23Tribological Properties

of Metallic andCeramic Coatings

23.1 Introduction23.2 Tribology of Coated Surfaces

Contact Mechanisms and Stresses in the Surface • Tribological Contact Mechanisms in Coated Surfaces • Macromechanical Friction and Wear Mechanisms • Micromechanical Tribological Mechanisms • Tribochemical Mechanisms of Coated Surfaces • Nanophysical Contact Mechanisms

23.3 Macromechanical Interactions: Hardness and GeometrySoft Coating on Hard Substrate • Hard Coating on Softer Substrate • Load-Carrying Capacity of Coated Surfaces

23.4 Micromechanical Interactions: Material ResponseMaterial Response • Energy Accommodation in the Surface • The Influence of Fracture Toughness in Material Response

23.5 Material Removal and Change Interactions: Debris and Surface LayersSurface and Subsurface Cracks • Tribological Influence of Wear Debris • Debris Agglomeration • Transfer Layers • Reaction Layers

23.6 Multicomponent CoatingsMultilayer Coatings • Nanocomposite Coatings • Functionally Graded Coatings • Adaptive Coatings

23.7 Concluding Remarks

23.1 Introduction

The physical and chemical phenomena that govern both the friction and wear in moving contacts takeplace at the surfaces of the solids in relative motion. Usually, the interaction mechanisms at the outermostsurfaces are the most crucial for tribological performance.

The introduction of a fluid between the surfaces has long been the most common way to change thetribological behavior of two surfaces moving relative to each other. Another increasingly popular possi-bility is to apply a thin surface layer or coating on one or both of the surfaces. This approach is promotedby the new coating techniques that have been developed over the last few decades. In particular, physicalvapor deposition (PVD) and chemical vapor deposition (CVD) offer considerable possibilities to tailorthe thin surface layers with regard to their material and structure as well as their mechanical and chemicalproperties. These deposition techniques are presented in Chapter 24.

Kenneth HolmbergVTT Manufacturing Technology

Allan MatthewsThe University of Hull

The first major commercial application of tribological thin surface coatings was the use of titaniumnitride and titanium carbide coatings on cutting tools. Other early applications were the use of leadcoatings for dry lubrication of rolling bearings in near-vacuum conditions in space, and gold coatingsfor electrical contacts. Following the early rapid adoption of advanced coatings on cutting tools, therehas been a gradual but accelerating interest in the use of vapor deposition technologies in other appli-cations sectors (e.g., automotive and aerospace) and in other branches of manufacturing industry (e.g.,food processing) (Leyland and Matthews, 1994; Cselle and Barimani, 1995; Bull and Jones, 1996; Zabinskiet al., 1996; Enomoto and Yamamoto, 1998; Theunissen, 1998; Verkammen et al., 1999).

Two issues that have constrained their wider adoption include their relatively higher cost and concernover repeatability of properties such as thickness, composition, hardness, and adhesion. Such reliabilityissues are now being solved through improvements in process monitoring and control. Also, develop-ments in continuous and semicontinuous high-throughput coating equipment have reduced the coatingcost per component and opened up high-volume markets such as automotive parts. These includecomponents used in fuel injection systems, air conditioning units, engines and even transmission systemsin high-performance vehicles. Thus, the nature of the tribological contacts to be protected by the coatingsis becoming ever more wide-ranging and demanding. Hence, there is a need for a full and systematicunderstanding of both the mechanisms occurring and the response behavior of coatings to these chal-lenging applications.

Today, there is a huge variety of metallic and ceramic coatings available and used for tribologicalapplications, as well as soft polymer coatings, lamellar coatings such as molybdenum disulfide, andextremely hard coatings such as diamond and cubic boron nitride that are receiving research attention.This chapter focuses on the tribological behavior and properties of metallic and ceramic coatings, as theothers are covered in elsewhere in this book.

Another constraint applied is to focus on thin tribological coatings, that is, those that are sufficientlythin that the substrate material plays a role in determining the friction and wear performance. Thus,coatings that are so thick that there is little or no substrate influence on the tribological behavior — thecoating in effect acts as a bulk material — are excluded. The tribology of bulk materials is presented inSection I of this Handbook.

This chapter primarily concerns the tribology of dry contacts. The addition of a lubricant will, inmany cases, lead to a lessening of the severity of the tribological mechanisms, but it may also change thefriction and wear mechanisms involved, for example, through its influence on the formation of surfacelayers. This aspect has received little attention in the literature.

The authors’ purpose is to review the present understanding of tribological mechanisms of metallic-and ceramic-coated surfaces, as well as to deepen the understanding of the crucial micromechanicalinteractions by introducing a new energy accommodation concept that includes elastic, plastic, andfracture mechanics considerations. The approach is to try to explain in qualitative terms the tribologicalmechanisms and interactions, as it has been difficult to produce precise quantitative generic data, dueto the large scatter in reported experimental methods and conditions in the literature.

This chapter focuses primarily on the most recent literature published, as the earlier works can befound in previous reviews of the subject, such as Holmberg and Matthews (1994) and Holmberg et al.(1998).

23.2 Tribology of Coated Surfaces

23.2.1 Contact Mechanisms and Stresses in the Surface

The phenomena that take place in a tribological contact are influenced by the force pressing the twosurfaces together. Calculation methods for the stress fields and deformations in a coated surface havebeen reviewed in the references cited above.

A useful approach to calculating the coating/substrate interface stresses has been used by Ramalingamand Zheng (1995) to assess the problems that can be encountered when hard coatings are applied on

substrates with lower or higher compliance than the coating. Numerical solutions are obtained fordifferent stress distributions on the coated surface. They show how tensile stresses in the wake of thecontact tend to separate the coating from the substrate, and how this can be solved by adjusting the filmthickness and choice of coating conditions to ensure that the coating-to-substrate adhesion strength canwithstand the stresses. The calculations were later extended to also include interface stresses and multilayercoatings (Zheng and Ramalingam, 1995, 1996).

In the analysis of interfacial stresses, it is necessary to include the influence of parameters such ascoating thickness, interface topography, and elastic mismatch, as well as the effects induced by pores,edges, or scratches (Wiklund et al., 1999). Coatings that are thin compared to the interface topographyhave been found to be less sensitive to residual stress-induced failure. At a critical coating thickness, thestress across the interface can be of a magnitude sufficient to initiate coating delamination. However,residual stresses can also be beneficial; for example, the influence of high compressive stresses in hardand smooth diamond coatings has been studied by Gunnars and Alaheliste (1996) and Gåhlin et al.(1996). They found that highly stressed coatings correspond with smoother surfaces and wear rates,which are only 5 to 20% of those of the stress-free coatings. The internal stresses were shown to influencethe crack propagation direction in diamond coatings. For other coatings, such as TiN, it should be notedthat the process parameters that control film stresses (e.g., ion current density, temperature, and biasvoltage) also influence other properties such as crystallographic structure, density, and even the texture.It is not usually possible to separate out the specific effect of the residual stress on performance. Never-theless, it is generally believed that the residual compressive stresses in the plane of the coatings, whichare usually produced by plasma-assisted PVD methods, are beneficial to tribological performance.

The surface profiles, both undeformed and deformed, the contact pressures, and the von Mises stressesbeneath the surface for a rough sphere in contact with a smooth plane are shown in Figure 23.1. Finite-element methods have been applied to evaluate the stress field in the hard coating and the substrateunder frictional loads (Diao et al., 1994a; Wong and Kapoor, 1996; Wong et al., 1997; Bouzakis et al.,1998; Lowell, 1998; Njiwa et al., 1998a,b). Diao and Kato analyzed the von Mises stress distributions inhard coatings and in elastic sliding (Diao and Kato, 1994a,b; Diao et al., 1994a). An elliptical distributionof normal and traction contact pressures was assumed for the analysis of von Mises stress for differentcoating thicknesses, friction coefficients, and elastic moduli of the coating and the substrate. The positionof yield could be calculated and they introduced a local yield map showing the yield strength ratio inrelation to the ratio of the coating thickness to the contact half-width. The local yield maps showed thatyield at the coating-substrate interface on the substrate side is the most common case under a wide rangeof contact conditions.

The surface roughness is an important parameter influencing the contact stresses, especially at thecoating-substrate interface. This was first shown by Sainsot et al. (1990), who also concluded that in hardcoatings with thicknesses less than 15 µm on softer substrates, the maximum von Mises stresses both inthe coating and in the substrate will be located just at their interface. This clearly confirms the importanceof analyzing the stresses at the coating-substrate interface and comparing them to the coating adhesionstrength. In rolling contact fatigue situations (e.g., rolling bearings), the surface roughness will result insmall-scale contact stress spikes that initiate cracks and result in failure (Polonsky et al., 1997; Néliaset al., 1999).

A numerical technique, including integral transform and finite-element methods for solving contactproblems in layered rough surfaces, has been developed by Mao et al. (1994; 1995) and Bell et al. (1998).They apply their numerical model for the two-dimensional dry sliding contact of two elastic bodies onreal rough surfaces, where an elastic body contacts with a multilayer surface under both normal andtangential forces. Of special interest is that their model uses surface profile data directly recorded with astylus-measuring instrument. They analyze the contact pressure distribution for layers with differentcoefficients of friction, thicknesses, and elastic moduli.

The magnitude of internal stresses can vary but, for example, for TiN coatings on steel deposited bydifferent PVD techniques, they are compressive. Hardness and bonding strength have an obvious effecton internal stresses in TiN coatings (Daming et al., 1998) and this influences their tribological behavior.

FEM modeling of a TiN-coated steel substrate subject to rolling with a surface crack perpendicular tothe surface shows that the stress intensities at the surface increased with layer thickness and crack length(Eberhardt and Kim, 1998). Crack face friction reduced the stress intensity more significantly for thickerlayers and longer cracks, and for higher coating/substrate stiffness ratios.

23.2.2 Tribological Contact Mechanisms in Coated Surfaces

The tribology of a contact involving surfaces in relative motion can be understood as a process withcertain input and output data (Holmberg and Matthews, 1994). Input data that are used as a starting

FIGURE 23.1 (a) Profiles of a rough sphere, undeformed, and deformed loaded against a plane surface.(b) Normalized contact pressures for a smooth and a rough sphere loaded against a plane surface. (c) Lines of constantnormalized von Mises stresses for a rough sphere loaded against a smooth plane surface with no friction. (a = Hertziancontact radius, h = height, z = depth, p = pressure, po = maximum Hertzian contact pressure.) (Adapted from Lee,S.C. and Ren, N. (1994), The subsurface stress field created by three-dimensionally rough bodies in contact withtraction, Tribol. Trans., 37, 615-621.)

point for the analysis of a tribological contact include the geometry of the contact, both on a macro-and microscale, the material properties based on the chemical composition and structure of the differentparts involved, and the environmental parameters, as shown in Figure 23.2. Other input data include theenergy parameters such as normal load, velocity, tangential force, and temperature.

The tribological process takes place as the two surfaces are moving in relation to each other, and bothphysical and chemical changes occur in accordance with the physical and chemical laws with respect tothe input data. As a function of time, the tribological process causes changes in both the geometry andthe material composition and results in energy-related output effects: friction, wear, velocity, temperature,sound, and dynamic behavior.

