Critical Problems with Using the Asphalt Ductility Test as a Performance Index for Modified Binders

The final version of this paper was published in:

Transportation Research Record: Journal of the Transportation Research Board, 2014

VOL 2370, Pages 84-91. DOI: 10.3141/2370-11

http://trb.metapress.com/content/b135086hu5444244/?genre=article&id=doi%3a10.3141%2f2370-11

Critical Problems with Using the Asphalt Ductility Test as a Performance

Index for Modified Binders

Paper #13-4125

Revised Version Submission date: February 28, 2013

Word count: 5386 plus 8 Table and Figures (250 words)

Total number of words: 7386

Hassan A. Tabatabaee, Ph.D. (Corresponding Author)

Research Associate

Department of Civil and Environmental Engineering

University of Wisconsin - Madison, 53706

Email: [email protected]

Phone: (608)-890-3321

Cristian Clopotel, Ph.D.

Graduate Research Assistant




Amir Arshadi

Graduate Research Assistant




Hussain Bahia, Ph.D.

Professor




Submitted for publication in the

mailto:[email protected]




Tabatabaee, Clopotel, Arshadi, Bahia, 2013 1

1

Transportation Research Record: Journal of the Transportation Research Board

February, 2013


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ABSTRACT

Despite the adaptation of advanced binder rheology characterization methods by many

agencies, the asphalt ductility test is still under use in some specifications in the USA and

a few countries as a performance indicator for asphalt modification. In the present study a

set of binders modified with two types of commonly used elastomeric polymer modifiers

were characterized in terms of fundamental, well-defined, binder properties known to

reflect rutting and fatigue resistance, results of which showed no correlations with binder

ductility.

Additionally, a test procedure was developed using a Dynamic Shear Rheometer (DSR) as a

surrogate to the conventional ductility test, results of which showed that both elastomeric

modified binders are much more ductile than conventional binders, even when conventional

ductility showed a loss of ductility. Finite Element Modeling (FEM) was used to show the

significant effects of decreasing true strain rates with elongation on the sample’s stress and strain

state due to the constant cross head speed in the conventional ductility test. Due to the well-

known dependency of failure stress and strain of viscoelastic material on strain-rate and

temperature, comparing binders with varying ductility values measured in the conventional test

is essentially flawed as it is equivalent to comparing them at different temperatures, thus

fundamentally unreliable as an indicator of the asphalt’s performance in the pavement. It is

therefore strongly recommended that the practice of using low temperature conventional ductility

be removed from modified binder specifications, or the DSR procedure proposed in this paper be

used to evaluate modified asphalts as a preferable test method.


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INTRODUCTION

The asphalt ductility test, as currently carried out in accordance to the ASTM D113-07

specification (1), is a measure of the ultimate elongation of an asphalt binder sample at a

constant rate and a specific temperature before breaking. Despite the adaptation of advanced

binder rheology characterization methods by many agencies, the asphalt ductility test is still

under use in some specifications in the USA and a few countries as a performance indicator for

asphalt modification, especially at lower service temperatures.

The significance of the results of the ductility test as a measure of asphalt performance

has been debated due to its empirical nature, low reproducibility, and unclear relationship

between the measured results and any fundamental material properties (2). Historically in

absence of advance characterization methods, some researchers have used the ductility test as a

qualitative measure for determining the level of aging in asphalt binders after laboratory aging or

extraction from the field (3,4,5,6,7). Doyle tested four asphalt binders recovered from field

section using the ductility at 13°C and noticed that binders that cracked more seemed to have a

lower ductility at this temperature, although the results did not correlate with any other low

temperature properties in the binders. Doyle suggested that elongation rate as low as 1°C/min

may still be too high to relate to low temperature shrinkage in pavements (4).

Halstead performed an extensive research on the relation of asphalt ductility to pavement

performance. Although he noticed that in some cases combining ductility and penetration data

may be useful for characterization asphalt binders, especially with regards to age hardening, he

was unable to define the exact role of the ductility measure on asphalt pavement durability,

concluding that more research was needed to define the significance of ductility for pavement

durability. Based on the results he further concluded that “it is most likely that the ability of

asphalt to undergo elongation is not the primary characteristic affecting durability.” He

suggested that ductility may be a function of the internal phase relationship of the asphaltic

constituents (5).

