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Invited review paper
The effect of friction on scratch adhesion testing: application to a sol±gelcoating on polypropylene
M.H. Blees*, G.B. Winkelman1, A.R. Balkenende, J.M.J. den Toonder
Philips Research Laboratories Eindhoven, Prof. Holstlaan 4, 5656 AA Eindhoven, The Netherlands
Received 20 November 1998; received in revised form 17 August 1999; accepted 14 September 1999
Abstract
The scratch test has long been used to study the adhesion of coatings. In this test an indenter is drawn across the surface of a coating under
an increasing (continuous or stepwise) load. The load (normal to the surface) at which detachment of the coating occurs is termed the critical
load. Usually, the magnitude of the critical load is related to the adhesion between the substrate and the coating by some theoretical model. It
is well known that apart from the adhesion the critical load depends on several other parameters including the friction coef®cient. In this
paper a review of theoretical models applicable to scratch adhesion testing is given. Experimental data is used to compare the ability of these
theoretical models to describe the effect of friction between the indenter and the coating on the critical load. We applied the scratch test to a
model system consisting of a (hybrid) sol±gel coating deposited on polypropylene. The friction coef®cient between indenter and coating was
varied by a short plasma modi®cation of the surface of the coating, while all other relevant parameters (i.e. interfacial adhesion, layer
thickness, E-modulus of the coating, etc.) remained constant. The critical load (normal to the surface) showed a pronounced decrease of more
than an order of magnitude with increasing friction coef®cient. Several models are discussed and compared to the experimental data. In
addition, the effect of substrate pretreatment on coating adhesion was studied. The adhesion of the sol±gel coating induced by microwave
oxygen plasma modi®cation of polypropylene is considerably better than the adhesion obtained by wet-chemical modi®cation in chromo-
sulfuric acid at room temperature. The adhesion induced by immersion in chromosulfuric acid is shown to be independent of the immersion
time between 1 and 10 min. q 2000 Elsevier Science S.A. All rights reserved.
Keywords: Adhesion; Plasma processing and deposition; Surface and interface states; Tribology
1. Introduction
Coatings are widely used in optical, microelectronic,
packaging, biomedical and decorative applications. The
coating is designed to impart favorable mechanical (i.e.
low friction, abrasion resistance), chemical (i.e. barrier for
gasses), optical, magnetic, and electrical properties to
various substrates.
In general, the functional behavior of these coatings
depends on the bulk or surface properties of the coating
material. Evidently, the durability and functionality of coat-
ings is critically dependent on the adhesion between the
coating and the underlying substrate.
A wide range of methods is used to assess to adhesion of
the coating [1]. These methods include the peel method (i.e.
`Scotch tape test'), the direct pull-off method, measurement
of the abrasion resistance, and the scratch test.
Despite the lack of a fully satisfactory analytical model of
its mechanics [2], the scratch test is widely used to quanti-
tatively evaluate the adhesion of coatings to substrates. In
this method a stylus having a well-de®ned tip is moved into
the substrate±coating laminate, while at the same time the
tip is moved (over much larger distances) tangential to the
surface. The normal force at which `failure' occurs is called
the critical load. The critical load is used to qualitatively
discriminate between differences in adhesion. In principle,
apart from the adhesion between substrate and coating, the
critical load depends on a large number of parameters
including the tip radius, loading rate, mechanical properties
of the substrate and coating, the thickness of the coating,
and the friction between indenter and coating.
The ®rst models for scratch adhesion testing were
published almost forty years ago [3]. Since then, a number
of alternative models have been proposed. In this paper a
review of theoretical models applicable to scratch adhesion
testing is given. Experimental data is used to compare the
Thin Solid Films 359 (2000) 1±13
0040-6090/00/$ - see front matter q 2000 Elsevier Science S.A. All rights reserved.
PII: S0040-6090(99)00729-4
www.elsevier.com/locate/tsf
* Corresponding author. Tel.: 131-40-27-42809; fax: 131-40-27-
43352.
E-mail address: [email protected] (M.H. Blees)1 Present Address: Cooperative Research Centre for Alloy and Solidi®-
cation Technology (CAST), CSIRO Manufacturing Science and Technol-
ogy, Locked Bag No. 9, Preston, Victoria 3072 Australia.
ability of these theoretical models to describe the effect of
friction between the indenter and the coating on the critical
load.
Polypropylene is an attractive engineering material due to
its low cost, ease of processability, good mechanical proper-
ties, resistance to organic solvents, and favorable environ-
mental aspects [4]. The major drawbacks of polypropylene
are the poor scratch resistance and the lack of adhesion [5].
Coatings deposited on polymers are widely used in many
industrial applications. The coating can impart various
desirable properties to the polymer such as scratch resis-
tance, barrier properties, chemical resistance or esthetical
properties.
Sol±gel coatings are frequently used to improve the
scratch resistance of polymeric materials (e.g. polycarbo-
nate). On polymeric substrates, hybrid coatings are
employed which are produced by hydrolysis of alkoxysi-
lanes usually containing a non-hydrolysable organic func-
tionality producing organically modi®ed silicate ®lms [6].
Films can be deposited from water/alcohol mixtures by
simple industrial methods such as spraying. However, the
internal stress in such coatings is relatively high compared
with conventional polymeric coatings due to the high degree
of crosslinking. Therefore, the application of well-adhering
highly crosslinked coatings on polypropylene poses a chal-
lenging problem, which cannot be solved without modi®ca-
tion of the polymer surface [5,7±10].
In this paper the adhesion of a sol±gel coating on poly-
propylene is studied by the scratch test. The main focus is on
the effect of friction between indenter and coating on the
critical load. The effect of friction has been studied by
chemical modi®cation of the coating surface, without chan-
ging the mechanical properties of the coating, the substrate
or the coating±substrate interface. The effect of chemical
surface modi®cation of the polymer substrate on the adhe-
sion between coating and substrate is also studied.
2. Review of previous work
2.1. General remarks
The adhesion strength of metal/ceramic, metal/polymer,
polymer/polymer and ceramic/polymer interfaces has been
characterized by the microscratch technique [11±13].
The scratch test is generally limited to the assessment of
hard coatings on softer substrates [14]. In a scratch experi-
ment a well-de®ned tip is drawn across a coating deposited
on a substrate. During the scratch either the position (z) or
the force (Fz) in the direction normal to the surface is
increased in a controlled way. Alternatively, a number of
scratches can be performed at different ®xed normal forces
[15]. At a certain normal force the ®lm is debonded from the
substrate. The force at this point is called the critical load
(Lcrit) and is associated with the adhesion strength between
substrate and coating. Recently however, an attempt was
also made to establish a direct relation between the geome-
try of the spall and the interfacial toughness in the case
where extensive spalling occurs ahead of the indenter [16].
