An Experimental Investigation of the Scratch Behaviour of Polymers: 1. Influence of Ratedependent Bulk Mechanical Properties
Pinar Kurkcu, Luca Andena*, Andrea Pavan
Dipartimento di Chimica, Materiali e Ingegneria Chimica “Giulio Natta”, Politecnico di Milano
Abstract
The scratch behaviour displayed by a series of polymers in constant penetration depth scratch tests conducted by a
microscratch tester was examined in relation to their intrinsic rate-dependent mechanical properties. Scratch
hardness, which is considered as a measure of the scratch performance, was assessed by using a recently proposed
scratch model. The adequacy of the model to explain deformation recovery and pile-up characteristics of the material
in the scratch test was questioned by comparing the predictions of the model with residual profile measurements.
The study demonstrated that the scratch phenomenon is dominated by the yield stress of the material. Besides,
scratch hardness was found to be independent of the loading conditions. Indentation hardness and scratch hardness
are comparable if the testing conditions are similar.
Keywords
Polymers, scratch testing, hardness, rate-dependence
1. Introduction
Polymers are more and more frequently being utilised in applications demanding a high surface quality such as in the
automotive industry, data storage and optical products. “Surface quality” may refer to different characteristics
depending on the specific application. From a mechanical point of view, one of the essential features is the resistance
to scratch. The performance of polymer components is greatly influenced by the presence of scratches. Indeed, they
are not just aesthetically undesirable: they may impair the component’s functionality or even its structural integrity.
Therefore, understanding scratch behaviour of polymers [1-5] is critical to both maintain their good appearance and
enhance their durability.
Various surface morphology evaluation techniques and analytical models are utilised to qualify and quantify the
surface damage taking place during a scratch test. The most common methods are observing the change in optical
properties such as gloss, reflectivity, grey level [6-9], creating scratch deformation maps which demonstrate the
change in deformation modes with varying indenter geometry and testing conditions [5,10-12], analysing the
remaining scratch groove either by a profilometer or microscopy techniques [3,13-14] and determining scratch
hardness [2,15].
Scratch behaviour of polymers depends on a variety of factors, which can be classified into two primary groups: (i)
indenter geometry [13,16-17] and testing conditions such as normal load [16,18-20], temperature [21-22] and sliding
tip velocity [16,20,22] and (ii) material properties such as modulus [6,23-24], yield stress [23], ductility [25-26],
crystallinity [13,27], hardness [24,28], surface roughness [17], surface tension [29-30] etc. Several authors correlated
modulus/hardness ratio of the materials with the recovery and pile-up characteristics [3] and penetration depth
recorded during scratch test [24].
The aim of this study was to examine the relationship between the intrinsic mechanical properties of the polymeric
material and its scratch behaviour and investigate the effect of strain rate on the scratch phenomenon. For this
purpose, different classes of materials, namely amorphous, semicrystalline and crosslinked polymers were
considered. Their scratch behaviour was characterised in terms of scratch hardness and residual profile parameters.
2. Experimental
2.1 Materials
Commercial grades of poly(methyl methacrylate) (PMMA), polycarbonate (PC), polyethylene (PE) and melamine
formaldehyde (MF) were provided in the form of plates with a thickness of 3 mm. Isotactic polybutene-1 (PB)
supplied by Basell Polyolefins in the form of pellets and a styrene-acrylonitrile copolymer (SAN) supplied by ANIC
S.p.A in the form of powder were compression moulded into sheets with a thickness of 10 mm. Epoxy sheets were
prepared by casting and curing two different types of bisphenol-A based epoxy resins (Duroglass P5/1, called Epoxy-
1
1 and Ampreg 26, called Epoxy-2) with their amine-based curing agents. 2 mm thick plates of an AA 2034
aluminium aerospace alloy (Al/4.5Cu/1.6Mg/1Mn, simply called Al in the present work) were also tested for
comparison purposes.