The complete tribological process in a contact in which two surfaces are in relative motion is verycomplex because it simultaneously involves friction, wear, and deformation mechanisms at different scalelevels and of different types. To achieve a holistic understanding of the complete tribological processtaking place and to understand the interactions, it is useful to separately analyze the tribological changesof four different types: the macro- and microscale mechanical effects, the chemical effects, and thematerial transfer taking place, as shown in Figure 23.3. In addition, there has recently been an increasinginterest in studying tribological behavior on a molecular level; that is, nanophysical effects (Bhushan,1995, 1999).

A better and more systematic understanding of the mechanisms involved in a tribological contact isnecessary for the optimization of the properties of the two contacting surfaces in order to achieve therequired friction and wear performance. Approaches to the tribological optimization of surfaces havebeen presented by Matthews et al. (1993) and Franklin and Dijkman (1995), who introduce eight weardesign rules in an expert system for assisting the selection of metallic materials, surface treatments, andcoatings during the initial stages of engineering design.

23.2.3 Macromechanical Friction and Wear Mechanisms

The macromechanical tribological mechanisms describe the friction and wear phenomena and are influ-enced by the stress and strain distribution in the entire contact, the total elastic and plastic deformationsthey result in, and the total wear particle formation process and its dynamics. In contacts between twosurfaces of which one or both are coated, four main parameters can be defined that control the tribologicalcontact behavior. They are:

FIGURE 23.2 The tribological process in a contact between two surfaces includes mechanical and tribochemicalchanges as well as material transfer. The changes are determined by material and energy input parameters and resultin a similar set of partly changed output parameters.

1. coating-to-substrate hardness relationship2. thickness of the coating3. surface roughness4. size and hardness of any debris in the contact, which may originate from external sources or be

produced by the surface wear interactions themselves

The relationship between these four parameters will result in a number of different contact conditionscharacterized by specific tribological contact mechanisms. Figure 23.4 shows schematically 12 such verytypical tribological contacts, with different mechanisms influencing friction, when a hard spherical slidermoves on a coated flat surface (Holmberg, 1992). The corresponding wear mechanisms have beendescribed in a similar way (Holmberg et al., 1993). The macromechanical influencing parameters arediscussed in more detail in Section 23.3.

23.2.4 Micromechanical Tribological Mechanisms

The origins of the friction and wear phenomena observed on the macrolevel are found in the mechanismsthat take place at the microlevel. The micromechanical tribological mechanisms describe the stress andstrain formation at an asperity-to-asperity level, the crack generation and propagation, material libera-tion, and particle formation, as shown in Figure 23.3(c). In typical engineering contacts, these phenomenaare at a size level of about 1 µm or less, down to the nanometer range.

FIGURE 23.3 Tribological contact mechanisms: (a) macromechanical, (b) material transfer, (c) micromechanical,(d) tribochemical, and (e) nanophysical.

Shear and fracture are two basic mechanisms for the first nucleation of a crack and for its propagationuntil it results in material liberation and formation of a wear scar and a wear particle. These mechanismshave been discussed by, for example, Argon (1980) and Suh (1986), but still today there is only a poorunderstanding of these quite fundamental phenomena. Another approach is to study the tribologicalmicromechanical mechanisms using the velocity accommodation concept developed by Berthier et al.(1989). This chapter presents a new energy accommodation approach to the micromechanical tribologicalmechanisms, which is described in Section 23.4.

23.2.5 Tribochemical Mechanisms of Coated Surfaces

The chemical reactions taking place at the surfaces during sliding contact, and also during the periodsbetween repeated contacts, change the composition of the outermost surface layer and its mechanicalproperties. This has a considerable influence on both friction and wear because they are determined toa great extent by the properties of the surface, where phenomena such as shear, cracking, and asperityploughing take place (Gee and Jennett, 1995). The chemical reactions on the surfaces are stronglyinfluenced by the high local pressures and the flash temperatures, which can be over 1000°C, at spotswhere asperities collide.

Very low coefficients of friction, down to 0.1, have been reported for a hard titanium nitride coatingsliding against itself (Mäkelä and Valli, 1985) and even lower values, down to about 0.01 but more typically0.05, have been measured for diamond-like carbon (DLC) coatings sliding against different counterfacematerials (Donnet, 1995; Holmberg et al., 1994; Donnet et al., 1994; Donnet, 1996) and diamond coatingssliding against diamond and ceramics (Erdemir et al., 1996b; Hayward et al., 1992; Zuiker et al., 1995;Habig, 1995; Gardos, 1994). This can be explained by the formation of low-shear microfilms on the hardcoating or perhaps only on the asperity tips of the coating.

Thus, if one considers such a contact on a microscale, there is effectively a soft coating on a hardsubstrate (see Figure 23.3(d)), although now the coating (e.g., TiN or diamond) plays the role of hardsubstrate, and the soft microfilm formed plays the role of a coating. It is obviously advantageous if thesubstrate under the hard coating is as hard as possible, to avoid fracture of the brittle coating bydeformation, to improve the load support, and to decrease the real area of contact. The very lowcoefficients of friction of polished diamond and diamond-like coatings are further explained by the

FIGURE 23.4 Macromechanical contact conditions for different mechanisms that influence friction when a hardspherical slider moves on a coated flat surface

extreme smoothness of the surface, excluding effects such as interlocking and asperity ploughing, as wellas the hard coating reducing the ploughing component of friction.

The chemical reactions and the chemically formed reaction layers are treated in more detail inSection 23.5.5.

23.2.6 Nanophysical Contact Mechanisms

Emerging technologies such as atomic force microscopy and other surface force methods (Israelachvili andTabor, 1972; Bhushan, 1999) have opened the possibility to study friction and wear phenomena on amolecular scale and to measure frictional forces between contacting molecules at the nano-Newton level.Increased computational power has made it possible to study friction and associated phenomena by molec-ular dynamic simulations of sliding surfaces and to investigate the atomic-scale contact mechanisms, asshown in Figure 23.3(e). Only some aspects of these complex nanophysical phenomena have thus far beeninvestigated; an example is the friction that arises from slippage between solid-to-solid interfaces (Thompsonand Robbins, 1989) and between closely packed films in sliding contact (McClelland and Glosli, 1992). Theatomic-scale mechanisms of friction when two hydrogen-terminated diamond surfaces are in sliding contacthas been studied and the dependence of the coefficient of friction on load, crystallographic sliding direction,and roughness have been investigated (Harrison et al., 1992, 1993). Work on molecular-scale viscoelasticeffects and viscous flow has been reported (Wahl and Unertl, 1998; Zhang and Tanaka, 1998). The presentunderstanding of nanoscale tribology is presented in Section II of this Handbook.

The increased understanding of the origin of friction at the atomic scale and even why friction existshas resulted in an examination of the relationship between the commonly used laws of friction at amacroscale and the molecular frictional behavior on a nanoscale. There have been suggestions that frictionarises from atomic lattice vibrations occurring when atoms close to one surface are set into motion bythe sliding action of atoms on the opposing surface. Thus, some of the mechanical energy needed toslide one surface over the other would be converted to sound energy, which is then eventually transformedinto heat (Krim, 1996).

In an interesting investigation, Tervo (1998) found that friction correlates to some extent with Rayleighsurface waves. This would suggest that friction arises primarily from the elastic interactions (i.e., latticevibrations) on the surfaces due to sliding motion. The velocity of a Rayleigh surface wave is dependenton the Poisson ratio, shear modulus, and density. Tervo found that nitrogen alloying of stainless steelslightly increases the friction.

In a molecular-level study of Amontons’ law, Berman et al. (1998) found that at the molecular level,the projected area is not necessarily proportional to the load and that the shear strength is not constant.Despite this, Amontons’ laws are obeyed and the friction force is still proportional to load on a macrolevel.They offer a physical model, based on intermolecular forces and thermodynamic considerations, thatexplains why the friction force is proportional to the applied load, and why the case of adhering surfaces —where the friction force is found to be proportional to the molecular contact area — is quite differentfrom that of nonadhering surfaces.

Today, we are only at the very beginning of the understanding of the nanomechanical tribologicalcontact effects that explain the origin of friction and wear, and there is no doubt that in the near futuremany new theories and explanations for the origin of tribological phenomena will become available.

The scaling up of the nanomechanical explanations of contact mechanisms to practically usefulconclusions on a macroscale is a most challenging and complex task and will take many years. Therealready are practical applications on a nanoscale where the increasing knowledge of tribological nano-mechanisms can be used. This has resulted in the development of Micro Electro Mechanical Systems(MEMS) such as motors, transducers, gears, and bearings of sizes in the micrometer range (Bhushan,1998a). For these extremely small components, silicon has been used in the early applications forproduction reasons, but studies have shown that tribological improvements can be achieved usingpolycrystalline diamond or MoS2 thin coatings or hydrogenated diamond-like carbon coatings (Donnetet al., 1995; Beerschwinger et al., 1995; Gardos, 1996).

23.3 Macromechanical Interactions: Hardness and Geometry

The macromechanical tribological mechanisms describe the friction and wear phenomena by consideringthe stress and strain distributions in the entire contact, the total elastic and plastic deformations theyresult in, and the total wear particle formation process and its dynamics. In contacts with one or twocoated surfaces, four main parameters that control the tribological process must be defined (Holmbergand Matthews, 1994). They are the coating and bulk hardness, coating thickness, surface roughness, anddebris and tribolayers in the contact. The effects of these parameters are further discussed.

The influence on the macromechanical friction mechanisms of coating and substrate hardness, coatingthickness, surface roughness, and debris present in the contact is illustrated in Figure 23.4 and thecorresponding wear mechanisms in Figure 23.5. The mechanisms involved are very different, dependingon whether the coating and/or the substrate is soft or hard (Holmberg and Matthews, 1994; Donnet,1995; Ramalingam and Zheng, 1995). Consider first the case of soft coatings on harder substrates.

23.3.1 Soft Coating on Hard Substrate

23.3.1.1 Decreased Friction by Shear in Soft Top Layer

For soft coatings, the thickness of the coating influences the ploughing component of friction. When thefilm is thin enough, the effect of ploughing on the film is small (see Figure 23.4(b)). The friction is thusdetermined by the shear strength of the film and the contact area, which is related to the deformationproperties of the substrate (Roberts, 1989, 1990). The formation of grooves in the coated surface byplastic deformation (see Figure 23.5(a)) is the main wear effect; but with thinner films, continuous slidingcan result in coating compaction as well as adhesive and fatigue wear (see Figure 23.5(b)).

Using a low cycle fatigue model, Tangena and co-workers correlated the wear volume with the calcu-lated von Mises stresses averaged over the contact width and the depth of a gold surface layer (Tangena,1987, 1989; Tangena and Wijnhoven, 1988; Tangena et al., 1988). The effect of decreased shear stress atthe contact due to a soft film was investigated by Erdemir et al. (1991a). They applied a thin 2-µm-thicksilver film on a hard ceramic plate sliding against a hard ceramic ball; the coefficient of friction decreased

FIGURE 23.5 Macromechanical contact conditions for different mechanisms that influence wear when a hardspherical slider moves on a coated flat surface.

from 0.8 to 0.4 and the wear of both the coating and the counter-surface decreased by several orders ofmagnitude. A soft coating not only reduces the coefficient of friction, but can also reduce the surfacetensile stresses that contribute to undesirable subsurface cracking and subsequently to severe wear (Spal-vins and Sliney, 1994).