Studies conducted by Kandhal in the 1960’s (8,9) showed that a pavement with ductility

above 10 cm at 15.6°C had relatively good performances in terms of thermal cracking. In

addition it was observed that when the ductility decreased to 3-5 cm at 15.6°C serious cracking

was observed.

An elaborate study regarding the correlation between the properties of the asphalt binders

and the properties of the asphalt concrete mixes was conducted by Goodrich (10). The author

showed that temperature susceptibility, forced ductility, toughness-tenacity and low temperature

ductility do not correlate with mixes performances.

Vonk et al (11), suggested that bitumen’s homogeneity is a condition for good ductility,

and the lower the asphaltene content, the better ductility. In addition the authors compared

ductility with Frass breaking point and concluded that ductility cannot be used as a surrogate for

the test. Citing Reese’s work (12), Vonk et al showed inconsistencies between ductility ranking

and reflective cracking ranking. The authors also showed that ductility highly sensitive to testing

temperature and rankings can be changed by simply altering the test temperature.


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An alternative methodology to the ductility was proposed by Ruan et al. (13,14), using

the Dynamic Shear Rheometer (DSR) by measuring G’/(η’/G’) for the binder. The parameter

correlated well with ductility when unmodified binders were tested and ductility was below 10

cm. However, the author showed a weak correlation between the ductility and G’/(η’/G’) for

modified binders.

In recent years researchers have focused on the applicability of ductility specification to

modified asphalt binders. Vonk and Korenstra in a detailed study reported that due to the different

structure in polymer modified binders; the ductility test would be inherently measuring different

properties of the asphalt binders depending on the presence of modification, thus concluding that

ductility is not a suitable performance indicator and binder selection tool (11).

Review of literature since the 1940s shows that limited work has focused on relating

ductility results to pavement performance, especially in terms of fundamental material

properties, with widely varying results and conclusions. The general lack of repeatable and

reliable correlation between ductility results and pavement performance in previous studies,

especially at lower temperatures, has limited the practical usage of this tool to specify suitability

of asphalt binders for pavement applications. On the other hand almost all studies in the literature

relating ductility to performance have been based on unmodified binders.

What is of utmost importance is that no fundamental analysis and interpretation of the

ductility test results has been presented to justify the use of this test, and the resulting ductility

measure, as a performance index. The objective of the present study was to identify and explain

the fundamental limitations of this test, to study the stress and strain development in the model

during elongation through finite element modeling, and to compare the ductility results to other

more fundamental binder damage resistance characterization parameters to further investigate the

applicability of this test as a reliable performance indicator, as well as introduce a preferable

alternative surrogate test method using the DSR.

EXPERIMENTAL DESIGN

In the present study a relatively large set of neat binders were modified with different dosages of

two commonly used elastomeric polymer modifiers, namely Elvaloy® and Styrene-Butadiene

Styrene (SBS). The studied materials are defined in TABLE 1.

Damage resistance testing of asphalt binders is recognized as a more direct method of

assessing performance of pavements based on engineering properties of the constituents. These

tests to some extent side step the complexity of non-linear viscoelasticity and stress dependency

of asphalt binder modulus by ranking binders based on ultimate failure parameters instead of

stiffness and modulus. This becomes especially advantageous when characterizing modified

asphalt binder for which non-linearity may be significantly affected by the modification type and

level. Such specification methods were suggested by researchers during the NCHRP 9-10 project

(15), and later used as a basis for the development of many new asphalt binder testing procedures

such as the Multiple Stress Creep and Recovery (MSCR) (16), and the Linear Amplitude Sweep

(LAS) (17). In the present study MSCR and LAS were performed at the high and intermediate

performance grade temperatures of the binders and compared to the ductility results to

investigate the potential relationship between ductility values and elongation failure through

ductility testing at 4°C.


