Mechanical characterization of concrete-asphalt ...

European Journal of Environmental and Civil Engineering

ISSN: 1964-8189 (Print) 2116-7214 (Online) Journal homepage: http://www.tandfonline.com/loi/tece20

Mechanical characterisation of concrete-asphalt interface in bonded concrete overlays of asphalt pavements Angel Mateos, John Harvey, Julio Paniagua, Fabian Paniagua & Angela Fan Liu To cite this article: Angel Mateos, John Harvey, Julio Paniagua, Fabian Paniagua & Angela Fan Liu (2017): Mechanical characterisation of concrete-asphalt interface in bonded concrete overlays of asphalt pavements, European Journal of Environmental and Civil Engineering, DOI: 10.1080/19648189.2017.1311808 To link to this article: http://dx.doi.org/10.1080/19648189.2017.1311808

Published online: 18 Apr 2017.

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Date: 31 August 2017, At: 17:39

Mechanical Characterization of Concrete-Asphalt Interface in Bonded Concrete Overlays of Asphalt Pavements Bonded Concrete Overlay of an Asphalt pavement (BCOA) is a rehabilitation technique consisting of 50-175 mm thickness concrete overlay on an existing asphalt pavement. This technique, known as thin (minimum 100 mm) or ultrathin whitetopping (thinner than 100 mm) in the past, relies on the composite action of concrete and asphalt layers acting together with a third phase of the system being the interface between the two materials. For this study, the stiffness and strength/fatigue resistance of this interface have been characterized by means of a series of laboratory tests conducted on asphalt and composite cylindrical specimens under different loading and environmental conditions. Testing conditions included wet and dry, and temperatures between 5 and 40°C, a range applicable to BCOA asphalt bases in California. Experimental results from this study indicate the mechanical nature of the concrete-asphalt interface is strongly related to that of the asphalt. The interface stiffness showed clear time dependency, and it significantly softened under wet conditions. Experimental results from this study are in line with common belief that water is one of the critical factors leading to failure of BCOA sections, but do not support the common belief that concrete does not bond well to new asphalt. Keywords: bonded concrete overlay, asphalt pavement, rehabilitation, concreteasphalt bonding

1

Introduction

A Bonded Concrete Overlay of an Asphalt pavement (BCOA) is a rehabilitation technique consisting of 50 to 175 mm thickness concrete overlay on an existing flexible, semi-rigid or composite pavement. This technique, that has also been known as thin (minimum 100 mm) or ultrathin whitetopping (thinner than 100 mm) in the past, relies on the composite action of the concrete and asphalt layers acting together with a third phase of the system being the interface between the two materials (Mu and Vandenbossche, 2011). This new concept results in a much stronger pavement structure than would result if the layers were to act alone, in the same way that two independent

beams placed one on top of the other cannot carry as much load as a single beam of double thickness. When well bonded, the underlying asphalt lowers the neutral axis of the concrete overlay, thus reducing tensile stresses at the bottom of the concrete and, consequently, significantly increasing the fatigue life. This composite action requires a good bonding between the two materials, which constitutes the main factor controlling the successful conception, design, and construction of this type of pavement. Numerous studies agree on the critical importance of concrete-asphalt bonding for BCOA performance and present numerous experimental results that support this conclusion (Rasmussen and Rozycki, 2004; Li et al., 2013). However, there has been very little research investigating the mechanics and performance of the concrete-asphalt interface, which are essentially unknown at this time (Mateos et al., 2015). The objective of the experimental study presented in this paper is to provide insight into the mechanical nature of the concrete-asphalt interface and to provide the basis for developing a laboratory protocol for mechanical characterization of the interface. Preliminary laboratory results are presented from a three-year research project supported by the California Department of Transportation and the concrete pavement industry, one of whose specific goals is understanding the performance of the concreteasphalt interface in BCOA. The primary goal of the project is to develop recommendations and guidance on the use of BCOA as a rehabilitation alternative, with particular emphasis in Californian materials, climatic, and traffic conditions.

2

Experimental Design

Laboratory tests to characterize the mechanical properties and performance of the interface have been conducted in order to achieve the goal of this experimental study.