23.3.1.2 Ploughing Friction

The friction usually increases with coating thickness for soft coatings, due to plastic or elastic deformationof the film and to the increased contact area at the interface between the sliding counterface and thecoating where the shear takes place (see Figure 23.4(a)). This increase in friction has been experimentallydemonstrated for lead films (Tsuya and Takagi, 1964; Sherbiney and Halling, 1977); for gold films (Takagiand Liu, 1967); for silver films (El-Sherbiny and Salem, 1986; Yang et al., 1999); and for MoS2 films(Aubert et al., 1990; Wahl et al., 1998).

For very thick soft coatings, the mechanism of ploughing will be very similar to that of soft bulkmaterials scratched by a hard indenter. These tribological mechanisms have been described by severalauthors, for example, Hokkirigawa and Kato (1988), who in addition to the ploughing mechanism, alsoidentify a wedge-forming and a cutting mechanism, depending on the degree of penetration and theshear strength of the contact interface.

23.3.1.3 Influence of Surface Roughness

The substrate surface roughness has an almost negligible influence on friction if the roughness is con-siderably smaller than the thickness of the soft coating and the coating is stiff enough to carry the load,as shown in Figure 23.4(e). However, when the roughness of the slider is higher than the coating thickness,coating penetration will take place (Figure 23.4(f)) and the friction is considerably increased due toscratching of the substrate material. This has been described both experimentally and theoretically bySherbiney and Halling (1977), Halling and El-Sherbiny (1978), and El-Sherbiny (1984). A model forpredicting the wear rate when hard asperities penetrate the soft coating and produce grooves in thesubstrate (see Figure 23.5(c)) was developed by El-Sherbiny and Salem (1984).

23.3.2 Hard Coating on Softer Substrate

The use of hard thin coatings on softer substrate materials is currently very popular in many tribologicalapplications. The hard coating can provide good wear protection and, with a suitable choice of materialand surface layer design, the friction can be very low as well.

23.3.2.1 Hard Coating Can Reduce Wear

With a very thin hard layer on top of a softer substrate (see Figure 23.5(e)), it may be that neither thecoating nor the substrate is able to support the load. However, the function of the coating is to separatethe substrate from the counter-surface and to prevent ploughing by increasing the hardness of the toplayer of the surface. The latter effect has been considered to be very important by Shepard and Suh (1982),Suh (1986), and Bull and Rickerby (1990). For very thin carbon nitride layers with thicknesses of 10 to200 nm on a silicon substrate, Wang and Kato (1999a,b) reported an increase in nano-indentationhardness with increasing coating thickness. The increase in wear resistance with increased coating hard-ness has recently been reported by Kodali et al. (1997) and Wiklund (1999). Increased substrate hardnessresults in decreased contact area, where the shear takes place, and decreased friction, which is indicatedin the results obtained by Ronkainen et al. (1999a).

23.3.2.2 Hard Coatings Without Microfilm Can Have High Friction

The prevention of ploughing reduces both friction and wear; however, the higher shear strength intro-duced at the contact interface by the hard coating can, on the other hand, have the effect of increasingfriction in sliding if no microfilms are formed as described above (see Figure 23.3(d)). That is why veryhigh coefficients of friction often occur in sliding contacts with hard coatings (Holmberg and Matthews,1994; Voevodin et al., 1995d). The increase in friction caused by increased shear strength generally seems

to be more dominant than the reduction in friction due to decreased ploughing. Here, the recentlydeveloped diamond and diamond-like carbon coatings are an exception. The friction may then beextremely low because of the graphitic microfilm produced, as described in Section 23.6.

23.3.2.3 Deflection Influence on Stresses and Cracks

For very soft substrates, the indentation deformation will be considerable and this will thus add aploughing or hysteresis effect on friction (see Figure 23.4(d)). The deflection increases the stresses in thecoating and also at the interface between the coating and the substrate, possibly resulting in fracture orfatigue cracks that may harm the coating or the substrate material. With a soft substrate, cracks mayoccur in the coating, both within the contact area and outside at the substrate material pile-up area, asshown in Figure 23.6 (Burnett and Rickerby, 1987b; Miyoshi,1989; Page and Knight, 1989; Bull et al.,1994; Guu et al., 1996a,b; Lin et al., 1996a,b; Lee and Jedong, 1998; Yuan and Hayashi, 1999). The harderthe substrate, the higher the loading that the coating can resist without failure by fracture (Hedenqvistet al., 1990).

23.3.2.4 Thick Hard Films Are Better Able to Carry the Load

Increased loads can also be resisted with thicker hard coatings because of their load-carrying capacity,which reduces deflection, as shown in Figure 23.4(c) (Rabinowicz, 1967; Roth et al., 1987). A thick hardcoating will have a modifying effect on the size and shape of the stress zone beneath the coating, as hasbeen shown for a hardness indenter by Burnett and Rickerby (1987a).

23.3.2.5 Cracks Generated by the Deflection

When a loaded sphere is sliding on a hard coating, the friction originating from both shear and ploughingwill result in tensile stresses behind the sphere and compressive stresses and material pile-up in front ofthe sphere. This can result in several different cracking patterns, depending mainly on the geometry andhardness relationships (Buckley, 1981; Je et al., 1986; Hedenqvist et al., 1990; Hills et al., 1990). Thecontact response of a coated system with large deflections has been modeled by Ramsey et al. (1991),showing the nonlinear contribution that stretching of the coating has on stiffness of the surface whenthe deflections are of the order of the coating thickness or greater.

23.3.2.6 Fatigue Life of Coatings

The fatigue life of a thin coating may be considerably longer than that of a thick coating for severalreasons. Under similar deformation conditions, the thicker coating will experience higher bend stresslevels. Because coatings typically have columnar growth morphologies, any crack normal to the surfacewill be large in a thick coating, and may exceed the critical crack length; whereas in a thin coating, thismay not be the case. It has been shown experimentally that in rolling contact fatigue tests, hard TiNcoatings with a film thickness well below 1 µm have up to 2 orders of magnitude longer lifetime than2- to 3-µm-thick similar coatings (Erdemir and Hochman, 1988; Chang et al., 1990, 1992; Erdemir, 1992).

In a more recent paper, Polonsky et al. (1997) claim, without giving experimental details, that theyhave been able to demonstrate that the rolling contact fatigue life increase achieved with less than 1-µmTiN coatings was entirely due to polishing of the steel loading balls by the significantly harder TiN coating.

FIGURE 23.6 Fracture of a hard brittle coating on a soft substrate takes place in the contact area and at the materialpileup area around the contact. Directions of material flow are indicated by arrows.

They changed the test procedure and eliminated the polishing effect and observed a somewhat negativeeffect of the TiN coating on the rolling contact fatigue life. In their paper, they present a theoretical modelof near-surface rolling contact fatigue initiation in coated rolling contact elements and conclude that atruly effective hard coating that can resist near-surface rolling contact fatigue must be relatively thick(>3 µm), adherent to the substrate, and have a fine microstructure to resist cohesive fatigue failure inthe coating. This has not been experimentally verified but they report that such work is ongoing.

23.3.2.7 Stresses at the Coating-Substrate Interface

It has been shown by Sainsot et al. (1990) and others that high stresses are concentrated at the interfacebetween the hard coating and the soft substrate (see Figure 23.1). The ability to resist the high stressesat the interface depends on the strength of adhesion between the two. Single or repeated loading canresult in breaking of the adhesive bonds and liberation of parts of the coating in flake-like wear debris(Miyoshi, 1989; Hedenqvist et al., 1990).

23.3.3 Load-Carrying Capacity of Coated SurfacesThe term “load-carrying capacity,” or load-bearing capacity, has been used to express the ability of acoated surface to withstand the action of a loaded tribological contact without failure. More specifically,the load-carrying capacity of a coating/substrate system has been determined as the ability of the systemto bear the normal and tangential forces applied on the coated surface without losing the functionalproperties of the system (Ronkainen et al., 1998b).

There is no standardized or generally agreed and defined method to measure the load-carrying capacity.Even its definition is still a subject for scientific discussion. However, many investigators have used thescratch test method (Valli, 1986; Valli et al., 1986) for this purpose, ostensibly to assess the effectiveadhesion of the coating under conditions in which bulk plastic deformation occurs.

Experiments have shown that the load-carrying capacity of a coated surface can be improved byincreased substrate hardness and increased coating thickness, in the coating thickness range of 0.5 to9 µm (Bell et al., 1998; Lee and Jeong, 1998; Ronkainen et al., 1999; Wang and Kato, 1998, 1999a,b), asshown in Figure 23.7. Correlating results of TiN coating breakdown with softer substrates in fretting testshave been reported (Shima et al., 1999). In a microscale scratch and wear resistance test, Sundararajanand Bhushan (1999) showed that very hard diamond-like carbon coatings can have a reasonable load-carrying capacity and wear resistance at a coating thickness even down to 5 nm. A thinner, 3.5-nm-thickcoating had poor load-carrying capacity. Lee and Jedong (1998) concluded from Rockwell hardness testresults that the load-carrying capacity is improved by decreased surface roughness, but these resultsshould not be generalized because they were not in correlation with scratch test results.

The mechanism when the load-carrying capacity is exceeded and the coating fails can be very different,depending on the materials involved and surface parameters. Wang and Kato (1998) have shown thatwhen a diamond tip slides over very thin carbon nitride coatings (h = 10–200 nm) on a bare siliconsubstrate, this will result in a “no-grooving” mode at low loads, a “grooving with no material removal”mode at higher loads, and finally in a “grooving with material removal” mode at high load (up to 300 mN)with a linear sliding speed of 10 µm/s. In this case, the failure occurs as grooves in the coating due toplastic deformation. In more severely loaded situations, the failure may be more due to brittle fractureof the coating (Lee and Jedong, 1998; Kodali et al., 1997) and flaking of the coating. The pattern ofcoating fracture and failure due to exceeding the load-carrying capacity has been explained and classifiedby Olsson (1989), Hedenqvist (1991), and Wiklund (1999).

Scratch testing is a fast and easy way of evaluating the load-carrying capacity of a substrate/coatingcombination that gives reasonably relevant results. One should be very careful when using the Rockwellindentation test for this purpose because it has been found that the results can be contradictory to thescratch test results (Lee and Jedong, 1998; Ronkainen et al., 1998b, 1999a). It has been suggested byBennet et al. (1994) that a more lightly loaded multipass scratch testing method gives better correlationwith practical situations than single-pass testing and would thus be more suitable as a quality assurancemethod for coatings. Ronkainen et al. point out that the stress situation under the diamond tip in a

scratch test is much more severe than generally observed in practice and thus a more suitable methodfor evaluating the load-carrying capacity of a coated surface would be a ball-on-disc configuration, astypically used in sliding tests.

23.4 Micromechanical Interactions: Material Response

The micromechanical tribological mechanisms describe the friction and wear phenomena by consideringthe stress and strain at the asperity level, the crack generation and propagation, material liberation, andsingle particle formation. In typical engineering contacts, these phenomena are at the size level of about1 µm or less, down to the nanometer level.

23.4.1 Material Response

In a tribological contact, one surface moves over another, typically by a sliding, rolling, cutting, or impactingkind of motion. The interactions between the surfaces and the resultant material changes will govern thefriction and wear behavior. Friction can be generated by adhesion between the two surfaces, by ploughing,or by asperity deformation, as shown in Figure 23.8. In a rolling contact situation, the friction is induced bythe hysteresis of the repeatedly elastically deformed rolling component (Holmberg and Matthews, 1994).

