Abrasive Wear by Hard Particles: Mechanisms & Resistance

Tribology International Vol. 31, No. 10, pp. 587–596, 1998
 1999 Elsevier Science Ltd. All rights reserved
Printed in Great Britain
0301–679X/98/$19.00 ⫹ 0.00
PII: S0301–679X(98)00079–6
Wear by hard particles
K.-H. Zum Gahr
Abrasive wear can be caused by hard particles sliding on a softer
solid surface and displaying or detaching material. Different types
of interactions are distinguished between the sliding particles and
the wearing surface of the solid. Frequently, resistance against
abrasive wear is only considered as a function of hardness of the
wearing material. However, a more general model shows that,
depending on the interaction, the capability of deformation or the
fracture toughness of the wearing material is very important in
addition to hardness. Abrasive wear resistance can substantially
be improved by second phases embedded in a hard or soft
matrix. The theoretical models are supported by a lot of
experimental results from studies on metallic or ceramic
materials.  1999 Elsevier Science Ltd. All rights reserved.
Keywords: abrasive wear, two-body abrasion, hardness
Introduction
Wear by hard particles occurs in many different situations such as on earth-moving equipment, slurry
pumps or pipelines, rock drilling, rock or ore crushers,
pneumatic transport of powders, dies in powder metallurgy, extruders or chutes. According to Fig 1, the
wear processes may be classified by different modes
depending on the kinematics and by mechanisms
depending on the physical and chemical interactions
between the elements of the tribosystem which result
in detaching of material from the solid surfaces. Compared with unlubricated sliding wear, the value of the
wear coefficient k, i.e. the dimensionless quotient of
the amount of volumetric wear WV times the hardness
of the wearing material H divided by the normal load
FN and the sliding distance s, estimated from practical
experience can be substantially greater in abrasive or
erosive wear1–5. Fig 1 can only represent a very rough
estimation of the wear coefficient because of wide
variation of the wear mechanisms occurring in an
actual tribosystem as a function of the operating conditions and properties of the triboelements involved,
which can result in changes of the k value by some
orders of magnitude.
In abrasive wear, material is displaced or detached
from a solid surface by hard particles, or hard particles
between or embedded in one or both of the two solid
surfaces in relative motion, or by the presence of hard
University of Karlsruhe, Institute of Materials Science II and
Karlsruhe Research Center, Institute of Materials Research I, P.O.
Box 3640, 76021 Karlsruhe, Germany
protuberances on a counterface sliding with the velocity
v relatively along the surface. Two-body abrasion is
caused by hard protuberances or embedded hard particles while in three-body abrasion the hard particles
can move freely (roll or slide) between the contacting
surfaces. According to Refs. 6–8, the rate of material
removal in three-body abrasion can be one order of
magnitude lower than that for two-body abrasion,
because the loose abrasive particles abrade the solid
surfaces between which they are situated only about
10% of the time, while they spend about 90% of the
time rolling. Hard particles striking a solid surface
either carried by a gas or a liquid stream can cause
erosive wear whereby the wear mechanism depends
strongly on the angle of incidence of the impacting
particles. The interaction between hard particles and a
solid surface can generally be accompanied by events
of adhesion, abrasion, deformation, heating, surface
fatigue and fracture.
Fig 2 shows schematically some general trends of wear
loss of materials depending on properties of the abrasive particles and the wearing materials as well as the
operating conditions9. With increasing hardness of the
abrasive particles, wear loss can increase by about one
to two orders of magnitude from a low to a high level
(Fig 2a). This transition depends on the ratio of the
hardness of the abrasive particles to the hardness of
the material being worn (Fig 2b). The increase from
the lower to the higher wear level occurs for singlephase material when the hardness of the abrasive particles is equal to the hardness of the material worn.
