability of the fatigue binder tests and the dependency of the results on the geometry. Although the dependency on the geometry is a universal problem in fatigue testing, it was important to study this issue. Detailed evaluation of the geometry effects and the stress level indicated that the test is a good candidate for measuring actual fatigue resistance. The results also indicated that the test is repeatable and could be designed to give effective results in less than 60 minutes for a wide range of binder types. Figure 2.46 depicts an example of the test measurements at six stress levels conducted to verify the test is measuring classic fatigue. The results show uniform behavior that conforms to the classic fatigue behavior of semisolid materials such as asphalt. All fatigue curves in Figure 2.46 show an initial linear portion that is indicative of no damage and constant dissipation in viscoelastic damping. Damping is used here to describe the conditions in which energy dissipated per cycle is constant, which indicates no damage. This initial phase is followed by the gradual change in energy ratio as the damage accumulation in the binder takes place. It is clear that the onset of this phase is very sensitive to the stress level because as the crack initiation starts, the effective stressbearing area of the sample decreases, and the crack propagates rapidly in proportion to the effective radius of the uncracked sample. 1.00E E+05 Dissipated Energy Ratio 1.00E E E Pa 300 Pa 1000 Pa 1.00E Pa Pa Pa 1.00E E E E E E E E E+07 Number of Cycles Figure 2.46 Effect of increasing stress level on the fatigue behavior of asphalt binders.