2.2 Methods
Fundamental mechanical properties of the polymeric materials such as tensile modulus and yield stress - and their
time-dependence - were determined by means of dynamic mechanical analysis and uniaxial compression tests at
different frequencies and strain rates, respectively. Data on PMMA, PC and MF had already been reported in [40].
Young modulus and yield stress of Al were characterized by tensile tests and were found to be 72.6 GPa and 316
MPa, respectively. Scratch performance and scratch hardness were evaluated by scratch tests. In addition,
indentation hardness was measured in microindentation tests.
2.2.1 Dynamic Mechanical Analysis (DMA)
The tests were carried out using a TA Rheometric Series RSA III instrument in a three-point bending configuration.
Specimens measured 7 mm in width, 2 mm in thickness, with a span of 40 mm. The measurements were carried out
at constant temperature (23°C) in the frequency range 0.01-80 Hz. Maximum applied strain was 0.02%. Each test
was repeated on at least three specimens.
2.2.2 Compression Tests
Yield behaviour of the samples was characterised by uniaxial compression tests since compressive stresses are
dominant in the scratch process. The tests were performed on prismatic specimens with a square cross-section of 3
mm × 3 mm and 6 mm height. The tests were performed by an Instron 1185R5800 dynamometer fitted with
compression plates, at constant crosshead speeds of 0.1, 1 and 10 mm/min and 23°C. To reduce friction and avoid
barrelling of the tests specimens a polytetrafluoroethylene film with 50 µm thickness was put between the sample
faces and the metal plates. Deformation between the two plates was measured using a strain-gage extensometer.
Each test was repeated on at least three specimens. Yield data on PMMA and PC was taken from [38-39]. The
maximum of the stress-strain curves was used to determine a conventional value of the yield stress.
2.2.3 Scratch Tests
Specimens with dimensions of 15 mm × 30 mm and 3mm in thickness were prepared for scratch tests. Prior to
testing, the test surface of all samples was cleaned with commercial detergents and dried in vacuum oven at 50ºC for
12 hours. Tests were performed in load-controlled mode adjusting the load so as to obtain a constant depth of either
40 µm or 80 µm. The instrument used is a CSM Microscratch tester equipped with a conical indenter having an apex
angle of 120° and a spherical tip with a radius of 200 µm. Total scratch length was 4 mm in each test. Constant
sliding tip velocities varying between 0.25 and 250 mm/min were applied. All tests were carried out at 23°C. At the
end of each test, the profiles of transverse and longitudinal sections of the scratch tracks were obtained by using a
CONSCAN contactless profilometer.
2.2.4 Microindentation Tests
Measurements of indentation hardness were conducted by a Fischerscope H100VP HY Helmut Fischer
Microindentation instrument equipped with a Vickers indenter having an apex angle of 136° using specimens with
dimensions of 15 mm × 30 mm × 3 mm. Three loading-hold-unloading-hold cycles were performed at different
locations on the sample surface. The normal load was increased to 250 mN in 10 s, held at that level for 10 s,
removed in 10 s and held at zero for another 10 s. The contact depth and the contact area were estimated from the
unloading curve using the widely known Oliver and Pharr’s method [31], based on the original work by Doerner and
Nix [32].
3. Theory
Scratch hardness is one of the most commonly used parameters to evaluate the scratch performance of materials. It
represents an average or global response of the material to the contact load exerted by the indenter. Similar to
2
indentation hardness, scratch hardness is calculated as the normal load divided by the projected load bearing area.
The difficulty of evaluating scratch hardness arises from the difficulty in determining the true contact area. The
classical approach assumes that only the front half of the indenter is in contact with the material, that the shape of the
contact is circular for a conical or spherical indenter and the contact radius can be taken as half the residual scratch
width. These assumptions are applicable in case of materials which deform just plastically such as metals, but are not
accurate for viscoelastic-viscoplastic materials like polymers, in which the rear side of the indenter is also partly in
contact with the material due to a significant elastic + viscoelastic recovery of the deformation impressed by the
indenter.