At equal bulk hardness of a multiphase material, the
Tribology International Volume 31 Number 10 1998
587
Wear by hard particles: K.-H. Zum Gahr
Fig. 1 Values of wear coefficient k as a function of wear mode and wear mechanism without lubricative media
Fig. 2 Schematic representation of wear loss by hard particles as a function of material properties and operating
parameters such as (a) hardness of abrasive particle; (b) ratio of hardness of abrasive particle and hardness of
the wearing material; (c) abrasive particle size, normal load and impact velocity, and (d) impact angle without
considering lubricative media
matrix of the material containing hard reinforcing
phases is softer than the matrix of the single-phase
material. Hence, the transition from low to high wear
level of the multiphase material starts and ends when
the hardness of the matrix and the hardness of the
reinforcing phase are respectively exceeded by the
588
hardness of the abrasive particles (Fig 2b). At the high
wear level, the wear loss of the multiphase material
can be greater than that of the single-phase material if
the reinforcing phase can be detached from the matrix
and acts abrasively on the matrix in addition. Important
operating parameters are the average size of the abras-
Tribology International Volume 31 Number 10 1998
Wear by hard particles: K.-H. Zum Gahr
ive particles, the applied normal load, or during erosive
wear the velocity of impinging hard particles. Wear
loss of ductile and brittle materials depends on these
parameters according to power laws exhibiting different
values of the exponent n (Fig 2c). For ductile materials
the velocity exponent can be expected to be between
2 and 3, while on brittle materials 3 to 4 are more
likely. A transition from low to high wear level can
occur on brittle materials such as alumina with increasing particle size and/or normal load or applied contact
pressure, respectively. In general, spherical particles
cause a lower amount of wear loss than angular particles at a given average particle size. However, the
effect of particle shape should be smaller in threebody abrasion because the loose particles can reorient
themselves during sliding and rolling contact compared
with two-body abrasion. During erosive wear, the
dependence of wear loss on the impact angle is influenced by the size of the impinging particles, the impact
velocity and the target material. A trend has been
observed that the wear loss increases, at a given impact
angle, with increasing size and velocity of the
impinging particles. Ductile metallic materials such as
plain carbon steels exhibit their maximum wear loss
at impact angles less of about 30° (Fig 2d). Brittle
materials such as ceramics or hardened tool steels show
increasing wear loss with increasing impact angles and
frequently maximum values for normal incidence. Wear
loss decreases continuously with increasing impact
angle for materials of very high elasticity such as
rubber. It has to be considered that Fig 2 can only
represent general trends because in a given tribosystem
many factors act in a complex manner simultaneously.
Modelling of two-body abrasive wear
Models for two-body abrasion have been developed to
a substantially greater depth than for three-body abrasion. Penetration of a sliding abrasive particle into a
metallic surface results in microploughing or microcutting depending on the attack angle10. Below a critical
attack angle, the metallic material is mainly elastic–
plastic deformed and flows around and beneath the
sliding particle but no material is removed from the
surface. Increasing the attack angle leads to a transition
from microploughing to microcutting, i.e. material
flows up the front face of the abrasive particle and it
is detached from the wearing surface in the form of
a chip.
This model of a critical attack angle can be replaced
by a criterion of a critical penetration depth of the
abrasive particle or a rigid indenter. Using a spherical
indenter, the degree of penetration as the ratio of
penetration depth divided by the radius of contact is
directly connected with the attack angle which
increases with penetration depth. According to Refs. 5
and 11, the degree of penetration Dp as a function of
the dimensionless shear strength f, i.e. the ratio between
the shear strength at the contact interface and the shear
yield stress of the wearing material, allows us to
distinguish between different abrasive mechanisms (Fig
3a). Wedge formation means that a wedge prow forms
at the front of the indenter and material is removed
by propagation of a crack. Fig 3b shows the increase
of the wear coefficient k with increasing degree of
penetration for metals under unlubricated sliding.
Hardness of a material can be defined as the resistance
to penetration by a hard indenter. Hence, hardness of
the wearing material affects the penetration depth of
abrasive particles but Fig 4 shows that it fails in
predicting abrasive wear resistance of different
materials. For example, cold working of metals can
increase hardness substantially but not abrasive wear
resistance and it may even reduce it. This general
relation between abrasive wear resistance and hardness
was measured in many studies12,13. The increment in
abrasion resistance with increasing hardness of
materials is substantially larger in pure metals than in
heat treated steels or ceramics. Hardness fails in predicting wear resistance because it cannot characterise
the interactions between abrasive particles and the
wearing materials sufficiently, which, however, determine the formation of wear debris.
According to an early model3, which predicts wear
volume WV only depending on the attack angle ␣, the
normal load FN, the hardness H of the material, the
geometry of the indenter (in this case a conical
indenter) and the sliding distance s, the wear volume
W V 2·tan ␣ F N
·
⫽
(1)
s
␲
H
can be calculated from the volume of the wear
groove produced.