The critical load can be detected by the accompanying
acoustic emission [17,18], by a sudden change of the
tangential force [19,20], or by microscopic examination of
the scratch track. In the case of position controlled scratch
experiments, a sudden change in the normal force can indi-
cate a critical event [11].
Occasionally, a geometry is used which is entirely differ-
ent from usual scratch testing [21,22]. In that case, an inden-
ter is moved across the surface of a cross-section of a coated
specimen from the substrate towards and across the
substrate/coating interface. Before scratch testing, the
coated article must be polished normal to the interface,
exposing a portion of the interface between coating and
substrate. Comparison of this method with the more
common geometry of scratch testing (where the indenter
is moved parallel to the surface of the coating) is very dif®-
cult.
It is known that, apart from the adhesion properties, the
critical load depends on a large number of other parameters
including the tip radius, loading rate, mechanical properties
of the substrate and coating, the thickness of the coating,
internal stress in the coating, ¯aw size distribution at the
substrate±coating interface, and the friction between inden-
ter and coating. Therefore the critical load measured cannot
be related in a simple way to the thermodynamic work of
adhesion [1], but it can allow a semiquantitative comparison
of different degrees of adhesion. The thermodynamic work
of adhesion is directly related to the interfacial energy (g sc)
and the surface energy of the coating (g c) and the substrate
(g s) by the Dupre equation [23]. Sound comparisons can
only be made by scratch testing if, apart from the adhesion
between substrate and coating, all other factors mentioned
above are equal for the samples tested, or if one has a
universal scratch model which relates the critical load to
the work of adhesion by taking into account the in¯uence
of all relevant parameters.
In several scratch models a quantity is used, usually
denoted by W, which is identi®ed with the thermodynamic
work of adhesion. However, none of these models can be
considered a universal scratch model in the sense that all
relevant parameters are taken into account. Therefore W is
better referred to as `practical' work of adhesion [1]. The
practical work of adhesion is a model-related quantity that
depends, besides the thermodynamic work of adhesion, on
other factors that were not taken into account explicitly in
the particular model of interest.
2.2. In¯uence of friction
Hamilton and Goodman made a careful analysis of the
scratching of a homogeneous body [24]. They showed that
both the tensile stress at the trailing edge and the compres-
sive stress at the front edge increased if the friction between
M.H. Blees et al. / Thin Solid Films 359 (2000) 1±132
the indenter and the coating was increased. Therefore, it is
expected that an increase of the friction will lead to an
increase of these stresses, which will stimulate delamination
on coated substrates.
The following experimenters observed the in¯uence of
friction on the critical load.
Steinman et al. [18] used two different indenter materials,
i.e. diamond and cubic boron nitride (CBN), to vary the
friction between the indenter and the surface during scratch-
ing. The geometry of the two indenters was identical. On
two of their coating substrate systems Steinman et al. found
a signi®cant difference in friction between the two inden-
ters. On a TiN/steel system the friction coef®cient increased
from 0.04 to 0.08 going from the diamond to the CBN tip.
This was accompanied by a decrease in critical load from 26
to 21 N. In the case of a TiN/cemented carbide system an
increase of the friction coef®cient from 0.03 to 0.10 resulted
in a decrease of the critical load from 67 to 46 N.
Bhansali et al. [25] also mention that the critical load is
dependent on the friction. However, differences in critical
load were observed under conditions where both the friction
coef®cient and the properties of the coatings varied at the
same time.
Valli [26] observed a threefold increase of the critical
load of a 2.3 mm TiN coating on steel after depositing an
additional 0.5 mm silver ®lm, which was attributed to the
effect of friction. Surprisingly, the friction of a TiN coating
lubricated by oil exhibited the same friction coef®cient, but
no signi®cant increase of the critical load.
Bull et al. [27], observed that TiN coated substrates that
had been stored in a plastic bag for 2 years gave much lower
critical loads than those measured just after deposition of the
coating. Bull et al. postulated that this phenomenon was due
to a change in friction caused by adsorption of contaminants
from the plastic bags or atmosphere. After careful cleaning
of the coating the critical load obtained just after deposition
was restored. Bull et al. studied the effect of friction on the
critical load by depositing thin layers of different metals on
top of the TiN coating of interest. The friction coef®cient
obtained varied from 0.2 to 0.6. For the majority of the
specimens they observed a reduction of the critical load as
the friction between indenter and coating was increased.
However, for several metal ®lms (Pb, W and Cr), the critical
load did not obey this trend, since these samples exhibited
both a low critical load and a low friction coef®cient. Bull et
al. attributed this deviating behavior to the metal ®lms being
scraped from the TiN surface during scratching, and subse-
quently being squashed out from under the indenter. It is
dif®cult to judge in what way these results are dependent on
the properties (e.g. stresses) of the metal ®lms and adhesion
of the additional metal ®lms to the coating±substrate system
of interest.
Coghill et al. [19] also stress the important role of friction
during the scratch test. For three different indenter tips, each
made from a different material, and having widely differing
tip radii, the resulting critical load could not be correlated to
the tip radius. It was therefore concluded that the friction
must have a dramatic effect. The observed trend is that the
critical load increases as the friction coef®cient decreases.
Although the references we have discussed above have
given some consideration of the role of friction in the
scratch test, it is clear that a systematic study of the effect
is still lacking. To draw ®rm conclusions, either not enough
data are available [18], or the in¯uence of other varying
factors is unclear so that the separate in¯uence of friction
cannot be established [19,25±27].
2.3. Existing scratch models for coated substrates
Below, we will brie¯y discuss several scratch models
proposed in the literature.
Benjamin and Weaver [3] were the ®rst to propose a
model describing the scratch test. Actually, they proposed
two independent models, one based on consideration of the
tangential force acting on the tip during scratching, and
another based on the force normal to the surface. In the
®rst model, the tangential force Fx is assumed to consist
of three components, a plowing force required to deform
the substrate, a force to remove the coating from the
substrate, and a plowing force required to push aside the
sheared ®lm. Hence, the model consists of three terms
Fx � d3
12RHs 1
p
4td2 1 dtHc �1�
In this equation, d is the scratch width, R is the tip radius,
Hs and Hc are the hardness of the substrate and coating,
respectively, t is the shear stress at the coating±substrate
interface, and t is the thickness of the coating. Eq. (1) can be
used to compute the critical shear stress from the measure-
ment of Fx and d at the occurrence of a critical event, by
expressing t in terms of the other parameters. This was done
by Coghill and StJohn [19], who found that their scratch
measurements on aluminum-coated glass were described
well by the model. Benjamin and Weaver themselves,
however, found that the model was not in agreement with
measurements of metal coatings on glass [3].