Tests are focused on either horizontal (shear) or vertical bonding mechanics and performance, and they consider two different time scales: the rapid loads caused by traffic and the slow loads caused by daily and seasonal changes in temperatures and moisture conditions. Loading times on the order of tenths of a second are important in the first case, so frequency sweep tests have been used for the characterization. Creep tests have been used to simulate the longer loading times (hours and months) of environmental changes. Shear performance has been evaluated with the Superpave Shear Tester (SST). This machine is able to apply different shear loading patterns, including sinusoidal, to a discshaped specimen. Shear deformation is measured by means of an LVDT attached to the specimen testing platens. A universal testing machine (UTM) has been used to evaluate vertical tensile performance of the interface. In this case, vertical deformation has been measured by using four LVDTs, two of them expanding across the interface and part of the asphalt, and two of them exclusively measuring asphalt strains, as reflected in Figure 1b. A summary of mechanical characterization tests conducted in this research is included in Table 1. The range of temperatures of interest for this study has been determined by using CalME temperature model (Lea and Harvey, 2012). This model includes a one dimensional combined finite difference and finite element procedure that estimates temperature profile through the entire depth of the pavement. Estimation is based on pre-calculated surface temperatures from the Enhanced Integrated Climate Model (EICM, Zapata et al., 2008) with 30 years of environmental data, and the thermal properties of pavement materials. Depending on geographic location and BCOA thickness, the expected range of concrete-asphalt interface temperatures in California is 5 to 45°C.

Specimens were produced in the laboratory by casting concrete overlays on asphalt slabs. Asphalt slabs were prepared using rolling-wheel compaction. Concrete overlays were produced with Type I/II Portland cement (US standard, mild sulfate resistance), and were covered with a wet burlap after casting to prevent moisture loss. In expected California practice, BCOA projects will use accelerated Type I/II cement in weekend construction closures, or Type III (high early strength) Portland cement or calcium sulfo-aluminate cement in overnight closures. Composite specimens were cored from the concrete-asphalt slabs, and then trimmed to the final height. SST and UTM composite specimens are shown in Figure 1. Concrete water/cement ratio was 0.35, and 28-day design flexural strength was 3.8 MPa (550 psi). The asphalt was a Superpave gap-graded mix with 12.5 mm nominal maximum aggregate size, 7.4% asphalt rubber binder content, compacted to 6% air voids, and was subjected to short-term oven aging. The asphalt rubber binder was obtained through the wet process-high viscosity (State of California Department of Transportation, 2006), by adding around 18% recycled tire rubber to the base binder. The base binder was plain asphalt with performance grading PG64-16 according to AASHTO M 320 (binder high and low performance grades, respectively, 64°C and 16°C). No surface texturing was applied to the asphalt before casting the concrete overlay. It should be remarked that BCOA is typically built on top of an old asphalt surface that is typically milled before receiving the concrete overlay. The resulting interphase may considerably differ from the interphase evaluated in this study. Specimens were temperature and moisture conditioned before testing. Temperature conditioning of the composite specimens did not differ from that followed for standard asphalt testing. Moisture conditioning presented several differences with AASHTO T 283-14 “Standard Method of Test for Resistance of Compacted Asphalt

Mixtures to Moisture Induced Damage”. In the first place, the freezing cycle was not applied, since chances of freezing temperatures below the concrete overlay are only applicable to the mountainous regions of California. Vacuum saturation time and negative pressure were applied to both the asphalt and composite specimens with the goal of achieving 70 to 80 percent saturation, as specified in the T 283 standard. Specimens were placed in a 60°C water bath for 24 hours after saturation, as specified in the T 283 standard. They were then wrapped in waterproof parafilmTM to prevent water from escaping during testing, glued to the testing platens, and immersed in water at the testing temperature for two hours before testing.