Briscoe et al [2] therefore introduced an adjusting parameter q and defined scratch hardness as follows:
Hs = q
Fn
(1)
π ac2
where Fn is the normal load and ac is the contact radius, with q varying between 1 and 2 according to the material
response.
The relative inadequacy of Briscoe’s approach is due to the fact that q depends on the material and the testing
conditions but it cannot be easily determined experimentally nor can it be easily predicted analytically. Usually it is
considered to be a constant value for all materials belonging to the same class independently of the testing
conditions, but this has never been assessed adequately.
Estimating the true contact area by using residual profiles of the scratch tracks observed after the test does not
provide accurate results for polymers because of significant recovery phenomena taking place after the contact.
Gauthier et al [22] constructed a microscratch tester equipped with an optical microscope, capable of observing and
measuring the contact area during the test; the important drawback of this in situ observation technique is its
applicability limited to transparent materials. Anyhow, the commercially available scratch testers do not possess such
a feature. Hence, there have been various attempts to predict the true contact area in different ways, by considering
indenter geometry, pile-up formation and recovery taking place at the rear side of the indenter.
Pelletier et al [33] suggested a method to predict the true contact area by considering the recovery occurring at the
rear side of the indenter and the pile-up forming in front of the indenter. The model takes as a basis the parameter X
that was defined by Johnson [34] as the ratio of the strain imposed by the indenter to the maximum strain a material
can accommodate before yielding. X, which was later referred to as the “rheological factor” by several authors, was
found to depend on the modulus/yield stress ratio of the material and the indenter geometry. For a conical indenter
the following expression is proposed:
X=
E
σY
tan β
(2)
where E is the Young modulus, σy is the yield stress and β is the attack angle defined as indicated in Figure 1.
The ratio of contact depth, hc, to penetration depth, h, also called “shape ratio” and indicated c2 was defined to
characterize the pile-up formation during indentation tests. By recourse to numerical analysis, Bucaille et al [35]
computed the value of the shape ratio for scratch tests conducted with a conical indenter for model materials having
the same yield stress but different rheological factors. Plotting the dependence of the shape ratio on rheological
factor they observed two different behaviours below and above a rheological factor value of 80, which can be
expressed with the equations:
c2 =
hc
= 0.25339 ln X + 0.5017,
h
c2 =
hc
= 0.0684 ln X + 1.2984,
h
X < 80
(3)
X ≥ 80
(4)
3
Figure 1. Schematic representation of deformation during scratch test
As to the recovery taking place at the rear side of the indenter it was defined by an angular parameter α as indicated
in Figure 2 and α was correlated with the rheological factor, X. According to their analysis,
α=
1
b+dX
(5)
with b = 8.54×10-3 and d = 4.3×10-3.
Figure 2. Schematic representation of recovery angle and contact area during scratch test
Pelletier et al [33] estimated the true contact depth, hc, and the recovery angle, α, with the help of Equations (3) and
(5) and by using geometrical considerations they calculated the contact radius and the contact area. The reader is
referred to their study for a more detailed outline of their analysis. In the present study, Pelletier’s model was used
instrumentally to evaluate scratch hardness, Hs, as the ratio between the applied normal load, Fn, and the predicted
contact area, Ac:
Hs =
Fn
Ac
(6)
4. Results and Discussion
In Pelletier’s paper [33] the rheological factor was estimated from nanoindentation test results and modulus/yield
stress ratio was deduced via Equation 2; in the present paper modulus/yield stress ratio was determined
experimentally and the dependence of these parameters on strain rate was taken into due consideration. The strain
rate corresponding to each sliding tip velocity was estimated as:
4
ε& = ε t
(7)
where ε is the average strain applied in the scratch test, which, for a conical indenter having a spherical tip is
according to Tabor [36]:
ε =K
hc
ac
(8)
and t is the loading time, given by:
t = ac ν
(9)
where ν is the sliding tip velocity. K is taken to be 0.2 as proposed in [36].