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DUCTILITY AND DAMAGE RESISTANCE CHARACTERIZATION

FIGURE 1(a) and (b) plots the number of cycles to fatigue failure, Nf, against ductility results at

4°C. Nf was calculated based on the LAS results at the intermediate grade temperature, by

means of the Viscoelastic Continuum Damage framework (VECD), based on a 35% decrease in

modulus as the failure criteria. This parameter can be calculated assuming any inputted cyclic

strain amplitude, which for this study was set at 2.5 and 5%. The use of 4°C for the ductility

testing was historically based on practical limitations of testing at lower temperatures, especially

using a water bath. In reality, 4°C is closer to the low end of the fatigue damage temperature

range, rather than low temperature cracking, although it must be noted that fatigue damage

results from repeated loading and unloading, unlike the monotonic loading procedure used in the

ductility test. If ductility at 4°C is any indicator of crack resistance at that temperature, some

correlation with fatigue damage resistance would be expected. Contrary to this expectation, Nf

values at both strain levels show no correlation with ductility results. This lack of relationship

highlights the potential lack of any fundamental relationship between ductility and cracking

damage in asphalt binders.

FIGURE 1(c) and (d) depicts the relationship observed for ductility at 4°C and MSCR

parameters at 3.2 kPa (Jnr and Percent recovery) at the binder high temperature grade for a

number of neat and elastomeric modified binders using both Elvaloy® and SBS. There does not

seem to be any correlation between the percent recovery and Jnr values and the corresponding

ductility values. Although stiffness and the ability to elongate are not fundamentally related, one

may expect that binders with higher ductility to tend to have lower stiffness (higher compliance).

FIGURE 1 shows that no such trend exists for the studied binders.

SAMPLE GEOMETRY EVOLUTION DURING ASPHALT DUCTILITY TEST

A point of concern when considering ductility of asphalt binders is the continuous change of

sample geometry during elongation. The change of geometry is governed by the Poisson ratio of

asphalt material, while the resulting stress in the sample will be affected by stiffness and rate of

relaxation in addition to the continuous cross sectional change of the sample. Binder

modification, especially with polymers, can potentially affect both of these aspects as the

elastomeric three dimensional networks may affect both the overall Poisson ratio of the asphalt

and the rate of stress relaxation as the sample elongates at the rate of 1 or 5 cm/min. Thus

comparison of the total elongation at any temperature using the ductility between different

modified and unmodified binders would be significantly affected by the varying stress states

between the samples.

The potential effect of change in ductility sample geometry in the test result ranking is

analyzed with two approaches. In one approach the dynamic shear rheometer (DSR) is used to

carry out a continuous rotation test on asphalt samples at 4°C. The test methodology follows that

of the Binder Yield Energy Test (BYET) (18). The strain rate experienced by the sample in the

ductility was calculated and converted to the appropriate strain rate to be used in the DSR

method. Such calculations were originally carried out for the development of the “Elastic


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Recovery – DSR” (ER-DSR) procedure at the University of Wisconsin-Madison (19). This DSR

BYET test was used as a comparison to the ductility test as it too imposes a continues

engineering strain on the sample to the point of apparent failure, while the sample geometry

evolution in this test may be better defined with a consistent effective cross section as opposed to

the asphalt ductility test. Multiple neat and two types of elastomer modified binders (Elvaloy®

and SBS) from different sources were tested using this method and compared to the ductility

results at 4°C. The results are shown in FIGURE 2. It should be mentioned that the DSR ductility

procedure was found to be very repeatable, with an average strain at peak stress coefficient of

variation of 8%.

The results shown in FIGURE 2 indicate that the conventional ductility test results show

erratic and inconsistent results for both elastomeric modification types for the six base binders

studied. However, conducting the ductility test on the same materials in a controlled geometry

condition in the DSR showed that both elastomeric modifiers significantly improved the ductility

at 4°C, regardless of the ductility test results. The addition of the elastomer clearly enhanced the

strain to peak stress by two to six times (200% - 600%), depending on base binder source, all

other conditions being identical. The difference in results between the conventional ductility and

the DSR ductility is attributed to the sample and load geometry difference.

Conventional ductility and DSR ductility results were plotted against each other in terms

of strain at respective failure criteria, as shown in FIGURE 3. A few notable observations are

made:

First, for the neat asphalt binders the strain at failure from the DSR-ductility and the

ductility test match up relatively well, confirming that the DSR procedure is capable

of capturing an equivalent response to that of the conventional ductility test.

Second, the DSR-ductility results from the Elvaloy-modified binders show much

higher strains at failure than that of the conventional ductility test, on the other hand,

SBS-modified binders seem to cover a range of responses from the conventional

ductility, while generally faring better in the DSR-ductility test.