However, it has to be considered that only a portion
of the volume of a wear groove formed on a metallic
surface is removed as wear debris and the remainder
of the groove volume is plastically displaced to the
sides of the groove14–16.
Hence, a more general model9,17 was developed which
describes abrasive wear by distinguishing four types
of interactions between abrasive particles and a wearing
material (Fig 5), namely microploughing, microcutting,
microfatigue and microcracking. In the ideal case,
microploughing due to a single pass of one abrasive
particle does not result in any detachment of material
from the wearing surface. A prow is formed ahead
of the abrading particle and material is continuously
displaced sideways to form ridges adjacent to the
groove produced. Volume loss can, however, occur
owing to the action of many abrasive particles or the
repeated action of a single particle. Material may be
ploughed aside repeatedly by passing particles and may
break off by low cycle fatigue, i.e. microfatigue. Pure
microcutting results in a volume loss by chips equal
to the volume of the wear grooves. Microcracking
occurs when highly concentrated stresses are imposed
by abrasive particles, particularly on the surface of
brittle materials. In this case, large wear debris is
detached from the wearing surface owing to crack
formation and propagation. Microploughing and
microcutting are the dominant processes on ductile
materials while microcracking becomes important on
brittle materials.
In a single scratch experiment, the ratio of volume of
material removed as wear debris to the volume of the
wear groove produced can be described by the fab
value (Fig 5) which is defined9,17 as
Tribology International Volume 31 Number 10 1998
589
Wear by hard particles: K.-H. Zum Gahr
Fig. 3 Modes of interaction and wear coefficient k owing to a hard spherical indenter sliding on metallic materials
as a function of (a) interfacial shear strength and (b) degree of penetration (from Hokkirigawa and Kato5,11)
where ␸s is the effective deformation on the wearing
surface and ␸lim is the capability of deformation of the
wearing material before microfracture occurs locally
during abrasion in a given tribosystem. ␤ is a factor
which describes the decay of deformation with increasing depth below the wearing surface and depends
mainly on the work-hardening behaviour of the wearing
material. Equation (3) is defined for ␸s ⱖ ␸lim only
while for ␸s ⬍ ␸lim microploughing or elastic/plastic
deformation occurs without any wear loss. Hence, the
fab value is not only a material property but it is also
a function of the operating conditions during abrasion
which affect both ␸s and ␸lim.
According to this model17 the linear wear intensity Wl/s
defined as the change in the length of a pin specimen
or the thickness of a plate per sliding distance is
given by
Fig. 4 Schematic drawing of abrasive wear resistance
(two-body abrasion, W−1
of different materials
v )
measured in the pin abrasion test as a function of
their bulk hardness
f ab ⫽
A V ⫺ (A 1 ⫹ A 2 )
AV
(2)
W l/s ⫽ ⌽ 1 ·f ab ·
p
H def
(4)
or inserting Equation (3):
冋 冉 冊册
W l/s ⫽ ⌽ 1 1 ⫺
␸ lim
␸s
2/␤
·
p
H def
(5)
where AV is the cross-sectional area of the wear groove
and (A1 ⫹ A2) represents the amount of material which
is pushed to the groove sides by plastic deformation.
In the case of microcracking, spalling occurs at the
edges of the wear grooves (Fig 5) which leads to
negative values of A1 and A2 and fab becomes greater
than 1. The fab value becomes equal to unity for ideal
microcutting and equal to zero for ideal microploughing. Values of fab ranging from 0.15 to 1.0 have
been measured experimentally on about 30 different
materials using metallographic taper sections9,18.
where Hdef is the hardness of the highly deformed
material, e.g. of wear debris, p is the applied surface
pressure and ⌽1 is a geometrical factor which depends
on the shape of the abrasive particles. From Equation
(5) it follows that abrasive wear loss is strongly influenced by the capability of deformation (␸lim) and the
work-hardening capacity (␤ ⬇ (H def /H) 1/3 as the first
approximation) and Hdef of the wearing material in
addition to the operating conditions (p, ⌽1 ␸s). This
theoretical model was strongly supported by a lot of
experimental results9,17.
