Introduction to Tribology (Bhushan/Introduction) || Wear

7Wear

7.1 Introduction

Wear is the surface damage or removal of material from one or both of two solid surfaces in asliding, rolling, or impact motion relative to one another. In most cases, wear occurs throughsurface interactions at asperities. During relative motion, first, material on the contactingsurface may be displaced so that properties of the solid body, at least at or near the surface, arealtered, but little or no material is actually lost. Later, material may be removed from a surfaceand may result in the transfer to the mating surface or may break loose as a wear particle. Inthe case of transfer from one surface to another, the net volume or mass loss of the interfaceis zero, although one of the surfaces is worn (with a net volume or mass loss). Wear damageprecedes the actual loss of material, and it may also occur independently. The definition ofwear is generally based on the loss of material, but it should be emphasized that damage dueto material displacement on a given body (observed using microscopy), with no net change inweight or volume, also constitutes wear.Wear, as friction, is not a material property, it is a system response. Operating conditions

affect interface wear. Erroneously it is sometimes assumed that high-friction interfaces exhibithigh wear rates. This is not necessarily true. For example, interfaces with solid lubricants andpolymers exhibit relatively low friction and relatively high wear, whereas ceramics exhibitmoderate friction but extremely low wear.Wear can be either good or bad. Examples of productive wear are writing with a pencil,

machining, polishing, and shaving, which require controlled wear. Wear is undesirable inalmost all machine applications such as bearings, seals, gears and cams. Components mayneed replacement after a relatively small amount of material has been removed or if the surfaceis unduly roughened. In well-designed tribological systems, the removal of material is a veryslow process but it is very steady and continuous. The generation and circulation of weardebris, particularly in machine applications where the clearances are small relative to the wearparticle size, may be more of a problem than the actual amount of wear.In this chapter, we describe various mechanisms of wear and types of particles present in

wear debris, followed by representative data of wear of materials.

Introduction to Tribology, Second Edition. Bharat Bhushan.© 2013 John Wiley & Sons, Ltd. Published 2013 by John Wiley & Sons, Ltd.

316 Introduction to Tribology

7.2 Types of Wear Mechanism

Wear occurs by mechanical and/or chemical means and is generally accelerated by frictionalheating (or thermal means). Wear includes six principal, quite distinct phenomena that haveonly one thing in common: the removal of solid material from rubbing surfaces (Burwell,1957/1958; Kragelski, 1965; Engel, 1976; Eyre, 1976; Rigney and Glaeser, 1978; Scott,1979; Peterson and Winer, 1980; Suh and Saka, 1980; Buckley, 1981; Rigney, 1981; Bhushanet al., 1985a, 1985b; Loomis, 1985; Suh, 1986; Zum Gahr, 1987; Blau, 1992; Hutchings,1992; Bayer, 1994; Rabinowicz, 1995; Bhushan, 1996, 2001a, 2001b, 2011; Shipley andBecker, 2002). These are: (1) adhesive; (2) abrasive; (3) fatigue; (4) impact by erosion andpercussion; (5) chemical (or corrosive); and (6) electrical-arc-induced wear. Other commonlyencountered wear types are fretting and fretting corrosion. These are not distinct mechanisms,but rather combinations of the adhesive, corrosive, and abrasive forms of wear. Accordingto some estimates, two-thirds of all wear encountered in industrial situations occurs becauseof adhesive- and abrasive-wear mechanisms. Wear by all mechanisms, except by fatiguemechanism, occurs by the gradual removal of material.Of the aforementioned wear mechanisms, one or more may be operating in one particular

piece of machinery. In many cases, wear is initiated by one mechanism and it may proceed byother wear mechanisms, thereby complicating failure analysis. Failed components are gener-ally examined to determine the type of wear mechanism(s) responsible for eventual failure.Microscopy and a variety of surface analytical techniques are generally used in failure analyses.

7.2.1 Adhesive Wear

Adhesive wear occurs when two nominally flat solid bodies are in sliding contact, whetherlubricated or not. Adhesion (or bonding) occurs at the asperity contacts at the interface, andthese contacts are sheared by sliding, which may result in the detachment of a fragment fromone surface and attachment to the other surface. As the sliding continues, the transferredfragments may come off the surface on which they are transferred and be transferred back tothe original surface, or else form loose wear particles. Some are fractured by a fatigue processduring repeated loading and unloading action resulting in formation of loose particles.Several mechanisms have been proposed for the detachment of a fragment of a material. In

an early theory of sliding wear (still well recognized), it was suggested that shearing can occurat the original interface or in the weakest region in one of the two bodies (Archard, 1953),Figure 7.2.1. In most cases, interfacial adhesion strength is expected to be small as comparedto the breaking strength of surrounding local regions; thus, the break during shearing occursat the interface (path 1) in most of the contacts and no wear occurs in that sliding cycle. Ina small fraction of contacts, break may occur in one of the two bodies (path 2) and a small

Figure 7.2.1 Schematic showing two possibilities of break (1 and 2) during shearing of an interface.

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fragment (the shaded region in Figure 7.2.1) may become attached to the other surface. Thesetransfer fragments are irregular and blocky shaped. In another mechanism, plastic shearing ofsuccessive layers of an asperity contact result in detachment of a wear fragment. According tothis theory, plastic shearing of successive layers based on a slip line field occurs in conjunctionwith the propagation of a shear crack, alongwhich the fragment detaches, Figure 7.2.2 (Kayaba

Figure 7.2.2 Schematic showing detachment of fragment of a material from plastic shearing of suc-cessive layers of an asperity contact. Reproduced with permission from Kayaba, T. and Kato, K. (1981),“Adhesive Transfer of the Slip-Tongue and the Wedge,” ASLE Trans. 24, 164–174. Copyright 1981.Taylor and Francis.


and Kato, 1981). This process results in thin wedge-shaped transfer fragments. The fragment isdetached from one surface and transferred to the mating surface because of adhesion. Furthersliding causes more fragments to be formed by either of the two mechanisms. These remainadhering to a surface, transfer to themating surface, or to another previously attached fragment;in the latter case a larger agglomerate becomes detached as a large loose wear particle. Theseparticles may be of roughly equal size in each dimension.Although the adherence of fragments presupposes a strong bond between the fragments and

the surface onto which they are transferred, the formation of the final loose particle impliesa weak bond. The formation of a loose particle often results from chemical changes in thefragment. The fragments have a large surface area and tend to oxidize readily, which reducesthe adhesive strength, and they readily break loose. A second mechanism responsible for theformation of loose particles involves the residual elastic energy of adherent fragments. Whensandwiched between two surfaces, the fragment is heavily stressed. As the other surface moveson, only residual elastic stresses remain. If the elastic energy is larger than the adhesive energy,a fragment breaks loose as a wear particle.In material combinations with dissimilar materials, wear particles of both materials are

formed, although more wear particles of the softer material are formed, and are usually largerthan that of the harder counterpart. Because of defects and cracks within the harder material,there are local regions of low strength. If the local regions of low strength of the hardermaterial coincide with local regions of high strength of the softer material at a strong contact,the fragment of the harder material is formed. Formation of fragments of the harder materialmay also be produced by detachment of the material transferred by adhesion to the hardersurface by a fatigue process as a result of a number of loading and unloading cycles.The transfer of material from one surface to another has been studied by several inves-

tigators. In the early 1950s, an autoradiography technique was used in which one slidingmaterial was made radioactive and the transfer of the radioactive material to the mating sur-face during sliding was demonstrated by placing a photographic film in contact with themating surface after rubbing, and later developing the film to obtain an autoradiograph of anytransferred material (Rabinowicz and Tabor, 1951; Rabinowicz, 1953; Kerridge and Lancaster,1956; Bhushan et al., 1986). Black impressions on the developed film are produced by eachfragment.A scanning electron microscope (SEM) micrograph of a stainless steel shaft surface after

adhesive wear by sliding in a stainless steel journal bearing under unlubricated conditions isshown in Figure 7.2.3 (Bhushan et al., 1985b). Evidence of adhesive debris pullout can beclearly seen. During sliding, surface asperities undergo plastic deformation and/or fracture.The subsurface also undergoes plastic deformation and strain hardening. The SEMmicrographof the cross section of the shaft surface from adhesive wear shows visible plastic deformationwith a 25 μm thick layer, Figure 7.2.4. (A copper plate was applied before sectioning toprotect the worn surface.) Selected area electron diffraction studies in a transmission electronmicroscope of the cross sections of worn samples showed that extensive structural changeshad occurred. We believe that material close to the worn surface (∼15 μm thick) may haverecrystallized from an instantaneous surface temperature rise during sliding. Microhardnessmeasurements of the cross-section ofworn samples showed that a 10–80% increase of hardnessin the worn layer had occurred (Cook and Bhushan, 1973; Bhushan et al., 1985b).Severe types of adhesive wear are often called galling, scuffing, welding or smearing,

although these terms are sometimes used loosely to describe other types of wear.

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Figure 7.2.3 SEMmicrograph of 303 stainless steel shaft surface after adhesivewear under unlubricatedconditions. Sliding direction is along the vertical axis.

7.2.1.1 Quantitative Equations

Based on experimental data of various unlubricated material pairs, the vast majority beingmetallic, it is possible to write the rules of adhesive wear as follows. The amount of wear isgenerally proportional to the applied load W and sliding distance x and generally inversely

Figure 7.2.4 SEM micrograph of cross section of 303 stainless steel shaft after adhesive wear. Slidingdirection is along the vertical axis.


Figure 7.2.5 Schematic of a hypothetical model of generation of a hemispherical wear particle duringa sliding contact.

proportional to the hardness H of the surface being worn away. That is, the volume of wearbeing worn away (Holm, 1946) is

v = kW x

H(7.2.1)

where k is a nondimensional wear coefficient dependent on the materials in contact and theircleanliness.Archard (1953) presented the theoretical basis for the expression in Equation 7.2.1. Consider

two surfaces in a sliding contact under applied load W. Assume that during an asperityinteraction, the asperities deform plastically under the applied load and that at each unit eventthere is a definite probability that a wear particle will be produced. Further assume that contactis made up of asperities with an average radius of a, Figure 7.2.5. If the material has yieldedunder the maximum normal load dW, supported by an asperity,

dW = πa2H (7.2.2a)

where H is the mean contact pressure under the condition of full plasticity, flow pressure, orhardness of the softer material. We now assume that this asperity contact results in a wornparticle of volume dv. The dimension of this worn particle will be directly proportional to thecontact size. Physical examination of the wear particles shows that particles are generally ofroughly equal lengths in three dimensions rather than, say, layers. Thus, dv is expected to beproportional to a3. If a particle is assumed to be hemispherical in shape with radius equal tothe contact radius, then,

dv = 2

3π a3 (7.2.2b)

Finally, contact is assumed to remain in existence for a sliding distance dx equal to 2a afterwhich it is broken and the load is taken up by a new contact,

dx = 2a (7.2.2c)

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From Equations 7.2.2a, 7.2.2b and 7.2.2c,

dv

dx= 1

3

dW

H(7.2.3a)

If only a fraction α (= 3k) of all encounters produce wear particles, then the volume of wearby all asperities is

v α1

3

W x

H= kW x

H(plastic contacts) (7.2.3b)

Equation 7.2.3b is identical to Equation 7.2.1 and is commonly referred to as Archard’sequation of adhesive wear. Equation 7.2.3b is generally considered to give the amount of wearremoved from the softer of the two surfaces. This equation can also be used to calculate theamount of wear of the harder surface by using its hardness. Therefore, in the calculation ofk of either surface in Equation 7.2.1 or 7.2.3b, the hardness of the surface which is wearingaway should be used. The term k is usually interpreted as the probability that transfer of amaterial fragment occurs or a wear particle is formed to a given asperity encounter. The valueof k ranges typically from 10−8 to 10−4 for mild wear and from 10−4 to 10−2 for severe wearfor most material combinations, dependent on the operating conditions.Archard’s analysis suggests that there should be two simple rules of wear, i.e., that the wear

rate is independent of the apparent area and is directly proportional to the applied load. Theserules are analogous to Amontons’ equations of friction discussed in Chapter 5. Further, thewear rate is constant with sliding distance (or time) and independent of sliding velocity.Equation 7.2.1 suggests that the probability of decohesion of a certain volume of material

and/or formation of a wear particle (worn volume) increases with each asperity interaction,i.e., an increase in the real area of contact, Ar (Ar = W/H for plastic contacts) and the slidingdistance. For elastic contacts which occur in interfaces with one of the materials with a lowmodulus of elasticity or with very smooth surfaces (such as in magnetic recording interfaces,Bhushan, 1996), Equation 7.2.1 can be rewritten as (Bhushan’s equation of adhesive wear)(Bhushan, 1996)

v = k ′W x

E∗ (σp/Rp

)1/2 (elastic contact) (7.2.4a)

or

v = k ′W x

E∗ (σ/β∗)(7.2.4b)

where E∗ is the composite or effective modulus of elasticity, σ p and 1/Rp are the compositestandard deviation and composite mean curvature of the summits of the mating surfaces, re-spectively, and σ and β∗ are the composite standard deviation of surface heights and correlationlength, respectively (Chapter 3), and k’ is a nondimensional wear coefficient. In an elastic con-tact, though the normal stresses remain compressive throughout the entire contact (Chapter 5),strong adhesion of some contacts can lead to generation of wear particles. Repeated elasticcontacts can also fail by surface/subsurface fatigue. In addition, in all contacts, contact first


occurs on the nanoasperities which always deform by plastic deformation regardless of thedeformation on the microscale (Chapter 3), and the plastic contacts are specially detrimentalfrom the wear standpoint.For a designer who is interested in the rate of wear depth, Equations 7.2.1 and 7.2.4a can

be rewritten as

d = kpV

H(plastic contacts) (7.2.5a)

and

d = kpV

E∗ (σp/Rp

)1/2 (elastic contacts) (7.2.5b)

where d is the rate of wear depth (d/t) (mm/s) (where t is the sliding time or duration), p is theapparent normal pressure (= W/Aa, where Aa is the apparent area) and V is the sliding velocity.Note that the wear rate is proportional to the pV factor or the life of an interface is inverselyproportional to the pV factor. The pV factor is generally used in the selection of materials fordry bearings, see discussion later.As discussed in Chapter 5, flow pressure or yield pressure under combined normal and shear

stresses, pm, is lower than that under a static normal load pm (= H)

pm = H(1+ α μ2

)1/2 (7.2.6)

where α is a constant (about 9) and μ is the coefficient of friction. This expression for thehardness may be used in Archard’s wear equation.Rabinowicz (1995) has suggested that average diameter of a loose wear particle,

d = 60,000Wad

H(7.2.7)

where Wad is the work of adhesion, as described in Chapter 4. The size of the particles inmetallic contact typically ranges from submicrons to tens of microns.The adhesive wear mechanism may be the only mechanism in which there may be some

correspondence between the coefficient of friction and the wear rate for metals and nonmetalssince the same adhesion factors affect friction and wear.

Example Problem 7.2.1

The flat face of a brass annulus having an outside diameter of 20 mm and an inside diameter of10 mm is placed on a flat carbon steel plate under a normal load of 10 N and rotates about itsaxis at 100 rpm for 100 h. As a result of wear during the test, the mass losses of the brass andsteel are 20 mg and 1 mg, respectively. Calculate the wear coefficients and wear depths for thebronze and the steel. (Hardness of steel = 2.5 GPa, density of steel = 7.8 Mg/m3, hardness ofbrass = 0.8 GPa, and density of brass = 7.5 Mg/m3.)

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Solution

In a brass–steel contact at the test load, the load is expected to be supported by plastic contacts.Therefore, wear coefficients are given by Equation 7.2.3.

Given

W = 10 N

Mass loss of brass (mb) = 20 mg

Mass loss of steel (ms) = 1 mg

Rotational speed = 100 rpm

Test duration = 100 h

Now,

vb = 2× 10−2

8.5× 106 m3 = 2.35× 10−9 m3

vs = 10−3

7.8× 106 m3 = 1.28× 10−10 m3

Average contact diameter = 10+ 202

= 15 mm

and average sliding distance,

x = π × 15× 10−3 × 100× 100× 60 m= 2.82× 104 m

kb = vb Hb

Wx= 2.35× 10−9 × 0.8× 109

10× 2.82× 104 = 6.65× 10−5

ks = vs Hs

Wx= 1.29× 10−10 × 2.5× 109

10× 2.82× 104 = 1.14× 10−5

db = vb

Ab= 2.5× 10−9

π(102 − 52) × 10−6 m = 10.6 μm

ds = vs

As= 1.28× 10−10

π(102 − 52) × 10−6 m = 0.54 μm

7.2.1.2 Experimental Evidence

Adhesive wear equations, Equations 7.2.1 and 7.2.4, imply that if k is a constant for a givensliding system, then the volume of worn material should be inversely proportional to H or E∗

and interface roughness and proportional to the normal load (pressure) and the sliding distance.


Figure 7.2.6 Wear resistance of self-mated pure metals under unlubricated conditions as a function ofVickers hardness. Reproduced with permission from Zum Gahr, K.H. (1987), Microstructure and Wearof Materials, Elsevier, Amsterdam. Copyright 1987. Elsevier.