Although the ductility test is used by some agencies as an indicator of the presence of

elastomeric modification, the results shown in FIGURE 2 and FIGURE 3 for two types

of elastomeric modification and various base binder types show that the conventional

ductility test is unable to consistently detect the presence of an elastomeric modification. On

the other hand, the DSR-Ductility methods consistently indicated the presence of elastomers

in the tested binders for both types of elastomeric modification and all base binder types.

The main finding from this comparison is important since it shows that the

conventional ductility test could be misleading due to the confounding effects of the

change in geometry during testing. This effect seems to bias results toward specific

types of modification systems, yielding results that cannot be supported by that of the

more robust DSR-ductility methodology. The DSR ductility results clearly show that

both elastomeric modifiers can indeed significantly improve the ductility of neat

binders.


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STRAIN RATE IN ELONGATION RHEOLOGY AND FAILURE PROPERTIES OF

VISCOELASTIC MATERIAL

Uniaxial elongational flows are believed to be strongly aligning flows in that a high degree of

orientational anisotropy in polymer chain segments can be induced (20). Thus the extreme test

conditions may result in widely varying responses for different polymer modification types, with

little relevance to realistic stress and strain states experienced by asphalt binders in the pavement.

Uniaxial elongation at constant strain rate is also classified as a strong flow because such a

deformation history has a tendency to stretch polymer chain segments. The high degree of chain

segment orientation and stretch induced by uniaxial elongation can result in rather dramatic

rheological responses including the phenomenon known as strain hardening, where the

elongational viscosity increases well above the linear viscoelastic prediction (20). In polymer

science, early attempts at extensional rheometry of polymer solutions were met with

disappointment due to the strong strain-rate and strain sensitivity of the extensional stresses

exhibited in polymer solutions and the widely disparate design of the instruments which resulted

in nonhomogeneous kinematics (21).

Material science studies have shown that true material failure properties can only be

achieved under constant-true-strain conditions (22,23,24). When experiments are carried out at a

constant crosshead speed condition (constant engineering strain), such as in the asphalt ductility

test, the specimen experiences a decreasing true strain rate during the test, thus making physical

interpretation of results difficult, especially in the non-elastic regimes (22,23). Thus material

scientists have gone to great length since early 1970s to maintain constant true strain rate in the

samples during elongational and tensile tests (25), such as detachable extensional rheometer

fixtures for torsional rheometers (21).

The study of stress and strain response in viscoelastic materials in not a new topic, and

has been investigated at length for many decades. Theoretical solutions and constitutive

equations have been developed for almost any loading geometry and time-temperature condition,

taking into account temperature and rate dependency, and more recently non-linear

viscoelasticity. Researchers have shown that for linear viscoelastic material time temperature

superposition is not only applicable to measurements of viscosity and stiffness, but to various

other physical properties, provided they involve the same molecular motions governing the

viscoelastic stiffness properties (26,27). Such observations have enabled researchers to plot

“failure strength envelope” as a function of ultimate strain for viscoelastic material (28). A

failure envelope is developed by plotting failure stress against strain at reduced strain rates

(26,28). Furthermore, the applicability of time temperature superposition to failure stress and

strain in viscoelastic material implies that the viscoelastic response of the amorphous chains

dominates failure properties (26). Many researchers studying viscoelastic fracture have shown

that there is a close dependency between failure stress and strain rate, such that the stress

decreases monotonically as the strain rate is decreased. On the other hand, researchers have

observed the existence of a “definite maximum strain at failure” for viscoelastic material tested

to failure at different strain rates (26,27,29,28). The strain at failure increases as the strain rate

decreases down to a certain strain rate, after which any decrease in strain rate will cause the

strain at failure decrease (27).

It has been shown that failure envelope in many instances works sufficiently well for

asphalt mixtures (30,31). Monismith et al. also showed that asphalt concrete failure stress at


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different temperatures, strain rates and test conditions behaved in a fashion similar to that

described by Smith (28,31), if tests are conducted at constant strain rates. In the failure envelope

for asphalt pavements tested at different temperatures and conditions, the higher failure stress

levels corresponded to lower test temperatures and higher strain rates. It was also be observed

that failure of samples tested under constant strain can be plotted on the same failure envelope as

those tested at the same temperature but under constant stress.