The elasticity, hardness, and fracture toughness of the material influence friction and wear, both on amacroscale and on a microscale. On a macroscale, the rough surface of an elastic material deforms withthe applied load and the load is thus spread more widely over the two surfaces, resulting in a lowersurface pressure for the materials to respond to, as shown in Figure 23.9(a) and (b). When the load isremoved, the surface will return to its original shape.

FIGURE 23.7 Influence of (a) substrate hardness, (b) coating thickness, and (c) surface roughness on load-carryingcapacity measured by scratch test critical load for TiN coating on a steel substrate. (Data from Lee, Y.Z. and Jeong,K.H. (1998), Wear-life diagram of TiN-coated steels, Wear, 217, 175-181.)

In sliding and rolling, the hardness of the material indicates its ability to carry a load. When acounterface such as a sphere slides on a soft flat surface, indentation and ploughing may take place, withhigh friction and wear as a result. For a harder material the ploughing action decreases, with less frictionand wear as a result, as shown in Figures 23.9(c) and (d). It was previously suggested that the E:H ratioshould be used as a parameter for estimating the relative wear rate of materials, and in several cases thiscorrelates well with results from experimental wear tests (Oberle, 1951).

It has been reported that in an abrasive wear situation, the wear is inversely proportional to the elasticmoduli (Spurr, 1957; Lancaster, 1963). However, this may be true only for surfaces with preexisting flawsor cracks because, in general, tough and elastic materials (i.e., with low E) are able to resist abrasion veryeffectively. If flaws are present, then a stiffer (i.e., higher E) surface may resist crack-opening forces, andthis is the reason why brittle fracture mechanics models prescribe high E values to produce “tougher”surfaces. However, the balance of evidence now suggests that a coating with a lower E value will be morewear resistant than one with a similar hardness but higher E (Leyland and Matthews, 2000; Matthewsand Leyland 2000a,b). Coatings with a higher H/E ratio have a longer elastic strain to failure, and arethus more able to deform (without yielding) as the substrate deforms.

23.4.2 Energy Accommodation in the SurfaceWhen one surface is loaded against another, the two surface geometries interact and influence thepressures on the surfaces and the stresses generated in the two solids close to their surfaces. The mech-anisms taking place related to friction and wear have been reviewed by Holmberg and Matthews (1994).

FIGURE 23.8 The three components of sliding friction are (a) adhesion, (b) ploughing, and (c) asperity deformation.

FIGURE 23.9 The deformation and contact pressure for a rough sphere on a rough plane in the case when bothare (a) rigid and (b) elastic. The substrate deformation when a hard sphere is sliding on (c) a hard substrate and(d) a soft substrate.

a) Adhesion b) Ploughing c) Asperity deformation

Godet (1989) presented a velocity accommodation approach in which he pointed out that the elastic,plastic, fracture, and rolling behavior in the contact are important for friction and wear. The followingdevelops and presents a new energy accommodation concept for the wear behavior of coated metallicor ceramic surfaces, which emphasizes the parameters of elasticity, plasticity, and fracture toughness.

The basic wear mechanisms can be described by studying the material response effects in a simplifiedsituation when a sphere is sliding over a flat surface, as in Figure 23.10. At stage 1, point A on the uppersurface is in contact with point B at the lower flat surface. When the upper surface moves to stage 2 somedistance to the right, point B will follow point A because of the adhesion or interlocking between thesurfaces. This movement of point B is allowed by the material response in the lower surface.

This material response can be split up into three components: elastic deformation, plastic deformation,and brittle fracture. In the elastic case, B will follow A; and when the contact is released, B will elasticallymove back to its original position. The work in this process, the load × distance moved, has beentransferred to heat. The main parameter indicating the material response is the Young’s modulus ofelasticity, (also called the elastic modulus, E).

In the situation of plastic behavior, point B will follow point A in the same way, but after the contactrelease, the deformation of the flat surface will not return to its original shape but remain deformed.The work of the rubbing process has been transferred to heat and permanent material modification. Inthis case, a key parameter indicating the material response is the shear strength, τ.

Third, in the situation of brittle fracture behavior, point B again moves with point A and now themovement is allowed by the material cracking on the left-hand side of the contact. The moving contacthas created a crack in the surface. The work of the process has now been transferred to heat and surfaceenergy needed for the formation of new surfaces in the crack. The parameter indicating the materialresponse is the fracture toughness or, more specifically, the critical stress intensity for microstructurallyshort cracks, KIc. The fracture toughness can more conveniently be expressed in the form of critical strainenergy release rate, Gc.

Thus far, it has been assumed that both materials are homogenous and defect-free. Now assume thatthe flat material is not defect-free, but has cracks at the surface perpendicular to the surface, created, forexample, from earlier sliding situations that are now repeated. Figure 23.11 shows how the surface in theelastic case is modified and returns after each contact to its original shape. However, after a certainnumber of repeated contacts, fatigue cracks will be created in the surface that result in material release.A load that is sufficiently large to cause plastic deformation in the early cycles of the loading mayaccommodate purely elastic deformation in the later stages due to a shake-down material effect of thesurface (Wong et al., 1997).

FIGURE 23.10 The three modes of energy accommodation in a sliding contact are elastic deformation, plasticdeformation, and brittle fracture.

In the fully plastic situation, the contact will result in a modified surface; and when this is repeated,the modification will continue, and finally this process will also result in material release from the surface.In the brittle fracture situation, the repeated contacts will result in more new cracks. One can see thatin all three cases, material will be released from the surface and wear particles will be formed.

The discussion has thus far been limited to a sphere sliding on a smooth surface. The situation isbasically the same even if the lower surface is not considered flat due to surface roughness, curvature, orploughing, as shown in Figure 23.12. In the case of rolling, the situation will be similar, but now thetangential friction force will be much lower and the stress field in the surface will have a different shape.

The authors’ conclusion from this discussion is that the Young’s modulus of elasticity, shear strength,and fracture toughness are basic important material parameters influencing the friction and wear behav-ior of the contact. To control the wear and friction situation, one needs to know the values of theseparameters in the coating, at the asperities, and in the bulk of the substrate, as illustrated in Figure 23.13.

FIGURE 23.11 Material release and particle formation mechanisms in the elastic, plastic, and brittle energy accom-modation modes.

FIGURE 23.12 Energy accommodation geometry in (a) sliding and (b) rolling, with curved surfaces or asperities.

23.4.3 The Influence of Fracture Toughness in Material Response

The influence of elasticity on wear has been recognized in several papers. The plasticity effect has alsobeen widely studied (Bowden and Tabor, 1950, 1964; Oberle, 1951; Lancaster, 1963; Halling and Arnell,1984). Hardness, H, is a measure of the contact pressure needed to cause yield and material flow in asurface, that is, plastic deformation.

The commonly used parameter indicating the response of brittle materials to loading is the fracturetoughness or, more specifically, the critical stress intensity ψ factor, KIc. The fracture toughness can bemore conveniently expressed in the form of critical strain energy release rate, Gc. Gc is a measure of thefracture toughness of a material and, as such, it is a material property just like Young’s modulus and theyield stress (Anderson et al., 1985). Gc can be expressed as:

where Kc is the critical stress intensity factor, E the Young’s modulus of elasticity, and ν the Poisson ratio.The models and measurement methods developed in fracture mechanics, such as the linear elastic

fracture mechanics (LEFM) techniques for expressing the fracture toughness of a material, are in mostcases applicable for large cracks longer than 1 mm, or with special modifications for physically shortcracks in the range of 100 µm to 1 mm.

Cracks less than 100 µm are considered microstructurally short cracks because their dimensions areon the order of the grain size in metals. At these scales, the material no longer behaves as a homogenouscontinuum because the crack growth is strongly influenced by microstructural features. In this region,crack growth is influenced by grain size and boundaries, inclusion spacing, precipitation spacing, micro-asperities, macro-asperities, etc. The crack growth is often very sporadic, such that it may grow rapidlyin certain intervals and then be followed by crack arrest when it meets barriers such as grain boundariesand second-phase particles (Haddad et al., 1979; Meguid, 1989; Anderson, 1995). Microstructurally shortcracks in a plain specimen may not affect its fatigue limit. Fracture mechanical aspects on short cracksand their behavior have been collated by Miller and de los Rios (1986).

The three regions of crack growth are illustrated in a diagram with log stress range, log ∆σ, as afunction of log crack lengths (log a), as first suggested by Kitagawa and Takahashi (1976) and shown inFigure 23.14.

In contrast to long fatigue crack propagation, there exists no adequate definition of the growthmechanism, the crack tip stress/strain fields, or the interaction with microstructural features for shortfatigue crack propagation, which is the more likely case in thin metallic or ceramic coatings.

FIGURE 23.13 Basic material response parameters in energy accommodation are the Young’s modulus of elasticity(E), the shear strength (τ), and the fracture toughness (indicated by KIc or Gc) in the coating (c), at the asperities(a) and in the bulk of the substrate (b).

G K E

G K E

c c

c c

=

= −( )2

2 1

for plane stress and

ν

The different features that can influence how short cracks are initiated and subsequently grow inmetallic components are shown in Figure 23.15. The process is influenced by the material parameters,the geometrical features, and the state of stress, both applied and residual.

The three fundamental mechanical properties — the elastic modulus, hardness, and fracture toughness —can be measured using a sharp indenter (Pharr, 1998). If they are to be measured close to the surface,as is relevant when evaluating tribological properties, a low-load nano-indenter should preferably beused. In indentation measurements, the effects of temperature, environment, and geometrical scale(Ramsey and Page, 1988), as well as mechanisms like the pile-up effect (Bolshakov and Pharr, 1998),should be considered.

23.5 Material Removal and Change Interactions: Debris and Surface Layers

The process of wear and friction results in both geometrical and structural changes in the surfaces of thecontacting bodies. These changes will influence future contact conditions and friction and wear generatedin the same contact. The changes range from nanoscale to macroscale. At the nanoscale, outer surfacemolecules are released and they react chemically with adjacent molecules; at the microscale, cracks areinitiated and debris released; and at the macroscale, wear products are agglomerated and surface layersformed and deformed.

23.5.1 Surface and Subsurface Cracks

The initiation of cracks at the surface or in the material is the starting point of a process that may resultin material detachment, debris generation, and the formation of transfer layers.

In most wear situations, in addition to hardness, the ability to elastically recover from deformationand the fracture toughness of the material are very important parameters (Zum Gahr, 1998; Wiklund,1999). The modern fracture mechanics approach to the initiation and generation of cracks is presentedby Andersson (1995). For the case of a loaded, very thin coated surface, there does not exist any generaltheory or model for crack initiation or generation.

FIGURE 23.14 The Kitagawa–Takahashi curve, illustrating the regions for different crack propagation mechanisms.(Adapted from Meguid, S.A. (1989), Engineering Fracture Mechanics, Elsevier Applied Science, London. 397 p.)

Based on macroscale experimental observations in rolling contact tests with bearing steel materials,Nélias et al. (1999) report that the cracks are either initiated from subsurface inclusions or surfacemicrocracks. They claim that the initiation and propagation of inclusion-initiated cracks is the mainmode of sub-surface-originated cracks in rolling bearings. Under the load is a stress concentration builtup around inhomogeneities like inclusions and primary carbides. This is due to the elasto-plastic mis-match between these inhomogeneities and the martensitic matrix. Once the yield stress is reached, aplastic strain is induced in a small volume surrounding the inclusion. Under repeated action, the dislo-cations will move back and forth and accumulate. This process leads to localized changes in the steelmicrostructure and a crack is nucleated in this domain once a critical density is reached.