3

Analysis of Laboratory Results

3.1 Shear Stiffness of the Interface Shear stiffness of the interface under rapid loads (traffic loads) was determined by conducting frequency-sweep dynamic modulus testing in the SST. In this testing, a sinusoidal shear displacement is applied to the specimen, while shear load and shear deformation are measured. The specimen is maintained at a constant height. Dynamic shear modulus, |G*|, is determined as the ratio between peak-to-peak shear stress and peak-to-peak shear strain. |G*| is a material property, so it is not applicable to a composite specimen. In this case, an equivalent |G*| can be determined as the ratio between shear stress and total shear strain, which includes asphalt and interface. It also includes concrete shear strain, but deformation of the concrete is negligible compared to the asphalt. Total shear strain is determined as the ratio between horizontal relative displacement (top to bottom platen) and thickness of the asphalt part of the composite specimen (total thickness minus concrete thickness). Results of the frequency-sweep testing are presented in Figure 2 for

both asphalt and composite specimens. Note that |G*| values for different temperatures have been shifted along the horizontal axis according to the time-temperature correspondence principle, which is applicable to asphalt. At intermediate and low temperatures, |G*| of the asphalt is higher than the equivalent |G*| of the composite specimens. This indicates that the concrete-asphalt interface can be thought of as an interlayer that behaves like a relatively soft material. The stiffness of this interlayer can be deduced after considering that complex compliance (inverse of complex modulus) of a composite specimen is the sum of the complex compliances of the asphalt and the interlayer. Using this approach, the calculated shear compliance of the interlayer was around 0.025 mm/MPa for the range of temperatures and frequencies applied in this experiment. This is equivalent to 100 MPa shear modulus if a theoretical thickness of 2.5 mm is assumed for the interlayer. This is an assumption, however it provides a simple way to picture the stiffness of the interlayer versus the stiffness of the asphalt. The calculated shear stiffness of the interlayer showed little temperature and frequency susceptibility, as can be seen in Figure 2. At high temperatures and low-intermediate frequencies, the equivalent |G*| of the composite specimens was higher than |G*| of the asphalt. This outcome may be related to the reinforcing effect of the concrete and cement slurry that penetrate the asphalt surface texture and voids. Asphalt and composite specimens were subjected to shear creep and recovery test after the frequency-sweep test. A constant shear stress (250, 100, 35 kPa, respectively, at 4, 20 and 40 °C) was applied for 10 seconds, while shear strain was measured. Then, the load was removed and the strain recovery was recorded for 100 additional seconds. Strain recovery was used to determine the shear creep compliance

function. This function defines, for a viscoelastic material, the strain developed under a constant stress. In this study, the creep compliance was determined as the ratio between recovered strain and the applied shear stress. Creep compliances at the three testing temperatures are shown for the asphalt in Figure 3. Data corresponding to different temperatures have been shifted along the horizontal axis according to the time-temperature correspondence principle. The shear creep compliance of the composite specimens is shown in Figure 3 for 20°C testing. For this particular temperature, deformation of the composite specimens was higher than deformation of the asphalt specimens. This result is in line with a soft interlayer between concrete and asphalt. The creep compliance of this interlayer can be determined as the difference between the creep compliances of the composite and asphalt specimens. Interlayer equivalent compliance, after assuming a theoretical thickness of 2.5 mm, is shown in Figure 3. As explained above, an assumed theoretical thickness for the interlayer is used to assist in the comparison of interlayer versus asphalt stiffness values. It should be noted that interlayer equivalent creep compliance is around 5x105 µɛ/MPa (1.25 mm for 2.5 mm interlayer thickness) after 100 seconds. Even for 4°C testing, interlayer creep compliance was around 105 µɛ/MPa after 100 seconds. These values represent a tremendous capacity for relaxation of shear stresses caused by concrete expansion and contraction. This means, in practice, the concrete overlay will be able to move horizontally under daily, seasonal, and long-term environment-related actions with minimal stress coming from the underlying asphalt base because of the slow loading times. This will significantly decrease the risk of early age cracking in the concrete due to tensile stresses caused by restraint from the base and will also improve long-term performance.

The effects of water on interface stiffness were evaluated by testing asphalt and composite specimens after moisture conditioning. Results, presented in Figure 4, show a large decrease in asphalt stiffness due to moisture, and an even greater decrease for the composite specimens. Interface equivalent stiffness (2.5 mm theoretical thickness) reached values between 1 and 4 MPa, which reflects a very weak bonding between concrete and asphalt when subjected to water.

