BCOA interface friction

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Mechanical characterisation of concrete-asphalt interface in bonded concrete
overlays of asphalt pavements
Article in European Journal of Environmental and Civil Engineering · April 2017
DOI: 10.1080/19648189.2017.1311808
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University of California, Berkeley
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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
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3. 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, concrete-
asphalt 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

4. 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 concrete-
asphalt 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.

5. 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 disc-
shaped 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.

6. 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

7. 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

8. 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

9. 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.

10. 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

11. 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

12. 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

13. 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 air-
voids to decrease permeability and increase moisture resistance. It should be remarked

14. 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

15. 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.

16. 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.

17. 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.

18. Tables
Table 1. Summary of mechanical characterization tests.
Stiffness Strength/Fatigue Resistance
Traffic loading Env. loading 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)
No test was
conducted
creep test
20 °C
dry
No test was
conducted
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.

19. 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

20. 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.
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.
a b
Figure 1. SST (a) and UTM (b) composite specimens
SST testing frame
Top platen
LVDT

21.
Figure 2. Dynamic shear modulus master curves of asphalt and composite specimens, and
asphalt/concrete interface
Figure 3. Shear creep compliance of asphalt and composite specimens, and asphalt/concrete
interface
10
100
1 000
10 000
1.E‐05 1.E‐03 1.E‐01 1.E+01 1.E+03 1.E+05
Dynamic shear modulus (MPa)
Reduced frequency (Hz)
4 °C
20 °C
40 °C
master curve
4 °C
20 °C
40 °C
master curve
4 °C
20 °C
40 °C
Asphalt
Composite
Interface
Interface shear modulus is determined
assuming a theoretical thickness of 2.5 mm
100
1000
10000
100000
1000000
1.E‐05 1.E‐03 1.E‐01 1.E+01 1.E+03 1.E+05
Shear Creep Compliance (/MPa)
Reduced time (s)
4 °C
20 °C
40 °C
Creep Func.
Composite @ 20°C
Interface @ 20°C
Composite
Interface
Asphalt
Interface compliance is
determined assuming a
theoretical thickness of 2.5 mm

22.
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)
1
10
100
1 000
10 000
1.E‐05 1.E‐03 1.E‐01 1.E+01 1.E+03 1.E+05
Dynamic shear modulus (MPa)
Reduced frequency (Hz)
wet asphalt
@ 20 °C
wet composite
@ 20 °C
wet interface @
20 °C
Interface shear modulus is determined
assuming a theoretical thickness of 2.5 mm
master curve (20°C)
of dry Asphalt
master curve (20°C)
of dry Composite
0°
5°
10°
15°
20°
25°
30°
35°
40°
0
50
100
150
200
250
1.E+00 1.E+01 1.E+02 1.E+03 1.E+04 1.E+05 1.E+06
Phase Angle, ϕ
Dynamic Shear modulus (MPa)
number of cycles
|G*| dry AC
|G*| dry Comp.
|G*| wet AC
|G*| wet Comp.
ϕ dry AC
ϕ dry Comp.
ϕ (wet AC)
ϕ wet Comp.

23.
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
10
100
1 000
10 000
1.E‐05 1.E‐03 1.E‐01 1.E+01 1.E+03 1.E+05
Dynamic shear modulus (MPa)
Reduced frequency (Hz)
dry AC master curv.
dry Comp. master curv.
dry AC (after fatigue)
dry Comp. (after fatigue)
dry AC (after rest)
dry Comp. (after rest)
wet AC (after fatigue)
wet Comp. (after fatigue)
wet AC (after rest)
wet Comp. (after rest)
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