Surface microcracks are typically generated in the discontinuities of the surface topography, such asat grinding marks or large-wavelength surface roughnesses (Nélias et al., 1999). The process of initiationof surface microcracks and formation of microspalls along grinding marks is shown schematically inFigure 23.16(a). Microcracks can also propagate parallel to the surface at a few micrometers depth alongthe machining direction, leading to the formation of microspalls, as shown in Figure 23.16(b). Theseprocesses have mainly been observed in pure rolling. Sliding may play a significant role, as it producestransverse microcracks at an angle of 30° to 40° to the friction direction. The cracks are located mainlyon the top of large roughness peaks, as shown in Figure 23.16(c). There are indications that the local

FIGURE 23.15 Classification of different sources for short crack initiation and propagation. (Adapted from Meguid,S.A. (1989), Engineering Fracture Mechanics, Elsevier Applied Science, London. 397 p.)

friction coefficient at the top of asperities may be large enough to initiate transverse microcracks. Suchmicrocracks are at the origin of microspalls, usually wider and deeper than those observed under purerolling conditions. Depending on the combination of load movement, friction direction, and crackorientation, a lubricant may help propagate surface cracks by hydraulic pressure effects.

Recent studies of the stress intensity factor at the tip of a surface breaking crack in a layered surfaceindicate that compressive stresses may tend to prevent crack propagation (Eberhardt and Peri, 1994).The crack patterns of ceramic coatings in indentation have been analyzed and correlated to coatingthickness and load, and the critical normal loads have been estimated by considering the deformationof the substrate (Diao et al., 1994b). There is a strong influence on both the fracture load and the fracturepattern by the mismatch of elastic properties between the layer and the substrate (Oliveira and Bower,1996). Variations in coating properties and coating thickness may change the fracture loads by up to afactor of ten.

In ceramic coatings, the thickness and grain size are important parameters for crack propagation. Inindentation tests of 1- to 10-µm-thick Al2O3 coatings on cemented carbide substrates, Yuan and Hayashi(1999) found that the critical load for generation of radial cracks decreases with increasing coatingthickness, and that the grain size of the coatings decreases with the thickness. A more coarse-grainedmicrostructure generally displays a lower cohesive strength as compared with a fine-grained microstruc-ture. Thus, the fracture strength of the coating increases with decreasing grain size and thickness.

Nanocrystalline-amorphous TiC-amorphous carbon (a-C) composite coatings have been producedby Voevodin and Zabinski (1998a,b) with a combination of high hardness and fracture toughness. Intheir coating design, nanocrystals of 10 to 50 nm size were encapsulated in an amorphous phase, thenanocrystals being separated from each other by 5 nm. This permitted generation of dislocations andgrain boundary microcracks and nanocracks, which terminated in the surrounding amorphous matrix.This surface design achieved a self-adjustment in composite deformation from hard elastic to plastic atloads exceeding the elastic limit, as opposed to deformation from hard elastic to brittle fracture.

FIGURE 23.16 Schematic stages of (a) microspall formation along grinding marks (transverse cut), (b) microspallformation below longitudinal roughness of large wavelength (transverse cut), (c) microcrack and microspall forma-tion (longitudinal cut), and (d) surface initiated deep spall formation (longitudinal cut). (Adapted from Nélias, D.,Dumont, M.L., Champiot, F., Vincent, A., Girodin, D., Fougères, R., and Flamand, L. (1999), Role of inclusions,surface roughness and operating conditions on rolling contact fatigue, J. Tribol., Trans. ASME, 121, 240-251.)

5 µm5 µm

10 µm 400 µm

40 µm

a) b)

c) d)

For amorphous diamond-like carbon (DLC) coatings in the thickness range of 0.7 to 2 µm, it wasshown that the hardness and apparent fracture toughness of the coating/substrate combination, measuredby Vickers indentation, depend on the thickness of the coating (Kodali et al., 1997). The hardness andthe apparent fracture toughness of the coating (both of which influence the wear) increased with increas-ing coating thickness. The hardness and the thickness influenced the initiation of cracks, whereas theresidual stress in the film influenced the propagation of cracks.

In a study of very thin DLC coatings in the thickness range of 3.5 to 20 nm, Sundarajan and Bhushan(1999) conclude that the formation of cracks depends on the hardness and fracture toughness of thecoating. They suggest that the observed nonuniform failure depends on spatial variation in the coatingproperties. Surface cracks are developed in weaker regions with lower fracture toughness. As crackspropagate, they are forced to expand within the weak region, as the neighboring strong regions inhibitextensive lateral crack growth. Because of this, cracks propagate down to the interface, where, aided byinterfacial stresses, they get diverted along the interface just enough to cause local delamination of thecoating. Simultaneously, the weakened regions experience excessive ploughing. Thus, weaker regions failwhile stronger regions remain wear-resistant. The propagation of cracks along the coating/substrateinterface is suppressed due to the strong adhesion of the coatings because, otherwise, coating delaminationwould take place.

There have been a number of papers published wherein the crack pattern has been experimentallydemonstrated and studied both by indentation and scratch testing. Such studies have been published byOlsson (1989), Hedenqvist (1991), Prasad and Zabinski (1997) for nanocrystalline ZnO, Sundararajanand Bhushan (1998) for SiC, Sue and Kao (1998) for TiN/TiC multilayer coatings, Wiklund et al. (1999)for TiN and CrN coatings, and Yuan and Hayashi (1999) for Al2O3 coatings.

23.5.2 Tribological Influence of Wear Debris

Debris that has been generated by the wear process or loose particles originating from the surroundingenvironment may be present in a tribological contact. It has been observed that the coefficient of frictionrises significantly once wear particles are formed at the sliding surface, and particles present in the contactaffect the instantaneous coefficient of friction and the wear. The coefficient of friction can be altered byremoving wear particles or by inserting particles in the interface, as shown by Hwang et al. (1999), whoalso found that the particle size influences friction but not so much the number of particles present inthe contact. In experiments with sliding surfaces of materials with different hardnesses (Pb, Zn, Al, Cu,Ni, Ti, and AISI 1045 steel), they found a clear difference between soft and hard surfaces. Soft and ductilesurfaces produced larger wear particles with a stronger tendency to agglomerate, while hard surfacesproduced smaller wear particles with weaker agglomeration tendency.

For coated surfaces, the influence of debris in the contact on friction and wear is, under someconditions, considerable, depending on the particle size and shape; coating thickness and surface rough-ness relationship; and the particle, coating, and substrate hardness relationship (Holmberg and Matthews,1994).

23.5.2.1 Particle Embedding

In the sliding situation shown in Figure 23.4(i), hard particles are present in the contact, the particleshaving a diameter considerably smaller than the thickness of the soft coating on a hard substrate. Theparticles are pressed into the soft coating and embedded into it without any further contact with theslider as long as the soft coating remains thicker than the particle diameter. In this case, the particleshave no great effect on friction, which is primarily controlled by the ploughing mechanisms describedearlier. The slider will produce a main groove by ploughing in the soft coating, and the surface asperitiesor trapped debris may cause microploughing and microgrooves within the main groove (Hwang et al.,1999). El-Sherbiny and Salem (1979) showed with a theoretical model that the wear appears to bedependent on the surface texture of the system rather than on material properties during the initial wearwhen hard conical asperities are ploughing a soft surface coating.

23.5.2.2 Particle Entrapping

For thin surface coatings where the dimensions of the particles are of the same magnitude or larger thanthe coating thickness and the surface roughness, as shown in Figure 23.4(j), their influence on frictioncan be considerable. If the particles are harder than the coating but softer than the substrate, then theyare easily caught by the roughness of the counterface or partly sink into it and scratch grooves in thesoft coating, as in the case of asperity penetration. The friction will increase because of particles ploughingthe coating (Hwang et al., 1999). The influence of ploughing wear particles on friction in sliding wearcontacts has been shown to be considerable (Kim and Suh, 1991; Komvopoulos, 1991a,b).

An increase in friction may follow if the slider and the substrate are of equal hardness and the looseparticles in the contact have a higher hardness (Suh, 1986). Then the particles may partly sink throughthe coating into the substrate, as well as into the counterface by a kind of anchoring mechanism resistmotion. Even if the particles are fairly spherical in shape, it is not probable that they will decrease frictionby rolling because they will be stuck into the soft coating. Sin et al. (1979) have shown that the contri-bution of ploughing on the friction coefficient is very sensitive to the ratio of the radius of curvature ofthe particle to the depth of penetration. The particles are free to rotate in the contact and they are oftenconsiderably sharper than the asperity angles of engineering surfaces. The wear rate depends on theparticle size.

If the particles in the contact are soft, then their tribological effect is quite different. Soft particles withlow shear stress trapped in the contact can carry part of the load and inhibit direct substrate-to-countersurface contact, thus reducing both wear and friction (Grill, 1997; Voevodin et al., 1999d). Asimilar effect has been observed by Yamada et al. (1990), who found that polymer particles can act veryeffectively to reduce wear and friction.

In contacts with MoS2 coatings or when Mo and S are present, MoO2 wear debris is produced thathas a very low shear stress and thus can act as a solid lubricant, reducing the friction and wear (Donnetet al., 1995; Wahl and Singer, 1995; Singer et al., 1996a,b; Donnet, 1996, 1998; Grossiord et al., 1998;Singer, 1998; Xu et al., 1999; Wahl et al., 1999). The tribological behavior of molybdenum disulfidecoatings is treated in detail in Chapter 22.

23.5.2.3 Particle Hiding

The introduction of small particles into the sliding contact of a hard and rough surface, as shown inFigure 23.4(k), does not necessarily make the tribological contact more severe. The particles can be hiddenin the valleys formed by the asperities, while the sliding takes place at the asperity tops. Thus, the particleswill have no great effect on either friction or wear. It is important to notice that reduced surface roughnesscan increase both friction and wear if the particles cannot hide in the valleys and instead interact withthe surfaces by scratching and interlocking.

23.5.2.4 Delamination

In the very common sliding situation when a hard rough slider moves over a hard rough surface, asshown in Figure 23.5(g), the sliding action takes place at the top of the contacting asperities. They mainlydeform plastically, although the overall contact stress may be less than the yield stress of the contactingmaterials, because the local stresses at the small asperity areas are much higher.

High contact stresses can generate dislocations, pile-up of dislocations, and crack nucleation very nearthe surface, as described by Suh (1973, 1986) in the delamination theory of wear and shown inFigure 23.17. Because of the small plastic deformation of the surface, a large number of cracks must benucleated before a loose particle can form. The delamination particles are flake-like and may be somehundred micrometers long. This kind of flake-like wear debris has been observed in the sliding contactbetween a steel slider and a hard titanium nitride coating by Sue and Troue (1990) after about 200 slidingcycles, agglomerated particles being observed after 500 cycles. This behavior correlates with the obser-vations of Fu et al. (1998) for plasma-sprayed Ni,Cr coatings sliding against a steel ball. The generationof flake-like wear debris, probably by some kind of delamination mechanism, has also been observed forsofter materials like polyimide polymers (Yamada et al., 1990). The wear rate in delamination can be

calculated from the equation developed by Suh (1986) where it is given as a function of hardness, contactlength, depth, average crack propagation rate, load, spacing of asperity contacts, and crack spacing.