For a given material combination with primarily plastic contacts, the wear rate generallydecreases with an increase in hardness, Figure 7.2.6; more data will be presented in the nextsection on abrasive wear. However, in the case of extremely hard and/or brittle materials, thefracture toughness generally affects the wear rate, but is not included in the wear equations.For a material combination with primarily elastic contacts, the wear rate generally decreaseswith an increase in the modulus of elasticity, Figure 7.2.7. (Contacts in magnetic head-mediuminterfaces are primarily elastic, Bhushan, 1996.) For a material combination with primarilyelastic contacts, the wear rate, in adhesive wear mode, should decrease with an increase insurface roughness. However, if the wear occurs by other wear modes, such as abrasive wear,the wear rate may increase with an increase in the surface roughness.In many material combinations, wear rate increases linearly with the load (pressure) over

a limited range; wear rate may either increase or decrease abruptly at some critical loads(Archard and Hirst, 1956). It is the apparent pressure which determines the critical value ofthe load. This can be explained by the breaking or formation of oxide layers as a function ofpressure or frictional heating. For example, the wear rate of a brass pin sliding on a tool steelring increases linearly with load (according to Archard’s wear equation), however, the wearrate of the ferritic stainless steel pin increases linearly at low loads and increases rapidly abovea critical load, Figure 7.2.8. Tests conducted with a steel cone sliding on a steel plate showthat the wear coefficient increases rapidly above an apparent pressure equal to one-third of theindentation hardness, Figure 7.2.9 (Burwell, 1957/1958). It is generally observed that at thesehigher loads, large-scale welding and seizure occur.Experimental data suggest that the wear volume increases with the sliding distance or

sliding time at a constant velocity (Archard and Hirst, 1956). At the start of sliding, during

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Figure 7.2.7 Wear rate of ceramic thin films as a function of their Young’s modulus of elasticity slidingagainst magnetic rigid disk heads of A�2O3 − T iC , in the elastic-contact regime. Reproduced withpermission from Tsukamoto, Y., Yamaguchi, H., and Yanagisawa, M. (1988), “Mechanical PropertiesandWear Characteristics of Various Thin Films for RigidMagnetic Disks,” IEEE Trans. Magn. MAG-24,2644–2646. Copyright 1988 IEEE.

the so-called running-in period, the wear rate may be either higher or lower, followed bysteady-state wear rate until failure of the interface. Figure 7.2.10 shows the wear data obtainedfrom pin-on-ring tests for a wide range of material combinations under unlubricated conditionsin air. In each case, the steady-state wear rate (wear volume per unit distance) is essentiallyconstant for each material combination.Next, wear rate should be independent of the sliding velocity according to wear Equations

7.2.1 and 7.2.4. For many sliding combinations, this assumption holds for a range of values ofsliding velocity. However, sharp transitions in wear rate are seen at critical sliding velocitiesand apparent pressures, which are described using wear maps, to be discussed later.

7.2.1.3 Role of Metallurgical Compatibility

Rabinowicz (1980, 1995) has argued that the tendency of the sliding metals to adhere stronglyto each other is indicated by their metallurgical compatibility, which is the degree of solidsolubility when the two metals are melted together. The increasing degree of incompatibilityreduces wear, leading to lower value of the wear coefficients. This is also true for the coefficient


Figure 7.2.8 Wear rate as a function of load (logarithmic scales) for brass and ferritic stainless steelpins sliding against tool steel counterfaces in unlubricated pin-on-ring tests. Reproduced with permissionfrom Archard, J.F. and Hirst, W. (1956), “The Wear of Metals Under Unlubricated Conditions,” Proc. R.Soc. Lond. A 236, 397–410, by permission of the Royal Society.

of friction. Table 7.2.1 shows typical values of wear coefficients of metal on metal andnonmetal on metals with different degrees of lubrication at the sliding interface. Both degreesof metallurgical compatibility and lubricant significantly affect wear. The wear coefficientvaries by up to two orders of magnitude depending on the degree of compatibility and by upto three orders of magnitude depending on the extent of lubrication at the sliding interface. Itis clear that identical metal pairs must be avoided for low wear and friction.

7.2.1.4 Structural Effects

Hexagonal close packed (HCP) metals exhibit lower wear (an order of magnitude less) andfriction than cubic metals (Rabinowicz, 1995). A material pair involving two hexagonalmetals behaves the same way as a pair with only one hexagonal metal. As stated in Chapter 5,hexagonal metals have a limited number of slip planes, responsible for low wear and friction.

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Figure 7.2.9 Wear coefficient/hardness ratio as a function of the average pressure for SAE 1095 steelhaving hardness of 223 Brinell against a 120◦ conical slider. Reproduced with permission from Burwell,J.T. (1957/1958), “Survey of Possible Wear Mechanisms,” Wear 1, 119–141. Copyright 1957/1958.Elsevier.

Figure 7.2.10 Wear volume removed from the specimen pin sliding against a tool steel ring (unlessotherwise indicated) as a function of total sliding distance from unlubricated pin-on-ring tests on thematerials indicated. Reproduced with permission from Archard, J.F. and Hirst, W. (1956), “The Wearof Metals Under Unlubricated Conditions,” Proc. R. Soc. Lond. A 236, 397–410, by permission of theRoyal Society.


Table 7.2.1 Typical values of wear coefficients (k) for metal on metal (both with non-hexagonalstructure) and nonmetal on metal (both with non-hexagonal structure) combinations under differentdegrees of lubrication.

Metal on metalk (× 10−6)

Nonmetal on metalCondition Like Unlike∗ k (× 10−6)

Clean (Unlubricated) 1500 15–500 1.5Poorly lubricated 300 3–100 1.5Average lubrication 30 0.3–10 0.3Excellent lubrication 1 0.03–0.3 0.03

∗ The values are dependent upon themetallurgical compatibility with increasing degree of incompatibilitycorresponding to lower wear.

7.2.1.5 Grain Boundary Effects

As stated in Chapter 5, grain boundary regions are high-energy regions at the surface. Forpolycrystallinematerials, the presence of grain boundaries in thematerials influences adhesion,friction, surface fracture, andwear. Sliding friction experiments conducted byKehr et al. (1975)show that wear rate of Ni-Zn ferrite sliding against two magnetic tapes decreases with anincrease in the grain size, Figure 7.2.11. Bhushan (1996) also reported that single-crystal Mn-Zn ferrites generally have lowerwear rate (by about 10–25%) than the polycrystallinematerials.These observations suggest that polycrystallinematerials with high grain boundary densities

(finer grains) would exhibit higher wear rates than those with lower grain boundary densities(coarser grains) or single-crystalline materials.

7.2.2 Abrasive Wear (by Plastic Deformation and Fracture)

Abrasive wear occurs when asperities of a rough, hard surface or hard particles slide on asofter surface and damage the interface by plastic deformation or fracture. In the case of

Figure 7.2.11 Wear depth of an Ni-Zn ferrite rod in a sliding contact with γ − Fe2O3 and CrO2 tapesas a function of ferrite grain size. Reproduced with permission from Kehr, W.D., Meldrum, C.B., andThornley, R.F.M. (1975), “The Influence of Grain Size on the Wear of Nickel-Zinc Ferrite by FlexibleMedia,” Wear 31, 109–117. Copyright 1975. Elsevier.

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Figure 7.2.12 Schematics of (a) a rough, hard surface or a surface mounted with abrasive grits slidingon a softer surface, and (b) free abrasive grits caught between the surfaces with at least one of the surfacessofter than the abrasive grits.

ductile materials with high fracture toughness (e.g., metals and alloys), hard asperities or hardparticles result in the plastic flow of the softer material. Most metallic and ceramic surfacesduring sliding show clear evidence of plastic flow, even some for ceramic brittle materials.Contacting asperities of metals deform plastically even at the lightest loads. In the case ofbrittle materials with low fracture toughness, wear occurs by brittle fracture. In these cases,the worn zone consists of significant cracking.There are two general situations for abrasive wear, Figure 7.2.12. In the first case, the hard

surface is the harder of two rubbing surfaces (two-body abrasion), for example, in mechanicaloperations, such as grinding, cutting and machining; and in the second case, the hard surfaceis a third body, generally a small particle of abrasive, caught between the two other surfacesand sufficiently harder, that it is able to abrade either one or both of the mating surfaces(three-body abrasion), for example, in free-abrasive lapping and polishing. In many cases, thewear mechanism at the start is adhesive, which generates wear particles that get trapped at theinterface, resulting in a three-body abrasive wear (Bhushan et al., 1985b).In most abrasive wear situations, scratching (of mostly the softer surface) is observed as a

series of grooves parallel to the direction of sliding (plowing). A scanning electron micrographof a stainless steel surface after abrasive wear by sliding against a stainless steel journal surfacein the presence of alumina particles under unlubricated conditions is shown in Figure 7.2.13.Scratching in the sliding direction can be seen. An SEM examination of the cross-section of asample from abrasive wear showed some subsurface plastic deformation, but not as much as


Figure 7.2.13 SEM micrograph of 303 stainless steel shaft surface after abrasive wear under unlubri-cated conditions. Sliding direction is along the vertical axis.

in adhesive wear (Bhushan et al., 1985b). However, a 10–80% increase in microhardness ofthe worn surfaces was observed.Other terms for abrasive wear also loosely used are scratching, scoring or gouging, depend-

ing on the degree of severity.

7.2.2.1 Abrasive Wear by Plastic Deformation

Material removal from a surface via plastic deformation during abrasion can occur by severaldeformation modes which include plowing, wedge formation and cutting, Figure 7.2.14.Plowing results in a series of grooves as a result of the plastic flow of the softer material. In theplowing (also called ridge formation) process, material is displaced from a groove to the sideswithout the removal of material, Figure 7.2.14a. However, after the surface has been plowedseveral times, material removal can occur by a low-cycle fatigue mechanism. When plowingoccurs, ridges form along the sides of the plowed grooves regardless of whether or not wearparticles are formed. These ridges become flattened, and eventually fracture after repeatedloading and unloading cycles, Figure 7.2.15 (Suh, 1986). The plowing process also causessubsurface plastic deformation and may contribute to the nucleation of surface and subsurfacecracks (Bhushan, 1999). Further loading and unloading (low-cycle, high-stress fatigue) causethese cracks and pre-existing voids and cracks to propagate (in the case of subsurface cracks topropagate parallel to the surface at some depth) and join neighboring cracks which eventuallyshear to the surface leading to thin wear platelets, Figure 7.2.15. In very soft metals, such asindium and lead, the amount of wear debris produced is small and the deformed material isdisplaced along the sides of the groove. This plowing wear process should not be confusedwith rolling contact fatigue (to be described later) which develops macroscopic pits and resultsdue to the initiation of subsurface, high-cycle, low-stress fatigue cracks at the level at whichHertzian elastic stresses are a maximum.

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Figure 7.2.14 Schematics of abrasive wear processes as a result of plastic deformation by threedeformation modes.

In the wedge formation type of abrasive wear, an abrasive tip plows a groove and develops awedge on its front. It generally occurs when the ratio of shear strength of the interface relativeto the shear strength of the bulk is high (about 0.5–1). In this situation, only some of thematerial displaced from the groove is displaced to the sides and the remaining material showsup as a wedge, Figure 7.2.14b.In the cutting form of abrasive wear, an abrasive tip with large attack angle plows a groove

and removes the material in the form of discontinuous or ribbon-shaped debris particles similarto that produced in a metal cutting operation, Figure 7.2.14c. This process results in generallysignificant removal of material and the displaced material relative to the size of the groove isvery little.The controlling factors for the threemodes of deformation are the attack angle (Figure 7.2.17,

to be presented later) or degree of penetration, and the interfacial shear strength of the interface.In the case of a sharp abrasive tip, there is a critical angle for which there is a transition fromplowing and wedge formation to cutting. This critical angle depends on the material being


Figure 7.2.15 Schematics of plowed groove and formation of wear particle due to plowing as a resultof fracture of flattened ridge and propagation of surface and subsurface cracks.

abraded. The degree of penetration is critical in the transition from plowing and wedgeformation to cutting as the coefficient of friction increases with an increase in the degree ofpenetration (Hokkirigawa and Kato, 1988). For ductile metals, the mechanisms of plowing,wedge formation, and cutting have been observed, Figure 7.2.16.

Quantitative Equation (Plowing)To obtain a quantitative expression for abrasive wear for plastic contacts, we consider asimplified model, in which one surface consists of an array of hard conical asperities sliding ona softer and flat surface and plows a groove of uniform depth (Rabinowicz, 1995). Figure 7.2.17shows a single conical asperity, with a roughness angle (or attack angle) of θ (apex semi-angle

Figure 7.2.16 SEMmicrographs observed of wear process during wear of unlubricated brass by a steelpin. Reproduced with permission from Hokkirigawa, K. and Kato, K. (1988), “An Experimental andTheoretical Investigation of Ploughing, Cutting and Wedge Formation During Abrasive Wear,” Tribol.Inter. 21, 51–57. Copyright 1988. Elsevier.

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Figure 7.2.17 A hard conical asperity in sliding contact with a softer surface in an abrasive wear mode.

of asperities 90◦ – θ ), creating a track through the softer surface with a depth of d and widthof 2a. We assume that the material has yielded under the normal load dW; therefore

dW = 1

2πa2H (7.2.8)

where H is the hardness of the softer surface. The volume displaced in a distance x is,

dv = a2x(tan θ ) (7.2.9)

From Equations 7.2.8 and 7.2.9, we get

dv = 2dWx(tan θ )

π H(7.2.10a)

The total volume of material displaced by all asperities is,

v = 2W x tan θ

π H(7.2.10b)

where tan θ is a weighted average of the tan θ values of all the individual conical asperities,called the roughness factor.The derivation of Equation 7.2.10 is based on an extremely simple model. For example, the

distribution of asperity heights and shapes and any material build-up ahead of the asperitiesare not taken into account. An equation of the form similar to Archard’s equation for adhesivewear is found to cover a wide range of abrasive situations, and is

v = kabrW x

H(7.2.11)

where kabr is a nondimensional wear coefficient that includes the geometry of the asperities(tan θ for a simple case of conical asperities) and the probability that given asperities cut(remove) rather than plows. Thus, the roughness effect on the volume of wear is very distinct.The value of kabr typically ranges from 10−6 to 10−1. The rate of abrasive wear is frequentlyvery large – two to three orders of magnitude larger than the adhesive wear.


Note that in the elastic contact regime, the real area of contact and consequently thecoefficient of friction decreases with an increase in surface roughness, whereas the abrasivewear rate in the plastic contact regime increases with roughness.The wear equation for two-body abrasive wear is also valid for three-body abrasive wear.

However, kabr is lower, by about one order of magnitude, because many of the particles tend toroll rather than slide (Rabinowicz et al., 1961). It seems that the abrasive grains spend about90%of the time rolling, and the remaining time sliding and abrading the surfaces, Figure 7.2.12.(The coefficient of friction during three-body abrasion is generally less than that in two-bodyabrasion, by as much as a factor of two.) In some cases, such as in free abrasive polishing,a surface can elastically deform sufficiently to allow the particles to pass through, whichminimizes the damage. In this instance, Young’smodulus has a direct bearing on abrasive wear.During wear, some blunting of the hard asperities or abrasive particles occurs, thus re-

ducing the wear rate. However, a brittle abrasive particle can fracture which would result inresharpening of the edges of the particle and an increase in wear rate.


A hard steel surface consisting of an array of conical asperities of an average semi-angle of60◦ slides on a soft lead surface (H = 75 MPa) under a load of 10 N. Calculate the volume oflead displaced in unit slid distance. Given that the volume of lead material removed is 10−6 m3

for a sliding distance of 1 km, calculate the wear coefficient of lead.

Solution

Given,

Roughness angle, θ = 30◦

W = 10 N

νlead = 10−6 m3

Hlead = 75 MPa

The volume of material displaced by all asperities in unit slid distance is

2W tan θ

π H= 2× 10× tan 30

π × 75× 106 m3/m

= 4.9× 10−8 m3/m

The wear coefficient of lead material,

kabr = vlead Hlead

Wx= 10−6 × 75× 106

10× 103= 7.5× 10−3

Wear 335

Experimental EvidenceThere is significant experimental evidence that the wear rate in two-body abrasion is generallyinversely proportional to the hardness and proportional to the normal load and sliding distancefor many pure metals; alloys often exhibit more complex behavior (Kruschov, 1957, 1974;Goddard andWilman, 1962; Mulhearn and Samuels, 1962; Misra and Finnie, 1981). Hardnessis an important parameter for abrasive wear resistance.Wear resistance (proportional to 1/wearrate) of annealed pure metals is generally directly proportional to their hardness but is morecomplex for alloys, Figure 7.2.18 (Kruschov, 1957, 1974; Kruschov and Babichev, 1958).These authors reported that prior work hardening of the pure metals and alloys had no effecton the wear rate. Cold working of the 0.4% carbon steel resulted in a significant increase inbulk hardness but had no effect on its wear resistance. These and other experiments showthat a metal surface strain hardens by plastic flow during abrasion to a maximum value, andit is this value of hardness which is important for abrasion resistance. Also note that if amaterial is hardened, it generally becomes more brittle. Brittle materials can produce largerparticles, resulting in high wear rates. During three-body abrasion with alumina particles, thewear resistance of metals is also found to be proportional to the hardness of the workpiece(Rabinowicz et al., 1961).

Figure 7.2.18 Relative wear resistance of pure metals and heat treated and cold worked steels as afunction of hardness in two-body abrasion. Reproduced with permission from Kruschov, M.M. (1957),“Resistance of Metals to Wear by Abrasion, as Related to Hardness,” Proc. Conf. Lubrication and Wear,pp. 655–659, Instn Mech. Engrs, London, UK. Copyright 1957. Institution of Mechanical Engineers.


Figure 7.2.19 Mass loss of three ductile metals as a function of applied normal load subjected totwo-body abrasion by 115 μm SiC abrasive paper. Reproduced with permission from Misra, A. andFinnie, I. (1981), “Some Observations on Two-Body Abrasive Wear,”Wear 68, 41–56. Copyright 1981.Elsevier.

The volume of wear generally increases linearly with an increase in applied normal load;often the linear relationship is not maintained at high loads. For example, in Figure 7.2.19,mass loss for three ductile metals increases with applied load when abraded against an abrasivepaper. Bhushan (1985, 1996) has also reported that the wear rate increases with applied loadfor abrasive magnetic tapes sliding against ceramic heads.The wear rate changes as a function of the sliding velocity and the particle size of an abrasive

paper or roughness of the abrading surface. The effect of the particle size on the wear rate intwo-body and three-body abrasion is discussed by Xie and Bhushan (1996a). Figure 7.2.20shows the dependence of the sliding velocity and the abrasive grit size for copper being abradedagainst an abrasive wear. Note that the wear rate increases by a few percent for an increasein the sliding velocity by three orders of magnitude, which suggests that the wear rate is notvery sensitive to the sliding velocity, as expected. An increase in the wear resistance withsliding velocity is due presumably to the increase in the strain rate which increases the yieldstress of the material. At very high sliding velocities, high interface temperatures as a result offrictional heating result in a decrease in the yield stress of the material being abraded, whichcounteracts the effect due to the increased strain rate. Wear rate increases with an increasein grit size up to about 100 μm and beyond this size the wear rate becomes less sensitive tothe particle size. In explaining this behavior, we note that in the wear equation, although theabrasive particle size does not enter explicitly, it is conceivable that the roughness factor tan θ

may be size-dependent. With large abrasive particles, the shape of the abrasive particles doesnot depend on the particle size. The surface roughness of the abrasive surface with the sametype and density of abrasive particles has an effect on the wear rate. Wear rate increases withan increase in the surface roughness of the abrasive tape with CrO2 magnetic particles sliding(partially flying) on an Ni-Zn ferrite head, Figure 7.2.21.