As a first approximation estimate of the strain rate, the contact radius, ac, was assumed to be equal to half the
residual scratch width. Residual scratch width was measured by the help of the profilometer. Modulus and yield
stress values were taken at that strain rate and the true contact radius, contact area and strain rate were calculated
according to Pelletier’s procedure [33]. Then, modulus and yield stress values for the strain rate so calculated were
substituted and the true contact radius and area were recomputed. This iteration was continued up to obtaining
consistency between predicted and computed strain rate values.
Values of scratch hardness determined from the calculated contact area were used to assess the scratch performance
of the materials. Moreover, pile-up and recovery characteristics of the materials were analysed from the residual
profiles of the scratch tracks and these data were compared with the predictions of the scratch model.
4.1 Rheological Factor, X
Storage tensile modulus and uni-axial yield stress of the materials evaluated by dynamic mechanical tests and
compression tests, respectively, displayed the strain rate dependence shown in Figures 3 and 4.
Figure 3. Strain rate dependency of the modulus of the polymers characterised
The strain rate range covered in the compression tests was between 10-4 and 10-1 s-1, which is a few decades lower
than that of scratch experiments. In order to estimate the yield stress values at higher strain rates Eyring’s Equation
[37] was utilised:
5
σY
T
=
ε&
2 ⎛ ΔH
⎜
+ 2.303 R log
ε&0
V * ⎜⎝ T
⎞
⎟⎟
⎠
(10)
where T is the temperature, V* is an “activation volume”, ΔH is the activation enthalpy and ε&0 is a reference strain
rate.
According to Eyring’s Equation, at any constant temperature the yield stress increases linearly with the logarithm of
strain rate.
Figure 4. Strain rate dependency of the yield stress of the polymers characterised
According to Pelletier’s model, both the recovery at the rear side of the indenter and the pile-up height in front of it
can be estimated by just using the rheological factor X. In the present study, the penetration depth was relatively
high, enough to consider the indenter as purely conical neglecting the effect of the small spherical tip. The values for
the rheological factor of all materials examined in this study are shown in Figure 5.
Figure 5. Rheological factor of the materials characterised calculated for a strain rate of 10-2 s-1
The model suggests that the materials having rheological factor values close to each other should have similar
recovery and pile-up characteristics. Furthermore, according to the model, the higher the rheological factor, the more
plastic the response is. Therefore, from Figure 5, it is expected that the polymers examined, with the exception of PE
6
and PB have a similar scratch performance. Aluminium, which has a much higher rheological factor compared to
those of polymers, is expected to be deformed mainly plastically.
4.2 Recovery Angle and Shape Ratio
As to the recovery occurring at the rear side of the indenter, which according to the theory recalled above is defined
by the angle α (Figure 2), it should be noted that the model does not make a distinction between elastic and
viscoelastic recovery: an elasto-plastic contact is assumed. Viscoelastic behaviour will in general cause a timedelayed recovery for polymeric materials. Here we will use the term “recovery” to denote the recovery observed
right after the scratch test. Thus (90-α) represents the difference between the recovery that would have occurred if
the deformation had been purely elastic (90) and the recovery actually taking place (α). In other words, the greater
the (90-α), the more plastic the deformation is. Figure 6 shows the model’s prediction based on X for (90-α) value of
each material scratched at a constant penetration depth of 40 µm with a velocity of 5 mm/min. The difference
between the expected recovery characteristics of most polymeric materials and that of aluminium is obvious. In the
case of aluminium the recovery angle is very small, thus the expected behaviour is very close to that of a fully plastic
contact whereas with polymers a significant amount of recovery is expected to occur. Figure 7 shows the profiles of
the transverse sections of the scratch tracks of the materials under consideration, tested under the same conditions,
penetration depth of 40 µm and velocity of 5 mm/min. The profiles were recorded after the test, whereas the model
estimates the recovery during the test. However, these two data are comparable since a considerable portion of the
recovery occurs quite rapidly. The elastic component occurs of course immediately while the viscoelastic component
takes time. Under the experimental conditions adopted here we may retain that the major portion of the viscoelastic
component of the recovery in polymeric materials had already occurred before the profiles were recorded after the
test as assessed in separate experiments (not shown here). Thus, since the scratch test is performed at a constant
penetration depth, smaller recovery means larger residual depth and pile-up height after the test. On comparing the
profiles with the recovery prediction of the model, it can be observed that these two data are not in good agreement.