A theoretical model9 for calculating the fab value owing
to a combination of microploughing and microcutting
results in
Fig 6 shows a good agreement between experimental
results measured on a lot of different metallic materials
and the prediction of Equation (4). Abrasive wear
resistance (Wl/s)−1 of these materials against flint abrasive grits (80 mesh) increased linearly with the ratio
(fab/Hdef)−1 in the pin abrasion test used. The deviation
f ab ⫽ 1 ⫺
590
冉 冊
␸ lim
␸s
2/␤
(3)
Tribology International Volume 31 Number 10 1998
Wear by hard particles: K.-H. Zum Gahr
Fig. 5 Schematic representation of different interactions between sliding abrasive particles and the surface of
materials, scanning electron micrograph of a wear groove on an austenitic steel and a draft of a taper section
through a wear groove
penetrating abrasive particles, and the type of internal
notches. Using models of fracture mechanics, a critical
applied load or surface pressure can be estimated,
above which flaking of material occurs9,17. For a
material containing cracks or thin graphite lamellae the
critical surface pressure pcrit is given by
p crit ⫽ ⌽ 2 ·
␭·K 2IIc
D 2ab ·H·␮2
(6)
and for the linear wear intensity owing to microcutting
(fab ⫽ 1) and microcracking it is obtained
W l/s ⫽ ⌽ 1 ·
with
Fig. 6 Abrasive wear resistance on pure metals, singlephase alloys, precipitation-hardened ferritic and austenitic alloys, austenitic and ferritic steels and plain
carbon steels against 80 mesh flint measured in a pin
abrasion test with p ⫽ 0.71 MPa contact pressure
versus the ratio of hardness of wear debris to fab
values measured in a scratch test by using a sliding
diamond loaded by 2 N
of the test data from the straight line at high ratios
was caused by a change in wear mechanism from
microploughing/microcutting to portions of microcracking on the steels of high hardness.
Smaller or greater portions of microcracking can occur
during abrasion of inherent brittle materials such as
highly hardened steels or ceramics or materials containing internal notches such as cracks, pores, graphite
lamellae, embrittled grain boundaries, inclusions or
large carbides, which promote crack formation and/or
crack propagation. The extent of cracking depends on
the fracture toughness (KIc, KIIc) of the stressed
material, loading conditions, the size and shape of the
p3/2 ·H 1/2 2
p
⫹ ⌽ 3 ·A f ·D ab
␮ ·⍀
H def
K 2Ic
(7)
再 冉 冊冎
⍀ ⫽ 1 ⫺ exp ⫺
p
p crit
1/2
⌽2 and ⌽3 are geometrical factors which depend on
the shape of the abrasive particles and the shape of
the cracking during abrasive wear. These factors can
be calculated for special cases9. ␭ is the mean free
path between and Af the area fraction or density of
the inherent defects, e.g. cracks or graphite lamellae.
␮ is the coefficient of friction at the leading face of
the abrasive particles, Dab is the effective size of the
abrasive particles and KIc, KIIc the fracture toughness
of the wearing material according to mode I (tension)
or II (shear) loading.
Other models19,20 using fracture mechanics for describing removal of material by lateral cracking result in
similar equations for linear wear intensity such as
W l/s ⫽ ⌽·
5/4
D 1/2
ab ·p
3/4
K c ·H 1/2
(8)
where ⌽ is a constant. As many studies have shown,
fracture toughness KIc decreases with increasing hard-
Tribology International Volume 31 Number 10 1998
591
Wear by hard particles: K.-H. Zum Gahr
Fig. 7 Schematic drawing of the relation between fracture toughness and abrasive wear resistance (Wl/s)−1 of
different metallic and ceramic materials (two-body abrasion, pin abrasion test)
Fig. 8 Interactions between sliding hard or soft abrasive particles and reinforcing phases
ness or yield stress of the materials tested, i.e. fracture
toughness and hardness of a given material can be
connected to each other. These models suggest that
wear loss of brittle materials increases more rapidly
than linearly with applied surface pressure or normal
load (FN ⫽ p·A, where A is the area of contact) in
accordance with Fig 2c.
Fig 7 summarises the general effect of fracture toughness and hardness of materials on their abrasive wear
resistance in a given tribosystem. Brittle materials such
as ceramics show under severe wear conditions an
increasing wear resistance with increasing fracture
toughness despite simultaneously decreasing hardness.