23.5.2.5 Particle Crushing

When particles, large in relation to the surface roughnesses, are introduced between two hard surfaces,the result can be particle crushing, scratching, or rolling, as shown in Figure 23.4(i). If the particles havelower hardness than the surfaces, then they will be crushed and destroyed under the load of the contact,with smaller debris and some increase in friction as a result.

If the particles have a higher hardness than the surfaces, they will be caught by the roughness of thesurfaces, resulting in ploughing and scratching. The scratching particles carry part of the load, whichresults in concentrated pressure peaks on both surfaces as they try to penetrate them. The high pressurepeaks may well be the origin of crack nucleation in the coating (De Wit et al., 1998).

The presence of hard particles between hard surfaces may in some cases even reduce the coefficientof friction. If the particles are fairly round in shape, hard enough to carry the load, and at least one ofthe surfaces is smooth, the particles may act as rollers and reduce the friction (Blomberg, 1993; Fu et al.,1998).

The debris in the contact can also change in its properties during the sliding action. In the slidingcontact of a TiN coating against a corundum ball, De Wit et al. (1998) observed that the wear debrisproduced by fretting was transformed during the sliding from amorphous debris into nanocrystallinerutile-like particles with a drop in the friction as a result. Wear debris formed in a contact with a multilayer(Ti,Al)N/TiN coating on steel substrate sliding reciprocatingly against a corundum ball was observed tobecome progressively lubricious with increasing humidity. The alumina debris produced was originallyquite abrasive, but increasing humidity lowered the abrasiveness of the alumina particles, probably dueto the formation of hydrated oxides (Huq and Celis, 1999).

The sliding process will have an influence on the material properties of the counterface surface material.Because of work hardening, phase transitions, or third-body formation, the microhardness of a steel wearsurface can be about 3 times larger than the initial bulk hardness. In sliding abrasive contacts withuncoated steel surfaces, most of the wear particles are often less than 1 µm in size (Kato, 1992). Abrasivewear and oxidative wear have been observed in steady-state sliding of a steel slider against a hard TiNcoating by Sue and Troue (1990), Malliet et al. (1991), Franklin and Beuger (1992), and Xu et al. (1999).

23.5.3 Debris Agglomeration

Particles that have been liberated from a surface by wear may still have an influence on the future frictionand wear behavior of the contact. In sliding contacts with materials of different hardness (lead, zinc,aluminum, copper, nickel, titanium, and steel), Hwang et al. (1999) observed different wear particleagglomeration behavior, depending on hardness and sliding direction. Smaller wear particles had astronger tendency to join together and form larger ones, and particles of soft and ductile metals had a

FIGURE 23.17 The initial stage of the delamination process of wear particle formation by shear deformation ofvoids in the material near the surface. (Adapted from Suh, N.P. (1973), The delamination theory of wear, Wear, 25,111-124.)

stronger tendency to agglomerate than those of hard and brittle metals. There was a difference inagglomeration behavior, depending on whether the sliding was unidirectional or reciprocating, and lowerfriction and wear were measured in the cases of reciprocating sliding. Particle agglomeration was notjust limited to one location but occurred simultaneously over a distributed area, and the observed particleor flake sizes were in the range of 100 to 600 µm. The mechanism of agglomeration is complicated andprobably due to adhesion and/or mechanical interlocking. The agglomerated particles can act as largerparticles in the contact, detach to either of the surfaces, or be rejected from the contact (Yuan et al., 1999).

23.5.4 Transfer Layers

The material transfer mechanism (Figure 23.3b) is well-known for polymers (e.g., polytetrafluoroethylene[PTFE]), sliding on steel. Surface material from the polymer will wear off and attach by adhesion to thesteel counterface to form an extensive PTFE film. This means that after some time of sliding, thetribological pair is actually PTFE sliding against a thin PTFE film on steel, which has very low friction.Similar mechanisms have been observed for sputtered, <1-µm-thick PTFE coatings on steel substrates,resulting in film-like PTFE wear debris transfer to the steel counterface; and for sputtered polyimide(PI)-coated steel, resulting in fine flake-like wear debris transfer to form a polymer layer on the steelcounterface (Yamada et al., 1990). In rolling contact bearings in different gas environments, PTFE filmshave transferred from the retainer onto CaF2 layers on the races and balls, effectively protecting themfrom wear and increasing the endurance life (Nosaka et al., 1999).

A typical building-up process of a transferred surface layer in contacts with steel and hard coatingssuch as TiN, CrN, and (TiAl)N has been described by Huang et al. (1994). As a result of the ploughingaction of hard coating asperities, slider material fragments were first removed and then adhered to somepreferential sites on the sliding track of the coating. The preferential sites were the highest asperities thatmade earliest contact with the counter-surface. Repeated sliding resulted in accumulation of fragments,which then united and formed discontinuous layers on the coating surface. After some sliding, the highestasperities were covered with transferred material, which was deformed and flattened under continuoussliding and a transferred layer built up. Similar processes of transfer layer buildup have been observedand reported for different ceramic and steel contacts (e.g., by Andersson and Holmberg, 1994).

Depending on the contact condition, the formation of transfer films in the sliding contact of a titaniumnitride coating against a steel slider may be very complex. Sue and Troue (1990) have described theprocess of minute wear fragments adhering to both surfaces, plastic deformation and strain hardeningof the transferred layers and patches, cracking and oxidation of the layers, removal of the layers andpatches by fracture, and finally the formation of very thin films, possibly oxides, on both surfaces. Inpin-on-disk experiments with a steel ball sliding on TiN coatings, Wilson and Alpas (1998) observed aload effect. At low loads of about 20 N, transfer and buildup of oxidized counterface material on thecoating surface and minimal damage to the coating took place; at higher loads up to 100 N, increaseddebris transfer, polishing, and brittle spallation occurred; but at higher loads, plastic deformation andductile ploughing or smearing of the TiN coating prevailed.

The material transfer layers generated in sliding coated contacts are generally some few micrometersin thickness, but may vary in the range of 0.01 to 50 µm (Ko et al., 1992; Blomberg, 1993).

The excellent tribological properties of molybdenum disulfide coatings, often in combination withother materials as a compound, is basically also due to a process of MoS2 debris generation and transferlayer formation. This process has been described by several authors (Fayelle et al., 1990; Donnet et al.,1995; Wahl and Singer, 1995; Donnet, 1996; Singer et al., 1996a,b; Grossiord et al., 1998; and Singer,1998) and is treated in detail in this book in Chapter 22.

23.5.5 Reaction Layers

In a tribological contact, the sliding process brings energy and often high local temperatures into thecontacts and at the same time the wear process results in exposure of pure uncontaminated material

surfaces to the environment. This situation is very favorable for chemical reactions to take place on thenewly formed or deformed surfaces. The chemical reactions are dependent on what kind of additionalmaterial is brought to the contact in gaseous or liquid form (Ruff et al., 1995; Singer et al., 1996c,Ronkainen et al., 1998a).

Based on observations from sliding contacts with MoS2 and TiN coatings, Singer (1998) presents athree-stage nanoscale model for the formation of reaction layers and for the buildup of transfer layers.First, a thin layer is removed from the coating and transferred to the counterface. Meanwhile, both thesurface and the transfer layer can react with possible surrounding gases, forming new compounds. Thefirst films to be transferred to the counter-surface may be very thin, of only molecular dimensions. Asthe transfer film thickens, it is extruded from the contact area and may break up to form new wear debris.This process is then repeated and a layer-by-layer buildup takes place. It has been observed for MoS2

coatings that particle detachment appears to be preceded by deformation, reorientation, and sometimescrack propagation within the first 20 to 50 nm of the coating.

In environments containing oxygen, such as air, a thin (about 1- to 10-nm-thick) oxide layer is formedvery quickly on most metal surfaces. Some oxide layers (e.g., copper oxide) are sheared more easily thana metal, while others (e.g., aluminium oxide) form a very hard layer. Erdemir et al. (1998) have studiedthe shear properties and lubricity of a number of oxides for high-temperature tribosystems. They con-cluded that a complex system, including the kinetics of oxidation and cation diffusion rates, heats offormation, electrostatic electronegativity, surface energy, and other fundamental crystallochemical prop-erties may influence the adhesion and shear rheology of an oxide film forming on a sliding surface.Mechanistically, the crystallochemical properties relate strongly to the melting point of an oxide and itsviscosity, the lowering of the surface energy and melting point of an oxide when a second oxide is presentin the system, and the solubility, chemical interactions, and compound-forming tendencies between twoor more oxides. Tribologically, these phenomena can play a significant role in shear rheology, adhesiveinteractions, and hence the frictional properties in a sliding contact.

There are a number of reports describing studies on different aspects of oxide layer formation ontribological properties. In contacts with alumina (Al2O3) surfaces, Gee and Jennett (1995) found that thetribologically formed hydroxide films are much softer than the remaining alumina, and that the filmsare liable to fracture under concentrated loading conditions. The hydroxide films were composed of smallparticles, about 20 to 50 nm in size. Erdemir et al. (1995, 1997) investigated the formation and self-lubricating mechanism of thin boron oxide (B2O3) and boric acid (H2BO3) films on the surfaces of boronoxide substrates. They developed a short-duration annealing procedure resulting in the formation of aglass-like boron oxide layer for which they measured very low coefficients of friction (down to 0.05)when sliding against a steel ball. The tendency of formation of an oxide layer on the surface of different(Ti, Al, Zr, Si)N coatings has been reported by Rebouta et al. (1995).

The formation of environmentally stable and smooth third-body films on the wear surface was believedto be responsible for the improvement in room-temperature tribological performance of nanocrystallinezinc oxide (ZnO) films observed by Prasad and Zabinski (1997). The friction coefficient of the ZnO filmsagainst a steel ball was between 0.16 and 0.34. Friction, in the beginning, was sensitive to load and slidingspeed; but once a smooth wear scar was formed, it was very stable and the test conditions had nosignificant effect on it.

Experiments with steel sliding against a chromium nitride (CrN) coating resulted in the formation ofa chromium oxide (Cr2O3) surface film with very good wear resistance (Lin et al., 1996b). Increases inthe applied load or the sliding velocity helped to form a thicker Cr2O3 film, thus reducing the wear.However, the growth of a thick TiO2 surface film in a TiN-coated contact under the same conditions didnot help reduce the wear loss.

Additional material, active in the chemical surface reaction, may also be brought to the contact inliquid form. Li et al. (1998) investigated the sliding contact of a plasma-sprayed Cr3C2-NiCr surfaceagainst a TiO2 coating under water and ethanol lubricated conditions in a block-on-ring test. They foundthat water deteriorated the tribological properties of both surfaces by accelerating cracking and fracture.Ethanol reduced the friction coefficient and wear of the Cr3C2-NiCr surface, which could be attributed

to the formation of a smooth surface film consisting mainly of Cr2O3. However, ethanol also increasedthe wear coefficient of the TiO2 coating by adsorption-induced cracking as a result of the low fracturetoughness of the coating.