The second model proposed by Benjamin and Weaver [3]
describes scratching in terms of a shear stress at the lip of
the indentation t s
ts � Hsa����������R2 2 a2p �2�
where a is the contact radius between indenter and coating,
which may readily be approximated by d/2. The model gives
a measure of adhesion in terms of a critical t s, by substitut-
ing a measured at the critical load. According to Benjamin
and Weaver, their experimental data could be satisfactorily
described by this simple relation. Ollivier and Matthews
[15] replaced Hs in Eq. (2) by Fz /pa2, resulting in a critical
shear stress given by
M.H. Blees et al. / Thin Solid Films 359 (2000) 1±13 3
tcrit � Lcrit
pacrit
������������R2 2 a2
crit
q �3�
where Lcrit stands for the critical load and acrit denotes the
contact radius at the critical load. They concluded that the
resulting model was able to give semi-quantitative compar-
isons for diamond-like carbon ®lms on polymer substrates.
Laugier [28] expressed the total compressive stress (s x)
under the leading edge of the indenter as the following sum
sx � Fz
2pa24 1 ns
ÿ � 3pm
82 1 2 2ns
ÿ �� ��4�
In this equation n s is Poisson's ratio of the substrate and m(m � Fx=Fz) is the friction coef®cient between the indenter
and the coating. The ®rst term originates from the compres-
sive stress at the leading edge of the indenter induced by the
friction during sliding. The second term of Eq. (4) describes
the radial surface stress on the edge of the contact circle
induced by the force normal to the surface. Laugier uses the
elastic Hertz formula to compute the contact radius
a3 � 3
4FzR
1 2 n2s
Es
11 2 n2
c
Ec
!�5�
in which n c is Poisson's ratio of the coating, Es is Young's
modulus of the substrate, and Ec is Young's modulus of the
coating. In practice, the contact radius a could be approxi-
mated with the help of the measured scratch width d, as was
mentioned above. For a ,, R, the shear stress t acting on
the coating±substrate interface at the lip of the indentation
was approximated by Laugier [28] in the following way
t < sxa=R �6�The value of this shear stress at the critical load is consid-
ered a measure of adhesion, i.e. after substitution of Lcrit, acrit
and mcrit (which denotes the friction coef®cient at the critical
load) in the foregoing equations.
In later publications [29,30], Laugier introduced an
energy-based description for coating removal. The practical
work of adhesion W was expressed as
W � s2crit
2Ec
t �7�
The stress s crit consists of an externally applied compo-
nent and an internal stress component
scrit � sx;crit 1 sint �8�in which s x,crit is given by Eq. (4) at the critical load. The
analysis of Laugier is purely elastic, and it is assumed that
a .. t. The model (Eq. (7)) was applied to TiN and TiC
coatings on WC±Co-based cemented carbides [30] to give
reasonable results.
Burnett and Rickerby [31,32] state that the driving forces
for coating loss in the scratch test may be viewed as the
summation of three components: (i) an elastic±plastic
indentation stress; (ii) an internal stress component; (iii) a
tangential frictional force. Burnett and Rickerby [32]
analyze the possible in¯uence of each of these components.
They discuss the effect of the hardness, thickness of the
coating, and the friction coef®cient. They conclude that
the relative contributions from the three driving components
depend on the precise scratching conditions. Burnett and
Rickerby do not give an explicit scratch model, but remain
at a qualitative level. Both Attar and Johannesson [33], and
Staia et al. [34] claim to have extracted the following equa-
tion from the papers of Burnett and Rickerby, labeling this
the `Burnett and Rickerby model'. The equation reads
Lcrit � pd2crit
8
2EcW
t
� �1=2
�9�
in which W denotes the work of adhesion, and dcrit represents
the width of the scratch track at the critical load. In this
simpli®ed model, the elastic±plastic indentation stress is
considered dominant, which, according to Burnett and Rick-
erby, is only the case for a relatively low friction coef®cient
and large coating thickness. According to our knowledge,
Eq. (9) was ®rst proposed by Bull and Rickerby [35].
Surprisingly neither Attar and Johannesson [33] nor Staia
et al. [34] refer to this paper. To avoid confusion with
another model proposed by Bull et al., we will refer to Eq.
(9) as the `Burnett and Rickerby model'.
In another paper, Bull et al. [27] modi®ed the theory
proposed by Burnett and Rickerby [31,32] by expressing
the stresses in terms of their contributions to the coef®cient
of friction. The friction coef®cient can be obtained experi-
mentally if both the normal and tangential forces are
measured during a scratch experiment. It is assumed that
coating detachment occurs when the tangential compressive
stresses in the coating in front of the indenter induce critical
tensile stresses normal to the coating±substrate interface due
to Poisson's effect. The critical load in their model is given
by
Lcrit � Acrit
ncmcrit
2EcW
t
� �1=2
�10�
where Acrit denotes the cross-sectional area of the track
which is given by
Acrit � R2sin21 dcrit
2R
� �2
dcrit
2R2 2
dcrit
2
� �2" #1=2
�11�
This model is designed for thin hard coatings on soft
substrates, since it is assumed that no shear tractions occur
at the interface. Bull et al. estimate that the penetration
depth must be at least twice the coating thickness t for the
model to be applicable. This would correspond to a mini-
mum scratch width of
dcrit � 4 t R 2 t� �f g1=2 �12�The exact value of the somewhat arbitrary numerical
prefactor in the criterion given in Eq. (12) seems to be
prompted by their experimental data of TiN on steel having
M.H. Blees et al. / Thin Solid Films 359 (2000) 1±134
an additional Ti metal layer which did not obey the trend
exhibited by other additional metal layers consisting of W,
Ag, Al, Cr.
Attar and Johannesson [33] modi®ed the model of Bull et
al. [27] by assuming that the tangential force responsible for
coating removal does not act on the total scratch track cross-
sectional area, but on the cross-section of the coating only.