Surfaces of these materials worn by hard particles
exhibit large or small portions of microcracking that
can result in spalling of individual grains of the wearing material and in addition microcutting in accordance
with Equation (7). Materials with values of fracture
toughness greater than a critical value, which depends
on the operating conditions (load, size and acuity of
the abrasive grits, sliding speed etc.), are worn by a
592
combination of microploughing and microcutting.
Hence, the wear loss of these materials is determined
with others by the hardness (see Equation (5)) but is
independent of fracture toughness. In this regime, wear
resistance decreases with decreasing hardness due to
enhanced penetration depth and experimentally a large
variety of values of wear resistance can be measured
on different materials but equal hardness owing to
differences in the capability of deformation ␸lim.
The aim of the foregoing modelling of two-body abrasion is to exhibit the most important parameters influencing wear resistance of homogeneous materials but
not calculating exact values of wear losses. The theoretical models discussed above do not consider the
interaction between parallel scratches or grooves which
can result in enhanced material removal22,23, temperature24 or liquid media25. It is evident and supported by
many experimental results that hardness of a wearing
material alone is insufficient for describing resistance
to abrasive wear. Microstructural parameters such as
grain size26–28 or reinforcing phases can substantially
Tribology International Volume 31 Number 10 1998
Wear by hard particles: K.-H. Zum Gahr
affect the amount of wear. Abrasive wear tests on
different oxide ceramics showed an increase in wear
resistance with an increase in the reciprocal of the
square root of the grain size to a first approximation.
Effect of second phases on abrasive wear
Composites can offer the answer for achieving high
hardness and sufficient fracture toughness to avoid
brittle fracture. Second phases, e.g. hard ceramic particles or fibres, can be incorporated into a softer and
more ductile matrix. The abrasive wear resistance of
such composites depends on different microstructural
parameters such as the hardness, shape, size, volume
fraction and distribution of the embedded phases, the
properties of the matrix and the interfacial bonding
between the second phase and the matrix.
Fig 8 shows different interactions between abrasive
particles and a reinforcing phase. Hard and soft abrasive particles, i.e. harder or softer than the reinforcing
phase and also small and large sizes of the reinforcing
phase are distinguished. Hard abrasive particles can
easily dig out small phases and cut or crack larger
ones. Soft abrasive particles are able to dig out small
phases or produce large pits. The indentation depth of
soft abrasive particles is substantially reduced by hard
reinforcing phases if the mean free path between them
is smaller than the size of the abrasive particles. Large
phases deficiently bonded to the matrix can be pulled
out. However, large phases strongly bonded to the
matrix can blunt or fracture soft abrasive particles.
Fig 9 shows wear grooves caused by a sliding SiC tip
(attack angle and wedge angle of 90°) on the surfaces
of TiC and TaC steel composite layers which were
produced on the die steel 90MnCrV8 (0.9% C) by
laser cladding29. The reinforcing TiC (⬇ 2600 HV) and
TaC (⬇ 1800 HV) particles were embedded in the
martensitic steel matrix (810 HV30). The soft TaC
phase was grooved (microcutting primarily) by the
hard SiC tip (Fig 9a) and was smashed under high
Fig. 9 Scanning electron micrographs of polished surfaces of heat treated composite layers scratched by a sliding
silicon carbide tip: (a) 3 ␮m TaC, FN ⫽ 1 N; (b) 3 ␮m TaC, FN ⫽ 2 N; (c) 30 ␮m TaC, FN ⫽ 5 N; (d) 3 ␮m
TiC, FN ⫽ 5 N; (e) 30 ␮m TiC, FN ⫽ 1 N; (f) 30 ␮m TiC, FN ⫽ 5 N
Tribology International Volume 31 Number 10 1998
593
Wear by hard particles: K.-H. Zum Gahr
normal loads (Fig 9c). Resulting debris of the TaC
were impressed into the matrix under the sliding SiC
tip (Fig 9b). The harder TiC particles of an average
size of about 30 ␮m were not or only moderately
grooved by the SiC (Fig 9e,f). With increasing normal
loads, the large TiC particles were broken. TiC particles of a size smaller than that of the wear groove
were dug out of the matrix (Fig 9d).