Wear 337

Figure 7.2.20 Wear rate of copper, subjected to two-body abrasion by SiC abrasive paper, as a functionof abrasive particle size at two different sliding velocities. Reproduced with permission from Misra, A.and Finnie, I. (1981), “Some Observations on Two-Body Abrasive Wear,” Wear 68, 41–56. Copyright1981. Elsevier.

The dependence of the abrasive wear rate as a function of the sliding distance is morecomplex. If wear takes place with fresh abrasive paper (in two-body wear) or fresh abrasiveparticles (in three-body wear), wear continues at a steady rate (Rabinowicz, 1995; Bhushan,1985, 1996). However, if a limited amount of abrasive is used as the sliding continues, the

Figure 7.2.21 Wear rate of Ni-Zn ferrite as a function of RMS surface roughness of a magnetic tapewith CrO2 magnetic particles. Reproduced with permission from Bhushan, B. (1985), “Assessment ofAccelerated Head-Wear Test Methods andWear Mechanisms,” in Tribology and Mechanics of MagneticStorage Systems, Vol. 2 (B. Bhushan and N.S. Eiss, eds), pp. 101–111, special publication SP-19, ASLE,Park Ridge, Illinois. Copyright 1985. ASLE. (Source: Bhushan, 1985).


Figure 7.2.22 Wear volume of steel as a function of sliding distance subjected to two-body abrasionby 220 grade silicon carbide paper. Reproduced with permission fromMulhearn, T.O. and Samuels, L.E.(1962), “The Abrasion of Metals: AModel of the Process,”Wear 5, 478—498. Copyright 1962. Elsevier.

wear rate generally decreases as a function of time. For example, in Figure 7.2.22 (Mulhearnand Samuels, 1962), the wear rate decreases as a function of the sliding distance when steelis abraded on silicon carbide abrasive paper. Mulhearn and Samuels (1962) reported that thedata fit the following form:

v = v∞[1− exp (−αx)

](7.2.12)

where v∞ is the total volume of metal removed if the sliding is continued indefinitely andα is a constant. Similar results for abrasive magnetic tapes sliding against ceramic heads oranother abrasive tape have been reported by Bhushan (1985, 1996).A decrease in the wear rate as a function of sliding distance is believed to occur as a result

of blunting of the abrasive surfaces in two-body wear or abrasive particles in three-body wear,Figure 7.2.23a. In addition, clogging of the abrasive surface by abraded debris occurs duringwear, Figure 7.2.23b (Rabinowicz, 1995). If at any instance, the wear debris is larger thanabrasive particles, it may leave the material being abraded above the level of the abrasivegrains and result in no additional wear. One can see that abrasive action should cease muchmore rapidly in wear with fine grades of abrasive paper than with coarse grades.

Effect of Relative Hardness of Abrasive Medium to WorkpieceIn two-body (Aleinikov, 1957; Richardson, 1968) and three-body (Rabinowicz, 1977, 1983)abrasive situations, if the abrading medium is softer than a workpiece, the wear coefficientdoes not remain constant. It is known that when the hardness ratio of the workpiece to theabrasive particles is less than unity, the wear coefficient remains approximately constant;however, if the ratio is equal to or greater than unity, the wear coefficient decreases rapidlywith an increase in the hardness ratio, Figure 7.2.24 (plotted by Rabinowicz, 1983 based ondata by Aleinikov, 1957 and Richardson, 1968). When the hardness of the workpiece is of thesame order of magnitude as the hardness of the abrasive particles, the wear of the workpieceis not rapid, since deformation occurs both in the abrasive particles and the workpiece, andwear generally occurs in both. When the workpiece is significantly harder than the abrasive

Wear 339

Figure 7.2.23 Schematics (a) of an abrasive surface before and after wear, showing blunting, and (b)of an abrasive surface clogged by wear debris. Reproduced with permission from Rabinowicz, E. (1995)Friction and Wear of Material, Second edition, Wiley, New York. Copyright 1995. Wiley.

Figure 7.2.24 Relative abrasive wear coefficient of materials, covering a range of hardnesses, as afunction of workpiece hardness to abrasive hardness quotient.


Table 7.2.2 Typical hardnesses of commonly used abrasives.

Abrasive material Hardness (GPa)

Diamond 80Cubic boron nitride 40Silicon carbide 25Alumina 21Quartz, silica 8Magnesium oxide 8

particles, negligible deformation, and consequently, wear occur. Accompanying the change inhardness ratio from less than unity to greater than unity, there is a significant change in theroughness of the worn surface because the wearmechanism has changed. Thewear coefficientsfrom the wear tests with workpieces with varying hardnesses against an abrasive can be usedto estimate the hardness of the abrasive (Rabinowicz, 1977; Xie and Bhushan, 1996b). Next,we note that if the abrasive wear is required, the abrasive material must be harder that thesurface to be abraded; it does not have to be much harder. The desired criteria for high wearare hardness and sharpness. Thus it is advantageous if the abrasive is brittle so that it resultsin sharp corners when it is subjected to high stresses; many nonmetals meet this criterion.Hardnesses of commonly used abrasives are listed in Table 7.2.2.

7.2.2.2 Abrasive Wear by Fracture

Quantitative EquationTo obtain a quantitative expression for abrasive wear of brittle solids by brittle fracture, weconsider an asperity with sharp geometry on a flat surface of a brittle solid, Figure 7.2.25(Evans and Marshall, 1981). At low loads, a sharp asperity contact will cause only plasticdeformation and wear occurs by plastic deformation. Above a threshold load, brittle fractureoccurs, and wear occurs by lateral cracking at a sharply increased rate. The threshold loadis proportional to (Kc/H)3 Kc (Lawn and Marshall, 1979). The H/Kc is known as the indexof brittleness, where H is hardness (resistance to deformation) and Kc is fracture toughness(resistance to fracture).Lateral cracks in amorphous materials develop from the residual stresses associated with

the deformed material (Lawn, 1993). The maximum extension of the crack is thus realizedwhen the penetrating asperity is removed. As a sharp asperity slides over the surface, lateralcracks grow upward to the free surface from the base of the subsurface-deformed region andmaterial is removed as platelets from the region bounded by the lateral cracks and the freesurface.The lateral crack length c for a sliding asperity contact is given by (Evans and Marshall,

1981)

c = α1

[(E/H )3/5

K 1/2c H 1/8

]W 5/8 (7.2.13)

Wear 341

Figure 7.2.25 Schematic of the mechanism of wear by a sharp asperity sliding on the flat surface ofa brittle material by lateral fracture. Reproduced by permission from Evans, A.G. and Marshall, D.B.(1981), “Wear Mechanisms in Ceramics,” in Fundamentals of Friction and Wear of Materials (D.A.Rigney, ed), pp. 439–452, Amer. Soc. Metals, Metals Park, Ohio. Copyright 1981. ASM International.

where α1 is a material-independent constant that depends on the asperity shape. The depth, d,of the lateral crack is given by (Evans and Marshall, 1981)

d = α2

(E

H

)2/5 (W

H

)1/2(7.2.14)

whereα2 is anothermaterial-independent constant. Themaximumvolume ofmaterial removedper asperity encounter per unit sliding distance is 2 dc. If N asperities contact the surface witheach carrying the loadW, then from Equations 7.2.13 and 7.2.14, the volume of wear per unitsliding distance of the interface is given by (Evans and Marshall, 1981),

v = α3N(E/H )W 9/8

K 1/2c H 5/8

(7.2.15)

where α3 is a material-independent constant. The ratio (E/H) does not vary by much fordifferent hard brittle solids. Therefore, wear rate is inversely proportional to the (fracturetoughness)1/2 and (hardness)5/8. Wear rate is proportional to (normal load)9/8 which impliesthat wear rate by lateral fracture increases more rapidly than linearly with the applied normalload as in plastic deformation. This implies that the wear coefficient in the wear equation isnot independent of load. Based on the method of calculation of c and d for the wear model,variations to Equation 7.2.15 have been reported in the literature.


Figure 7.2.26 Correlation between the reciprocal of wear rate at constant load with the materialproperty quantity K 1/2

c H 5/8 for several ceramic materials. Reproduced by permission from Evans, A.G.and Marshall, D.B. (1981), “Wear Mechanisms in Ceramics,” in Fundamentals of Friction and Wear ofMaterials (D.A. Rigney, ed), pp. 439–452, Amer. Soc. Metals, Metals Park, Ohio. Copyright 1981. ASMInternational.

Experimental EvidenceThere is some experimental evidence that wear rate in two-body abrasion is inversely propor-tional to (Kc1/2 H5/8). Figure 7.2.26 shows a good correlation in grinding experiments betweenwear rate at constant loads and material properties for various ceramic material combinationsas predicted by the lateral fracturemodel. For additional wear data, for example, see Yamamotoet al. (1994).

7.2.3 Fatigue Wear

Subsurface and surface fatigues are observed during repeated rolling (negligible friction) andsliding, respectively. The repeated loading and unloading cycles to which the materials areexposed may induce the formation of subsurface or surface cracks, which eventually, aftera critical number of cycles, will result in the breakup of the surface with the formation oflarge fragments, leaving large pits in the surface, also known as pitting. Prior to this criticalpoint (which may be hundreds, thousands, or even millions of cycles), negligible wear takesplace, which is in marked contrast to the wear caused by an adhesive or abrasive mechanism,

Wear 343

where wear causes a gradual deterioration from the start of running. Therefore, the amount ofmaterial removed by fatigue wear is not a useful parameter. Much more relevant is the usefullife in terms of the number of revolutions or time before fatigue failure occurs.Chemically enhanced crack growth (most common in ceramics) is commonly referred to

as static fatigue. In the presence of tensile stresses and water vapor at the crack tip in manyceramics, a chemically induced rupture of the crack-tip bonds occurs rapidly, which increasesthe crack velocity. Chemically enhanced deformation and fracture result in an increased wearof surface layers in static and dynamic (rolling and sliding) conditions.

7.2.3.1 Rolling Contact Fatigue

Adhesive and abrasive wear mechanisms are operative during direct physical contact betweentwo surfaces moving relative to each other. If the two surfaces are separated by a fluid film(and abrasive particles are excluded), these wear mechanisms do not operate. However, inan interface with nonconforming contact, the contact stresses are very high and the fatiguemechanism can be operative. In these cases, although direct contact does not occur, the matingsurfaces experience large stresses, transmitted through the lubricating film during the rollingmotion. Well-designed rolling element bearings usually fail by subsurface fatigue.From a Hertz elastic stress analysis, the maximum compressive stresses occur at the sur-

face, but the maximum shear stresses occur some distance below the surface, Figure 7.2.27.As rolling proceeds, the directions of the shear stresses for any element change sign. Time tofatigue failure is dependent on the amplitude of the reversed shear stresses, the interface lubri-cation conditions, and the fatigue properties of the rolling materials (Lundberg and Palmgren,

Figure 7.2.27 Variation of principal shear stress at various depths directly below the point of contactof two hard surfaces in pure rolling (μ = 0), pure sliding (μ, high value), and combined contact(μ, moderate value). The z is the distance below the surface in the vertical direction and a is half of theHertzian diameter.


Figure 7.2.28 Spalling of a 52100 ball bearing race from subsurface fatigue. Reproduced with permis-sion from Tallian, T.E., Baile, G.H., Dalal, H., and Gustafsson, O.G. (1974), Rolling Bearing Damage,SKF Industries Inc., King of Prussia, Pennsylvania. Copyright 1974. SKF USA Inc.

1947, 1951). When a fatigue crack does develop, it occurs below the surface, until a regionof metal is separated to some extent from the base metal by the crack and ultimately becomesdetached and spalls out. By the time cracks grow large enough to emerge at the surface andproduce wear particles, these particles may become large spalls or flakes. A typical example ofspalling of a ball bearing race due to subsurface fatigue is shown in Figure 7.2.28. Sincemateri-als in rolling contact applications are often surface hardened, the surface could be brittle.Hence,cracks may also initiate at the surface as a result of tensile stresses and lead to surface fatigue.The earliest method of rating rolling element bearing is based on the stochastic life predic-

tion methods based on the Weibull distribution (Weibull, 1951; Bhushan, 1999). The Weibulllife model was developed by Lundberg and Palmgren (1947, 1951). This method was laterstandardized by ISO in 1962 (281) and is recommended by the Antifriction Bearing Manu-facturers Association (AFBMA). It is still widely used. The life of rolling element bearings(normally referred to as L10 or B10) inmillions of revolutions for 90%of the bearing populationis determined from

L10 = (C/W )p (7.2.16a)

Wear 345

where

p = 3 for ball bearings (point contacts)

= 10

3for roller bearings (line contacts).

Here, C is the bearing’s basic load capacity and W is the equivalent radial or thrust load forthe radial or thrust bearing, respectively. The basic load capacity of a bearing is the load that90% of the bearings can endure for 1 million revolutions under given running conditions.The AFBMA method for determining bearing load rating as published in bearing manufac-

turer’s catalogs is based on bearing tests conducted in the 1940s. There have been significantimprovements inmaterials and processing. The bearing life also depends on the lubrication andoperating conditions. In other words, the bearing life depends on the bearing geometry (sizeand accuracy of manufacture), the physical properties (such as the modulus of elasticity andfatigue strength), the metallurgy of bearing materials, and the lubrication conditions (such asviscosity, speed of rotation, and surface roughness). It is assumed that various bearing designfactors, as a first approximation, are multiplicative. Then the expected bearing life for givenoperating conditions, L A, can be related to the calculated rating life L10 by using the variouslife adjustment factors (Bamberger et al., 1971):

L A = (D)(E)(F)(G)(H ) L10 (7.2.16b)

where D is the bearing material factor, E is the bearing processing factor, F is the lubricationfactor, G is the speed effect factor and H is the misalignment factor. Typical values of lifeadjustment factors are presented by Bamberger et al. (1971), Harris (1991) and Zaretsky(1992) and these can be less than or greater than 1.Because of the continued improvement in bearing materials and better understanding of

bearing behavior, in many cases, bearings manufactured from clean, homogeneous steel, vir-tually infinite life can be obtained. However, Lundberg-Palmgren theory, based on probabilityof survival from subsurface-initiated fatigue, predicts a finite life. Ioannides and Harris (1985)developed a modified theory in which they introduced the fatigue limit as the lower limit of thefatigue behavior, i.e., they assumed that no failure can occur if the stress in a volume elementis less than or equal to the endurance limit of the material. This modification is able to predictinfinite L10 life.Based on statistical methods (e.g., Johnson, 1964; Nelson, 1982), the life of a system with

multiple bearings of lives L1, L2, . . . for a given probability of survival is

L =[(

1

L1

)β

+(1

L2

)β

+ · · ·]1/β

(7.2.17)

where β is the Weibull slope for a Weibull distribution. β is about 1.5 for rolling elementbearings. The probability of the survival of multiple bearings with probabilities 1–Pi in fraction(i = 1, . . . n) for n bearings is (1− P1) (1− P2), . . . (1− Pn). For n bearings with identical(1 − P), the probability of survival of the system is (1 – P)n.



The basic load capacity of a radial ball bearing is 8 kN. Calculate its life based upon a 90%probability of survival for the bearing operating at 600 rpm and at radial loads of 6 kN and12 kN.

Solution

For a ball bearing,

L10 = (C/W )3

For W = 600 N, L10 =(8

6

)3× 106 rev.

= 2.37× 106 rev.Bearing life in hours = 2.37× 106/(600× 60) h = 65.8 h

For W = 1200 N, bearing life in hours = 65.8 (1/2)3 = 8.3 h


A ball bearing spindle with two radially loaded ball bearings and belt driven at 600 rpm, drivestwo pumps. The radial load on each bearing is 10 kN. Given that the basic load capacity ofeach bearing is 25 kN, calculate the L10 life of the system.

Solution

L10 = (C/W )3 × 106 rev.= (25/10)3 × 106 rev.= 15.63× 106 rev.

= 15.63× 106600× 60 = 434 h

Since the spindle has two bearings, the system probability of survival for a life of 434 h is0.92 or 81%. Assuming β = 1.5, the life of the system at a 90% probability of survival, fromEquation 7.2.17, is

L10 of the system = 434

21/1.5h = 273 h

Wear 347

Rolling/Sliding Contact FatigueRolling contact is frequently accompanied by slip or sliding. The complex motions in mostrolling contact situations produce at least a small fraction of slip or sliding (on the order of1 to 10%) such as in rolling bearings, hypoid gear teeth, can roller followers, and wheel-railcontacts. The friction stresses due to sliding cause the maximum shear stresses to be nearerthe surface, Figure 7.2.27, and the failure occurs by near surface fatigue. (See Hertz analysisof an elastic sphere acting on an elastic semi-infinite solid presented in Chapter 3.) Slip canresult in severe adhesive wear (scuffing) damage to the mating surfaces. Proper lubrication isimportant to minimize the deleterious effects of slip in these rolling-contact situations.

Sliding Contact FatigueWe have seen that, when sliding surfaces make contact via asperities, wear can take place byadhesion and abrasion. However, it is conceivable that asperities can make contact withoutadhering or abrading and can pass each other, leaving one or both asperities plasticallydeformed from the contact stresses. As the surface and subsurface deformation continues,cracks are nucleated at and below the surface. Once the cracks are present (either by cracknucleation or from pre-existing voids or cracks), further loading and deformation cause cracksto extend and propagate. After a critical number of contacts, an asperity fails due to fatigue,producing a wear fragment. In a sliding contact, friction is generally high compared to arolling contact, and the maximum shear stress occurs at the surface (Figure 7.2.27), whichleads to surface fatigue. This may be the situation in a boundary lubrication system in whichone or several monolayers of lubricant or absorbed surface layers at the interface separate theasperities but contact stresses are still experienced by the asperities.