In the asphalt ductility test, due to the constant cross head speed conditions, the true-

strain-rate decreases with time and elongation. This decrease can cause the strain rate to change

by more than an order of magnitude only after a few cm of elongation. Equations 1 and 2 show

the formulation for calculation of “engineering strain” (also known as Cauchy strain) and “true

strain” (also known as natural strain). Mathematically, the true strain is the sum of all

instantaneous engineering strains and is calculated as follows:

[eq. 1]

∫

[eq. 2]

In which:

and are the “engineering” and “true” strain, respectively, and

l is the sample length, with l0 being the initial length.

At small strain levels the engineering strain and true strain do not significantly differ,

thus usable results may be derived from tests performed using constant engineering strain rates,

provided the strain levels remain low. But for the case of the asphalt ductility test, for which the

sample is elongated to many times its original length, such approximations lead to significant

error. This is demonstrated by using equations 1 and 2 to calculate the average engineering and

true strains and strain rates, assuming a 10 mm/min crosshead deflection rate and a 40 mm gage

length (defined as the uniform mid-section of the ductility dog-bone sample). FIGURE 4 shows

the results of the strain calculation for the ductility test. For a test run at a constant crosshead

speed, as is the case in the ductility test, the engineering strain increases proportionally to the

crosshead displacement rate, thus varies linearly as the sample elongates. It can be seen in

FIGURE 4(a) that when the elongation is small (i.e. less than 10 mm) the true and engineering

strain are very similar, but they rapidly diverge as the sample elongates. The linear increase of

engineering strain indicates a constant engineering strain rate as the sample elongates. On the

other hand, the true strain non-linearly varies and immediately decreases as the sample length

increases, as shown in FIGURE 4 (b).

In order to run an elongation test such as the asphalt ductility at a constant true strain rate

one must exponentially increase the crosshead speed to counteract the decreasing true strain rate.

This was done for the case of the asphalt ductility test by calculating the crosshead speed profile

needed to achieve constant true strain rates over 260mm of elongation (resulting in a 300 mm

sample mid-section). Three true strain rate levels, corresponding to the engineering strain rate,

the average true strain rate, and the terminal true strain rate, were targeted in these calculations.

The resulting crosshead deflection and strain rate profiles are shown in FIGURE 5. It can be seen

that there is a dramatic difference between strain at a constant cross head speed and the constant


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strain rates, potentially leading to significant variation in response of the viscoelastic medium at

different elongations, and consequently different strain rates.

Using the calculated crosshead deflection profiles, stress and strain curves were derived

for each true strain rate scenario using a finite element simulation. A model of a ductility sample

was developed in the FEM environment of ABAQUS, a powerful commercial FEM package.

Viscoelastic material properties of a typical asphalt binder were inputted using a 4-element

Prony series. FIGURE 6 shows the model geometry and simple stress distribution under a

constant crosshead speed.

Using the finite element model the four crosshead speed profiles shown in FIGURE 5(a)

were imposed upon the sample to assess the difference in stress-strain curves. As seen in

FIGURE 7, when testing at a constant crosshead speed (as in the ductility test), the slope of the

stress-strain curve will rapidly decrease as the true strain rate decreases, becoming relatively

constant at higher elongations where rate change becomes less pronounced. On the other hand,

performing the test at constant true strain rates using the cross head speed profiles shown in

FIGURE 5(a), the stress-strain curves build-up at a much more constant rate. Furthermore, the

simulation results presented in FIGURE 7 highlight the dramatic change in the rate of stress

build-up in the ductility sample for different true strain rates. For the constant crosshead

condition of the asphalt ductility test the rate of stress build-up is initially equal to that of the

high constant strain rate, but as the strain rate decreases with elongation the stress build-up rate

tends toward the lower constant strain level, as seen in the higher strains in the Figure.

An important consequence of the changing true strain rate in the ductility tests is the

continuous change of critical failure stress and strain in the sample due to strain-rate dependence

of failure criteria in viscoelastic material. Thus when comparing asphalt binders failing at

different elongation levels, one can’t simply rank the materials’ damage resistance based on

ranking of ductility results. This is due to the strain-rate dependency of failure criteria in

viscoelastic material, which results in significantly different failure conditions for samples failing

at different elongations and consequently different strain rates at the time of failure. Recall that

for asphalt and other viscoelastic materials failure stress and strain vary similarly with increasing

temperature or decreasing true strain rate, thus comparing the failure strain (or ductility) of two

binders at different elongation and strain rates will in essence be similar to comparing failure

properties of binders tested at two different temperatures. Thus the varying true strain-rate

conditions in the asphalt ductility test make any comparison between results for different binder

types and reliable interpretation of the results extensively difficult and fundamentally unsound.