Hence, they replace the Acrit of Bull et al. [27], as given in
Eq. (11), by Acrit � tdcrit, which results in the following
expression for the critical load
Lcrit � dcrit
ncmcrit
2tEcWÿ �1=2 �13�
In this model, the indenter force is concentrated on the
coating, corresponding with the situation that interface fail-
ure and ¯aking occur simultaneously.
Venkataraman et al. [36] proposed a model assuming that
coating removal is caused by the release of strain energy
produced by the contact of the indenter. The work of adhe-
sion is assumed to be equal to the strain energy released.
Contributions to the strain energy are modeled with elastic
stresses due to the combined action of a normal point force
and a tangential force. The strain energy released during
®lm debonding is equal to an integral over the debonded
area of expressions containing squares of the stresses. The
debonded area is an important parameter in the model of
Venkataraman et al., implying that the application of this
model is restricted to the case where a clear debonded area
appears during scratching. A substantial ¯aw of the model
seems to be that it is based totally on elasticity theory for
bulk materials. Nevertheless, the model has been success-
fully applied to Pt±NiO coating±substrate system. For coat-
ings on soft substrates, only qualitative results could be
obtained with the model [11,37].
If the simpli®cations used in the analytical models are
compared to the real complexity of the scratch test, it is
clear that none of these models give a complete description
of the mechanics during scratching. This is also apparent
from the considerable differences between the mathematical
formulation of these models. For example, an explicit
account of the effect of friction is given in the model of
Benjamin and Weaver (Eq. (1)), Laugier (Eq. (4)), Bull et
al. (Eq. (10)), and Attar (Eq. (13)). In the other models the
effect of the friction coef®cient on the critical load is
contained implicitly.
Despite the obvious limitations associated with the analy-
tical models, it is interesting to assess their capability to
properly explain the in¯uence of a change in parameters
(such as an increase of the coating thickness or a change
of the friction coef®cient) on the critical load observed in
experiments. For this purpose, experiments are required in
which the parameters are varied in a systematic way. Few
attempts to do this have been published (e.g. Coghill et al.
[19], Staia et al. [34]). Unfortunately, in many cases it is
impossible to vary just one parameter, so that one measures
the in¯uence of a mixed effect. For example, an increase of
the thickness of the coating is often accompanied by an
increase of the forces at the coating±substrate interface
due to internal stress in the coating. It is possible that
previous studies, in which the thickness of the coating was
varied, were disturbed by this effect.
3. Experimental
3.1. Substrate
`Polypropylene' (Stamylan P48M40) was obtained from
DSM (The Netherlands) and was injection molded into
2 mm thick sheets. The polymer falls into the category of
thermoplastic ole®ns and is a blend of polypropylene and
ethylene-propylene rubber. The surface of injection molded
thermoplastic ole®ns of this kind consists entirely of poly-
propylene [38].
3.2. Contact angle measurement
Contact angle measurements were performed by the
sessile drop method using a video camera. The video
image was fed to a computer and processed by a software
package (Image-Pro Plus 1.3). Advancing and receding
contact angles were determined by depositing a water
droplet on the surface through a syringe needle. Liquid
was added (advancing) or removed (receding) from the
droplet using a micropump at a rate of 6 ml/s. The tip of
the needle was not removed from the droplet between the
advancing and receding contact angle measurements. The
experimental error of measurement was about 38.
3.3. Surface modi®cation
The polymer substrates were rinsed with deionized water,
2-propanol, and n-heptane, and were blown dry in a stream
of nitrogen prior to modi®cation. One of the treatments
consisted of immersing the polymer substrates in chromo-
sulfuric acid (K2Cr2O7:H2O:H2SO4 7:12:150 by weight) at
ambient temperature for various times, followed by thor-
ough rinsing with deionized water.
In another treatment, a Tepla 300E microwave plasma
apparatus was used to modify the surfaces at 2.45 GHz.
The volume of the cylindrical quartz reactor chamber was
18 dm3. Oxygen gas was supplied at a rate of 60 cm3/min
(STP). The pressure used was 0.2±0.25 mbar, and the micro-
wave power was 400 W. After modi®cation up to 1 min, the
temperature of the reactor did not exceed 458C. Polypropy-
lene was treated in this way prior to coating for 20 s.
Microwave oxygen plasma was also used for the
enhancement of the friction between the tip of the indenter
and the surface of the sol±gel coating. The plasma treatment
increased the friction coef®cient from about 0.5 to almost 3.
In that case, the exposure time was only 1 s. At the same
time the wettability was also increased. The plasma modi-
®cation reduced the advancing and receding water contact
M.H. Blees et al. / Thin Solid Films 359 (2000) 1±13 5
angles (measured approximately 5 min after treatment)
from 78 and 638, to 43 and 188 respectively.
It was observed that the friction coef®cient between the
indenter and the surface of the coating decreased with
increasing time between the plasma treatment of the coating
and scratch testing. This phenomenon allowed us to vary the
friction coef®cient by simply changing the time interval
between plasma modi®cation and scratch measurement.
In order to reduce the coating±substrate adhesion in a
controlled way, some samples were also exposed to deep
UV radiation (160±190 nm, maximum at 172 nm) from an
incoherent Xe excimer light source (Hereaus Noblelight
Excimer Laboratory System) in a nitrogen atmosphere.
The estimated maximum intensity of the lamp unit is 20±
40 mW/cm2. A function generator connected to the power
supply of the lamp controlled the duration and intensity of
the exposure.
3.4. Coating
All silane precursors were obtained from ABCR
(Germany). 7.69 g ethyltriethoxysilane, 3.26 g 1,6-bis[tri-
methoxysilyl]hexane, 1.92 g tetraethoxysilane, and 5.0 g
ethanol (Merck) were added to a small reaction vessel at
ambient temperature (293 K). Under stirring using a
magnetic bar 3.95 g 0.1 M HCl (Merck) was added. Within
a few minutes the milky suspension became transparent, and
the reaction was allowed to continue for a total time of 1 h.
The reaction mixture was then transferred via a pre®lter
(Millipore, SJCN 013 NS) to a series of 2 ml cryogenic
vials (Nalgene), which were tightly sealed and stored in
liquid nitrogen until further use.
3.5. Coating deposition
A vial of coating liquid was removed from the liquid
nitrogen depository and was allowed to stand at ambient
temperature for 15 min. Subsequently, the liquid was spin-
coated on the polypropylene substrates at 400 rpm for 1 min
in an open spinner, and was cured for 1 h at 1008C. All
samples were allowed to age for at least 24 h before further
experiments.