3.2 Shear Fatigue Resistance of the Interface Repeated shear sinusoidal loading (1200 µɛ peak to peak) was applied on asphalt and composite specimens in order to reproduce damage accumulation under traffic loads at high temperatures (40°C). Tests were conducted on dry and wet conditions (parafilmTM wrapped, saturated specimens), as reflected in Figure 5. This testing revealed that the evolution of dynamic modulus and phase angle during fatigue testing was almost identical in both asphalt and composite specimens. Furthermore, stiffness recovery after one month at rest was also almost identical in both asphalt and composite specimens. These results seem to indicate that fatigue damage primarily occurred in the asphalt, and not in the interface. This outcome is in line with experimental results obtained with the Fabac accelerated pavement testing facility at Nantes (Chabot et al. 2008), that concluded that asphalt damage took place instead of debonding (damage to the interface) in one section tested during summer. Another important observation after repeated shear testing is that specimens that were fatigued in dry conditions recovered most stiffness after the rest period, indicating that it was not permanent damage (instead it was thixotropy or healing depending on the theory assumed), while recovery of the specimens fatigued under wet conditions was negligible, as can be seen in Figure 6. This result is in line with the conclusions of some other research (Vandenbossche and

Barman, 2010; Burnham, 2006) that considered exposure of the interface to water as one of the main factors leading to failure of BCOA projects.

3.3 Tensile Stiffness and Strength of the Interface Tensile stiffness of the interface was evaluated by means of a creep test. A tensile stress of 200 kPa was applied on a composite specimen, 100 mm diameter, at 20°C. Strain was measured in the asphalt and also across the interface. This last measurement was conducted by placing two LVDTs anchored to the concrete in one side, and to the asphalt, 10 mm below the interface, on the other side (Figure 1b). Measured asphalt strain was used to remove the contribution of the 10 mm asphalt from the total interface LVDTs opening. Creep compliance functions were calibrated using the unloading part of the creep tests. The creep function defines the time-dependent strain that takes place under constant unitary stress. Equation of the creep function was D*t^n, where t is time, * and ^ denote multiplication and power, respectively, and D and n are parameters. The exponent of the power law (n) was very similar for the asphalt and the interface, 0.477 for both. This indicates the mechanical nature of the interface is strongly related to that of the asphalt. Equivalent compliance of the interface (assuming a theoretical thickness of 2.5 mm) was around 8 times bigger than the creep compliance of the asphalt, i.e., its stiffness would be around 8 times smaller. This ratio is in line with results presented in Figure 2 and Figure 3. Tensile strength of the interface was evaluated at 20°C by conducting constant load rate tensile strength tests. A single rate of 3 kPa/min was used for final testing as several specimens were lost during the test setup process. One of the two tested specimens failed at the interface, while the other one failed in the asphalt as reflected in

Figure 1b. The fact that failure can take place in the asphalt as well as in the interface contradicts common belief that concrete does not bond well to new asphalt (Sheehan et al., 2004; Yu and Tayabji, 2007) and that the interface strength is always weaker than the asphalt strength in tension.

4

Discussion of Results

Experimental data presented above indicate that the mechanical nature of the concrete-asphalt interface is strongly related to that of the asphalt. One of the facts that supports this conclusion is the clear time dependency of the interface stiffness, as shown in Figures 3 and 4. Another fact supporting the asphalt-related nature of the interface is the softening it experiences due to moisture conditioning (Figure 4), a phenomenon that is widely known to happen for asphalt mixes (Kanitpong et al., 2003). Finally, the exponent of the creep function fitted after tensile creep tests was the same for both the asphalt and the interface. The time-dependent nature of both the asphalt and the interface has important implications in terms of expected BCOA performance, which can likely be extended to all concrete slabs on asphalt bases. This is because the interface-asphalt system will react differently under rapid traffic loads and under slow environment-related actions. In the latter case, the asphalt and the interface will tend to flow and, consequently, will produce low resistance to concrete thermal and drying shrinkage related contraction, which decreases tensile stresses in the concrete. In the former case, the asphalt and interface will be stiff, which increases the bending resistance of the composite system and reduces tensile stresses in the concrete caused by bending. Experimental data from creep and frequency sweep tests, as reflected in Figures 2 and 3, indicate that the stiffness difference is at least one order of magnitude between the two loading conditions. This means that the interface-asphalt system has the “ideal” characteristics