7.2.3.2 Static Fatigue

Static fatigue results from a stress-dependent chemical reaction between water vapor and thesurface of the ceramic. The rate of reaction depends on the state of stress at the surface andthe environment. The stress is greatest at the roots (or tips) of small cracks in the material,and consequently the reaction proceeds at its greatest rate from these roots. The small cracksgradually lengthen and failure occurs when the cracks are long enough to satisfy the Griffithfailure criteria for fracture (Wiederhorn, 1967). Thus, two stages of crack growth can bevisualized: (1) slow crack motion occurs because of chemical attack at the crack tip; and(2) a catastrophic stage of crack motion is initiated when the crack is long enough to satisfythe Griffith criteria. The time to failure is the time required for the crack to grow from thesubcritical to critical Griffith size (Wiederhorn, 1967, 1969; Westwood, 1977; Lawn, 1993).This moisture-assisted crack propagation and fracture is called static fatigue.The stresses (residual stresses produced during machining and stresses introduced during

static or dynamic contact) at the crack tip control the rate of crack growth. The kinetics offluid flow from the environment also control the rate of rupture of the crack-tip bonds. Fora start, the gas molecules do not have direct access to the crack-tip bonds. The mean freepath for intermolecular collisions at ambient conditions is typically 1 μm, which will clearlyexceed the crack-wall separation some distance behind the tip of a brittle crack. Thus, as thegas molecules migrate along the crack interface, a point will be reached where collisions withthe walls become more frequent than with other gas molecules. The gas then enters a zone of


Figure 7.2.29 (a) Gaseous flow along interface to tip of brittle crack. Flow changes from that of a low-viscosity fluid to that of a dilute gas as the crack-wall separation diminishes below the intermolecularmean free path. (b) Two-dimensional representation of water-induced bond rupture in silica glass.Large circles, oxygen; intermediate circles, silicon; small circles, hydrogen. Solid circles denote speciesoriginally from the environment. Reproduced with permission from Lawn, B.R. (1993), Fracture ofBrittle Solids, Second edition, Cambridge University Press, Cambridge. Copyright 1993. CambridgeUniversity Press.

free molecular flow within which diffused molecular scattering at the walls may considerablyattenuate the flow rate (Figure 7.2.29a).One of the best studied examples of chemically enhanced crack growth is that of silica glass

in the presence of a water environment. The basic crack-tip reaction is as follows:

(H − O − H )+ (−Si − O − Si−) → (−Si − OH HO − Si−) (7.2.18)

Wear 349

That is, an incident water molecule hydrolyzes a siloxane bridging bond at the crack tip toform two terminal silane groups. A two-dimensional representation of the process is given inFigure 7.2.29b.The dependence of crack velocity on applied stress and water vapor for soda-lime glass

(amorphous structure) is shown in Figure 7.2.30a. A similar effect is also seen for a single-crystal sapphire in water vapor, Figure 7.2.30b. As the level of applied stress is raised, the crackvelocity shows an initial, rapid increase, followed by a rather abrupt saturation. Humidity alsoplays a significant role; crack velocity changes by three orders of magnitude (0.1–100 μm/s)when the relative humidity is changed by three orders of magnitude (0.1–100% RH) forthe same nominal stress. We note that the moisture-assisted fracture occurs for both crystallineand amorphous ceramics. For a given material and the environment, wear due to static fa-tigue can be reduced with lower residual stresses and a lesser degree of microcracking of thesurface (Bhushan, 1996).Wear of metals and ceramics in three-body abrasion with ceramic abrasive particles (Larsen-

Basse, 1975) and of ceramics in two-body abrasion (Larsen-Basse and Sokoloski, 1975;Wallbridge et al., 1983; Takadoum, 1993; Bhushan, 1996; Bhushan and Khatavkar, 1996) isreported to increase sharply with an increase in relative humidity of the environment above theambient conditions while there is little effect at low humidities for example, see Figure 7.2.31.Moisture-assisted fracture of the abrasive particles in three-body abrasion, at high humidi-

ties, bringsmore and sharper cutting asperities into contact with the abrading surface. However,moisture-assisted fracture of one of the sliding ceramic bodies in two-body abrasion producessharp particles; some of which become trapped at the interface and result in high wear ratesfrom three-body abrasion. Even in the absence of relative sliding, loose particles can be gen-erated from ceramic bodies with high density of cracks placed in tension and exposed to highhumidities. In an experiment, an Mn-Zn ferrite (MnO-ZnO-Fe2O3) ceramic rod under a statictensile stress was exposed to close to 100% RH for seven days. Debris particles could becollected in a dish placed under the rod, Figure 7.2.32. In a sliding interface, some of theseparticles would be trapped at the interface resulting in high wear rates.

7.2.4 Impact Wear

Two broad types of wear phenomena belong under this heading: erosive and percussive wear.Erosion can occur by jets and streams of solid particles, liquid droplets, and implosion ofbubbles formed in the fluid. Percussion occurs from repetitive solid body impacts. Repeatedimpacts result in progressive loss of solid material.

7.2.4.1 Erosion

Solid Particle ErosionSolid particle erosion occurs by impingement of solid particles, Figure 7.2.33. It is a form ofabrasion that is generally treated rather differently because the contact stress arises from thekinetic energy of particles flowing in an air or liquid stream as it encounters a surface. Theparticle velocity and impact angle combined with the size of the abrasive give a measure ofthe kinetic energy of the impinging particles, that is, of the square of the velocity. Wear debrisformed in erosion occurs as a result of repeated impacts.


Figure 7.2.30 (a) Crack velocity for (amorphous) soda-lime glass tested in moistened nitrogen gas(relative humidities indicated) at room temperature. Some data points are included to demonstrate thescatter of the data between runs. Reproduced with permission fromWiederhorn, S.M. (1967), “Influenceof Water Vapor on Crack Propagation in Soda-Lime Glass,” J. Amer. Cer. Soc. 50, 407–414. Fig. 3 p.409. Copyright 1967. Wiley, (b) Crack velocity for single-crystal sapphire tested in moistened nitrogengas (relative humidities indicated) at room temperature. (Source: Wiederhorn, 1969).

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Figure 7.2.31 Wear rate and coefficient of friction as a function of relative humidity for Mn-Znferrite sliding against a magnetic tape with CrO2 magnetic particles. Reproduced with permission fromBhushan, B. and Khatavkar, D.V. (1996), “Role of Water Vapor on the Wear of Mn-Zn Ferrite HeadsSliding Against Magnetic Tapes,”Wear 202, 30–34.

As in the case of abrasive wear, erosive wear occurs by plastic deformation and/or brittlefracture, dependent upon material being eroded away and upon operating parameters. Wear-rate dependence on the impact angle for ductile and brittle materials is different, as shownin Figure 7.2.34 (Bitter, 1963). Ductile materials will undergo wear by a process of plastic

Figure 7.2.32 Optical micrographs of particles shed by stressed Mn-Zn ferrite rod placed at 22◦C and100% RH for seven days. Reproduced with permission from Bhushan, B. and Khatavkar, D.V. (1996),“Role of Water Vapor on the Wear of Mn-Zn Ferrite Heads Sliding Against Magnetic Tapes,”Wear 202,30–34.


Figure 7.2.33 Schematic of a jet of abrasive particles hitting a surface at a high velocity.

deformation in which the material is removed by the displacing or cutting action of the erodedparticle. In a brittle material, on the other hand, material will be removed by the formation andintersection of cracks that radiate out from the point of impact of the eroded particle (Finnie,1960). The shape of the abrasive particles affects that pattern of plastic deformation aroundeach indentation, consequently the proportion of the material displaced from each impact.In the case of brittle materials, the degree and severity of cracking will be affected by theshape of the abrasive particles. Sharper particles would lead to more localized deformationand consequently wear, as compared to the more rounded particles.Two basic erosionmechanisms have been observed for erosion of ductile materials (Bellman

and Levy, 1981; Soderberg et al., 1983): cutting erosion and deformation (plowing) erosion. Incutting erosion, the detachment of crater lips occurs by one or several impacts of the microma-chining, plowing or lip formation type. Cutting erosion is, in many respects, similar to abrasive

Figure 7.2.34 Rate of erosive wear as a function of angle of attack (with respect to the material plane)of impinging particles.

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Figure 7.2.35 SEM micrograph of 303 stainless steel surface after solid particle erosion.

wear. In deformation erosion, the detachment of the material occurs by surface fragmentationdue to several impacts of the indentation type. The effect of a single-indentation-type crateris to raise small lips of material around the impact type. The effect of successive impacts isto flatten and to strain further the lips, creating thin platelets of highly stressed metal that arefinally knocked off the surface by succeeding particles. The relative importance of the twoerosion mechanisms at multiple impacts is strongly dependent on the angle of impingement.Cutting erosion and deformation erosion dominate at grazing and normal incidence, respec-tively (Hutchings and Winter, 1974; Hutchings et al., 1976). Surface hardness and ductilityare the most important properties for cutting and deformation erosion resistance, respectively.An SEM micrograph of 303 stainless steel (microhardness of 320 kg/mm2) eroded by sandblasting is shown in Figure 7.2.35. An SEM of the cross-section of worn samples showed thepresence of a very thin layer (∼2 μm) of visible plastic deformation. From Figure 7.2.35, itappears that erosion results primarily from a deformation erosion mechanism.Solid particle erosion is a problem in machinery such as ingested sand particles in gas

turbine blades, helicopter and airplane propellers, the windshields of airplanes, the nozzles forsand blasters, coal turbines, hydraulic turbines and the centrifugal pumps used for coal slurrypipelines. It has useful application in processes such as sand blasting, abrasive deburring, anderosive drilling of hard materials.

Quantitative EquationWe first consider erosion, involving plastic deformation, by a single hard particle striking asofter surface at normal incidence, Figure 7.2.36. Based on Hutchings (1992), assume that the


Figure 7.2.36 Schematic of erosion by a single hard particle striking a softer surface at normalincidence.

particle does not deform and the deformation of the surface is perfectly plastic with a constantindentation pressure (hardness), H. At time t after initial contact, the particle, of mass dmwith an initial velocity V, indents the surface to a depth x such that cross-sectional area ofthe indent impression is A(x), which is dependent upon the shape of the particle. The upwardforce decelerating the particle is due to the plastic pressure acting in the contact area A(x). Theequation of motion of the particle is given as

− H A(x) = dmd2x

dt2(7.2.19)

If the particle comes to rest at a depth d after time t0, the work done by the retarding force isequal to the initial kinetic energy of the particle

∫ d

0H A(x)dx = 1

2dmV 2 (7.2.20)

or

dv = dm V 2

2H(7.2.21a)

where dv is the volume of material displaced from the indentation. If there are particles of atotal mass m, then

v = mV 2

2H(7.2.22b)

All of the displaced material does not end up as wear debris. If k is the proportion of thedisplaced material result as wear debris, then

v = kmV2

2H(7.2.23a)

Summation of Equation 7.2.23a, over many impacts gives the volume of wear for a periodover which erosion takes place. The erosion wear equation is normally written in terms of

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dimensionless erosion ratio (E), the mass of material removed divided by the mass of erosiveparticles striking the surface. Equation 7.2.23a can be rewritten as

E = kρV 2

2H(7.2.23b)

where ρ is the density of the material being eroded.Comparedwith abrasivewear equations, the volume of erosivewear is inversely proportional

to the hardness as in abrasive equations. The normal load in abrasive wear is replaced by mV 2

in erosive wear (Hutchings, 1992). The derivation of Equation 7.2.23 is based on an extremelysimple model. It does not include the effect of the impact angle and the shape and size ofparticles. k depends on the impact angle and the shape and size of particles. The value of ktypically ranges from 10−5 to 10−1.The erosive wear rate by brittle fracture depends, in addition, on the fracture toughness of

the material being eroded (Hutchings, 1992).

Experimental EvidenceThe erosion of puremetals shows strong sensitivity to particle impact velocity, and the hardnessof metals. The erosion is related to the particle velocity by E ∝ V n , where n ranges between2.3 and 3. For erosive wear data on copper as a function of impact velocity at two impactangles, see Figure 7.2.37. For pure annealed metals, erosion decreases with an increase inthe hardness, where as observed for abrasive wear, dependence for work hardened metals isnot linear. A better correlation is found between the erosion resistance and the hardness of asurface after work hardening by erosion.

Liquid Impingement ErosionWhen small drops of liquid strike the surface of a solid at high speeds (as low as 300 m/s), veryhigh pressures are experienced, exceeding the yield strength of most materials. Thus, plasticdeformation or fracture can result from a single impact, and repeated impact leads to pittingand erosive wear. In many cases, the probable impact velocities and impact angles are such thatpure liquid impingement erosion is an unlikely mechanism; an erosion-corrosion mechanismusually does more damage (Preece, 1979). The damage by this process is important in theso-called moisture erosion of low-pressure steam turbine blades operating with wet steam,rain erosion of aircraft or missile surfaces and helicopter rotors, nuclear power plant pipes,and heat exchangers.Based on Haymann (1992), the high-velocity impact of a liquid drop against a solid surface

produces high contact pressure in the impact region followed by liquid jetting flow along thesurface, radiating out from the impact area. In ductile materials, a single intense impact mayproduce a central depression, with a ring of plastic deformation around it where the jetting-outflow may remove the material by a tearing action. In brittle materials, circumferential cracksmay form around the impact site caused by tensile stress waves propagating outward along thesurface. In subsequent impacts, material can spall off the inside surface due to the compressivestress wave from the impact reflecting there as a tensile wave.


Figure 7.2.37 Erosion ratio of copper as a function of impact velocity for two impact angles. Re-produced with permission from Ives, L.K. and Ruff, A.W. (1979), in Erosion: Prevention and UsefulApplications (W.F. Adler, ed), pp. 5–35, Special Tech. Pub. ASTM, Philadelphia. Copyright 1979. ASTMInternational.

Cavitation ErosionCavitation is defined as the repeated nucleation, growth, and violent collapse of cavities orbubbles in a liquid. Cavitation erosion arises when a solid and fluid are in relative motion,and bubbles formed in the fluid become unstable and implode against the surface of the solid.When bubbles collapse that are in contact with or very close to a solid surface, they willcollapse asymmetrically, forming a microjet of liquid directed toward the solid. The solidmaterial will absorb the impact energy as elastic deformation, plastic deformation or fracture.The latter two processes may cause localized deformation and/or erosion of the solid surface(Preece, 1979). Damage by this process is found in components such as ships’ propellers andcentrifugal pumps.All liquids contain gaseous, liquid and solid impurities, which act as nucleation sites for the

bubbles or vapor-filled voids. When a liquid is subjected to sufficiently high tensile stresses,bubbles are formed at weak regions within the liquid. Subsequently if this liquid is subjectedto compressive stresses, i.e. to higher hydrostatic pressures, these bubbles will collapse. Inpractice, cavitation can occur in any liquid in which the pressure fluctuates either becauseof flow patterns or vibration in the system. If, at some location during liquid flow, the local

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pressure falls below the vapor pressure of the liquid, then cavities may be nucleated, grow toa stable size and be transported downstream with the flow. When they reach the high-pressureregion, they become unstable and collapse (Hansson and Hansson, 1992). The stability ofa bubble is dependent on the difference in pressure between the inside and outside of thebubble and the surface energy of the bubble. The damage created is a function of the pressuresproduced and the energy released by collapse of the bubble. Thus, reduction of surface tensionof the liquid reduces damage, as does an increase in vapor pressure.Materials that are resistant to fatigue wear, namely, hard but not brittle materials, are also

resistant to cavitation. Resistance to corrosive attack by the liquid, however, is an additionalrequirement for cavitation resistance.

7.2.4.2 Percussion

Percussion is a repetitive solid body impact, such as experienced by print hammers in high-speed electromechanical applications and high asperities of the surfaces in a gas bearing. Inmost practical machine applications, the impact is associated with sliding; that is, the relativeapproach of the contacting surfaces has both normal and tangential components known ascompound impact (Engel, 1976). Percussive wear occurs by hybrid wear mechanisms whichcombines several of the following mechanisms: adhesive, abrasive, surface fatigue, fracture,and tribochemical wear.To model compound impact, Figure 7.2.38a shows a slug of radius R normally directed at a

speed V against a tangentially moving platen with a speed u (Engel, 1976). The slug of massm may be idealized as a point mass, supported by a tangential spring with a stiffness of k.After a time ts (called the slipping time), the slug comes up to the horizontal speed u of theplaten. They then travel together in the horizontal direction for the rest of the impact durationti , after which the two bodies separate. The Hertz impact force during the impact duration isshown in Figure 7.2.38b. To simplify contact analysis, we can assume the Hertz impact force,F(t), to have a sinusoidal distribution,

F(t) = F0 sin

(π t

ti

)(7.2.24)

The approximation retains the peak force F0 while raising the rest of the force-time curvefrom the bell shaped one, typical of Hertz impact, to sinusoidal. The peak force F0 dependson geometrical, material, and normal impact parameters. For an elastic contact of a slug on aninfinitely massive flat platten with an infinitely high modulus of elasticity (Figure 7.2.38a),

F0

∣∣∣∣elastic =(5

3mV 2

) 3/5[

4E

3(1− ν2

) R1/2]2/5

(7.2.25)

We now write the equation of tangential motion for an infinitely massive platen,

mx = μ

∫ t

0F(t) dt, 0 ≤ t ≤ ts (7.2.26a)


Figure 7.2.38 Schematic of (a) impact of a slug on a tangentially moving platen, and (b) impact forcecycle.

where μ is the coefficient of friction. Using Equations 7.2.24 and 7.2.26, we calculate theslipping time, ts

ts ∼ tiπcos−1 (1− S) (7.2.27a)

where

S = πmu

μF0ti(7.2.27b)

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S is called the slip factor. If S = 0, normal impact occurs. For larger S, slipping persists fora longer time during impact (compound impact). If S ≥ 2, slipping persists during the entirecontact time (ti ).The impact wear is proportional to the slip factor because wear primarily occurs during

the portion of the impact spent in relative sliding. Normal impact on a harder substrate canproduce fracture, and repeated impacts can give rise to a subsurface fatigue wear mechanism.An impact with associated sliding (compound impact) gives rise to surface fatigue and/oradhesive/abrasive wear; specific wear mechanisms depend on the geometrical, material, andoperative conditions. For materials with high toughness, the contribution due to surface fatigueis negligible.We apply the impact wear analysis to a print head striking on a paper covering the platen

(Engel, 1976). Abrasive wear of the print head occurs during the slipping time. The abradedvolume v with respect to the sliding distance x of the print head relative to the paper is

dv(t)

dx(t)= kF(t)

H(7.2.28)

where k is the abrasive wear coefficient and H is the hardness of the print head.Therefore, the total wear volume per impact cycle is

v =∫ ts

0

k

HF(t)dx(t) = kuti F0S

2π H, 0 ≤ S ≤ 2 (7.2.29a)

= kmV2

2 μH= 2kuti F0

π H

(1− 1

S

), S ≥ 2 (7.2.29b)

Once the wear volume per cycle has been determined, the total wear after N cycles can bepredicted by multiplying the unit wear by N cycles.