The effects of size and volume fraction of a hard
reinforcing phase were studied on SiC dispersoides
of varying size (5–200 ␮m) which were galvanically
embedded in a nickel matrix. Fig 10 shows the wear
intensity (amount of linear wear divided by length of
the wear path) of the composites related to that of the
pure nickel matrix as a function of the volume fraction
of the SiC reinforcing dispersoides. Wear intensities
were measured using 80 mesh (grit size of about
200 ␮m) SiC abrasive paper. Small dispersoides were
easily dug out and resulted in enhanced wear loss with
increasing volume fraction owing to their own abrasive
action as loose particles. At a given volume fraction,
wear intensity decreased with increasing size of the
dispersoides. A minimum of wear intensity was
measured for 15 and 120 ␮m SiC dispersoides as a
function of volume fraction. Microcracking connected
with spalling of dispersoides occurred at high volume
fraction. Composites containing the large 200 ␮m SiC
dispersoides exhibited continuously decreasing wear
intensity with increasing volume fraction. In this case,
the ratio of size of wear grooves to size of the dispersoides had been substantially smaller than unity.
In agreement with the results of Fig 10, studies on SiC
or Al2O3 particle reinforced aluminium alloys showed
decreasing abrasive wear loss with increasing volume
fraction and size of the reinforcing particles30,31. In
studies on a polyester resin which was reinforced by
continuous steel fibres perpendicularly oriented to the
wearing surface, wear intensity decreased also continuously with increasing area fraction of the reinforcing
fibres up to about 45 vol%32. However, in other studies
on polymeric composites containing dispersed Al2O3
particles33 or short fibres34 of glass or carbon, wear
loss decreased only below a critical volume fraction
or increased even continuously with increasing amount
of reinforcing phase.
From the foregoing results and also studies on white
cast irons9,35 it follows that under mild operating conditions abrasive wear loss decreases with increasing
volume or area fraction of the reinforcing phase. This
means that the reinforcing phase is strongly embedded
in the matrix and spalling does not occur. However,
under severe conditions with microcracking and
spalling of the reinforcing phase, wear loss can increase
with increasing volume fraction of reinforcing particles
or short fibres. These conclusions are also in agreement
with orientation effects of fibre reinforced materials.
Fibres perpendicular oriented to the wearing surface of
polymeric composites lead to the lowest wear loss34.
Fibres parallel to the surface may be dug out more
easily than those perpendicular to it, if the indentation
depth of the abrasive particles is larger than about a
half of the fibre diameter.
For brittle materials such as ceramics, resistance to
severe abrasion is influenced with others by fracture
toughness of the wearing material according to Equations (6)–(8), Fig 7 and experimental results9,20,21,36.
Fracture toughness of brittle materials can generally be
enhanced by mechanisms of energy dissipation. Such
toughening mechanisms are stress-induced phase transformation, stress-induced microcracking, crack pinning
Fig. 10 Relative linear wear intensity (related to that of pure nickel) of SiC dispersion-hardened galvanic nickel
versus SiC volume fraction. Average size of SiC dispersoides 5, 15, 120 and 200 ␮m
594
Tribology International Volume 31 Number 10 1998
Wear by hard particles: K.-H. Zum Gahr
Fig. 11 Volumetric wear loss in the abrasive wheel test against 120 mesh SiC abrasive and hardness versus the
volume fraction of a soft eutectic phase of an Al2O3 ceramic surface alloyed by HfO2
with crack deflection, crack bridging or pullout of
reinforcing phases. However, all tribologically induced
interactions between abrasive particles and the ceramics
are concentrated on a relatively thin surface zone.
Hence, some of these toughening mechanisms can fail
under tribological contact because they act in larger
volumes only. Also an appropriate fracture toughness
value of the relevant surface zone of the wearing
material which can differ from the bulk value has to
be inserted in the theoretical models.
Laser surface alloying of ceramics can be very effective
in increasing fracture toughness and wear resistance of
brittle ceramics37,38. Fig 11 shows the volumetric wear
of a HfO2 alloyed alumina measured using an abrasive
wheel covered with 120 mesh SiC abrasive paper and
sliding back and forth across the ceramic surface. The
alloyed alumina surface contained a cellular eutectic
HfO2–Al2O3 phase with a hardness only about half of
that of the alumina crystallites. The volume fraction
of this soft phase was varied up to an amount of about
50%. Despite decreasing hardness of the composite,
the volumetric wear loss decreased substantially with
increasing amount of the soft phase. This behaviour
can be explained by enhanced fracture toughness compared with that of the monolithic alumina owing to
grain refinement, elimination of porosity and introduction of the soft phase between the hard and brittle
Al2O3 crystallites.