To determine the thickness of the coating an organic
lacquer dissolved in toluene was used to cover part of the
coating. After etching for 30 min in 25% aqueous HF at
room temperature, the organic lacquer was removed in
toluene and the resulting pro®le was analyzed by a Tencor
Alpha-step 200 pro®ling instrument. A thickness of
4:4 ^ 0:3 mm was obtained.
3.6. Microscratch tester
Scratch tests were performed using a home built appara-
tus equipped with a needle topped with a sapphire sphere of
300 mm in diameter. After cleaning with acetone, the needle
was lowered to the surface at a rate of 0.5 mm/s until a force
of 2 mN was measured. After a waiting period of a few
seconds the indenter was moved tangential to the surface
at 10 mm/s over a length of 2 mm. At the same time, the
force normal to the surface was increased linearly at a rate of
1.2 mN/s, reaching 242 mN at the end of the scratch. During
the scratch the forces both tangential and normal to the
surface were recorded digitally. At the same time the posi-
tion in both directions was recorded. The resolution of the
positions of the indenter parallel to the scratch direction and
normal to the surface is equal to 0.1 mm and 0.01 mm,
respectively. The resolution of the forces parallel to the
scratch direction and normal to the surface is 1.0 mN and
0.2 mN, respectively. All scratches were performed at ambi-
ent temperature.
3.7. Microscopy
Scanning electron micrographs were obtained on a
Philips 505-type instrument operated at 30 kV after coating
the samples with a few nanometers of gold by sputter
deposition. Optical examinations of the scratch pro®les
were made using a Leica DMR microscope. Measurements
of the track width were performed on images produced by a
video printer coupled to the microscope.
4. Results
Without surface modi®cation of the substrate the sol±gel
coating delaminated spontaneously during curing at 1008C.
This is caused by the lack of adhesion between coating and
substrate in combination with the stress built up in the coat-
ing during drying. An increase of the adhesion was achieved
by microwave oxygen plasma modi®cation (20 s) of the
polymer substrate prior to coating. This resulted in a stable
adherent crack-free coating on polypropylene after curing.
The plasma treatment is known to introduce polar oxygen-
containing moieties at the surface of the polymer [39],
which induce a better adhesion of the coating. The work
of adhesion of the coating to the substrate corresponding
to this treatment will be denoted by Wplasma.
An optical micrograph of a scratch experiment performed
on such a coating on polypropylene is displayed in Fig. 1a.
The force increases linearly up to 242 mN over a distance of
2 mm. The track pro®le shows plastic deformation without
failure phenomena. The forces exerted by the indenter on
the system were not large enough to overcome the adhesion
between coating and substrate. Delamination was not
observed and no critical load could be identi®ed. In Fig.
1b the friction coef®cient m (m � Fx =Fz) is given as a func-
tion of the tangential displacement. The friction constant
was almost independent of the normal load over the entire
scratch track.
In order to observe a critical load the stresses on the coat-
ing must be increased. We decided to increase the friction
between the indenter and the coating by treating the (fully
cured) coated surfaces with a very short (1 s) microwave
oxygen plasma. The polar functional groups at the surface
M.H. Blees et al. / Thin Solid Films 359 (2000) 1±136
of the coating introduced by this treatment, result in an
increase of the friction by the enhanced interactions with
the sapphire indenter. The effect of the oxygen plasma on
the surface properties was evident from the changes in the
wettability. The coating (24 h after curing) exhibited advan-
cing and receding water contact angles of 788 (^38) and 638(^38) respectively. The plasma modi®cation reduced these
contact angles (measured approximately 5 min after treat-
ment) to 438 (^38) and 188 (^38) respectively. The mechan-
ical properties of the coating are not signi®cantly affected by
such a short treatment (typical ashing times for 1 mm thick
organic coatings are of the order of hours). Fig. 2a shows an
optical micrograph of the scratch. Fig. 2b shows both fric-
tion coef®cient and the position normal to the surface as a
function of the tangential displacement for the scratch
shown in Fig. 2a.
Comparison of Fig. 2b with Fig. 1b shows that the friction
coef®cient is much higher in the case of the plasma-treated
surface compared to the non-treated surface. The ®rst part of
Fig. 2a clearly shows the effect of friction enhancement
(compared with Fig. 1a) induced by the plasma treatment
on the scratch behavior. Already at relatively low forces
normal to the surface, considerable damage occurred. Scan-
ning electron microscopic (SEM) examination (Fig. 3a)
reveals that these instances of coating damage were due to
tensile cracking behind the indenter [31], and total removal
of the coating did not occur at this stage. It has been shown
that in the case of coatings on a polymeric substrate, tensile
cracks may already exhibit delamination [40]. In our case,
tensile cracking at the trailing edge of the indenter is accom-
panied by some delamination in the area close to the crack
(Fig. 3a). Unfortunately, since the delamination is only
partial, it is rather dif®cult to establish whether, and in parti-
cular to what extent, delamination has occurred. Further-
more, in contrast to the point at which total delamination
occurs (vide infra), it is very dif®cult to relate the generation
of these tensile cracks (and possible partial delamination) to
the friction data. Therefore, in this study, we have chosen to
focus our attention to the point of total delamination.
Total delamination was observed at a normal force of
194 mN after a scratch distance of 1.6 mm (Fig. 2a),
M.H. Blees et al. / Thin Solid Films 359 (2000) 1±13 7
Fig. 2. (a) Optical micrograph of a scratch under identical conditions as in
Fig. 1a, but with higher friction between the tip of the indenter and the
surface of the coating induced by a very short microwave oxygen plasma.
The length of the bar is 0.5 mm. (b) Friction coef®cient and displacement
normal to the surface as a function of the displacement of the indenter
tangential to the surface corresponding to the scratch shown in Fig. 2a.
Due to the effect of curvature of the sample over the length of the scratch,
the normal displacement shown here should not be regarded as displace-
ment relative to the undisturbed surface of the coating.
Fig. 1. (a) Optical micrograph of a scratch created by an indenter (300 mm
diameter spherical tip) on a coated polypropylene substrate. The normal
force increases linearly from left (2 mN) to right (242 mN) over a distance
of 2 mm. The length of the bar is 0.5 mm. (b) Friction coef®cient as a
function of the displacement of the indenter tangential to the surface corre-
sponding to the scratch shown in (a).
which was accompanied by a large drop of the friction
coef®cient by a factor of about 3 (Fig. 2b). After delamina-
tion occurred, the friction coef®cient remained rather low
upon further increasing of the normal force, and the remain-
der of the scratch showed only plastic deformation without
failure phenomena. At the critical load a sudden change in
the position normal to the surface was also observed (Fig.