7.2.5 Chemical (Corrosive) Wear

Chemical or corrosive wear occurs when sliding takes place in a corrosive environment. In air,the most dominant corrosive medium is oxygen. Therefore chemical wear in air is generallycalled oxidative wear. In the absence of sliding, the chemical products of the corrosion (e.g.,oxides) would form a film typically less than a micrometer thick on the surfaces, which wouldtend to slow down or even arrest the corrosion, but the sliding action wears the chemical filmaway, so that the chemical attack can continue. Thus, chemical wear requires both chemicalreaction (corrosion) and rubbing. Machinery operating in an industrial environment or nearthe coast generally produces chemical products (i.e., it corrodes) more rapidly than whenoperating in a clean environment. Chemical wear is important in a number of industries, suchas mining, mineral processing, chemical processing, and slurry handling.Corrosion can occur because of the chemical or electrochemical interaction of the interface

with the environment. Chemical corrosion occurs in a highly corrosive environment and inhigh-temperature and high-humidity environments. Electrochemical corrosion is a chemicalreaction accompanied by the passage of an electric current, and for this to occur, a potentialdifference must exist between two regions. The region at low potential is known as an anode


Figure 7.2.39 SEM micrograph of 52100 quenched and tempered roller bearing after corrosive wear.Reproduced with permission from Tallian, T.E., Baile, G.H., Dalal, H., and Gustafsson, O.G. (1974),Rolling Bearing Damage, SKF Industries Inc., King of Prussia, Pennsylvania. Copyright 1974. SKFUSA Inc.

and the region at high potential is known as a cathode. If there is a current flow between theanode and cathode through an electrolyte (any conductive medium), at the anode the metaldissolves in the form of ions and liberates electrons. The electrons migrate through the metalto the cathode and reduce either ions or oxygen. Thus, electrochemical corrosion is equivalentto a short-connected battery with partial anodic and partial cathodic reactions occurring onthe two sliding members (commonly referred to as galvanic corrosion) or in a sliding memberon two regions atomic distances away. These regions may shift to different locations (Wagnerand Traud, 1938). Electrochemical corrosion is influenced by the relative electropotential.Electrochemical corrosion may accelerate in a corrosive environment because corrosive fluidsmay provide a conductive medium necessary for electrochemical corrosion to occur on therubbing surfaces. The most common liquid environments are aqueous, and here small amountsof dissolved gases, commonly oxygen or carbon dioxide, influence corrosion.A typical example of a corroded roller subsequent to running in a bearing is shown in

Figure 7.2.39. The corrosion left a multitude of dark-bottomed pits, the surroundings of whichare polished by running. The condition subsequently creates extensive surface-originatedspallings from a multitude of initiated points.

7.2.5.1 Tribochemical Wear

Friction modifies the kinetics of chemical reactions of sliding bodies with each other, andwith the gaseous or liquid environment, to the extent that reactions which occur at hightemperatures occur at moderate, even ambient, temperatures during sliding. Chemistry dealing

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with this modification of chemical reaction by friction or mechanical energy is referred to astribochemistry, and the wear controlled by this reaction is referred to as tribochemical wear(Heinicke, 1984; Fischer, 1988). The most obvious mechanism by which friction increases therate of chemical reaction (tribochemistry) is frictional heat produced at contacting asperities.Besides the friction heat, other mechanisms are: the removal of product scale resulting in freshsurfaces; accelerated diffusion and direct mechanochemical excitation of surface bonds. Thetribochemical reactions result in oxidative wear of metals, the tribochemical wear of ceramics,formation of friction polymer films on surface sliding in the presence of organics, and thedissolution of silicon nitride in water during sliding without fracture. Oxidative wear of metalsand tribochemical wear of ceramics will be described later in Section 7.4.Applications of tribochemistry include formation of friction polymer films for low friction

andwear and tribochemical polishingwithout fracture (also called chemomechanical polishingor CMP).The formation of organic films on sliding surfaces occurs as a result of the repolymerization

of organics (in the form of solid, liquid, or vapors) in the sliding contact due to a localincrease in surface temperature in the regions of greatest contact and possibly due to the addedcatalytic action of certain freshly exposed surfaces. These films are known as friction polymersor tribopolymers (Chaikin, 1967; Furey, 1973; Lauer and Jones, 1986; Bhushan and Hahn,1995). The chemical structure of the friction polymers must be different from the organicsource from which they are formed. In many cases, these films reduce friction and wear ofsliding surfaces operating in boundary or hydrodynamic lubrication regimes. These protectivepolymeric films may be intentionally produced by adding additives to the lubricating oil. Theprocess of forming friction polymers and degrading them is a dynamic one, that is, initiallyformed films will degrade and must be replenished. Degraded friction polymers can producehigh friction and wear. Also the thickness of the friction polymer films may become so largethat they begin to delaminate and generate wear debris. A second form of tribochemistry is thetribochemical polishing in ceramics in liquids. Friction in water enhances the tribochemicalreaction. For example, the tribochemical dissolution of silicon nitride in water occurs atcontacting asperities that are removed, and it results in extremely smooth surfaces (Fischer,1988). If the material removal is purely tribochemical, no microfracture or plastic deformationtakes place. Therefore, the polished surface is free of microcracks and other defects.

7.2.6 Electrical-Arc-Induced Wear

When a high potential is present over a thin air film in a sliding process, a dielectric breakdownresults that leads to arcing. During arcing, a relatively high-power density (on the order of1 kW/mm2) occurs over a very short period of time (on the order of 100 μs). The heat-affectedzone is usually very shallow (on the order of 50 μm). Heating is caused by the Joule effect dueto the high power density and by ion bombardment from the plasma above the surface. Thisheating results in considerable melting and subsequent resolidification, corrosion, hardnesschanges, and other phase changes, and even in the direct ablation ofmaterial (Guile and Juttner,1980; Bhushan and Davis, 1983). Arcing causes large craters, and any sliding or oscillationafter an arc either shears or fractures the lips, leading to three-body abrasion, corrosion, surfacefatigue, and fretting (Figure 7.2.40). Arcing can thus initiate several modes of wear resultingin catastrophic failures in electrical machinery (Bhushan and Davis, 1983).


Figure 7.2.40 SEMmicrographs of typical worn area by electrical-arc-induced wear of a 303 stainlesssteel surface. Arc craters and sheared lips can be seen.

When contacts are rubbing, as in the case of a copper commutator or copper slip ringsagainst graphite-based brushes (Johnson and Moberly, 1978), the sparking damage on thecopper commutator or slip ring can cause excessive wear of the brush by abrasion. In certainapplications of bearings in electrical machinery, there is the possibility that an electric currentwill pass through a bearing. When the current is broken at the contact surfaces between rollingelements and raceways, inner-race and shaft, or outer-race and housing, arcing results. Bothsurfaces should be in the path of least resistance to a potential difference. Electrical-arc-induced wear has been productively used as a method of metal removal in electrodischargemachining.

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Methods to minimize electrical-arc-induced wear are as follows: (1) to eliminate the gapbetween the two surfaces with a potential difference; (2) to provide an insulator of adequatedielectric strength (e.g., an elastomer orAl2O3 coating) between the two surfaces; (3) to providea low impedance connection between the two surfaces to eliminate the potential difference; or(4) to have one of the surfaces not ground. Bearing manufacturers recommend that bearingsshould be press-fitted to the shaft and conducting grease should be used to eliminate arcing,for example in the case of rolling-element bearings, shaft and the inner race, the inner andouter races, and rolling elements.

7.2.7 Fretting and Fretting Corrosion

Fretting occurs where low-amplitude oscillatory motion in the tangential direction (rangingfrom a few tens of nanometers to few tens of microns) takes place between contacting surfaces,which are nominally at rest (Anonymous, 1955; Hurricks, 1970;Waterhouse, 1981, 1992). Thisis a common occurrence, since most machinery is subjected to vibration, both in transit and inoperation. Examples of vulnerable components are shrink fits, bolted parts, and splines. Thecontacts between hubs, shrink- and press-fits, and bearing housings on loaded rotating shaftsor axles are particularly prone to fretting damage. Flexible couplings and splines, particularlywhere they form a connection between two shafts and are designed to accommodate somemisalignment, can suffer fretting wear.Basically, fretting is a form of adhesive or abrasive wear, where the normal load causes

adhesion between asperities and oscillatory movement causes ruptures, resulting in weardebris. Most commonly, fretting is combined with corrosion, in which case the wear modeis known as fretting corrosion. For example, in the case of steel particles, the freshly wornnascent surfaces oxidize (corrode) to Fe2O3, and the characteristic fine reddish-brown powderis produced, known as cocoa. These oxide particles are abrasive. Because of the close fit of thesurfaces and the oscillatory small amplitude motion (on the order of a few tens of microns),the surfaces are never brought out of contact, and therefore, there is little opportunity for theproducts of the action to escape. Further oscillatorymotion causes abrasive wear and oxidation,and so on. Therefore the amount of wear per unit sliding distance due to fretting may be largerthan that from adhesive and abrasive wear. The oscillatory movement is usually the result ofexternal vibration, but in many cases it is the consequence of one of the members of the contactbeing subjected to a cyclic stress (i.e., fatigue), which results in early initiation of fatigue cracksand results in a usually a more damaging aspect of fretting, known as fretting fatigue.Surfaces subjected to fretting wear have a characteristic appearance with red-brown patches

on ferrous metals and adjacent areas that are highly polished because of the lapping quality ofthe hard iron-oxide debris. Figure 7.2.41 shows the SEMmicrograph of the 303 stainless steelshaft after it underwent fretting corrosion.A rapid increase in wear rate occurs with slip amplitude over an amplitude range, Fig-

ure 7.2.42. For a given slip amplitude, the amount of wear per unit of sliding distance per unitof applied normal load linearly increases with the number of oscillating cycles up to an ampli-tude of about 100 μm. Above this amplitude, the wear rate per unit sliding distance becomesconstant, identical with unidirectional or reciprocating sliding wear rates. This then gives apossible upper limit for the slip amplitude for the case of true fretting. At small amplitudes,characteristic of fretting, the relative velocities are much lower, even at high frequencies,


Figure 7.2.41 SEM micrographs of 303 stainless steel shaft surface after fretting corrosion.

Figure 7.2.42 Volume wear rate per unit sliding distance per unit of normal load as a function of slipamplitude for mild steel against itself. Each curve is the result of a separate investigation. Reproducedwith permission from Waterhouse, R.B. (1992), “Fretting Wear,” in ASM Handbook, Vol. 18: Friction,Lubrication and Wear Technology, pp. 242–256, ASM International, Metals Park, Ohio. Copyright 1992.ASM International.

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compared with conditions in typical unidirectional sliding. The fretting wear rate is directlyproportional to the normal load for a given slip amplitude. In a partial slip situation, the fre-quency of oscillation has little effect on the wear rate per unit distance in the low-frequencyrange, whereas the increase in the strain rate at high frequencies leads to increased fatiguedamage and increased corrosion due to rise in temperature. However, in the total-slip situation,there is little effect of the frequency (Waterhouse, 1992).There are various design changes which can be carried out to minimize fretting wear. The

machinery should be designed to reduce oscillatory movement, reduce stresses or eliminatetwo-piece design altogether.

7.3 Types of Particles Present in Wear Debris

The size and shape of debris may change during sliding in dry and lubricated systems;therefore, the condition of a system can be monitored by debris sampling and maintenance canbe scheduled, known as condition-based maintenance. Mild wear is characterized by finelydivided wear debris (typically 0.01–1 μm in particle size). The worn surface is relativelysmooth. Severe wear, in contrast, results in much larger particles, typically on the order of20–200 μm in size, which may be visible even with the naked eye; the worn surface is rough.Particles are collected during sliding for analysis. Particles are collected from dry interfacesby sucking air from the sliding interface on a filter paper. Particles in a sample of lubricantfrom an oil-lubricated system are recovered by filtration, centrifuging, or magnetically (formagnetic particles). The technique commonly used for ferrous metals, known as ferrography,uses a magnetic field to sort particles by the size and shape of the magnetic particles (Scottand Westcott, 1977). The size, shape, structural, and chemical details of particles are analyzedusing various techniques including optical microscopy, scanning electron microscopy (SEM),transmission and scanning transmission electron microscopy (TEM/STEM), energy dispersiveandwavelength dispersive spectroscopy (EDS andWDS), Auger electron spectroscopy (AES),X-ray photoelectron spectroscopy (XPS), X-ray and electron diffraction. Size analysis ofairborne particles is also carried out using particle counters, generally based on a light-scattering method.Particles can be classified based on the wear mechanism or their morphology. Particles

collected from a wear test may not be in the same state in which these were first producedbecause of changes in subsequent sliding. Since it is difficult to identify the exact possible wearmechanism, particles are generally classified based on their morphology, and their descriptionfollows (Scott, 1975; Scott and Westcott, 1977; Ruff et al., 1981; Samuels et al., 1981;Hokkirigawa and Kato, 1988; Rigney, 1992; Glaeser, 2001).

7.3.1 Plate-Shaped Particles

Thin, plate-shaped or flake-type wear particles with an aspect ratio of 2–10, are commonlyfound in wear debris from dry and lubricated interfaces. These particles are produced as aresult of plowing followed by repeated loading and unloading fatigue, as a result of nucleationand propagation of subsurface cracks or plastic shear in the asperity contacts, Figures 7.3.1and 7.3.2. Note that particles produced in rolling-contact fatigue are large and their formationdevelops macroscopic pits.


Figure 7.3.1 Secondary electron images of (a) a wear track on single-crystal Si(111) after slidingagainst a diamond pin at a normal load of 0.5 N and a sliding velocity of 25 mm/s in vacuum, and(b) a flake-type debris particle. Reproduced with permission from Rigney, D.A. (1992), “The Role ofCharacterization in Understanding Debris Generation” in Wear Particles (D. Dowson, C.M. Taylor,T.H.C. Childs, M. Godet, and G. Dalmaz, eds), pp.405–412, Elsevier Science Publishers, Amsterdam.Copyright 1992. Elsevier.

7.3.2 Ribbon-Shaped Particles

Ribbon-shaped or cutting-type particles are frequently found with aspect ratios, on the orderof ten or more; and usually are curved and even curly. These are produced as a result of plasticdeformation. They have all the characteristics ofmachining chips: as a result, the ribbon-shaped

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Figure 7.3.2 SEM micrograph of flake-type debris particle generated by sliding a Cu-Be block on aM2 tool steel ring at a normal load of 133 N, a sliding velocity of 50 mm/s and a sliding distance of360 m in dry argon. Reproduced with permission from Rigney, D.A. (1988), “Sliding Wear of Metals,”Ann. Rev. Mater. Sci. 18, 141–163. Copyright 1988. Annual Reviews.

particles are referred to as microcutting chips or cutting chips. They are generally producedduring run-in, as a result of detachment of fin-like ridges generally present at the edges of theabrasion grooves in machined (e.g., ground) surfaces. These particles are produced with sharpasperities or abrasive particles digging into the mating surface with material flowing up thefront face of the asperity or abrasive particles and being detached from the wearing surface inthe form of a chip, Figure 7.3.3. Typically, changes in chemical composition are small.

7.3.3 Spherical Particles

Spherical particles are not common. Wear particles of various shapes may not escape fromthe interface to become loose debris. Some of them remain trapped and are processed furtheras in the spherical shape. Spherical particles have been observed in sliding (Rigney, 1992),fretting and rolling contact fatigue (Smith, 1980; Samuels et al., 1981). Spherical particleswith 1–5 μm in diameter are reported to be associated with rolling-contact fatigue just priorto fatigue failure, Figure 7.3.4.

7.3.4 Irregularly Shaped Particles

The majority of particles have an irregular morphology. Wear debris produced by detach-ment of the transferred fragment in adhesive wear and brittle fracture are irregularly shaped,Figure 7.3.5.


Figure 7.3.3 SEM micrograph of ribbon-shaped or cutting-type debris particle generated by sliding a304 stainless steel block on anM2 tool steel ring at a normal load of 67 N and sliding velocity of 50 mm/sin dry argon. Reproduced with permission from Rigney, D.A. (1992), “The Role of Characterization inUnderstanding Debris Generation” inWear Particles (D. Dowson, C.M. Taylor, T.H.C. Childs, M. Godet,and G. Dalmaz, eds), pp.405–412, Elsevier Science Publishers, Amsterdam. Copyright 1992. Elsevier.

Figure 7.3.4 SEM micrograph of spherical particles present on the surface of a crack produced as aresult of rolling contact fatigue of steel surfaces. Reproduced with permission from Smith, R.A. (1980),“Interfaces of Wear and Fatigue,” in Fundamentals of Tribology (N.P. Suh and N. Saka, eds), Figure 5,C© 1980 Massachusetts Institute of Technology, by permission of The MIT Press.

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Figure 7.3.5 SEM micrographs of (a) a wear surface of an austenitic steel showing irregular frag-mentation of the surface, and (b) typical debris particles produced from multiple fracturing in a brittlelayer produced on the surface during sliding. Reproduced with permission from Samuels, L.E., Doyle,E.D., and Turley, D.M. (1981), “Sliding Wear Mechanisms,” in Fundamentals of Friction and Wear ofMaterials (D.A. Rigney, ed), pp. 13–41, Amer. Soc. Metals, Metals Park, Ohio. Copyright 1981. ASMInternational.