Surface chemistry considerably influences the surfaces in sliding contacts involving hard carbon coat-ings such as diamond films or amorphous diamond-like carbon layers. The extremely low friction andthe wear properties observed are partly explained by grafitization of the top layer of the surface (Jahanmiret al., 1989; Erdemir et al., 1991b; Donnet et al., 1994; Ronkainen et al., 1994; Schouterden et al., 1995;Erdemir et al., 1995; Pimenov et al., 1995; Voevodin et al., 1996a; Liu et al., 1996; Erdemir et al., 1996a,b;Schmitt et al., 1998; Ronkainen et al., 1999b). The tribological mechanisms involved have been explainedby several authors (Gardos, 1994; Holmberg et al., 1994; Donnet et al., 1995; Voevodin et al., 1995b;Donnet, 1996; Hogmark et al., 1996; Voevodin et al., 1996a; Grill et al., 1997; Huu et al., 1998; Singer,1998; Voevodin et al., 1999d).

23.6 Multicomponent Coatings

It has been known for many years that it can be tribologically beneficial to mix together different materialsand phases in a coating to achieve enhancements in frictional behavior, as achieved by the addition ofPTFE or MoS2 to metallic coatings, or the wear behavior, as achieved by using mixed-phase ceramicssuch as TiAlN (Knotek et al., 1986; Jehn et al., 1986; Bienk et al., 1995), TiCN (Bienk et al., 1995; Sproul,1994), TiBN (Gissler, 1994), etc. The concept of utilizing mixed phases is now being taken forward usingthree main approaches: multilayer coatings, nanocomposite coatings, and functionally graded coatings.

23.6.1 Multilayer Coatings

The idea to use a laminated structure to enhance the toughness of materials frequently occurs in nature,for example, in the shells that protect animals. In man-made bulk materials, it has been used for manydecades, for example, in the pearlite structure in carbon steels or the “sandwich” construction of laminatedglass. The multilayering concept applied to coatings is now regarded as one of the most promising meansof performance enhancement. The mechanism of that enhancement can take many forms, including:

1. The use of one interlayer or several interlayers to reduce the mismatch in mechanical or chemicalproperties between coating and substrate, in order to enhance the adhesion of the coating

2. Multilayered structures to control the residual strain and therefore the stress within coatings, mostcommonly to enhance the effective adhesion

3. Alternating layers that can act as crack-stoppers either by introducing layer boundaries to stopcracks or providing a tough medium through which propagation is curtailed

4. Compliant alternate layers that allow harder and more brittle layers to slide over each other duringbending, thereby avoiding the occurrence of the significant stresses which can occur with thickhard coatings when they deflect under load

5. Extremely thin nanoscale multilayers that constrain dislocation movement — and thereby impartextreme mechanical properties to the coating

6. Layers that provide distinct individual physical properties, such as a diffusion barrier or a thermalbarrier

Often, a number of the above enhancement mechanisms are combined in a single coating. Some of thekey historical developments in this field are outlined below.

Work on PVD multilayers was carried out in the late 1970s and early 1980s by Springer and Catlett(1978) and Springer and Harford (1982). They attempted to build on the earlier models of Koehler(1970), which predicted that high-yield-strength materials could be fabricated by alternating thin layersof high-shear-modulus material with thin layers of low shear modulus. The model was based on theinhibition of dislocation formation and mobility. They investigated Al/AlαOγ films deposited by a pulsedgas process in an electron beam gun system, and showed that a Hall-Petch type relationship was obeyed

for yield stress based on the layer spacings. Bunshah and co-workers also deposited microlaminatecoatings by evaporation techniques, both metal/metal (Bunshah et al., 1984) and metal/ceramic (Sanset al., 1983) couples being studied. In general, an improvement in mechanical properties was observedfor decreasing layer thicknesses.

In recent years, the literature and, at a certain level, the theoretical understanding of multilayer filmshas grown, in particular with regard to compositionally modulated superlattice thin films (Barnett, 1993;Barnett and Shinn, 1994). However, many of the theories developed are applicable to highly epitaxial orsingle-crystal layers; and while considerable progress has been made in understanding the theoreticalaspects of superlattice hardening and strengthening mechanisms, their value in practical tribologicalsystems is somewhat limited. There are several reasons for this, not the least of which is the difficulty ingrowing true superlattices (i.e., epitaxial layers) on real polycrystalline engineering components.

A more practical approach is to consider how the requirements of a surface differ at different locationswithin it — that is, at the interface with the substrate, within the coating itself, and at its surface, asshown by Figure 23.18. Multilayer coatings offer the possibility of designing the surface according to suchrequirements (Holleck, 1986, 1990; Holleck and Schier, 1995; Voevodin et al., 1996b). This is the conceptbehind the functionally graded coatings discussed in Section 23.6.3.

Holleck tended to use multilayers of different ceramic materials (e.g., TiC and TiB2), selected on thebasis of their dominant bonding mechanism (i.e., metallic, covalent, or ionic). He demonstrated improve-ments in hardness, indentation toughness, adhesion, and wear performance under optimized layerthickness conditions (Holleck et al., 1990), and attributed this, in part, to the crack deflection and stressrelaxation mechanisms for the TiC/TiB2 system.

This will pertain to many kinds of contact, especially where cyclic, fatigue-inducing conditions prevail.The benefits are no longer defined in terms of the early Koehler-type arguments based on the maximi-zation of yield strength. Rather, they relate to the macro behavior of the multilayer stack. However, theconcept of the alternation of high- and low-shear-modulus layers does seem to provide benefits beyondthose described even by Holleck. One can cite, for example, the considerable benefits exhibited bymultilayer diamond-like carbon and metal carbide films (Matthews and Eskildsen, 1994) or multilayerTiN/Ti films (Leyland and Matthews, 1994; Bull and Jones, 1996; Ensinger et al., 1992). In these cases,a different argument must be developed to fully explain the benefits observed.

One way to consider the behavior of these films is to model the coating under a normal force, whichcan be a point or distributed load, causing the coating to deflect. This will be accompanied by deformationof the substrate, which ideally should remain elastic. If one considers the bending stresses induced in thecoating under such a normal load, the level of maximum stress will increase with thickness for a givendeflection and radius of curvature; see Figure 23.19.

FIGURE 23.18 Tribologically important properties in different zones of the coated surface.

The diagram is highly schematic and treats the coating in two dimensions as a beam, taking no accountof the influence of the substrate on the stress distribution. Nevertheless, it is illustrative because it showsthat many multilayers, if bent to a similar radius, would individually see a much lower stress than onethick layer. This concept only works if the layers can effectively slide over each other, that is, in the mannerof a sheaf of paper or a paperback book being bent. In effect, one sees another benefit of multilayercoatings; that is, that the role of one of the sets of alternate layers can be to offer a shear zone to permitthe harder and more brittle layers to deflect under load without fracture. Thus, a straight line scribedon such a cross-section would deflect as shown in Figure 23.20. This ensures that each of the hard layerswill be subject to a much lower maximum bending stress than would be the case for a thick film bentto the same radius, although clearly the deflection under the same load would be greater due to the lowerstiffness of the surface coating. That is not necessarily an issue, because for many tribological contactswith thin surface coatings, it is the substrate that provides the main load support; the coating is presentto provide a hard, low-friction outer layer to reduce abrasive and/or adhesive wear.

Thus, one can now see why multilayer coatings based on Ti/TiN and DLC/WC, for example, haveproved successful in many practical contact conditions. The former have shown promise in erosiveconditions (Leyland and Matthews, 1994), while the latter are now used on gears and bearing surfacesthat see cyclic contact conditions (Matthews and Eskildsen, 1994). Because this concept is relatively new,the development of models to permit layer optimization in terms of relative thickness and mechanicalproperties has been limited. The constraint on both the hard and soft layers may be that the elastic limitshould not be exceeded, although this criterion seems more important for the hard layers because if theyyielded, then rapid fracture would ensue.

The softer layers will be present to give an easy shear; that is, they should have a low shear modulus,which equates to a low elastic modulus. Also, these layers should have a long elastic strain to failure,meaning that yield should not occur. Although in this case plastic yielding is to be avoided, especially incontacts subject to cyclic loading and fatigue, the main purpose will be to prevent the harder layers fromexceeding their yield stress. There are some interdependent variables in the optimization of layer thick-nesses. For example, because the layers also carry a load normal to the surface, they will be subject to

FIGURE 23.19 (a) Thin hard coatings on soft substrate generate lower stresses in the coating and at the coating-substrate interface, compared with (b) a thick hard coating with the same deflection.

FIGURE 23.20 A multilayer coating with alternate hard and soft layers can allow deflection to occur under loadwithout yielding of the harder layers. They effectively slide over each other, with shear occurring in the soft layer.The pattern of shear is illustrated by the line through the film, which was initially straight in the unloaded condition.

H

H

H

S

S

S

H = hard layerS = soft layer

Load

Substrate

compressive stresses, which will tend to squash the layers, in particular the softer ones. This will have theeffect of limiting their maximum thickness; otherwise, they may collapse and provide inadequate supportfor the harder layers within the coating. While an increase in the thickness of the layers will increase therelative sliding distance of one hard layer over another, it will not necessarily increase the shear strain —which is a ratio between the sliding distance of one hard layer over the next and the thickness. The idealseems to be to incorporate a large number of thin layers, thereby ensuring that the load support is notcompromised. In the case of DLC/WC coatings produced commercially, layer thicknesses of 20 nm aretypical for sliding and rolling contacts, while for Ti/TiN films, layers 1 and 2 µm thick have proved effectivein erosion conditions (Leyland and Matthews, 1994). The optimization of layer thickness will thus dependon the application and the loading conditions. In erosive or abrasive contacts, for example, a greater overallcoating thickness is often needed, especially for coarse and hard third bodies.

The benefits of the multilayering of relatively hard and relatively elastic layers has recently beendemonstrated in a specially developed cyclic impact test (Bantle and Matthews, 1995; Voevodin et al.,1995c). In that work, a relatively soft substrate (316 stainless steel) was used. This deformed plasticallyunder repeated loads of 900 N applied at a frequency of 8 Hz by a 6-mm-diameter tungsten carbide ball,while coatings comprising multilayer stacks of Ti, TiN, TiCN, TiC, Ti/DLC, TiC, and Ti/DLC showedgood wear performance, although not optimized for adhesion. Further tests using machines such as this,which apply the kinds of impact load often encountered in machine applications, will allow the fullbenefit of multilayered coatings to be achieved through further optimization.

23.6.2 Nanocomposite Coatings

As mentioned earlier, the most significant driver in the development of new tribological coatings hasbeen the realization that increased hardness should not be the sole goal when seeking to optimize coatingsfor enhanced wear performance (Leyland and Matthews, 2000). It is now clear that in most tribologicalcontacts, the elastic modulus, hardness, and fracture toughness should be considered together in definingthe likely material response of a surface to a loading condition. This is, in fact, the secret behind thesuccess of many of the multilayered coatings previously discussed; that is, they provide a means ofcombining hardness with an ability to deform under load without cracking — they are tough.

Another key route to the production of coatings that combines toughness characteristics with hardnessis to control their grain size. Several researchers have, in recent years, concentrated on this aspect ofcoating development (Veprek, 1999; Musil, 1999; Niedorfer et al., 1999; Mitterer et al., 1999; Hauert et al.,1999; Vaz et al., 1999; Musil et al., 1999a,b; Musil et al., 1998; Musil and Vlcek, 1999; Musil and Regent,1998; Benda and Musil, 1999; Misina et al., 1998; Zeman et al., 1999). Thus, the understanding ofnanocomposite coatings represents both a challenge and an opportunity, not only to coating processdevelopers, but also to those involved in modeling the deformation and failure behavior of such films.