of a base for concrete, allowing the concrete slabs to freely move under temperature and moisture-related shrinkage actions without building up stresses and, at the same time, will providing stiff support under traffic loads, with the important caveat that the asphalt and interface are not wet or damaged. This stiffness duality is not considered in current BCOA design methods. Experimental data shown above indicate that water has a clear negative impact on concrete-asphalt bonding. In the first place, the estimated dynamic shear stiffness of the interface decreased by at least one order of magnitude after water saturation (equivalent shear stiffness modulus around 100 MPa in dry condition versus less than 10 MPa in wet condition, Figure 2 Figure 4). In addition, the stiffness decrease in wet fatigue testing had very little recovery due to thixotropy and/or healing after one month at rest (Figure 6). In comparison, the stiffness of the asphalt recovered more than 50 percent after dry fatigue testing (note y-axis in Figure 6 is in log scale). As explained above, the negative effects of water for the concrete-asphalt debonding process have been reported before. Results from this study indicate that the problem is not only debonding, but the inability of the low stiffness wet interface-asphalt system to provide the concrete overlay with much support to resist bending in the composite system. This negative effect would happen as soon as water enters the interface, without having to wait for fatigue damage to occur. This result indicates that moisture damage resistance is a primary performance criterion for design of asphalt bases for concrete pavement. Prior work at the UCPRC has produced mix designs for asphalt bases for concrete pavement that include lime treatment and increased asphalt contents as an aid to achieving field compaction to close to zero airvoids to decrease permeability and increase moisture resistance. It should be remarked

that the reason why moisture has a negative impact on the concrete-asphalt interphase is still unknown, as it is essentially unknown the nature if this interphase. A common belief exists that BCOA does not work well when placed on new asphalt mixes, and a few field experiences support this thinking (Sheehan et al., 2004). This has been typically attributed to poor bonding between concrete and new asphalt. Nonetheless, this laboratory experiment shows that the failure of a composite specimen can take place in the asphalt as well as in the interface , as reflected in Figure 1b above. Furthermore, fatigue testing presented in Figure 5 did not show differences between asphalt and composite specimens, which indicates that shear fatigue resistance of the interface was higher than shear fatigue resistance of the asphalt. As explained above, concrete was cast on slabs of new (rubberized) asphalt in this experiment, without applying any surface texturing. Iowa shear testing (Iowa DOT, 1991) was conducted to further compare shear strengths of asphalt versus interfaces. In this test, a composite specimen is sheared at a constant stress rate of 2.8 to 3.5 MPa/min. Asphalt and concrete parts of the specimen are held by metal rings, 15 mm below and above the interface, respectively. Two specimens were tested, resulting in an average strength of 740 kPa. This value is slightly over the 700 kPa (100 psi) reference strength that has been reported in the literature as indicating a good bond (Rasmussen and Rozycki, 2004). As shown in Figure 7, failure in the specimens occurred both at the asphalt and the interface, but the area that cracked was greater in the asphalt than in the interface. Again, this result indicates that bonding between concrete and new asphalt can be as strong as the asphalt itself. This suggests that it is likely that the bad performance that has been reported for BCOA on top of new asphalt was related to the stiffness and the strength of the new asphalt, rather than the bonding between concrete and asphalt. This

outcome widens the field of application of BCOA, that could be used to rehabilitate asphalt pavements in poor condition as well. A new asphalt layer would be placed before the concrete overlay. This asphalt mix should be engineered to maximize its stiffness and fatigue resistance.