Rules of mixtures may offer a simple method for a
first estimation of wear resistance of multiphase structures or composites if the reinforcing constituents are
strongly fixed in the matrix and are not pulled or dug
out by abrasively acting particles. Fig 12 displays wear
resistance of two-phase structures as a function of the
volume fraction of the reinforcing constituent. Wear
resistance of structures of type B increases strongly
with volume fraction of the reinforcing constituent
Fig. 12 Schematic representation of abrasive wear
resistance of two-phase structures as a function of
volume fraction of a reinforcing phase
even at small volume fractions. In contrast, wear resistance of structures of type A is substantially increased
only by large volume fractions of the reinforcing phase.
In the case of type B, the reinforcing constituent
determines wear behaviour predominantly and the
phases are worn consecutively. The effect of the reinforcing phase depends on the identity, size and the
distribution of this phase and also on size and identity
of abrasive particles used or more generally on the
wear system. The rule of mixtures of type B is supported by experimental results on polyester resin
reinforced by continuously steel fibres and the rule of
mixtures of type A by results on ferritic–perlitic steels9.
Experimental values of wear resistance of multiphase
Tribology International Volume 31 Number 10 1998
595
Wear by hard particles: K.-H. Zum Gahr
structures are frequently measured between the boundaries given by both rules of mixtures.
References
1. Burwell, J. T. and Strang, C. D., Metallic wear. Proc R Soc,
1952, A212, 470–477.
2. Archard, J. F., Contact and rubbing of flat surfaces. J Appl
Phys, 1953, 24, 981–988.
3. Rabinowicz, E., Friction and Wear of Materials. John Wiley,
New York, 1965.
4. Hutchings, I. M., Tribology: Friction and Wear of Engineering
Materials. Edward Arnold, London, 1992.
5. Hokkirigawa, K., Advanced techniques for in-situ observations
of microscopic wear processes. In Proceedings of Surface Modification Technologies VIII, eds T. S. Sudarshan and M. Jeandin,
Institute of Materials, London, 1995, pp. 93–105.
6. Rabinowicz, E., Dunn, L. A. and Russell, P. G., A study
of abrasive wear under three-body conditions. Wear, 1961, 4,
345–355.
7. Fang, L., Kong, X. L., Su, J. Y. and Zhou, Q. D., Movement
patterns of abrasive particles in three-body abrasion. Wear, 1993,
162–164, 782–789.
8. Xie, Y. and Bushan, B., Effects of particle size, polishing pad
and contact pressure in free abrasive polishing. Wear, 1996,
200, 281–295.
9. Zum Gahr, K. H., Microstructure and Wear of Materials. Tribology Series 10, Elsevier, Amsterdam, 1987.
10. Mulhearn, T. O. and Samuels, L. E., The abrasion of metals: a
model of the process. Wear, 1962, 5, 478–498.
11. Hokkirigawa, K. and Kato, K., An experimental and theoretical
investigation of ploughing, cutting and wedge formation during
abrasive wear. Tribology International, 1988, 21, 51–57.
12. Richardson, R. C. D., The maximum hardness of strained surfaces and the abrasive wear of metals and alloys. Wear, 1967,
10, 353–382.
13. Moore, M. A., Abrasive wear. In Fundamentals of Friction and
Wear, ed. D. A. Rigney, ASM, Metals Park, OH, 1980, pp.
73–118.
14. Stroud, M. F. and Wilman, H., The proportion of the groove
volume removed as wear in abrasion of metals. Br J Appl Phys,
1962, 13, 173–178.
15. Zum Gahr, K. H., Formation of wear debris by the abrasion of
ductile metals. Wear, 1981/82, 74, 353–375.
20. Moore, M. A. and King, F. S., Abrasive wear of brittle solids.
Wear, 1980, 60, 123–140.
21. Evans, A. G. and Marshall, D. B., Wear mechanisms in ceramics.
In Fundamentals of Friction and Wear of Materials, ed. D. A.
Rigney, American Society of Metallurgy, Columbus, OH, 1981,
pp. 439–452.