7.4 Wear of Materials

Wear process is generally quantified by wear rate. Wear rate is defined as the volume ormass of material removed per unit time or per unit sliding distance. Other forms could bedimensionless, such as the depth of material per unit sliding distance, or the volume removedper apparent area of contact and per unit sliding distance.Wear rate is generally not constant. Ingeneral, wear rate is a complex function of time. Wear rate may start low and later rise, or viceversa, Figure 7.4.1. After a certain duration, the wear rate remains constant for a period and

Figure 7.4.1 Three hypothetical cases of wear volume as a function of sliding distance showing run-in,steady-state and failure regions.


may change if transition from one mechanism to another occurs during a wear test. The initialperiod during which wear rate changes is known as the run-in or break-in period. Wear duringrun-in depends on the initial material structure and properties and on surface conditions suchas surface finish and the nature of any films present. During this transition period, the surfaceroughness is modified to a steady-state condition by plastic deformation. Initial conditionsaffect the damage during the transition period and its duration.The wear rate, like friction, of a material is dependent upon the counterface or mating

material (or material pair), surface preparation, and operating conditions. The usefulness ofwear coefficients or wear data presented in the published literature lies more in their relativemagnitudes than in their absolute values. A number of handbooks present typical wear ratesof a variety of material pairs (Peterson and Winer, 1980; Blau, 1992; Bhushan and Gupta,1997, Bhushan, 2001a, 2001b, 2011). The wear rate for a material pair is normally presentedin terms of the nondimensional wear coefficient (k). Wear rate is also presented in terms ofa wear factor defined as the wear volume per unit applied normal load and per unit slidingdistance (mm3/Nm). Typical ranges of the coefficient of friction and the wear coefficientsof metals, alloys, ceramics, polymeric and solid lubricant pairs are presented in Table 7.4.1.Values of selected pairs are presented in Table 7.4.2.A self-mated steel pair exhibits high friction and wear. Pairss of dissimilar metals exhibit

moderate friction andwear. These are generally used in lubricated applications. Ceramic versusmetal or versus another ceramic or versus itself exhibits moderate friction but extremely lowwear. Self-mated ceramic pairs as opposed to self-mated metal pairs are desirable as they are

Table 7.4.1 Typical ranges of friction and wear coefficients of various material pairs.

Material combinationTypical range ofcoefficient of friction

Typical range of wearcoefficient Comments

Self-mated metalspairs

0.5– > 1 (high) 5 × 10−3 (high) Undersirable in drycontacts

Dissimilar metalspairs

Alloy-alloy pairs

0.3–0.9 (moderate)0.2–0.6 (moderate)

10−4–10−3 (moderate)10−6–10−3 (moderate)

Easy to fabricate, low cost

Ceramic-metal pairsCeramic-ceramicpairs

0.25–0.8 (moderate)0.25–0.7 (moderate)

10−7–10−4 (very low)10−7–10−5 (very low)

For high temperatureapplications, for lowestwear requirements,self-mated ceramic pairsdesirable, good forunlubricated conditions

Polymer-hard surfacepairs

0.05∗–0.6 (low) 10−6–10−3 (low) For corrosive environmentand low frictionapplications, good forlow loads, good forunlubricated conditions

Solid lubricant-hardsurface pairs

0.05–0.15 (very low) 10−4–10−3 (low) For lowest frictionrequirement

∗PTFE

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Table 7.4.2 Coefficient of friction and wear coefficients of softer material for various material pairs inthe unlubricated sliding at a normal load of 3.9 N and sliding speed of 1.8 m/s (Archard, 1980). Thestated value of the hardness is that of the softer (wearing) material in each example.

Materials

Wearing surface Counter surface

Vickersmicrohardness(kg/mm2)

Coefficient offriction Wear coefficient (k)

60/40 leaded brass Tool steel 95 0.24 6 × 10−4

Mild steel Mild steel 186 0.62 7 × 10−3

Ferritic stainless steel Tool steel 250 0.53 1.7 × 10−5

Stellite Tool steel 690 0.60 5.5 × 10−5

Tungsten carbide Tungsten carbide 1300 0.35 1 × 10−6

PTFE Tool steel 5 0.18 2.4 × 10−5

not abusive to the mating surface. Since they exhibit very low wear, these are used in bothunlubricated and lubricated conditions. Polymers and solid lubricants against hard surfacesexhibit very low friction but not very low wear.Metal pairs are most commonly used because of the ease of machinability and low cost.

Ceramics are used because they are somewhat inert, strong, and can be used at high tempera-tures. Polymers are inexpensive and ideal in corrosive environments. In the case of polymericmaterials and solid lubricants, their mechanical properties degrade and in some cases oxidize attemperatures somewhat higher than ambient, making them unusable at elevated temperatures.The temperature rise occurs as a result of friction heating, which is a function of a product ofpressure and velocity known as PV limit. These materials are classified based on the PV limit.Since polymeric materials and solid lubricants exhibit a low coefficient of friction and wearwhether self-mated or sliding against other materials, they are commonly used in unlubricatedapplications. These are commonly used against harder mating materials.In journal-bearing applications with soft liners, embeddability and conformability are im-

portant considerations. If particles are longer than the thinnest region of the oil film, the particlemay reside at the interface and may result in significant abrasive wear. One of the ways tominimize the damage is to select a hard journal shaft and a soft bearing alloy or a polymersuch that particles are embedded in the bearing material. The ability to embed the abrasivesin this way is referred to as “embeddability.” Further in the applications, the use of bearingswith significant misalignment, resulting in high loads, can lead to significant damage. Sideloads from the misalignment can be accommodated without severe damage by the use of softbearing alloys or polymers. The soft materials can deform plastically and accommodate anymisalignment.

7.4.1 Wear of Metals and Alloys

As indicated in Chapters 4 and 5, the clean metals and alloys in a solid contact exhibit highadhesion consequently high friction and wear. The wear rate of contacting metallic surfacescleaned in a high vacuum can be very high. The slightest contamination mitigates contact orforms chemical films which reduce adhesion, resulting in the reduction of friction and wear(Buckley, 1981; Rigney, 1988; Bhushan, 1996). In the case of soft metals, such as In, Pb, and


Table 7.4.3 Wear coefficient of softer material for various metal-metal pairs at a normal load of 20 Nand a sliding velocity of 1.8 m/s (Source: Archard, 1953). The stated value of hardness is that of thesofter wearing material in each example.

Metal pairVickers hardness(kg/mm2) Wear coefficient, k (× 10−4)

Cadmium on cadmium 20 57Zinc on zinc 38 530Silver on silver 43 40Copper on copper 95 110Platinum on platinum 138 130Mild steel on mild steel 158 150Stainless steel on stainless steel 217 70Cadmium on mild steel 20 0.3Copper on mild steel 95 5Platinum on mild steel 138 5Mild steel on copper 95 1.7Platinum on silver 43 0.3

Sn, the contact area is large, even at low loads, which results in high wear rates. Hexagonalmetals such as Co and Mg as well as other non-hexagonal metals such as Mo and Cr exhibitlow friction and wear; consequently Co, Mo, and Cr are common alloying elements in steels toreduce friction, wear and corrosion. Metallurgical compatibility determines the wear rates ofa given metal pair. Lead-based white metals (babbitts), brass, bronze, and cast iron generallyexhibit relatively low friction and wear in dry and lubricated conditions. In general, wear foralloys tends to be lower than that for pure components. Typical values of wear coefficients forvarious similar and dissimilar metals are presented in Table 7.4.3.Steels form the most commonly used family of materials for structural and tribological

applications. Based on chemical composition (percentage of alloying components and carbon)and processing, a variety of microstructures and physical properties of steel can be obtained.The wear resistance of different microstructures are summarized in Figure 7.4.2 (Moore, 1981;Zum Gahr, 1987; Glaeser, 1992).In metal-to-metal wear tests, high stresses can result in catastrophic galling and eventual

seizure even after a single cycle, so a committee of stainless steel producers of AISI devised abutton and block galling test. In this test, a small button specimen and a large block specimenare machined and polished to provide parallel contacting surfaces. The specimens are dead-weight loaded in a Brinell hardness tester, and the button is rotated 360◦ against the block.Specimens are then examined for galling at 10x magnification, with new specimens beingtested at progressively higher stress levels until galling just begins. This point is called theunlubricated threshold galling stress. Galling usually appears as a groove, or score mark,terminating in a mound of metal (Schumacher, 1977; Anonymous, 1978; Foroulis, 1984).Galling stress is a good measure of wear resistance of a given material pair.Galling data show that identicalmetal couples usually do poorly in terms of galling compared

with dissimilar metal couples. When stainless steels are coupled with each other, with theexception of some Nitronic steels, they exhibit worse galling resistance than all other steelsby a factor of 2 or more.

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Figure 7.4.2 Relative wear resistance as a function of hardness of different microstructures of steels.Reproduced with permission from Zum Gahr, K.H. (1987), Microstructure and Wear of Materials,Elsevier, Amsterdam. Copyright 1987. Elsevier.

Cobalt-based alloys such as T-400 and Stellite 6B have, in general, good galling resis-tance. However, several nickel-based alloys exhibit a very low threshold galling stress whenself-mated or coupled with other similar alloys (Bhansali, 1980). A nickel-based alloy suchas Waukesha 88 can be modified specifically for galling resistance. Waukesha 88 exhibitsextremely high galling resistance in combination with several stainless steels. It should also benoted that a high nickel content in steels has a detrimental effect on galling resistance. Whencompared with steels, with the exception of 316 stainless steel, the remaining steels exhibitmoderate threshold galling stress when coupled with the cobalt-based alloy Stellite 6B. Type316 steel probably exhibits a low galling stress because it has a higher nickel content than type304 stainless steel.Wear data (Anonymous, 1978; Bhansali, 1980; Foroulis, 1984) show that among the various

steels tested, types 201, 301, and hardened 440C and the proprietary Nitronic austenitic gradesprovide good wear resistance when mated to themselves under unlubricated conditions. High-nickel alloys generally are rated as intermediate between the austenitic andmartensitic stainlesssteels. Cobalt-based alloys also do well. Considerable improvement in wear resistance can beachieved when dissimilar metals are coupled, and this is especially true for steels coupled withsilicon bronze and Stellite alloys. The wear data further suggest that improvement in wearresistance can be achieved by altering the surface characteristics, such as by surface treatmentor by adding a coating.Operating conditions – normal load, sliding velocity and environment – have a significant

effect on the wear modes as well as wear rates. Their discussion follows.


7.4.1.1 Effect of Temperature (Oxidative Wear)

Interface temperatures produced at asperity contacts during sliding of metallic pairs undernominally unlubricated conditions result in thermal oxidation which produces oxide filmsseveral microns thick. The oxidation is generally a beneficial form of corrosion. A thick oxidefilm reduces the shear strength of the interface which suppresses the wear as a result of plasticdeformation (Quinn, 1983a, 1983b). In many cases, tribological oxidation can reduce the wearrate of metallic pairs by as much as two orders of magnitude, as compared with that of thesame pair under an inert atmosphere. Tribological oxidation can also occur under conditions ofboundary lubrication when the oil film thickness is less than the combined surface roughnessof the interface. The oxidation can prevent severe wear. In oxidative wear, debris is generatedfrom the oxide film.At low ambient temperatures, oxidation occurs at asperity contacts from frictional heating.

At higher ambient temperatures, general oxidation of the entire surface occurs and affects wear.In the case of steels, the predominant oxide present in the debris depends on the sliding condi-tions. At low speeds and ambient temperatures, the predominant oxide is α − Fe2O3, at inter-mediate conditions it is Fe3O4, and at high speeds and temperatures it is FeO (Quinn, 1983b).Oxygen and other molecules are adsorbed on clean metals and ceramic surfaces, and form

strong chemical bonds with them. The slow step inhibiting the continuation of this reactionis the diffusion of the reacting species through the film of reaction product. Oxidation of ironand many metals follows a parabolic law, with the oxide film thickness increasing with thesquare root of time,

h = Ct1/2 (7.4.1)

where h is the oxide film thickness, t is the average growth time, and C is the parabolic rateconstant at elevated temperatures.Since diffusion is thermally activated, growth rate in oxide film thickness during sliding

as a function of temperature, similar to thermal oxidation under static conditions, follows anArrhenius type of relationship

K = A exp(−Q/RT) (7.4.2)

where K is the parabolic rate constant for the growth of the oxide film, A is the parabolicArrhenius constant (kg2/(m4 s)) for the reaction,Q is the parabolic activation energy associatedwith oxide (kJ/mole), R is the universal gas constant and T is the absolute temperature of thesurface. It has been reported that the Arrhenius constant for sliding is several orders ofmagnitude larger than that for static conditions, which means that oxidation under slidingconditions is much more rapid than that in the static oxidation condition. Increased oxidationduring sliding may result from increased diffusion rates of ions through a growing oxide filmwhich generally consists of high defect density due to mechanical perturbations.

7.4.1.2 Effect of Operating Conditions (Wear-Regime Maps)

The wear-regime maps elucidate the role of operating environment on wear mechanisms. Nosingle wear mechanism operates over a wide range of conditions. There are several wear

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mechanisms which change in relative importance as the operating conditions are changed.The transitions in dominant wear mechanisms, consequently wear rates, occur as sliding loadsand velocities are changed. In some cases, changes also occur as a function of sliding time (ordistance). The dominant wear mechanisms are based on mechanical strength and interfacialadhesion. Increase in normal load results in an increase in mechanical damage due to high sur-face stresses. Increase in both normal load and sliding velocity results in a monotonic increasein interface temperature. High interface temperature results in the formation of chemical films,mostly the formation of oxide films in air. High temperatures result in a decrease in mechanicalstrength and in some cases in structural changes. At high load-velocity (PV) conditions, theremay even be localized melting near the surface.Various regimes of mechanical (plastically dominated) and chemical (oxidational) wear for

a particular sliding material pair are observed on a single wear-regime map (or wear-modemap or wear-mechanism map) plotted on axes of normalized pressure and normalized slidingvelocity on the macroscale (Lim and Ashby, 1987; Lim et al., 1987) and nanoscale (Tambeand Bhushan, 2008). Here we present maps for macroscale and maps for nanoscale will bepresented later in Chapter 10 on nanotribology. Normalized pressure is the nominal pressuredivided by the surface hardness (p/H) and normalized velocity is the sliding velocity dividedby the velocity of heat flow (given by the radius of the circular nominal contact area divided bythe thermal diffusivity). As an example, a wear regime map for steel sliding on steel in air atroom temperature in the pin-on-disk configuration is shown in Figure 7.4.3. The general formof the map would be similar for the sliding of most unlubricated metals in air. It can be seenthat, in principle, the map can be divided into areas corresponding to different wear regimes,with boundaries of sliding velocities and contact pressure beyond which oxidative wear wouldbe dominant, as compared to mechanical wear at low speeds. Prevailing wear mechanisms cangive mild or severe wear. Mild wear gives a smooth surface and severe wear produces a surfacethat is rough and deeply torn and the wear rate is usually high. The wear rates may differ by asmuch as two or three orders of magnitude. The transition between mild and severe wear takesplace over a wide range of sliding conditions. These are load-dependent, velocity-dependentor load- and velocity-dependent. In addition, some are sliding-distance dependent.The mild wear occurs because direct metal–metal contacts are minimized mostly by the

oxide layer produced as a result of frictional heating. Mild wear takes place under fourdistinct sets of conditions (Lim et al., 1987). In the first set (i), at low contact pressuresand sliding velocities, a thin (usually several nm thick) and ductile oxide film is formedwhich prevents direct metal–metal contact and is not ruptured at light loads. In the secondset (ii), at higher velocities, a thicker and more brittle oxide film is continuously generatedby high interface temperatures. Continuous oxidation replenishes the oxide film. In the thirdset (iii), at higher loads, a hard surface layer (martensite) formed on carbon–steel surfacesbecause of localized frictional heating followed by rapid quenching as the friction heat isdissipated. The higher interface temperatures also produce a thicker film of oxide, supportedby the hardened substrate. In the fourth set (iv), at yet higher sliding velocities, the increasedinterface temperature produces thick films. Insulating oxide films reduce the heat flow fromthe surface to the underlying conducting substrate resulting in severe oxidation.Severe wear occurs under conditions in which direct metal–metal contacts occur. Severe

wear takes place under three distinct sets of conditions (Lim et al., 1987). In the first set(i), at high contact pressures and low sliding velocities, contact pressures are high enough torupture the thin oxide layer [as described in mild wear – (i)] which leads to direct metal–metal


Figure 7.4.3 Wear-regime map for unlubricated steel sliding on steel in air at room temperature inthe pin-on-disk configuration. Reproduced with permission from Lim, S.C., Ashby, M.F., and Brunton,J.H. (1987), “Wear-Rate Transitions and Their Relationship to Wear Mechanisms,” Acta Metall. 35,1343–1348. Copyright 1987. Elsevier.

contact. In the second set (ii), at moderate contact pressures and sliding velocities, the loadis high enough to penetrate the thicker but brittle oxide films generated [as described in mildwear – (ii)]. In the third set (iii), at high contact pressures and sliding velocities, the slidingconditions are so severe that local temperatures reach the melting point of the steel, resultingin a liquid film in contact which leads to severe wear.All steels exhibit similar wear-regime maps as just discussed. Many other metals show

similar behavior. These wear maps are useful to provide guidance with respect to the properselection of materials and performance envelopes for metals.

7.4.2 Wear of Ceramics

Ceramics exhibit high mechanical strength, do not lose much mechanical strength or oxidizereadily at elevated temperatures and are resistant to corrosive environments; therefore, ceramiccouples are commonly used in extreme environmental applications, such as high loads, highspeeds, high temperatures and corrosive environments. High mechanical properties result invery low real area of contact responsible for low friction and very low wear. Under clean

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Table 7.4.4 Abrasive wear coefficient values of various ceramic pairs at roomtemperature (self-mated couples) in a three-body abrasive wear experiment at anormal load of 10 N, hydrocarbon-base lapping fluid and 600-grit SiC.