5

Summary and Conclusions A series of laboratory tests has been conducted on asphalt and composite

(concrete on asphalt) cylindrical specimens cored from composite slabs in order to characterize interface stiffness and strength/fatigue resistance under different loading and environmental conditions. Loading conditions were intended to reproduce the rapid pulses of shear from traffic and the slow tension and shear pulses caused by temperature and moisture-related changes in the concrete. Environmental conditions included wet and dry, and temperatures between 5 and 40°C, a range applicable to asphalt bases located under 100 to 175 mm thick concrete overlays in California. Concrete was produced with Type I/II cement, and the asphalt base was a new gap-graded mix with 7.4% asphalt rubber binder content. Preliminary conclusions from analysis of this testing are as follows: •

The mechanical nature of the concrete-asphalt interface was strongly related to that of the asphalt. The interface stiffness showed a clear time dependency, and it significantly softened –by at least one order of magnitude– under wet conditions. This behaviour needs to be considered in BCOA design procedures: the same asphalt and interface stiffness should not be assumed under both traffic (rapid) and environmental (slow) loading.

•

Experimental results from this study do not support the common belief that concrete does not bond well to new asphalt.

o Shear and tensile strengths of the concrete-asphalt bond were comparable to shear and tensile strengths of the asphalt. o Fatigue of the composite specimens tested in shear at 40°C primarily occurred in the asphalt, and not in the interface. •

Experimental results from this study are in line with the common belief that water is one of the critical factors leading to failure of BCOA sections, because it damages the interface and the asphalt. Nonetheless, results from this study indicate that the negative action of water happens even before debonding, as the interface and asphalt soften to a point where they no longer contribute significantly to the bending resistance of the composite system.

References Burnham, T. R. (2006). The effect of joint sealing on the performance of thin whitetopping sections at MnRoad, No. MN/RC-2006-18. Chabot, A., Balay, J. M., Pouteau, B., and De Larrard, F. (2008). FABAC accelerated loading test of bond between cement overlay and asphalt layers. In Sixth RILEM CP Conference, Chicago. In Taylor and Francis Group Proceedings (pp. 13-23). Iowa Department of Transportation, Highway Division (1991). Method of Test for Determining the Shearing Strength of Bonded Concrete, Test Method No. Iowa 406-C, Iowa DOT. Kanitpong, K., and Bahia, H. U. (2003). Role of adhesion and thin film tackiness of asphalt binders in moisture damage of HMA. In Association of Asphalt Paving Technologists Technical Sessions, 2003, Lexington, Kentucky, USA (Vol. 72). Lea, J. D., and Harvey, J. (2012). Simplified thermal modeling approach used in CalME. In Transportation Research Board 91st Annual Meeting (No. 12-2938). Li, Z., Dufalla, N., Mu, F., and Vandenbossche, J. M. (2013). Bonded Concrete Overlay of Asphalt Pavements Mechanistic-Empirical Design Guide (BCOA-ME). User’s Guide, FHWA TFP Study, 5, 165.

Mateos, A., Harvey, J., Paniagua, J. C., and Paniagua, F. (2015). Development of Improved Guidelines and Designs for Thin Whitetopping: Literature Review (No. UCPRC-TM-2015-01). Mateos, A., Harvey, J., Paniagua, J., Paniagua, F., and Fan, A. (2016). Role of Concrete-Asphalt Interface in Bonded Concrete Overlays of Asphalt Pavements. In 8th RILEM International Conference on Mechanisms of Cracking and Debonding in Pavements (pp. 489-494). Springer Netherlands. Mu, F., and Vandenbossche, J. M. (2011). Development of Design Guide for Thin and Ultra-Thin Concrete Overlays of Existing Asphalt Pavements, Task 2: Review and Selection of Structural Response and Performance Models(No. MN/RC 2011-25). Rasmussen, R. O., and Rozycki, D. K. (2004). Thin and ultra-thin whitetopping: A synthesis of highway practice (Vol. 338). Transportation Research Board. Sheehan, M. J., Tarr, S. M., and Tayabji, S. D. (2004). Instrumentation and field testing of thin whitetopping pavement in Colorado and revision of the existing Colorado thin whitetopping procedure (No. CDOT-DTD-R-2004-12,). Colorado Department of Transportation, Research Branch. State of California Department of Transportation, Materials Engineering and Testing Services (2006). Asphalt Rubber Usage Guide. Vandenbossche, J., and Barman, M. (2010). Bonded Whitetopping Overlay Design Considerations for Prevention of Reflection Cracking, Joint Sealing, and the Use of Dowel Bars. Transportation Research Record: Journal of the Transportation Research Board, (2155), 3-11. Yu, H. T., and Tayabji, S. (2007). Thin Whitetopping—the Colorado Experience (No. FHWA-HIF-07-025). Zapata, C. E., and Houston, W. N. (2008). Calibration and validation of the enhanced integrated climatic model for pavement design. NCHRP Report 602. Transportation Research Board.