22. Jacobson, S., Wallen, P. and Hogmark, S., Fundamental aspects
of abrasive wear studied by a new numerical simulation model.
In Wear of Materials, ed. K. C. Ludema, ASME, New York,
1987, pp. 595–606.
23. Xu, H. H. K., Jahanmir, S. and Wang, Y., Effect of grain size
on scratch interactions and material removal in alumina. J Am
Ceram Soc, 1995, 78(4), 881–891.
24. Berns, H., Fischer, A. and Kleff, J., Scratch tests on iron-,
nickel- and cobalt-based alloys at elevated temperatures. Wear,
1993, 162, 585–589.
25. Backmark, U., Wear testing in water slurries. Wear, 1993, 162,
1029–1032.
26. Rice, R. W., Micromechanics of microstructural aspects of ceramic wear. Ceramic engineering and science. Proceedings of the
9th Annual Conference on Composites and Advanced Ceramic
Materials, American Ceramics Society, Westerville, OH, 1985,
pp. 940–945.
27. Mukhopadhyay, A. K. and Mai, Y. Q., Grain size effect on
abrasive wear mechanisms in alumina ceramics. Wear, 1993,
162, 258–268.
28. Davidge, R. W. and Riley, F. L., Grain-size dependence of the
wear of alumina. Wear, 1995, 186, 45–49.
29. Axén, N. and Zum Gahr, K.-H., Verschlei␤ von TaC- und TiCLaserdispersions-schichten durch weiche und harte Abrasivstoffe.
Mat-wiss u Werkstofftech, 1992, 23, 360–367.
30. Zamzam, M. A., Abrasive wear of aluminium–matrix composites.
Metall, 1991, 45, 250–254.
31. Garcia-Cordovilla, G., Narciso, J. and Louis, E., Abrasive wear
resistance of aluminium alloy/ceramic particulate composites.
Wear, 1996, 192, 170–177.
32. Zum Gahr, K.-H., Einflu␤ des Makroaufbaus von Stahl/PolymerFaserverbund-werkstoffen auf den Abrasivverschlei␤. Z Werkstofftech, 1985, 16, 297–305.
33. Lu, X.-Ch., Wen, S. Z., Tong, J., Chen, Y.-T. and Ren, L.-Q.,
Wettability, soil adhesion, abrasion and friction wear of PTFE
(⫹ PPS) ⫹ Al2O3 composites. Wear, 1996, 193, 48–55.
34. Friedrich, K., Wear of reinforced polymers by different abrasive
counterparts. In Friction and Wear of Polymer Composites, ed.
K. Friedrich, Elsevier, Amsterdam, 1986, pp. 233–287.
16. Moore, M. A. and Swanson, P.A., The effect of particle shape
on abrasive wear: a comparison of theory and experiment. In
Wear of Materials, ed. K. C. Ludema, ASME, New York, 1983,
pp. 1–11.
35. Zum Gahr, K.-H. and Eldis, G. T., Abrasive wear of white cast
iron. Wear, 1980, 64, 175–194.
17. Zum Gahr, K.-H., Modelling of two-body abrasive wear. Wear,
1988, 124, 87–103.
18. Zum Gahr, K. H. and Mewes, D., Severity of material removal
in abrasive wear of ductile metals. In Wear of Materials, ed.
K. C. Ludema, ASME, New York, 1983, pp. 130–139.
37. Zum Gahr, K. H. and Schneider, J., Multiphase Al2O3 ceramic
with high resistance to unlubricated sliding wear. Proceedings
of the International Tribology Conference, Yokohama, 1995,
Vol. 1, Japanese Society of Tribologists, Tokyo, 1996, pp.
397–402.
19. Evans, A. G. and Wilshaw, T. R., Quasistatic particle damage
in brittle solids. I. Observations, analysis and implications. Acta
Metall, 1976, 24, 939–956.
38. Zum Gahr, K. H., Bogdanow, Ch. and Schneider, J., Friction
and wear reduction of Al2O3 ceramics by laser-induced surface
alloying. Wear, 1995, 181, 118–128.
596
36. Yamamoto, T., Olsson, M. and Hogmark, S., Three-body abrasive wear of ceramic materials. Wear, 1994, 174, 21–31.
Tribology International Volume 31 Number 10 1998