For instance, the model suggests that MF and SAN should have similar recovery characteristics owing to their
similar rheological factor values; however, their residual profiles show considerable difference. MF has the highest
recovery capacity among the materials investigated; its scratch behaviour is very close to that of a purely elastic
material. On the contrary, SAN has the smallest recovery capacity among the polymers investigated. According to
the predictions of the model, PE is expected to have the largest recovery among the polymers but the experiments
show a larger recovery in MF, PMMA and Epoxy-1. Evidently the rheological factor alone is not enough to represent
the variety of materials’ scratch behaviour.
Figure 6. The angle representing the area not in contact with the indenter
According to Pelletier’s model and Bucaille’s analysis, the front pile-up height during the test can be estimated by
using the relations between the shape ratio and the rheological factor (Equations 3 and 4) as mentioned earlier.
Figure 8 shows the comparison between the predicted values and the ones measured from the longitudinal profiles
recorded at the end of the test. According to the model, a larger pile-up formation is expected in materials having a
higher rheological factor. Both the predictions of the model and the profile measurements demonstrate that this is
7
indeed the case if we compare aluminium with the polymers as a whole. However, among the polymers, the model
does not rank the materials correctly. The model suggests that the materials having similar rheological factor values
such as PMMA, MF, PC, Epoxy and SAN should have similar pile-up characteristics; however, both the profiles
shown in Figure 7 and the experimental shape ratio values given in Figure 8 display significant differences among
these five polymers. It should be reminded once again that the profiles of the scratch tracks were obtained at the end
of the test, while the theory is intended to predict the pile-up height during the test. Yet, that cannot justify the large
discrepancies observed. The results suggest that the amount of accumulated material in front of the indenter does not
solely depend on modulus/yield stress ratio but also on the recovery capacity and flow characteristics of the material
(i.e. strain hardening).
Figure 7. Profiles of the transverse sections of the scratch tracks (penetration depth: 40 μm). The profile of the fully plastic
material was obtained by taking X = 1000
Figure 8. Comparison of experimental front pile-up height and the prediction of the model (penetration depth: 40 μm)
4.3 Contact Area
The contact area during the scratch test was estimated considering both the recovery taking place at the rear side and
the pile-up forming in front of the indenter, whose magnitude can be predicted by using X as mentioned in the
previous sections. Then, the contact area can be calculated as follows:
8
⎛π
⎞
Ac = ac2 ⎜ + α + sin α cos α ⎟
⎝2
⎠
(11)
Application of the model to the range of materials investigated (Figure 9) demonstrated that the contact area does not
change considerably among most of the polymers since they have similar modulus/yield stress ratio and hence
similar X. When the deformation is to a larger extent plastic, such as in case of PE and Al, the contact area is
relatively larger.
Figure 9. Area of contact during scratching of materials with a velocity of 5 mm/min (penetration depth: 40 μm)
4.4 Scratch Hardness
After calculating the contact radius and the projected contact area, scratch hardness was evaluated by Equation 6 for
each material at different sliding velocities (Figure 10). MF has the highest scratch hardness among polymers as
expected from its values of modulus and yield stress. On the other hand, PE and PB have very low resistance to
scratch owing to their modulus and yield stress values.
Figure 10. Strain rate dependency of scratch hardness
While scratch hardness of aluminium does not change significantly with strain rate, as expected in view of its purely
elasto-plastic nature, scratch performance of polymers is somewhat rate dependent because of their viscoelasticplastic nature. In an attempt to assess the correlation of the strain-rate sensitivity of scratch hardness of the polymers
9
with the strain rate dependency of their intrinsic mechanical properties, the strain-rate sensitivity of each property
was computed as follows:
dP
P0 d log ε&
(12)
where P is the property at strain rate ε& and P0 is the value of the property at the lowest strain rate.