Tables Table 1. Summary of mechanical characterization tests. Stiffness Traffic loading Env. loading

Strength/Fatigue Resistance Traffic loading Env. loading

Shear (*) (SST)

frequency sweep 4 / 20 / 40 °C dry / wet

creep test 4 / 20 / 40 °C dry

fatigue test, 1200 με 10 Hz, 40 °C dry / wet

No test was conducted

Tensile (*) (UTM)


creep test 20 °C dry


monotonic test 1 / 10 kPa/s, 20 °C dry

(*) A brief description of each test is presented below: • Shear frequency sweep test. A sinusoidal shear relative displacement was applied between top and bottom faces of the cylindrical specimen (150 mm diameter). Applied shear force and displacement were measured, and then used to determine average shear stress (force divided by specimen area) and shear deformation (shear relative displacement divided by specimen height). The dynamic shear modulus is determined as the ratio between the two previous variables, and the phase angle is determined as the lag between force and displacement signals. Due to the viscous nature of asphalt, both dynamic shear modulus and phase angle depend on the frequency of the sinusoidal displacement. A wide range of frequencies was applied in this test, starting with 10 Hz and finishing with 0.01 Hz. A dynamic modulus value was a obtained for each frequency, and the same applies to the phase angle. It must be indicated that specimen height was asphalt height for both asphalt and composite specimens. For the composite specimens, this is equivalent to assuming Portland cement concrete deformation is negligible compared to the deformation of the asphalt part. In any case (asphalt and composite specimens), relative displacement is measured between top and bottom platens to which the specimen is glued (Figure 1a). This is applicable to the three shear tests described here. • Shear creep test. A constant shear load was applied between top and bottom faces of the cylindrical specimen (150 mm diameter). Load was maintained during 10 seconds, and then it was released. Shear relative displacement, top to bottom faces of the specimen, was measured. This displacement was used to determine shear strain (displacement divided by specimen height). Due to the viscous nature of asphalt, the shear strain continuously increased during the 10-second loading. Recovery of the strain, after the 10-second loading, was measured for 100 additional seconds. The strain developed during the loading period can be used to fit the shear creep compliance function of the specimen. The shear creep compliance function can be also obtained from the strain recovery (after the 10-second loading), which was the approach followed in this study. The shear creep compliance function defines the time-dependent shear deformation of the specimen under a unit shear stress. This function depends on the test temperature. • Shear fatigue test. A 10-Hz sinusoidal shear relative displacement was applied between top and bottom faces of the cylindrical specimen (150 mm diameter). A large number of cycles was applied, 1 million. Magnitude of the peak to peak displacement was constant during the test (constant peak to peak shear strain). Dynamic modulus was determined as explained above for the shear frequency sweep test. Since the shear strain was high, 1200 µɛ peak to peak, the specimen experienced damage and, consequently, the dynamic shear modulus decreased. The output of this test is the evolution of the dynamic modulus versus the number of cycles or load applications. • Tensile creep test. A constant axial tensile force was applied to a cylindrical composite specimen (100mm diameter). Deformation was measured at the asphalt and across the concrete-asphalt interphase. The load was maintained for 10 seconds, and then it was released. Deformation was measured during loading, and also during 100 additional seconds recovery. Measured deformation was used to calibrate the creep tensile functions of asphalt and interface. These functions define the time-dependent tensile deformation of asphalt and interphase under a unit tensile stress. • Tensile monotonic test. An increasing axial tensile load is applied to a cylindrical composite specimen (100mm diameter). Load rate was set to produce in the specimen a constant tensile stress rate of either 1 or 10 kPa per second. Load was continuously increased until complete failure of the specimen. Axial and interphase deformations were measured during this test.