Material k (× 10−3)

SiC 1.1WC 3.3B4C 5.5Si3N4 43

environments, the coefficients of friction and the wear rates of ceramic pairs do not reach thevery high values observed in clean metals, especially in ultra-high vacuum or in the absenceof oxygen. Ceramic materials show only limited plastic flow at room temperature and muchless ductility than metals. The fracture toughness of ceramics is an important property in thewear of ceramics. Ceramic materials respond to conventional lubricants similar to metals.A significant amount of wear data of ceramic pairs is found in the literature (e.g., Bhushan

and Sibley, 1982; Anonymous, 1987; Jahanmir, 1994; Bhushan and Gupta, 1997). Wear dataof various ceramic pairs at elevated temperatures are presented by Zeman and Coffin (1960).It is desirable to use self-mated ceramic pairs, unlike in metals. Table 7.4.4 presents the weardata of ceramic pairs against themselves in a three-body abrasive wear experiment conductedusing a lapping machine (Bhushan and Gupta, 1997).Operating conditions – normal load, sliding velocity and environment – have asignificant

effect on the wear modes as well as the wear rates. A discussion of these follows.

7.4.2.1 Effect of the Operating Environment

Tribochemical interactions of ceramicswith the liquid or gaseous environment control thewearand friction of ceramics. Depending on the chemical reaction between the ceramic and theenvironment, wear and friction can decrease or increase (Lancaster, 1990). It can change boththe wear mechanisms and the wear rates. This chemical interaction can result in modificationof the surface composition and decrease in the purely chemical form of wear by dissolutionin the liquid environment (chemomechanical polishing or CMP), but it can induce chemicalfracture, which increases wear rates. The CMP process has been described earlier.

Tribochemical Wear of Non-oxide CeramicsThe formation of oxide films in the case of non-oxide ceramics exposed to an oxidizing envi-ronment and the formation of hydrated layers in all ceramics exposed to humid environmentsare responsible for a change in friction as a function of the environment (Fischer, 1988). Theformation of chemical films during sliding at interface temperatures is referred to as tribo-chemistry. Non-oxide ceramics such as silicon nitride, silicon carbide, titanium nitride andtitanium carbide are all known to form oxide films during sliding in an oxidizing environment.Oxygen may be derived from oxygen in the air or from water vapor, e.g.,

Si3N4 + 3O2 → 3SiO2 + 2N2 (7.4.3)


or

Si3N4 + 6H2O → 3SiO2 + 4NH3 (7.4.4)

Oxide ceramics react with water, whether it is present as a liquid or as a vapor. In the case ofnon-oxide ceramics, oxidation (Equations 7.4.3 or 7.4.4) can be followed by hydration,

SiO2 + 2H2O → Si(OH)4 (7.4.5)

During sliding at low temperatures (ambient and from frictional heating), the kinetics of chem-ical reactions does not allow a sufficient amount of oxide to form. The rate of oxidation is accel-erated by the simultaneous action of friction. The hydrated layer [Si(OH)4] exhibits low frictionand wear and also provides the source of the wear debris. The removal of the hydrated materialexposes the nascent surface and the hydrated layer is reformed from tribochemical reaction.As an example, Figure 7.4.4 shows the effect of the absorbed water on friction and wear. The

values of the coefficient of friction and the wear rates fall with the increasing availability ofwater for silicon nitride sliding against itself in dry nitrogen gas, air of two different humiditylevels, and liquid water. The reactions outlined in Equations 7.4.3 to 7.4.5 lead to the formationand hydration of a silica film at the interface, which is soft with low shear strength and reducesthe coefficient of friction and wear rate. A decrease in the wear rate of silicon nitride againstitself as a function of increase in the relative humidity has also been reported by Fischer andTomizawa (1985). A drop in the coefficient of friction and wear rate as a function of relativehumidity has also been observed for silicon carbide, Figure 7.4.5 (Kapelski, 1989).

Figure 7.4.4 Effect of environment on the coefficient of friction and wear rate of hot-pressed siliconnitride on itself at a normal load of 10 N and sliding velocity of 150 mm/s, in a pin-on-disk configuration.Reproduced with permission from Ishigaki, H., Kawaguchi, I., Iwasa, M. and Toibana, Y. (1986),“Friction and Wear of Hot Pressed Silicon Nitride and Other Ceramics,” ASME J. Trib. 108, 514–521.Copyright 1986. ASME.

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Figure 7.4.5 Coefficient of friction and wear rate of silicon carbide on itself as a function of relativehumidity at a normal load of 10 N and sliding velocity of 0.1 m/s after a sliding distance of 1 km in aball-on-disk configuration.

Figure 7.4.6 shows the effect of sliding velocity on the coefficient of friction and wear rateof hot-pressed silicon nitride on itself. Both the coefficient of friction and the wear rate arerelatively constant at low velocities but increase considerably at transition velocities of about200 mm/s. At low velocities, silicon nitride reacts with water vapor in the air and forms ahydrated film which is responsible for low friction and wear. At higher velocities, the interface

Figure 7.4.6 Coefficient of friction and wear rate as a function of sliding velocity of hot-pressed siliconnitride on itself at a normal load of 10 N and ambient air in a pin-on-disk configuration. Reproduced withpermission from Ishigaki, H., Kawaguchi, I., Iwasa, M., and Toibana, Y. (1986), “Friction and Wear ofHot Pressed Silicon Nitride and Other Ceramics,” ASME J. Trib. 108, 514–521. Copyright 1986. ASME.


Figure 7.4.7 Coefficient of friction as a function of sliding revolutions for single-crystal silicon onitself in ambient air and dry nitrogen environment in a pin-on-disk configuration.

temperature increases, which reduces the amount of water vapor in the air. Reduction in thewater vapor results in a reduced amount of tribochemical products, which is responsible forhigh friction and wear.Friction tests conducted with bare single-crystal silicon and thermally oxidized single-

crystal silicon sliding against themselves show that the coefficient of friction in dry nitrogenis about half that of in air, Figure 7.4.7 (Venkatesan and Bhushan, 1994). These data suggestthat the interaction between the surfaces is strong in air (promoted by either oxygen or watervapor present in the ambient air), resulting in higher friction and consequently high wear inair compared to nitrogen.

Chemically-Induced Fracture (Static Fatigue) in Oxide CeramicsIn the case of many oxide ceramics such as alumina (Wallbridge et al., 1983; Kapelski, 1989),and zirconia (Fischer et al., 1988), friction and wear of these ceramics also show strongsensitivity to water, but in these cases the coefficient of friction and the wear rate increase withan increase in relative humidity, Figure 7.4.8 (Kapelski, 1989). This increase inwear rate occursbecause of enhanced crack growth rate, which results from the attack of the bonds betweenthe neighboring metal and oxide ions at a crack tip by water. This chemically induced fracturephenomenon is known as stress-corrosion cracking or static fatigue. Exposure to humiditymay also increase surface plasticity as a result of change in the mobility of near-surfacedislocations, with consequent wear. This chemomechanical effect, in which the mechanicalproperties of many materials change as a result of exposure to many liquids, is also known asthe Joffe–Rehbinder effect (Rehbinder and Shchukin, 1972).The effect of temperature on the coefficient of friction and wear is shown in Figure 7.4.9

(Dong et al., 1991). At temperatures below 200◦C and at temperatures above 800◦C, thecoefficient of friction and the wear volume are low. In the temperature range of 200–800◦C,the coefficient of friction and the wear volume are very large.

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Figure 7.4.8 Coefficient of friction and wear rate of silicon carbide with alumina as a function ofrelative humidity at a normal load of 10 N and sliding velocity of 0.1 m/s after a sliding distance of 1 kmin a ball-on-disk configuration.

7.4.2.2 Effect of Operating Conditions (Wear-Regime Maps)

As stated for metals, no single wear mechanism operates for ceramics either over a widerange of conditions. Various regimes of mechanical (plastic deformation or brittle fracture)and chemical (or tribochemical) wear for a particular sliding materials pair are observed ona single wear-regime map plotted on axes of Hertzian pressure and sliding velocity (Hsuand Shen, 1996). Wear mechanisms dominated by plastic flow or tribochemical reactions

Figure 7.4.9 Coefficient of friction and wear volume as a function of temperature for α-alumina onitself at a normal load of 59 N and sliding velocity of 1.4 mm/s in a ball-on-a flat configuration.


Figure 7.4.10 Wear-regime maps for Al2O3 and Si3N4 sliding on themselves in air at room temperaturein a ball-on-three-flats geometry on a four-ball wear tester (a) under dry conditions and (b) under paraffinoil lubricated conditions. Wear volumes per unit time are also listed in the figures. Reproduced withpermission fromHsu, S.M. andShen,M.C. (1996), “CeramicWearMaps,”Wear 200, 154–175.Copyright1996. Elsevier.

generally result in mild wear with low wear rates and smooth surfaces. The wear debris isgenerally finely divided and may be chemically different from the bulk material, whereas wearmechanisms dominated by brittle intergranular fracture result in severe wear with high wearrates and rough surfaces. The wear debris is generally angular and not chemically differentfrom the substrate. Mild wear occurs at a combination of low pressures and velocities,whereas severe wear occurs at combinations of high pressures and velocities.The wear-regime maps for ceramics are material specific. Wear maps for Al2O3 and Si3N4

under dry and paraffin-oil-lubricated conditions are shown in Figure 7.4.10. The tests were

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conducted by using a ball-on-three-flats geometry on a four-ball wear tester with identicalmaterials. Wear volumes per unit time are also listed in the figures. Various wear mechanismslisted in the figures are self-explanatory. Note that the interaction of the lubricant with theceramics extends the pressure-velocity boundary toward the higher values for a transitionbetween mild to severe wear.

7.4.3 Wear of Polymers

Polymers include plastics and elastomers. Polymers generally exhibit low friction as com-pared to metal and ceramic couples but exhibit moderate wear. Most commonly used plas-tics in tribological applications include polytetrafluoroethylene (PTFE), acetal, high-densitypolyethylene (HDPE), polyamide (Nylon), poly (amide-imide), polyimide and polyphenylenesulfide (Bartenev and Lavrentev, 1981; Briscoe, 1981; Bhushan and Wilcock, 1982, Bhushanand Gupta 1997). Most commonly used elastomers include natural and synthetic rubber, buta-diene – acrylonitrile (Buna-N or nitrile) rubber, styrene-butadiene rubber (SBR) and siliconerubber (Bhushan and Winn, 1981; Bhushan and Gupta, 1997). These polymers are a family ofself-lubricating solids. The polymer composites, impregnated generally with fibers of carbongraphite or glass and powders of graphite, MoS2, bronze and PTFE, are used for their desirablemechanical and tribological properties. Polymers are also used as additives to nonpolymericsolids and liquid lubricants.The dominant wear mechanisms are adhesive, abrasive, and fatigue. If the mating surface is

smooth, then the wear primarily occurs from adhesion between the mating surfaces. As statedin Chapter 5, wear of many polymers occurs first by the transfer of polymer to the hardermating surface followed by removal as wear particles (Steijn, 1967; Lancaster, 1973; Bhushanand Wilcock, 1982). During the initial run-in period, a steady-state condition is reached. Ifthe steady-state condition is reached, the wear rate is generally small and stable. The transferfilm thickness for PTFE composites sliding against smooth, steel surfaces is on the orderof 0.5–2 μm, which is adherent and cannot be scraped off easily. For the cases of slidingof polymers against rough surfaces, the abrasive mechanism may be dominant. The fatiguemechanism is important in harder polymers such as many thermoset polymers sliding againstsmooth surfaces. Asperity deformation in polymers is primarily elastic and wear due to fatigueresults from the formation of cracks associated predominantly with elastic deformation. Wearparticles are produced by the propagation and intersection of cracks.Polymers flow readily at modest pressures and temperatures. Therefore polymers and poly-

mer composites are used at relatively low loads, speed and temperatures, lower than that in thecase of metals and ceramics. Polymers generally have low thermal conductivities, thereforethey result in high interface temperatures. The interface temperatures generated during slidingare a function of normal pressure x sliding velocity (PV), thus polymers and solid lubricantsare classified based on a PV limit. Beyond the PV limit, polymers start to melt at the inter-face even at ambient temperature and wear rate increases rapidly. Methods generally used forestablishing PV limit are described by Bhushan (1999).Polymers have a high tolerance to abrasive particles (embeddability), resilience in dis-

tributing the load under misaligned conditions (thus preventing seizure), low cost and easyavailability. Polymers are generally insensitive to corrosive environments unlike metals, butreact with many fluids; they swell with degradation in mechanical properties.


Table 7.4.5 The PV limits, wear coefficients, and coefficients of friction of various unfilled and filledplastics sliding on steel, under dry conditions. (These are approximate values taken from variouspublications.)

Material

PV limits at(V) and 22◦C,MPa·m/s(at m/s)

Maximumoperatingtemperature

(◦C)Wear coefficient, k(× 10−7 mm3/Nm)

Coefficientof friction

PTFE (unfilled) 0.06 (0.5) 110–150 4000 0.05–0.1PTFE (glass-fiber filled) 0.35 (0.05-5.0) 200 1.19 0.1–0.25PTFE (graphite-fiber filled) 1.05 (5.0) 200 – 0.1Acetal 0.14 (0.5) 85–105 9.5 0.2–0.3Acetal (PTFE filled) 0.19 (0.5) – 3.8 0.15–0.27

0.09 (5.0) – – –UHMW polyethylene 0.10 (0.5) 105 – 0.15–0.3UHMW polyethylene(glass-fiber filled)

0.19 (0.5) 105 – 0.15–0.3

Polyamide 0.14 (0.5) 110 38.0 0.2–0.4Polyamide (graphite filled) 0.14 (0.5) 150 3.0 0.1–0.25Polycarbonate 0.03 (0.05) 135 480 0.35

0.01 (0.5)Polycarbonate (PTFE filled) 0.06 (0.5) 135 – 0.15Polycarbonate (PTFE, glassfiber)

1.05 (0.5) 135 5.8 0.2

Polyphenylene sulfide 3.50 (0.5) 260–315 – 0.15–0.3Polyphenylene sulfide (PTFE,carbon fibers)

3.50 (0.5) 260–315 – 0.1–0.3

Poly(amide-imide) 3.50 (0.5) 260 – 0.15–0.3Poly(amide-imide)(PTFE,graphite)

1.75 (0.5) 260 – 0.08–0.3

Linear aromatic polyester(graphite filled)

1.75 (0.5) 260–315 – 0.2–0.4

Phenolic 0.17 (0.05) 260 – 0.9–1.1Phenolic (PTFE filled) 1.38 (0.5) – – 0.1–0.45Polyimide 3.50 (0.5) 315 30.0 0.15–0.3Polyimide (graphite filled) 3.50 (0.5) 315 5.0 0.1–0.3Epoxy (glass filled) 1.75 (0.5) 260 – 0.3–0.5

7.4.3.1 Plastics

The maximum operating temperature, wear coefficients, the coefficient of friction and the PVlimits of various unfilled plastic and plastic composites are presented in Table 7.4.5. The PVlimits of elastomers are generally lower than those of plastics. The PV limit of the polymersin the lubricated conditions (oils or water) can be up to an order of magnitude larger than thatin dry conditions. A liquid medium removes frictional heat from the interface, thus allowingoperation at high PV conditions. High-temperature polymers can be operated under lubricatedconditions with a PV of 17.5 MPa × m/s (500,000 psi × fpm), comparable to the PV limit ofcarbon-graphites (manufactured carbon), commonly used in wear applications.

Wear 385

Figure 7.4.11 Effect of carbon fiber orientation on wear volume of polymer in reinforced polyesterresin (25 wt. % fiber) slid against hardened tool steel. The coefficients of friction are given adjacentto each curve. Reproduced with permission from Lancaster, J.K. (1968), “The Effect of Carbon Fiber-Reinforcement on the Friction andWear Behavior of Polymers,” Br. J. Appl. Phys. 1, 549–559. Copyright1968. IOP Publishing.

In polymer composites, orientation of the fibers affects the wear rates. Figure 7.4.11 il-lustrates the effect of fiber orientation on carbon-fiber-reinforced polyester sliding against arelatively smooth hardened tool steel surface. Both the coefficient of friction and the wear rateof the polymers are lower when the fibers are oriented normal to the sliding surface. Certainfillers, such as glass and carbon, commonly used in polymer composites are harder than themating metals, such as mild steel and cast iron, and thus may cause damage to the mating metalsurface. Cumulative damage caused by the composite can be attributed to the abrasiveness ofthe particular filler used and is an important factor to consider when selecting a material for


Figure 7.4.12 (a) Wear rate as a function of surface roughness of PTFE filled with 10% glass fiberplus 15 wt. % CdO-graphite-Ag slid against nitrided SAE 7140 steel in reciprocating mode in a nitrogenatmosphere, (b) initial wear versus roughness of same combination in nitrogen atmosphere (R, random;L, longitudinal; andT, transverse orientation of roughness of thematingmetal with respect to the directionof sliding). Reproduced with permission from Bhushan, B. and Wilcock, D.F. (1982), “Wear Behaviorof Polymeric Compositons in Dry Reciprocating Sliding,” Wear 75, 41–70. Copyright 1982. Elsevier.

use against a relatively soft metal such as aluminum. The hardness of the mating metal playsa role in determining abrasiveness of a particular filler.The surface roughness of the mating metal and its orientation in relation to the direction of

sliding have a significant influence on wear rate, Figure 7.4.12 (Bhushan and Wilcock, 1982).The rougher mating metal surface results in a thicker transfer film buildup, which may beresponsible for the lower friction. Initial wear is high with high roughness because it takesmore polymer material to pack the roughness grooves. Once the adherent transfer film is builtup, the wear rate decreases and the subsequent wear rate of a rougher surface is lower than

Wear 387

that of a smoother surface. Transverse grooves have more initial and steady-state wear thanlongitudinal grooves owing to a more abrading (cutting) action. Mating metals with randomroughness show high initial wear and longer time for buildup of the transfer film. Therefore, amating surface with a relatively high surface roughness (∼0.8 μm) and roughness grooves onthe metal surface in the direction of sliding are recommended.