As shown in Figure 11, the correlation is not so clear with any of the three properties, E, σy and E/σy.
Figure 11. Relation between the strain rate sensitivities of scratch hardness and intrinsic mechanical properties of the
polymeric materials
Our experimental scratch hardness results are compared with the predictions obtained by applying the results of the
numerical analysis of Bucaille et al [35] in Figure 12. In that analysis a conical indenter with an apex angle of 140.6º
was considered and a frictionless contact was assumed; the plastic deformation was assumed to follow Von Mises
yield criterion with a tensile yield stress value of 1 GPa; the rheological factor was varied by changing the materials’
elastic modulus from 2.793 GPa to 2793 GPa. The outputs of the numerical analysis suggest that at low rheological
factors the response is mainly elastic. As the rheological factor increases, the recovery taking place at the rear side of
the indenter becomes less and less significant. When the rheological factor exceeds 10, scratch hardness/yield stress
ratio remains almost constant. Figure 12 shows thatOur experimental results are not in good agreement with the
predictions obtained from that analysis. This can be due to several reasons. Firstly, the indenter geometry: differently
from Bucaille’s analysis, our experiments were performed with a conical indenter having an apex angle of 120 and a
spherical tip. Second, the yield stress: for a constant penetration depth, the results of the numerical analysis are
expected to depend on the yield stress value whereas Bucaille’s analysis, a single value of 1 GPa was used for all the
model materials, which is at least one order of magnitude higher than the values measured on most of the materials
investigated in our study. Third, the rheological factor: as already remarked on considering the residual profiles the
rheological factor may be inadequate to describe alone the scratch phenomenon. Modulus/yield stress ratio can be a
practical parameter for determining the onset of ductile ploughing during the scratch test, but the formation of the
groove and the accumulation of the material in front of the indenter are influenced by the flow characteristics of the
polymer beyond yield (strain hardening phenomenon). Fourth, friction: material flow along the indenter can also be
strongly influenced by friction, which is neglected in Bucaille’s analysis.
10
Figure 12. Comparison of experimental Hs/σY results with the output of numerical analysis of Bucaille [35]
To assess the possible change in scratch hardness with varying loading histories, we also performed some scratch
tests at different constant penetration depths (40 μm and 80 μm) and at a constant normal load (15 N). As shown in
Figure 13, scratch hardness of the materials was turned out to be substantially the same in all three cases.
Figure 13. Scratch hardness at different penetration depths
4.5 Correlation between scratch hardness and bulk mechanical properties
As mentioned before, the contact area during scratching of polymers under conditions of constant penetration depth
does not vary considerably since they have similar modulus/yield stress ratio. Hence, the determining quantity of
scratch hardness is the normal load applied to impose the given penetration depth. As demonstrated in Figure 14, the
scratch hardness of the polymers having similar rheological factor values and therefore similar recovery and pile-up
characteristics bears quite a perfect linear relation to the normal load applied. The materials which deform plastically
to a greater extent and therefore have a larger contact area, i.e. PE and Al, deviate somewhat from this correlation,
however (Fig. 9).
11
It can also be observed that the normal load required to impose the same penetration depth is directly proportional to
the yield stress (Figure 14, lower part). This suggests that scratch hardness can be predicted by evaluating just the
yield stress.
Figure 14. Correlation between scratch hardness, normal load (penetration depth: 40 μm) and yield stress (strain rate: 10-1 s)
4.6 Relation between scratch hardness and indentation hardness
Indentation hardness and scratch hardness of the investigated materials are compared in Figure 15. Both hardness
measurements are known to be dependent on the indenter geometry and the loading time. Here, the testing
parameters of the two tests were chosen so as to be consistent: scratch hardness values reported in Figure 15 are for a
sliding velocity of 5 mm/min, which corresponds to a loading time of 3 s according to Equation 9, which is
reasonably close to the holding time of 10 s used in the indentation tests.