2b). Both the drop of the friction coef®cient and the sudden
change of the position were observed for all scratches at the
point where delamination occurred. The point of delamina-
tion of several scratches was examined by SEM. In all cases,
it appeared that failure occurred at the substrate/coating
interface (Fig. 3b). Furthermore, the magnitude of the
sudden change of the position normal to the surface at the
point of delamination was always between 3.7 and 4.7 mm,
and showed no correlation with the critical load at delami-
nation. The magnitude corresponds closely with the thick-
ness of the coating as determined by pro®lometry
(4:4 ^ 0:3 mm). We therefore suppose that the coating,
which is detached from the substrate, sticks to the indenter.
In the remainder of the scratch the combination of the inden-
ter and detached coating slides over the forthcoming coating
without causing further delamination. This also explains the
relatively low friction coef®cient during the remainder of
the scratch.
These results clearly show the importance of friction in
scratching experiments. The differences between Fig. 1 and
Fig. 2 also show that adhesion cannot be identi®ed in a
unique way by an experimentally determined critical load
normal to the surface without taking into account the effect
of friction.
A large number of scratches were performed in the
manner described above. The critical load was obtained
from the characteristic drop in the friction coef®cient (all
scratches were also inspected by microscopy). It was
observed that the friction coef®cient between the indenter
and the coating decreased with increasing time between the
microwave oxygen plasma treatment of the surface of the
coating and scratch testing. The decrease of the friction
coef®cient was accompanied by a decrease of the wettability
(Table 1). The change of the friction coef®cient is caused by
organic contamination from the atmosphere adsorbing on
the surface of the coating, or changes of the surface due to
reorientation or migration in the coating. This phenomenon
allowed us to vary the friction coef®cient by simply chan-
ging the time interval between plasma modi®cation and
scratch measurement, with all other parameters (like adhe-
sion, thickness of the coating, etc.) remaining constant.
Fig. 4 shows a graph of the critical load against the fric-
tion coef®cient obtained in this way. The value of the fric-
tion coef®cient used in Fig. 4 was measured at the point just
prior to the load drop associated with the critical load.
In another set of experiments we purposely reduced the
adhesion between the coating and the substrate by exposing
the plasma treated polypropylene for a short period to deep
UV radiation (2 s at 25% of the maximum available power)
in an inert atmosphere prior to coating. The UV radiation
eliminates part of the hydrophilic moieties introduced by the
oxygen plasma treatment. This is indicated by the advancing
contact angle of water on plasma treated polypropylene,
which increases from 63 to 768 after deep UV exposure.
The (practical) work of adhesion of the coating to the
M.H. Blees et al. / Thin Solid Films 359 (2000) 1±138
Fig. 3. (a) SEM micrograph of cracks in the coating induced by tensile
stresses behind the tip of the indenter. The direction of the movement of the
indenter is from right to left. (b) SEM micrograph showing coating dela-
mination.
Table 1
Time dependence of the water contact angles of the coating after plasma
modi®cation
Time between plasma
and measurement
u advancing ( ^ 38) u receding ( ^ 38)
No plasma modi®cation 78 63
, 5 min 43 18
1 h 47 30
2 h 50 37
3 h 54 42
10 h 56 46
98 h 56 47
substrate corresponding to this treatment will be denoted by
WUV. The results of scratch experiments on these samples
are also given in Fig. 4.
For both sets of experiments a pronounced decrease of the
critical load with increasing friction coef®cient is observed.
The reduced adhesion of the WUV samples is also evident
from Fig. 4. At the same friction level, the critical load of
the samples that received an additional deep UV exposure
prior to coating (WUV) is much lower than that of the non-
UV exposed samples denoted by Wplasma.
We will now compare the results summarized in Fig. 4
with the theoretical predictions described above. All
models require knowledge of the track width (or a closely
related parameter) at the critical load. Experimentally, we
did not observe signi®cant differences in track width for
different samples discussed here. We de®ne the track
width as the width of the depression induced by plastic
deformation. The track width was most easily determined
by measurements from optical micrographs from scratches
in cases where no damage other than plastic deformation
was observed because of a relatively low friction coef®-
cient (see Fig. 1). In cases where cracking of the coating
was observed the cracks extended outside the track width
(Fig. 2a), but no signi®cant differences in the track width
itself were noted compared to scratches exhibiting plastic
deformation only. In Fig. 5 the track width is given as a
function of the force normal to the surface during the
scratch.
At the start of a scratch, the scratch width increases
rapidly with the applied force. At forces above 50 mN the
track width increases linearly with the applied force. These
experimental data were ®tted to a sixth degree polynomial
(drawn line in Fig. 5). This polynomial function was used to
predict the scratch width at the (critical) normal load
observed experimentally. In this way, the lowest critical
load observed (7 mN) corresponds to a scratch width of
21 mm, whereas the highest critical load (232 mN) corre-
sponds to a scratch width of 67 mm.
5. Discussion
5.1. Effect of friction on the critical load
It should be noted that the internal stress of our sol±gel
coatings was not known, but since all results described here
were obtained at a ®xed thickness of the coating, the contri-
bution due to internal stress is a constant which is contained
in the experimentally obtained work of adhesion. The
capability of the models listed above to describe the change
in the observed critical load with the friction will be
discussed. We stress that a successful description of this
relation by a model does not imply that the model is realistic
in a universal sense. Many other variations (e.g. changes of
Young's modulus of substrate and coating) should be tested
as well to support such a model.
Fig. 4 shows that, with all other parameters identical, the
critical load decreases with increasing friction coef®cient.
The track width at the point of critical delamination
increases with increasing normal load (Fig. 5), and accord-
ingly, with decreasing friction coef®cient. Therefore, in the
models that do not contain the friction coef®cient (or a force
tangential to the surface) explicitly, there is still an implicit
dependence of the critical load on the friction coef®cient.
Consequently, in the analysis these models cannot be disre-
garded a priori.
According to the Ollivier et al. [15] modi®cation of the
Benjamin et al. [3] model, the critical shear stress t crit as
given in Eq. (3) should be independent of the friction coef®-
cient. A plot of t crit against the friction coef®cient m (Fig. 6),
shows a large decrease with increasing friction coef®cient.
Therefore, the `Benjamin±Ollivier' model disagrees with
the present experimental data.
In the `Burnett and Rickerby model', it was assumed that
the frictional force was equal to the critical (normal) load.