CONCLUSIONS

A review of literature was conducted to show that there is no agreement among experts regarding

the value of the ductility test results in terms of defining pavement performance, with many

studies failing to find any reliable or consistent trend for the relationship between ductility and

pavement performance. Furthermore, ductility and fatigue and rutting performance tests were

conducted for a set of binders to compare ductility values with binder properties known to reflect

rutting and fatigue resistance. No correlations could be found to indicate the relevance of the

ductility in terms of fatigue or rutting resistance of asphalt.


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A detailed analysis of the conventional ductility test indicated that one of the fundamental

problems with the test is the variation of sample geometry during the test. It is clear that this

confounding effect can cause significant problems leading to erroneous evaluation and ranking

of binder performance, in particular with respect to modified binders. To prove this point

specialized testing in the DSR was designed to measure true ductility at 4°C while carefully

controlling the geometry of the samples. Results of the DSR ductility test clearly show that

modified binders are more ductile than conventional binders, and that the results of conventional

ductility can be misleading for modified binder systems. Through this specialized testing it was

found that both elastomeric modifiers studied can improve ductility by a range of 200% to as

high as 600% while the conventional ductility is showing a loss of ductility for these modifiers.

The results for a large set of binders confirm this finding and show no correlation between the

conventional ductility test and the true ductility measured with the DSR.

Finite Element Modeling (FEM) was used to further highlight the fundamental

deficiencies of the use of the ductility test, especially with regards to the decreasing true strain

rate resulting from the constant cross head speed in the ductility test. A discussion of strain-rate

dependency of viscoelastic materials was used to conclude that comparing binders based on

ductility at failure is essentially similar to comparing failure strain between samples tested at

different temperatures, due to similar dependency of failure stress and strain of viscoelastic

material to strain-rate and temperature. Such conditions make any direct comparison of different

binders with different properties highly complex and fundamentally unreliable as an indicator of

the asphalt’s performance in the pavement.

Based on this study it is highly recommended that ductility test at 4 o C be removed from

modified binder specifications or requirements. If ductility of modified binders is still needed in

a specification, then the proposed method for using a test that can maintain a stable geometry,

such as the DSR, be used. The use of the conventional ductility test is misleading and could

unscientifically disqualify modified binders with superior ductility than the conventional binders.

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31. Monismith, C. L., Secor, G. A., Secor, K. E. Temperature Induced Stresses and Deformations in

Asphalt Concrete. Proccedings of the Association of Asphalt Pavement Technologists Annual

Meeting,, pp 248-286.


13

List of Table Titles:

TABLE 1 Crude sources and modification types used in present study


14

List of Figure Captions:

FIGURE 1 (a) and (b) show relationship between LAS and Ductility at 4°C; (c) and (d) show the

relationship between MSCR parameters at 3.2 kPa and Ductility at 4°C.

FIGURE 2 (a) Effect of elastomeric modification on ductility at 4°C for different base binders,

(b) effect of elastomeric modification on real ductility ( % strain at Peak stress) at 4°C for

different base binders], (c) DSR-ductility shear stress against shear strain at 4°C.

FIGURE 3 Relation between ductility engineering strain at failure (assuming a 4 cm gauge

length) and DSR-ductility strain at peak stress, both measured at 4°C.

FIGURE 4 (a) True strain and engineering strain variation with ductility sample elongation, (b)

true strain rate and engineering strain rate variation as ductility sample elongates.

FIGURE 5 (a) Crosshead deflection curve with test time, and (b) true strain rate profile,

calculated for the asphalt ductility tests

FIGURE 6 FEM of asphalt ductility sample under elongation

FIGURE 7 Stress vs. Strain curve from finite element simulation using different crosshead speed

profiles.