If the testing conditions are consistent, scratch hardness and indentation hardness appear to be quite similar for
almost all materials. At the present stage, this is a purely phenomenological observation. There is one showy
exception, represented by MF; MF behaves differently from the other materials as already noticed in Section 4.2.
Indentation hardness measurements are strongly influenced by recovery characteristics: in the Oliver and Pharr’s
method [31] the contact depth and therefore the contact area are determined from the unloading curve. As mentioned
earlier, during unloading MF exhibits a large elastic recovery. Consequently, the estimated contact depth and the
contact area are very small compared to those of other polymers. On the contrary, in scratch hardness determination,
contact area during the test was estimated based on the rheological factor. Therefore, the effect of recovery
characteristics of the material is not as obvious as in case of indentation. Furthermore, the penetration depth during
the indentation of MF is so small that the tip of the indenter may have a considerable effect; in the case of other
polymers, conversely, the penetration depth is large enough to ensure that contact mainly occurs on the conical part
of the indenter.
12
Figure 15. Indentation hardness vs. scratch hardness
5. Conclusions
This study aimed at investigating the correlation between the scratch behaviour of polymers and their intrinsic
mechanical properties and observing the influence of strain rate on the scratch phenomenon. To evaluate scratch
performance of the materials, scratch hardness was computed by using the recently proposed Pelletier’s model [33].
Even tough the strain rate sensitivity of the scratch hardness and yield stress in polymers do not appear to be strictly
correlated, the results demonstrated that the scratch phenomenon is dominated by the yield stress of the materials.
Tests performed at different normal loads showed that scratch hardness is substantially independent of the loading
conditions. Moreover, indentation hardness and scratch hardness are similar if not identical as long as the indenter
geometry and the loading time are similar in the two tests.
The residual profiles of the scratch tracks were analysed to question the adequacy of the suggested scratch model to
predict the recovery and pile-up characteristics of the materials. The model is sounder in explaining the scratch
behaviour of polymers compared to the classical approach which assumes a mainly plastic deformation during the
scratch test. However, the comparison between the predictions of the model and our experimental results revealed
that the rheological factor alone cannot describe the complexity of the scratch phenomenon accurately. Other factors
such as the flow properties of the material, strain hardening and the viscoelastic characteristics may have
considerable influence on the scratch behaviour.
Acknowledgements
Most of the research work reported in the present paper is part of Pinar Kurkcu’s PhD Thesis in Materials
Engineering, Politecnico di Milano (2011). The authors wish to thank Luca Nobili and Fabio Pagano for the
indentation tests and Oscar Bressan for specimen preparation.
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Figure Captions
Figure 1. Schematic representation of deformation during scratch test
Figure 2. Schematic representation of recovery angle and contact area during scratch test
Figure 3. Strain rate dependency of the modulus of the polymers characterised
Figure 4. Strain rate dependency of the yield stress of the polymers characterised
Figure 5. Rheological factor of the materials characterised calculated for a strain rate of 10-2 s-1
Figure 6. The angle representing the area not in contact with the indenter
Figure 7. Profiles of the transverse sections of the scratch tracks (penetration depth: 40 μm). The profile of the fully plastic
material was obtained by taking X = 1000
Figure 8. Comparison of experimental front pile-up height and the prediction of the model (penetration depth: 40 μm)
Figure 9. Area of contact during scratching of materials with a velocity of 5 mm/min (penetration depth: 40 μm)
Figure 10. Strain rate dependency of scratch hardness
Figure 11. Relation between the strain rate sensitivities of scratch hardness and intrinsic mechanical properties of the polymeric
materials
Figure 12. Comparison of experimental Hs/σY results with the output of numerical analysis of Bucaille [30]
Figure 13. Scratch hardness at different penetration depths
Figure 14. Correlation between scratch hardness, normal load (penetration depth: 40 μm) and yield stress (strain rate: 10-1 s)
Figure 15. Indentation hardness vs. scratch hardness
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