Therefore, the equation for the critical load (Eq. (9)) does
not contain the friction coef®cient explicitly. Accordingly,
Lcrit /dcrit2 should be independent of m. The pronounced
decrease of Lcrit /dcrit2 with increasing m (Fig. 7), both for
M.H. Blees et al. / Thin Solid Films 359 (2000) 1±13 9
Fig. 5. Track width as a function of the force normal to the surface. Solid
line: polynomial least squares ®t.
Fig. 4. Critical load (normal to the surface) as a function of the friction
coef®cient just prior to the critical load for delamination. (A), Good adhe-
sion (Wplasma); (W), reduced adhesion (WUV).
Wplasma and WUV, shows that this model also does not
describe the friction dependence of the critical load
correctly.
The model of Bull et al. [27] explicitly takes into account
the effect of friction. According to their equation for the
critical load (Eq. (10)), Lcritmcrit /Acrit should be proportional
to the square root of the work of adhesion and independent
of the friction coef®cient. Fig. 8 shows that for our data this
is not the case. In contrast to the predictions by the model, a
pronounced increase of Lcritmcrit /Acrit with mcrit is observed
for both Wplasma and WUV.
It should be noted that according to Bull et al., Eq. (10)
would not strictly apply here, since for the experimental
data discussed dcrit is between 20 and 70 mm, whereas
dcrit $ 100 mm would be required to satisfy Eq. (12).
However, it is discouraging that even the critical loads
corresponding to a scratch width close to 70 mm (WUV
data around m � 0:6) fail to show a trend towards becoming
independent of the friction coef®cient.
In Eq. (1) proposed by Benjamin et al. [3], the plowing
forces required to deform the substrate material and to push
aside the sheared ®lm, are also taken into account. Therefore
two additional parameters, the indentation hardness of the
substrate (Hs) and the coating (Hc), are required. For rela-
tively thin coatings on substrates exhibiting considerable
elasticity, reliable values for the indentation hardness are
dif®cult to obtain. Coghill et al. [19] also compared the
contribution of the plowing forces to the experimentally
observed tangential forces at failure for several stylus
sizes on the same coating/substrate laminate. These data
indicate that the contribution due to plowing is important
if the radius of the stylus is small (e.g. 30 mm) or if the
friction coef®cient is relatively small (#0.1). Since the
size of the stylus we used had a radius of 150 mm and the
friction coef®cients at the critical load are relatively large
( $ 0.5), the in¯uence of plowing forces on the experimental
critical load can be neglected in a ®rst approximation. The
observation that the friction coef®cient can be increased
from about 0.5 to around 3 by microwave plasma treatment
of the coating surface shows that especially at high friction
coef®cients the plowing contribution must be very small.
Neglecting plowing forces, the critical load following
from Eq. (1) can be written as
Lcrit � p
4
tcritd2crit
mcrit
�14�
Accordingly, Lcritmcrit /dcrit2 should be proportional to the
square root of the work of adhesion (since W is proportional
to t crit2 ), and independent of the friction coef®cient. In
contrast to these predictions, Fig. 9 shows that, especially
for the Wplasma data, Lcritmcrit /dcrit2 increases rapidly with the
friction coef®cient. In order to make sure that this discre-
pancy is not due to the approximations made in Eq. (14), we
used least squares analysis to compare our data with Eq. (1),
taking Hc and Hs as adjustable parameters. Both for Wplasma
and WUV, this resulted in unphysical results for Hc and Hs
(negative values were obtained for both parameters).
Furthermore, the largest contribution due to plowing is
M.H. Blees et al. / Thin Solid Films 359 (2000) 1±1310
Fig. 7. Lcrit /dcrit2 (from Eq. (9)) as a function of the friction coef®cient just
prior to the load drop associated with the critical load. (A), Good adhesion
(Wplasma); (W) reduced adhesion (WUV).
Fig. 8. Lcritmcrit /Acrit (from Eq. (10)) as a function of the friction coef®cient
just prior to the load drop associated with the critical load. (A), Good
adhesion (Wplasma); (W), reduced adhesion (WUV).
Fig. 6. Critical shear stress (Eq. (3)) as a function of the friction coef®cient
just prior to the load drop associated with the critical load. (A), Good
adhesion (Wplasma); (W), reduced adhesion (WUV).
expected for the WUV results. However, the largest discre-
pancy from Eq. (14) is for the Wplasma data. Therefore, it can
be concluded that the Benjamin et al. model is unable to
explain the friction dependence of the critical load for our
experimental system.
In the ®rst model proposed by Laugier the adhesion is
related to the (critical) shear stress (t crit) at the coating±
substrate interface which is obtained from Eq. (6) with s x
given by Eq. (4). Instead of using Eq. (5), we approximate
the contact radius a by d/2. Furthermore, Poisson's ratio (n s)
of the substrate is estimated to be 0.3. The values of the
critical shear stress calculated from Eqs. (6) and (4) should
depend only on the adhesion, since the dependence on the
friction coef®cient is already contained in Eq. (4). In Fig. 10
the critical shear stress calculated by this method is plotted
against the friction coef®cient. If compared to the models
discussed before, the model ®ts the experimental data
reasonably well, especially for the Wplasma measurements.
In contrast to the Benjamin et al. model, the plowing
effect is not taken into account in Eq. (6). The largest rela-
tive contribution from plowing is expected when both m and
W are relatively low. The plowing forces will contribute to
the applied tangential force (Fx � Lcritmcrit) and will lead to
an increase of Lcritmcrit compared with the situation where
plowing forces are negligible. Therefore, the systematic
differences observed between m � 0:5 and 1 for the WUV
data are possibly caused by plowing effects.
In the second model proposed by Laugier the work of
adhesion (Eq. (7)) is related to the square of the total stress
given by Eq. (8). Since, for our measurements, both
Young's modulus (Ec), the thickness (t) and the internal
stress (s int) of the coating are ®xed, s x as calculated from
the right-hand side of Eq. (4) should be independent of the
friction coef®cient. Fig. 11 shows that Laugier's energy
approach is not in agreement with our experimental results.
According to the model (Eq. (13)) proposed by Attar et
al., Lcritmcrit /dcrit should be proportional to the square root of
the work of adhesion, and be independent of the friction
coef®cient mcrit. A plot of Lcritmcrit /dcrit against mcrit is repre-
sented in Fig. 12. The experimental data ®t the model
reasonably well, especially for the Wplasma measurements.