7.4.3.2 Elastomers

Elastomers are generally used at PV limits lower than that for many plastics. Friction andwear properties of elastomers are modified by adding fillers such as carbon black, silica,graphite, MoS2, and PTFE powders and sometimes glass fibers. Friction and wear of elastomercomposites can be comparable to the plastic composites (Bhushan and Winn, 1981).

7.4.3.3 Effect of Operating Environment

Exposure to environment (gases and humidities) affects mechanical properties and frictionand wear of polymers. In the data shown in Figure 7.4.13, the wear rate of PTFE compositesliding against cast iron at 70◦C, exposed to different environments decreases with an in-crease of environmental humidity (Schubert, 1971). The exact trend depends on the particularenvironment. The wear rate in air and oxygen is up to 1000 times greater than in nitrogen.

Figure 7.4.13 Wear rate of PTFE composite sliding against cast iron at 70◦C as a function of waterconcentration in three different environments.


7.5 Closure

Wear is the surface damage or removal of material from one or both of two solid surfaces in asliding, rolling or impact motion relative to one another. Wear damage precedes actual loss ofmaterial, and it may also occur independently. The definition of wear is generally based on lossof material from one or both of the mating surfaces. Strictly, the wear-like friction is not aninherent material property and it depends on the operating conditions and surface conditions.Wear rate does not necessarily relate to friction. Wear resistance of a material pair is generallyclassified based on a wear coefficient, a non-dimensional parameter or wear volume per unitload per unit sliding distance.Wear occurs bymechanical and/or chemical means and is generally accelerated by frictional

heating. Principal types of wear mechanism include: (1) adhesive; (2) abrasive; (3) fatigue; (4)impact by erosion and percussion; (5) chemical; and (6) electrical-arc-induced wear. Other, notdistinct, mechanisms are fretting and fretting corrosion, a combination of adhesive, corrosive,and abrasive forms of wear. Wear by all mechanisms, except by fatigue mechanism, occurs bygradual removal of material. Of the aforementioned wear mechanisms, one or more may beoperating in one particular machine. In many cases, wear is initiated by one mechanism and itmay proceed by other wear mechanisms, thereby complicating failure analysis.Adhesive wear occurs because of adhesion at asperity contacts at the interface. These

contacts are sheared by sliding which may result in the detachment of a fragment from onesurface to another surface. As the sliding continues, the transferred fragments may come offthe surface on which they are transferred and be transferred back to the original surface, orelse form loose particles. Some are fractured by a fatigue process during repeated loading andunloading action resulting in formation of loose particles. During sliding, surface asperitieson or near undergo plastic deformation and/or fracture. The subsurface, up to several micronsin thickness also undergoes plastic deformation and strain hardening with microhardness asmuch as factor of two higher than the bulk hardness. Based on Archard’s equation, the volumeof wear of contacts going through plastic deformation is proportional to the normal load andsliding distance and is inversely proportional to the hardness of the surface being worn away.Based on Bhushan’s equation, the volume of wear of contacts going through primarily elasticdeformations is proportional to the normal load and sliding distance and inversely proportionalto the composite modulus of elasticity and roughness parameters ratio. Wear equations suggestthat the wear coefficient is independent of normal load and sliding velocity, but this assumptionholds only for a range of values of loads and velocities.Abrasive wear occurs when the asperities of a rough, hard surface or hard particles slide

on a softer surface, and damage the interface by plastic deformation or fracture in the case ofductile and brittle materials, respectively. In many cases, there are two general situations forabrasive wear. In the first case, the hard surface is the harder of two rubbing surfaces (two-bodyabrasion); and in the second case, the hard surface is a third body, generally a small particle ofabrasive, caught between the two surfaces and sufficiently harder that it is able to abrade eitherone or both of the mating surfaces (three-body abrasion). In many cases, the wear mechanismat the start is adhesive, which generates wear particles that get trapped at the interface, resultingin a three-body abrasive wear. In most abrasive wear situations, scratching is observed witha series of grooves parallel to the direction of sliding. During sliding, like adhesive wear,asperities on or near the surface undergo plastic deformation and strain hardening with anincrease in hardness. Abrasive wear rate is a function of surface roughness and, in contrast

Wear 389

to adhesive wear mechanism, it increases with an increase in surface roughness. The wearequation for two-body abrasive wear is also valid for three-body abrasive wear. However, thewear rate will be lower by about an order of magnitude because many particles will tend toroll rather than slide. If the wear takes place with a fresh abrasive medium, wear continuesat a steady rate, whereas, if a limited amount of abrasive medium is used as the slidingcontinues, the wear rate generally decreases as a function of time. A decrease in wear rate asa function of time is believed to occur primarily as a result of blunting of the abrasives. Therelative hardness of the abrasive medium to the workpiece affects the wear rate. When thehardness ratio of the workpiece to the abrasive particles is less than unity, the wear coefficientremains approximately constant; however, if the ratio is equal to or greater than unity, the wearcoefficient decreases rapidly with an increase in the hardness ratio.Subsurface and surface fatigue are observed during repeated rolling (negligible friction)

and sliding (coefficient of friction ≥ 0.3), respectively. The repeated loading and unloadingcycles to which the materials are exposed may induce the formation of subsurface and surfacecracks, which eventually, after a critical number of cycles will result in the breakup of thesurface with the formation of large fragments, leaving large pits on the surface. Prior to thiscritical point, negligible wear takes place, which is in marked contrast to the wear caused bythe adhesive or abrasive wear mechanism, where wear causes a gradual deterioration from thestart of running. Therefore, the material removed by fatigue wear is not a useful parameter.Much more relevant is the useful life in terms of the number of revolutions or time beforefailure occurs. Another difference between adhesive and abrasive wear and fatigue wear is thatfatigue wear does not require direct physical contact between two surfaces. Mating surfacesexperience large stresses, transmitted through the lubricating film during the rolling motionsuch as in well-designed rolling element bearings. The failure time in fatigue wear is statisticalin nature and is predicted based on Weibull analysis in terms of probability of survival.Chemical-induced crack growth (most common in ceramics) is commonly referred to as

static fatigue. In the presence of tensile stresses and water vapor at the crack tip in manyceramics, a chemically induced rupture of the crack-tip bonds occurs rapidly, which increasesthe crack velocity. Chemically enhanced deformation and fracture result in an increased wearof surface layers in static and dynamic (rolling and sliding) conditions.Impact wear includes erosive and percussive wear. Erosion can occur by jets and streams

of solid particles, liquid droplets, and implosion of bubbles formed in the fluid. Percussionoccurs from repetitive solid body impacts. Repeated impacts result in progressive loss of solidmaterial. Solid particle erosion is a form of abrasion that is generally treated rather differentlybecause the contact stress arises from the kinetic energy of particles flowing in an air or liquidstream as it encounters a surface. The particle velocity and impact angle combined with thesize of the abrasive give a measure of the kinetic energy of the impinging particles, that is, ofthe square of the velocity. As in the abrasive wear, erosive wear occurs by plastic deformationand/or fracture, dependent upon material being eroded away and operating parameters. Inliquid impingement erosion, with small drops of liquid striking the surface of a solid at highspeeds (as low as 300 m/s), very high pressures are experienced, exceeding the yield strengthof most materials. Thus, plastic deformation or fracture can result from a single impact, andrepeated impact leads to pitting and erosive wear. Cavitation erosion arises when a solid andfluid are in relative motion, and bubbles formed in the fluid become unstable and implodeagainst the surface of the solid. Cavitation erosion is similar to surface fatigue wear. Percussionis a repetitive solid body impact. Percussion wear occurs by hybrid wear mechanisms which


combine several of the following mechanisms: adhesive, abrasive, surface fatigue, fractureand tribochemical wear.Chemical or corrosive wear occurs when sliding takes place in a corrosive environment.

Corrosion can occur because of a chemical or electrochemical interaction of the interfacewith the environment. In air, the most corrosive medium is oxygen. Therefore, chemical wearin air is generally called oxidative wear. In the absence of sliding, the chemical productsof the corrosion (e.g., oxides) would form a film typically less than a micrometer thick onthe surfaces, which would tend to slow down or even arrest the corrosion, but the slidingaction wears the chemical films away, so that the chemical attack can continue. Thus chemicalwear requires both chemical reaction (corrosion) and rubbing. Frictional heating modifies thekinetics of chemical reactions of sliding bodies with each other, and with the gaseous or liquidenvironment, to the extent that reactions which normally occur at high temperatures occur atmoderate or even ambient temperatures during sliding. The wear controlled by this reaction isreferred to as tribochemical wear.When a high potential is present over a thin air film in a sliding process, a dielectric

breakdown results, leading to arcing. During arcing, a relatively high-power density occursover a very short period of time. The heat-affected zone is usually very shallow (on the orderof 50 μm) and the heating results in considerable melting and subsequent resolidification,corrosion, hardness changes, and other phase changes, and even in the direct ablation ofmaterial. Arcing causes large craters, and any sliding or oscillation after an arc either shearsor fractures the lips, leading to three-body abrasion, corrosion, surface fatigue, and fretting.Fretting occurs where low-amplitude oscillatory motion (a few tens of nanometers to a few

tens of microns) takes place between contacting surfaces, which are nominally at rest. A rapidincrease in wear rate occurs with slip amplitude over an amplitude range. Basically, fretting isa form of adhesive or abrasive wear, where the normal load causes adhesion between asperitiesand oscillatory movement causes ruptures, resulting in wear debris. Most commonly, frettingis combined with corrosion, in which case the wear mode is known as fretting corrosion.Regarding the particles present in wear debris, these are generally classified based on their

morphology: plate-shaped; ribbon-shaped’ spherical; and irregular-shaped.Finally, wear of a material is dependent on the mating material (or material pair), surface

preparation and operating conditions. Clean metals and alloys exhibit high adhesion, andconsequently high friction and wear. Any contamination mitigates contact, and chemicallyproduced films which reduce adhesion result in reduction in friction and wear. In dry sliding,identical metals, particularly iron on iron, are metallurgically compatible and exhibit highfriction and wear, so they must be avoided. Soft and ductile metals such as In, Pb, andSn exhibit high friction and wear. Hexagonal metals such as Co and Mg as well as somenon-hexagonal metals such as Mo and Cr exhibit low friction and wear. Lead-based whitemetals (babbitts), brass, bronze, and gray cast iron generally exhibit relatively low friction andwear, and are commonly used in dry and lubricated bearing and seal applications. For high-temperature applications, cobalt-based alloys are used which exhibit good galling resistance.(Galling resistance is a measure of the normal stress at which twomaterials loaded against eachother gall or weld.) Nickel-based alloys are poor in unlubricated sliding because of generallycatastrophic galling.In dry sliding conditions, similar or dissimilar ceramic pairs are commonly used which

exhibit moderate friction but maximum wear resistance. In ceramics, fracture toughness is animportant mechanical property which affects friction. Ceramics react with the humidity from

Wear 391

the environment. Non-oxides may form beneficial hydrides (from the tribochemical reaction)and result in low friction and wear. On the other hand, oxide ceramics, because of enhancedcrack growth at high humidity (static fatigue) result in high friction and wear. In both metalsand ceramics, no single wear mechanism operates over a wide range of operating conditions.Various regimes of mechanical and chemical wear for a sliding material pair are presented ina wear-regime map, plotted on axes of Hertzian pressure and sliding velocity.Polymers, which include plastics and elastomers, generally exhibit very low friction and

moderate wear. Among polymers, PTFE exhibits the lowest friction and low wear. Polymersflow at modest pressures and modest temperatures; therefore, polymer composites are com-monly used. Polymers react with fluids in the environment and can swell and lose mechanicalproperties. Since polymers soften at moderate temperatures, they are classified based on thePV limit, which is a measure of the interface temperature rise.

Problems

7.1 A cylindrical bronze pin of 1 mm radius rests on a rotating steel disk at a mean radius of25 mm. The normal load on the pin is 10 N. The rotational speed of the disk is 300 rpmand the test lasts for 10 hours. The mass losses of the pin and disk are 50 mg and 3 mg,respectively. Using the material data given below, calculate the wear coefficients andwear depths for the bronze pin and steel disk. (Hardness of bronze = 0.8 GPa, densityof bronze = 8.5 Mg/m3; hardness of steel = 2.5 GPa, density of steel = 7.8 Mg/m3).Calculate the wear coefficients for a cylindrical steel pin on a bronze disk under thesame test conditions.

7.2 A milling cutter was used to saw through a medium carbon steel bar (H = 3 GPa) of10 mm diameter with a width of cut of 0.5 mm. It took 10 minutes to saw and the energyexpended was 50 W (Nm/s). The coefficient of friction between the saw and the steelbar is 0.3. Calculate the wear coefficient of the steel bar during the cutting process.

7.3 A steel surface consisting of conical asperities with roughness angle of 10◦, reciprocateson a soft lead surface (H= 75 MPa) under a load of 1 N with a reciprocating amplitudeof 10 mm at 5 Hz. Given that the volume of lead material removed is 10−6 m3 in10 hours, calculate the abrasive wear coefficient of the lead material. Given that theroughness angle of the steel surface is 30◦, calculate the wear coefficient of the leadmaterial.

7.4 A nickel surface (hardness= 3 GPa) of a square block (10 mm× 10 mm) electroplatedwith ruthenium (hardness = 5 GPa) to a thickness of 5 μm is rubbed at a normalload of 5 N against an abrasive paper so that fresh abrasive paper always contactsthe ruthenium. Estimate what distance of rubbing is needed before the abrasive firstpenetrates the ruthenium, thus exposing the nickel substrate. Assume that kabr forruthenium rubbing against the abrasive paper is 10−6.

7.5 A cubic pin with a linear dimension of 1mm andwith hardness H of 0.2 GPa slides upona surface at a constant velocity V of 0.1 m/s and apparent pressure pa of 0.001 GPa. Thewear coefficient k is = 4 × 10−6. The failure occurs when the fraction of the volumeof 0.1% is worn. Calculate sliding time until failure.

7.6 A body slides upon another body with a plastic contact. For a distance× of 1000 m, thevolume v of 1 mm3 is worn during sliding and the wear coefficient k is 10−6. Calculatefriction force, if adhesion shear strength τ a is equal to 106 Pa.


7.7 In grinding of silicon carbide, the material removal rate by brittle fracture is 2 μm/h.Silicon carbide is processed to increase its fracture toughness by a factor of 1.5 withthe same hardness and modulus of elasticity. Calculate the material removal rate for theprocessed silicon carbide.

7.8 Based on a tribological test of various materials, the following values of coefficient offriction and wear coefficient were reported for graphite-steel, steel-steel, steel-bronze,and alumina-alumina. Enter name of these material pairs in Table P.7.1.

Table P.7.1

Material pair Coefficent of friction Wear coefficient

? 0.2 10−6

? 0.1 10−4

? 0.3 10−9

? 0.6 10−2

7.9 Using standard AFBMA calculations, the life, L10, of a roller bearing is 1000 h. Giventhat the material factor is 2.2, the processing factor is 3, and the lubrication factor is 4,calculate the expected bearing life.

7.10 The basic load rating of a roller bearing based upon the AFBMA calculations is 10 kN.Calculate the bearing catalog life for an applied radial load of 1 kN and shaft speed of6000 rpm.

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Further Reading

Bartenev, G.M. and Lavrentev, V.V. (1981), Friction and Wear of Polymers, Elsevier, Amsterdam.Bayer, R.G. (1994), Mechanical Wear Prediction and Prevention, Marcel Dekker, New York.Bhushan, B. (1996), Tribology and Mechanics of Magnetic Storage Devices, Second edition, Springer-Verlag, New

York.Bhushan, B. (2001a),Modern Tribology Handbook Vol. 1: Principles of Tribology, CRC Press, Boca Raton, Florida.Bhushan, B. (2001b), Fundamentals of Tribology and Bridging the Gap Between the Macro- and Micro/Nanoscales,

NATO Science Series II-Vol. 10, Kluwer, Dordrecht, The Netherlands.Bhushan, B. (2011), Nanotribology and Nanomechanics I – Measurement Techniques and Nanomechanics, II –

Nanotribology, Biomimetics, and Industrial Applications, Third edition, Springer-Verlag, Heidelberg, Germany.Bhushan, B. and Gupta, B.K. (1997),Handbook of Tribology: Materials, Coatings and Surface Treatments, McGraw-

Hill, New York (1991); reprinted by Krieger, Malabar, Florida (1997).

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Blau, P.J. (1992), ASM Handbook, Vol. 18: Friction, Lubrication, and Wear Technology, Tenth edition, ASM Interna-tional, Materials Park, Ohio.

Bruce, R.W. (2012),Handbook of Lubrication and Tribology, Vol. II: Theory and Design, Second edition, CRC Press,Boca Raton, Florida.

Buckley, D.H. (1981), Surface Effects in Adhesion, Friction, Wear, and Lubrication, Elsevier, Amsterdam.Hutchings, I.M. (1992), Tribology: Friction and Wear of Engineering Materials, CRC Press, Boca Raton, Florida.Kragelski, I.V. (1965), Friction and Wear, Butterworths, London.Loomis, W.R. (ed) (1985), New Directions in Lubrication, Materials, Wear, and Surface Interactions: Tribology in

the 80s, Noyes Publications, Park Ridge, New Jersey.Peterson, M.B. and Winer, W.O. (eds) (1980), Wear Control Handbook, ASME, New York.Rabinowicz, E. (1995), Friction and Wear of Materials, Second edition, Wiley, New York.Rigney, D.A. (ed) (1981), Fundamentals of Friction and Wear of Materials, Amer. Soc. Metals, Metals Park, Ohio.Rigney, D.A. and Glaeser, W.A. (eds) (1978), Source Book on Wear Control Technology, Amer. Soc. Metals, Metal

Park, Ohio.Scott, D. (ed) (1979), Wear Treatise on Materials Science and Technology, Vol. 13, Academic Press, San Diego,

California.Shipley, R.J. andBecker,W.T. (2002),Metals Handbook, Vol. 11: Failure Analysis and Prevention, ASM International,

Metals Park, Ohio.Suh, N.P. (1986), Tribophysics, Prentice-Hall, Englewood, New Jersey.Suh, N.P. and Saka, N. (1980), Fundamentals of Tribology, MIT Press, Cambridge, Massachusetts.Zum Gahr, K.H. (1987),Microstructure and Wear of Materials, Elsevier, Amsterdam.

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