List of Figures Figure 1. SST (a) and UTM (b) composite specimens Figure 2. Dynamic shear modulus of asphalt and composite specimens Figure 3. Shear creep compliance of asphalt and composite specimens Figure 4. Effect of water on shear stiffness of asphalt and composite specimens (Mateos et al., 2016) Figure 5. Stiffness and phase angle evolution during shear fatigue tests (1200 με, 40°C) Figure 6. Stiffness reduction after 106 loading cycles (1200 με shear strain), and recovery after 1 month at rest Figure 7. Composite specimens after Iowa shear test

Top platen

SST testing frame LVDT

Shear composite specimen, ф150mm and h75mm (concrete 25mm + asphalt 50mm). Shear deformation is measured with the LVDT attached to the top testing platen. LVDT target is attached to the bottom testing platen.

a Figure 1. SST (a) and UTM (b) composite specimens

Tensile test specimen, ф100mm and h190mm (concrete 100mm + asphalt 90mm). Left and right LVDTs in the picture measure asphalt strain, and LVDTs in the front and back of the specimen (the latter is not visible in the picture) measure interface opening plus 10 mm of asphalt.

b

Dynamic shear modulus (MPa)

10 000

4 °C

Asphalt

20 °C 40 °C master curve

1 000

4 °C Composite 20 °C 40 °C

100

master curve 4 °C

Interface

20 °C 10 1.E‐05

40 °C 1.E‐03

1.E‐01

1.E+01

1.E+03

Reduced frequency (Hz)

1.E+05

Interface shear modulus is determined assuming a theoretical thickness of 2.5 mm

Figure 2. Dynamic shear modulus master curves of asphalt and composite specimens, and asphalt/concrete interface

Shear Creep Compliance (/MPa)

1000000

Interface 100000

4 °C 10000

Asphalt

20 °C

Composite

40 °C Creep Func.

1000

Composite @ 20°C Interface @ 20°C

100 1.E‐05

1.E‐03

1.E‐01

1.E+01

Reduced time (s)

1.E+03

1.E+05

Interface compliance is determined assuming a theoretical thickness of 2.5 mm

Figure 3. Shear creep compliance of asphalt and composite specimens, and asphalt/concrete interface


10 000

master curve (20°C) of dry Asphalt 1 000

wet asphalt @ 20 °C

master curve (20°C) of dry Composite

wet composite @ 20 °C

100

wet interface @ 20 °C 10

1 1.E‐05

1.E‐03

1.E‐01

1.E+01


1.E+03

1.E+05

Interface shear modulus is determined assuming a theoretical thickness of 2.5 mm

Figure 4. Effect of water on shear stiffness of asphalt and composite specimens (Mateos et al., 2016)

40° 35°

200 30° 25°

150

20° 100

15° 10°

50 5°

|G*| dry AC

Phase Angle, ϕ

Dynamic Shear modulus (MPa)

250

|G*| dry Comp. |G*| wet AC |G*| wet Comp. ϕ dry AC ϕ dry Comp. ϕ (wet AC) ϕ wet Comp.

0 0° 1.E+00 1.E+01 1.E+02 1.E+03 1.E+04 1.E+05 1.E+06

number of cycles

Figure 5. Stiffness and phase angle evolution during shear fatigue tests (1200 με, 40°C)


10 000

dry AC master curv. dry Comp. master curv.

1 000

dry AC (after fatigue) dry Comp. (after fatigue) dry AC (after rest) dry Comp. (after rest)

100

wet AC (after fatigue) wet Comp. (after fatigue) wet AC (after rest) 10 1.E‐05

wet Comp. (after rest) 1.E‐03

1.E‐01

1.E+01

1.E+03

1.E+05


Figure 6. Stiffness reduction after 106 loading cycles (1200 με shear strain), and recovery after 1 month at rest

Figure 7. Composite specimens after Iowa shear test