15

TABLE 1 Crude sources and modification types used in present study

Crude Source Modification Types Base Performance Grade Modified Performance Grades

Russian Elvaloy 4170® (1 levels) PG 64-28 PG 70-28

Czech Republic Elvaloy AM® (2 levels) PG 64-22 PG 70-22, PG 76-22

Flint Hills SBS (2 Levels) PG 64-22 PG 70-22, PG 76-22

Valero Elvaloy 4170® (1 levels) PG 58-28 PG 64-28

Nustar SBS (2 Levels) Elvaloy

AM® (2 levels) PG 64-22 PG 70-22, PG 76-22

BP Elvaloy 4170® (2 levels) PG 58-28 PG 64-28, PG 70-28


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(a) (b)

(c) (d)

FIGURE 1 (a) and (b) show relationship between LAS and Ductility at 4°C; (c) and (d)

show the relationship between MSCR parameters at 3.2 kPa and Ductility at 4°C.

0.0E+00

2.0E+05

4.0E+05

6.0E+05

8.0E+05

1.0E+06

1.2E+06

0 10 20 30 40

LA

S N

f (2

.5%

Str

ain

)

Ductility @4⁰C (cm)

0.0E+00

5.0E+03

1.0E+04

1.5E+04

2.0E+04

2.5E+04

0 10 20 30 40

LA

S N

f (5

% S

tra

in )


0

10

20

30

40

50

60

70

80

0 20 40 60MS

CR

%R

eco

ver

y a

t 3

.2 k

Pa


0

1

2

3

4

5

6

0 20 40 60MS

CR

Jn

r a

t 3

.2 k

Pa

(1

/kP

a)



17

(a) (b)

(c)

FIGURE 2 (a) Effect of elastomeric modification on ductility at 4°C for different base

binders, (b) effect of elastomeric modification on real ductility ( % strain at Peak stress) at

4°C for different base binders], (c) DSR-ductility shear stress against shear strain at 4°C.

0

5

10

15

20

25

30

35

40

A B C D E F

Du

ctil

ity

at

4°C

(cm

)

Base Binder Type

Base Elastomeric ModificationX

0

200

400

600

800

1000

1200

1400

A B C D E FDS

R-D

uct

ilit

y P

eak

Str

ain

at

4°C

(-)

Base Binder Type

Base Elastomeric Modification

SBS

Modified Elvaloy

Modified

X

0

0.2

0.4

0.6

0.8

1

1.2

1.4

1.6

1.8

0 500 1000 1500 2000 2500

Str

ess

(MP

a)

Strain (%)

Nustar 76-22(ELV) Nustar 76-22(SBS)Nustar 70-22(SBS) Nustar

Elvaloy

Modified

SBS

Modified


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FIGURE 3 Relation between ductility engineering strain at failure (assuming a 4 cm gauge

length) and DSR-ductility strain at peak stress, both measured at 4°C.

0

200

400

600

800

1000

1200

1400

0 200 400 600 800 1000 1200 1400

DS

R S

tra

in t

o P

eak

Str

ess

(%)

Ductility Engineering Strain at Failure (%)

Line of Equality SBS Modified

Elvaloy Modified Neat


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(a) (b)

FIGURE 4 (a) True strain and engineering strain variation with ductility sample

elongation, (b) true strain rate and engineering strain rate variation as ductility sample

elongates.

0.0

1.0

2.0

3.0

4.0

5.0

6.0

7.0

0 100 200 300

Str

ain

(-)

Elongation (mm)

True Engineering

0.000

0.001

0.002

0.003

0.004

0.005

0 100 200 300

Str

ain

Ra

te (

1/s

)

Elongation (mm)

True Engineering


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(a) (b)

FIGURE 5 (a) Crosshead deflection curve with test time, and (b) true strain rate profile,

calculated for the asphalt ductility tests

0

50

100

150

200

250

300

0 500 1000 1500 2000 2500 3000 3500

Elo

ng

ati

on

(m

m)

Elapsed Test Time (sec)

Constant True Strain Rate (High)Constant True Strain Rate (Ave)Constant True Strain Rate (Low)Constant Crosshead Speed

0.000

0.001

0.002

0.003

0.004

0.005

0 50 100 150 200

Tru

e S

tra

in R

ate

(1

/s)

Elongation (mm)

Constant True Strain Rate (High)

Constant True Strain Rate (Ave)

Constant True Strain Rate (Low)

Constant Crosshead Speed


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(a)

(b)

FIGURE 6 FEM of asphalt ductility sample under elongation


22

FIGURE 7 Stress vs. Strain curve from finite element simulation using different crosshead

speed profiles.

Critical Problems with Using the Asphalt Ductility Test as a Performance Index for Modified Binders

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