The Wplasma data give �2tEWplasma�1=2 � 3:8�^0:5� £ 103 N/
m, whereas the WUV data (for m . 1) result in
�2tEWUV�1=2 � 1:11�^0:28� £ 103 N/m. Since for both
systems t, Ec and n c are identical, it follows that according
to the Attar model Wplasma=WUV < 12. It should be stressed
that this ratio does not simply apply to the thermodynamic
work of adhesion (as de®ned by the Dupre equation), since
the practical work of adhesion contains additional dissipa-
tion terms.
It is easily veri®ed that the critical load (Fz � Lcrit) in
Laugier's ®rst model (Eq. (4)) and in the Attar model (Eq.
(13)) exhibit the same friction dependence if the second
term in Eq. (4) is negligible compared to the ®rst term
which applies when the friction coef®cient is relatively
large. For the smallest friction coef®cient encountered in
our experiments m,0.5, and substituting ns � 0:3, the
second term of Eq. (4) is at least six times smaller than
the ®rst term. An increase of the friction coef®cient will
M.H. Blees et al. / Thin Solid Films 359 (2000) 1±13 11
Fig. 10. Critical shear stress (Eq. (6)) as a function of the friction coef®cient
just prior to the load drop associated with the critical load. (A), good
adhesion (Wplasma); (W), reduced adhesion (WUV).
Fig. 11. Total compressive stress (according to Eq. (4)) as a function of the
friction coef®cient just prior to the load drop associated with the critical
load. (A), Good adhesion (Wplasma); (W), reduced adhesion (WUV).
Fig. 9. Lcritmcrit /dcrit2 (from Eq. (14)) as a function of the friction coef®cient
just prior to the load drop associated with the critical load. (A), Good
adhesion (Wplasma); (W), reduced adhesion (WUV).
lead to a further increase of this factor. This explains the
minor differences between Laugier's ®rst model (Fig. 10)
and the Attar model (Fig. 12). The same reasoning applies to
the simpli®ed Benjamin et al. model (Fig. 9, Eq. (14)) and
Laugier's energy model (Fig. 11, Eq. (7)). We conclude that
in our experiments the former two models give the best
description of the friction dependence of the critical load.
5.2. Activation by chromosulfuric acid
Apart from plasma activation, there are several other
well-known methods to increase the adhesion of coatings
to polypropylene such as corona treatment, short exposure
to a ¯ame, or immersion in an oxidizing liquid. Since treat-
ment by an oxidizing liquid can be performed in a very
reproducible way, we decided to compare this activation
method to the results obtained by plasma treatment.
Fig. 13 shows a graph of the critical load against the
friction coef®cient for coated polypropylene samples
which were immersed in chromosulfuric acid for various
times prior to coating with the sol±gel liquid. After curing,
the coating surface was again treated with a very short (1 s)
microwave oxygen plasma to increase the friction between
the indenter and the coating. The friction coef®cient was
varied as before by exposing the samples to the ambient
atmosphere for different times (vide supra). Again, the criti-
cal load showed a pronounced decrease with increasing fric-
tion coef®cient. Within the accuracy of the experimental
data given in Fig. 13, no effect of the immersion time in
chromosulfuric acid could be observed between 1 and
10 min. In Fig. 14, these data have been analyzed by the
Attar model (it is clear from the discussion above that an
analysis by Laugier's equation for the critical shear stress
will reveal almost identical results). Systematic deviations
from the Attar model were observed between m � 0:5 and 1,
which are comparable in magnitude to the deviations
observed for the WUV samples (Fig. 12). This is consistent
with the value for �2tEW�1=2 � 1:28�^0:25� £ 103 N/m
obtained for m . 1, which is equal to WUV, within experi-
mental error.
A considerable increase of the adhesion (determined as
lap shear strength) was observed by Brewis et al. [41] by
immersing polypropylene in chromosulfuric acid for 6 h at
708C, instead of 1 min at room temperature. The results
presented above suggest that within 1 min a stationary
state is reached where surface oxidation and dissolution of
material into solution are in dynamic equilibrium. There-
fore, the increase of the adhesion resulting from the elevated
temperature (708C) combined with the long immersion time
may be due to an increase of the surface roughness due to
the etching process. Alternatively, the etching of a thin
surface layer, which is phase separated from the elastomer,
could be responsible for the increase of the adhesion.
6. Conclusions
A detailed study of the effect of friction on scratch adhe-
sion testing was made. Scratch testing was performed on a
sol±gel (hybrid) coating deposited on polypropylene. To
prevent spontaneous delamination of the coating during
M.H. Blees et al. / Thin Solid Films 359 (2000) 1±1312
Fig. 12. Lcritmcrit/dcrit (from Eq. (13)) as a function of the friction coef®cient
just prior to the load drop associated with the critical load. (A), Good
adhesion (Wplasma); (W), reduced adhesion (WUV). Dashed lines represent
3:8 £ 103 N/m and 1:11 £ 103 N/m (see text for explanation).
Fig. 13. Critical load (normal to the surface) as function of the friction
coef®cient just prior to the load drop associated with the critical load.
Immersion time in chromosulfuric acid: (S) 1 min, (W) 2 min, (A) 10 min.
Fig. 14. Lcritmcrit/dcrit (from Eq. (13)) as a function of the friction coef®cient
just prior to the load drop associated with the critical load. Immersion time
in chromosulfuric acid: (S) 1 min, (W) 2 min, (A) 10 min. Dashed line
represents 1:28 £ 103 N/m (see text for explanation).
curing, it was necessary to chemically modify the polypro-
pylene surface by microwave oxygen plasma or immersion
in chromosulfuric acid. The friction coef®cient between
indenter and coating was varied by chemical modi®cation
of the surface of the coating, while all other relevant para-
meters (adhesion, thickness, modulus of the coating, etc.)
remained constant.
The critical load (normal to the surface) during scratching
showed a pronounced decrease of more than an order of
magnitude with increasing friction coef®cient. Therefore,
in experiments where the tangential (friction) force is not
measured directly, consistent results can only be obtained if
all samples exhibit a highly reproducible and uniform fric-
tion coef®cient with the indenter.
Several scratch-adhesion models proposed in literature
were compared to the experimental data. Both the model
proposed by Attar and Johannesson [33] and Laugier's
analysis of the interfacial shear stress [28] give a reasonable
description of the dependence of the critical load on the
friction. Systematic deviations from these models were
observed in the case where both adhesion and friction coef-
®cient were relatively small.
The adhesion of the sol±gel coating induced by micro-
wave oxygen plasma modi®cation of polypropylene was
considerably higher than that obtained by wet chemical