Veprek (1999) and Musil (1999) have produced comprehensive review papers on the subject ofsuperhard and nanocomposite coatings. According to Veprek, superhard refers to materials whose hard-ness exceeds 40 GPa, and these can be divided into those with intrinsic hardness, such as diamond (70 to100 GPa), cubic boron nitride (~48 GPa); and those with extrinsic hardness, which is determined by theirmicrostructure. One group of these consists of the multilayer and superlattice structures discussedpreviously. The other group comprises the nanocrystalline (nc) composites. Examples are nc-MnN/a-Si3N4 (where M is Ti, W, V, and other transition metal nitrides, and a-Si3N4 is an amorphous siliconnitride), nc–TiN/BN, nc–TiN/TiB2, and others with a hardness over 50 GPa. In the case of nc-TiN/SiNα,the hardness is claimed to reach 105 GPa (Veprek, 1999).

While Veprek has tended to emphasize the nc-metal nitride/amorphous nitride route to the productionof superhard nanocomposites, Musil has concentrated on the nc-metal nitride/metal coating approach.Examples of coatings in this latter category are ZrCuN (Musil et al., 1999c) and TiAlN (Musil and Hruby,1999).

To date, much of the research on the ceramic/metal and ceramic/ceramic nanocomposite coatings hasemphasized investigations of their grain size and microstructure (e.g., through X-ray diffraction techniques)

and the evaluation of hardness using nanoindentation. Comparatively little work has been done on theirtribological properties. That is not the case for the nanocomposite coatings based on metal carbides andamorphous carbon, which have been extensively researched by Voevodin and co-workers (e.g., Voevodinet al., 1997a,b; Voevodin et al., 1998a,b; Voevodin et al., 1999a,b,c). In dry ball-on-disk tests, nanocrys-talline TiC-amorphous DLC composites showed a friction coefficient of 0.15 against steel, compared to0.5 to 0.6 for TiN and CrN, and 0.3 to 0.4 for TiC. The coating was also tough and withstood staticloading tests without cracking much more effectively than the single-phase TiC and even DLC (Voevodinand Zabinski, 1998b). Similarly, nanocrystalline WC/amorphous DLC composites showed low wear ratesand friction coefficients of about 0.2 against steel. A particularly encouraging feature of these coatingswas that the pulse laser deposition process used allowed them to be produced at less than 300°C. WC2

coatings with an extremely fine grain size of 1 to 3 nm could be obtained (Voevodin et al., 1999c).As pointed out by Mitterer and co-workers (Mitterer et al., 1999), it is not only the nanoscale grain

size that is important in providing the advantageous properties of nanocomposite materials. The inter-facial strength between the grains will also be important. To avoid unstable crack propagation, an interfaceof high cohesive strength is needed. Mitterer et al. state that this can be obtained from those compoundsthat show a high miscibility gap in the solid state, but a certain chemical affinity to each other to formhigh-strength grain boundaries. This can be achieved, for example, with the quasi-binary systems TiN-TiB3 and TiC-TiB2, both showing a eutectic type, as well as for other similar transition-metal nitride/borideor carbide/boride films.

Various authors have examined the Ti-B-C-N system for nanocomposite coating production. Furtherto the early work of Holleck (1986) and Holleck and Schultz (1988), other authors (Mitterer et al., 1990;Ronkainen et al., 1990, 1991, 1992; Gissler, 1994) have also investigated nanograined coating productionbased on the TiN/TiB2 and TiC/TiB2 quasi-binary phase systems, using both magnetron sputtering andelectron-beam PVD, with promising results in various tribological applications. While the recent work ofMitterer et al. (1998, 1999) has shown the importance of cohesive interfacial strength, other studies byRebholz et al. (1998, 1999b,d) of aluminum in the Ti-B-N phase system have demonstrated that coatingswith high hardness (≥30 GPa) but relatively low elastic modulus (≤300 GPa) can be produced, whichhave a high H:E ratio. Thus, the toughness is improved through better coating/substrate elastic propertymatching, showing enhanced tribological behavior in both laboratory wear tests and practical cuttingapplications, where a life twice that of TiN can be observed (Rebholz et al., 1999b). Further work bythese authors (Rebholz et al., 1999a) on adding small amounts of boron (≤5 at%) to TiN coatings toreduce coating grain size, and on metallic/intermetallic/ceramic Ti-Al-B coatings (Rebholz et al., 1999c)to reduce coating elastic modulus, has shown excellent results in laboratory wear tests.

23.6.3 Functionally Graded Coatings

In some ways, functionally graded coatings can be regarded as the logical progression from multilayeredcoatings.

In effect, the multilayer stack described at the end of Section 23.6.1 represents an excellent exampleof a functionally graded coating — where the coating is designed to utilize specific layers to impartdesired properties at specific levels within the coating. This is illustrated in Figure 23.21. Such a conceptcan even be achieved with just two layers; for example, an electroless nickel coating can be used as a loadsupporting layer for a hard PVD TiN or CrN coating, that improves the abrasive wear performance andenhances the corrosion resistance (Vetter et al., 1993; Leyland et al., 1993; Bin-Sudin et al., 1996). It iseven possible to include within this approach the concept of pretreating a substrate surface to producea hardened outer layer, which can then support the coating more effectively. Such approaches have beentermed hybrid or duplex coatings (Matthews et al., 1995; Matthews, 1997; Bell, 1997).

Many research groups are now examining the possibility of grading the composition, and even thestructure, of coatings to meet the needs of specific applications. For example, Savan et al. (1999) havedeveloped a coating that combines a relatively hard TiAlN phase and a relatively soft MoS2 phase withincreasing MoS2 toward the outer surface, a coating that might have benefits in dry machining applications.

Goller et al. (1999) have reported that TiN coatings with small amounts of MoSα (up to 8 at%) canextend coating drill lifetimes compared to TiN coatings alone.

One of the benefits of grading the composition of coatings has been shown to be the ability to enhancethe effective adhesion of DLC-based coatings that contain metal additions. Improvements have beenmeasured under local high contact loads (Voevodin et al., 1997a) or even under repetitive impact con-ditions (Bantle and Matthews, 1995), as well as under loaded sliding conditions (Voevodin et al., 1995b;Voevodin et al., 1996b).

23.6.4 Adaptive Coatings

The concept of “intelligent” materials is one that is gaining increasing attention in the research commu-nity, and this is equally true among those who are developing coatings. An intelligent coating can bedefined as one that adapts to its operating environment. While appearing at first sight to be a ratherexaggerated “blue-skies” concept, such adaptability is in fact a property that has existed in many coatingsfor some decades. For example, coatings designed to resist high-temperature corrosion in gas turbineengines are designed to form a stable oxide on their surface during operation.

A similar observation could be made for some of the most recently developed cutting tool coatings,based on, for example, TiAlN, which produces a stable oxide during high-temperature cutting operations.This effect has been further enhanced in TiAlN coatings that contain additions of yttrium and/orchromium to help to further stabilize the oxide film (Munz 1997; Savan et al., 1999). In some ways, thediamond and DLC films discussed in Chapter 24 can also be regarded as adaptive in the sense that theformation of a temperature-induced graphite-like layer in the contact is thought to be the source of thelow friction demonstrated by such coatings.

One of the main advocates of adaptive coatings that can provide advantageous properties across arange of operating temperatures and environments has been Zabinski who, together with co-workers,has been working on solid lubricant coatings capable of operating at temperatures from ambient up to600°C and beyond, even up to 800°C. At such temperatures, conventional oils and greases cannot beused, and graphite and metal dichalcogenides such as MoS2 and WS2 have historically been used asalternatives. Their lubricious nature is generally attributed to weak interplanar bond strength; thiscondition may not prevail at higher temperatures at which the materials may oxidize. Zabinski et al.(1995a,b) have reported that this oxidation may become a problem at 350°C, and have carried out researchinto the deposition of coatings with additives such as graphite fluoride (CFα) in order to extend thepermissible operating temperature up to 450°C. They utilized a laser-pulsed deposition method andshowed that the films were less sensitive to moisture than pure WS2, achieving friction coefficients of0.01 to 0.04 against a stainless steel counterface in dry air. Although not adaptive in the usual sense ofthe word, these coatings point the way toward coating materials and processes that can achieve a wideroperating range.

FIGURE 23.21 (a) A functionally graded multilayer coating design to utilize specific layers for distinct properties.(b) A particular multilayer coating tested by Voevodin et al. (1996).

To provide lubricity in the 500 to 800°C temperature range, other materials must be considered (Johnet al., 1999). Oxides such as ZnO, along with fluorides such as CaF2 and BaF2, are lubricious in thisregime, due to their low shear strength and high ductility. The oxides and, to a lesser extent, the fluoridesare also chemically stable in air at high temperatures. At ambient temperatures, however, these materialsare brittle and subject to cracking and high wear rates. According to John et al., no single material isknown to be lubricious over the entire range from ambient temperature to 800°C. Thus, the obviousanswer to producing a lubricant coating that can operate over a broad temperature range is to combinelow- and high-temperature lubricant materials into either a composite or layered structure. The objectiveis for the properties of both materials to be manifested in their respective temperature ranges. John et al.(1999) examined both composite and layered structures of CaF2 and WS2 grown by pulsed-laser depo-sition on steel and TiN-coated steel substrates, and found that the films were lubricious at up to 500°C.They found that the CaF2 and WS2 materials interacted to form CaSO4, among other compounds. Thus,the coatings, in effect, adapted to their environment by forming a low-friction outer layer.

The concept of combining coating materials to extend the operating range using several other com-pounds has been applied; for example, MoS2 with PbO, and WS2 with ZnO, to provide lubrication overa wide temperature range (Donley and Zabinski, 1992; Zabinski et al., 1996; Zabinski et al., 1992; Walcket al., 1994; Zabinski et al., 1995a,b: Walck et al., 1997; Zabinski and Donley, 1994). In these cases, themetal dichalcogenides reacted with the oxides to form PbMoO2 and ZnWO4, which were found to belubricious at high temperatures. The problem with these coatings was that they were adaptive only onheating; that is, the low-temperature lubricant materials reacted to form a high-temperature lubricant —but the reaction was irreversible. After returning to room temperature, the lubricious properties haddisappeared. Work is proceeding to control the reactions using layered structures with diffusion barriercoatings, and thereby to develop coatings that will function through multiple temperature cycles.

23.7 Concluding Remarks

This chapter began by setting the scene on the present state-of-play regarding advanced metallic andceramic coatings and their applications. The macroscopic and microscopic responses of surfaces toloading conditions were discussed, and, in the latter case, the energy accommodation concept wasintroduced as a means of understanding the critical property requirements that the surfaces of bulkmaterials and also coatings must exhibit in order to have adequate functionality. This highlighted theimportance of toughness, in addition to the usually emphasized property — hardness.

When a surface has a coating applied, the coating and substrate act together as a composite and thisallows each to provide specific benefits. For example, the coating can add a low friction layer, or a diffusionor thermal barrier, while the substrate can impart a load-supporting function or a bulk toughness, whicha monolithic component made just of the coating material might not possess. In effect, with newdevelopments in tribological understanding and in coating production methods, we are now at the dawnof a new era in effective tribological design, in which concepts and practices that were previously onlytheoretically possible can now be applied — giving enhanced and predictable levels of reliability andtherefore well-controlled and optimized life-cycle costs for coated components.

References

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