Mechanics of Coatings

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MECHANICS OF COATINGS

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TRIBOLOGY SERIES 17

MECHANICS OF COATINGS

edited by

D. DOWSON, C. M.TAYLOR and M. GODET

Proceedings of the 16th Leeds-Lyon Symposium on Tribology held at The lnstitut National des Sciences Appliquees, Lyon, France 5th - 8th September 1989

ELSEVIER

Amsterdam -Oxford - New York -Tokyo 1990 For the Institute of Tribology, Leeds University and The lnstitut National des Sciences Appliquees de Lyon

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QElsevier Science Publishers B.V., 1990

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CONTENTS

ix xi

Introduction Epitaph Session I

Session I1

Session I11

Session IV

Session V

Session VI

Session VII

Conference themes On the elastic constants of thin solid lubricant films M.N. GARDOS Frictional properties of lubricating oxide coatings M.B. PETERSON, S.J. CALABRESE, S.Z. LI and X.X. JIANG Elastic and viscoelastic analysis of two multiply layered cylinders rolling over each other with coulomb friction J.J. KALKER Theory Analysis of damage mechanism using the energy release rate P. DESTUYNDER and T. NEVERS Integrity of wear coating subjected to high-speed asperity excitation F.D. JU and J.-C. LIU Coating design methodology

and L. VINCENT Experiments Reduction in friction coefficient in sliding ceramic surfaces by in-situ formation of solid lubricant coatings A. GANGOPADHYAY, S. JAHANMIR and B.E. HEGEMANN In-situ engineered oxide coatings H-S HONG and W.O. WINER A morphological study of contact fatigue of TiN coated rollers H.S. CHENG, T.P. CHANG and W .D. SPROUL Soft coatings 1 A full solution to the problem of film thickness prediction in natural synovial joints D. DOWSON and J. YAO Finite element analysis of EHD lubrication of rubber layers A. GABELLI and B. JACOBSON Analyses of shear deformations between cylinders with and without surface films J.W. K A " E L and T.A. DOW Solid lubricants Effects of microstructure and adhesion on performance of sputter-deposited MoS, solid lubricant coatings P.D. FLEISCHAUER, M.R. HILTON and R. BAUER Role of transfer films in wear of MoS, coatings S. FAYEULLE, P.D. EHNI and I.L. SINGER Assessing the durability of organic coatings M.J. ADAMS, B.J. BRISCOE, A.L. CARTER and P.J. TWEEDALE Rough coated surfaces Effect of surface coatings in a rough normally loaded contact Ph. SAINSOT, J.M. LEROY and B. VILLECHAISE Elastic behaviour of coated rough surfaces J.I. McCOOL Hardness Scratch tests on hard layers A.G. TANGENA, S. FRANKLIN and J. FRANSE A theoretical approach of hardness distribution in rigid perfectly plastic coated materials M.L. EDLINGER and E. FELDER Damage mechanisms of hard coatings on hard substrates: a critical analysis of failure in scratch and wear testing R. REZAKHANLOU and J. von STEBUT

M. GODET, Y. BERTHIER, J.-M. LEROY, L. FLAMAND

3

15

27

37

45

53

63

73

81

91

I03

1 1 1

121

129

139

151

157

169

175

I83

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VI

Session VIII

Session IX

Session X

Session XI

Session XI1

Session XI11

Session XIV

Multilayer theory Stress determination in elastic coatings and substrate under both normal and tangential loads J.M. LEROY and B. VILLECHAISE A survey of cracks in layers propelled by contact loading D.A. HILLS, D. NOWELL and A. SACKFIELD A statistical approach for cracking of deposits : determination of mechanical properties A. MEZIN, R. RAMBUARINA and J. LEPAGE Youngs modulus Youngs Modulus of TiN and Tic coatings L. CHOLLET and C. BISELLI A method for in situ determination of Youngs modulus of deposits J.P. CHAMBARD and M. NIVOIT Biomechanics Use of a polyelhylene coating to improve the biotribology of a hemiarthroplasty implant M. LABERGE, G. DROUIN, J.D. BOBYN and C.H. RIVARD Comparison of theoretic 31 and experimental values for friction of lubricated elastomeric surface layers under transient conditions J.R. GLADSTONE and J.B. MEDLEY A preliminary investigation of the cushion bearing concept for joint replacement implants D.D. AUGER, J.B. MEDLEY, J. FISHER and D. DOWSON Soft coatings 2 The influence of elastic deformation upon film thickness in lubricated bearings with low elastic modulus coatings D. DOWSON and Z. JIN Frictional mechanism in uncoated and zinc-coated steel sheet forming - theoretical and experimental results V. SAMPER and E. FELDER Effect of nitrogen ion implantation on the friction and wear properties of some plastics M. WATANABE, H. SHIMURA and Y. ENOMOTO Failure mechanisms The effect of continuous Au sputter deposition on the enhancement of growth of transfer particles K. HIRATSUKA, L.L. HU and T. SASADA Elasto-plastic finite element analysis of axisymmetric indentation using a simple personal computer M. GUEURY, P. BAGUR, R. REZAKHANLOU, J. von STEBUT Oxides The oxide film and oxide coating on steels under boundary lubrication Y-W ZHAO, J-J LIU and L-Q ZHENG New tools and models A survey of research in acoustic microscopy applied to metallurgy

and A. SAIED Detection of interface defects in layered materials by photothermal radiometry M. HEURET, E. VAN SCHEL, M. EGEE and R. DANJOUX A low cycle fatigue wear model and its application to layered systems. A.G. TANGENA

J. ATTAL, R. CAPLAIN, H. COELHO-MANDES, K. ALAMI

195

203

209

217

223

233

24 1

25 1

263

27 1

28 I

289

295

305

315

323

329

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vii

Session XV Wear 1 The inter-relationship between coating microstructure and the tribological performance of PVD coatings S.J. BULL and D.S. RICKERBY Identification and role of phosphate coatings for tribological applications G.T.Y. WAN, R.J. SMALLEY and G. SCHWARM Protective coatings for application in seawater M.F. LIZANDIER, E. LANZA, A. SEBAOUN, A. GIROUD and P. GUIRALEDENQ

Lubrication influences on the wear of piston-ring coatings J.C. BELL and K.M. DELARGY Wear behaviour of Cr2 0 and Al, 0, plasma sprayed coatings under lubricated and non-lubricated conditions R. VIJANDE, F.J. BELZUNCE, J.E. F E R N ~ D E Z , M.C. PEREZ and A. RINC6N

Influence of viscoelastic parameters on coated bearing behaviour Y.T. SUN, B. BOU-SAID and B. FANTINO

Morpholigical aspects of the friction of hot-filament-grown diamond thin films P.J. BLAU, C.S. W S T , L.J. HEATHERLY and R.E. CLAUSING Factors affecting the sliding performance of titanium nitride coatings F.E. KENNEDY and L. TANG The mechanism of failure of coatings in roller tests J. VIZINTIN

Session XIX Deformation Measurements of thin films adhesion and mechanical properties with indentation curves J.L. LOUBET, J.M. GEORGES and Ph. KAPSA Soft metallic coatings in metal forming processes P. MONTMITONNET, F. DELAMARE, E. DARQUE-CERETTI and J. MSTOWSKI Deformation and fracture of hard coatings during plastic indentation D.M. ELLIOTT and I.M. HUTCHINGS

Coating evaluation methods : a round robin study. H. RONKAINEN, S. VARJUS, K. HOLMBERG, K.S. FANCEY, A.R. PACE, A. MATTHEWS, B. MATTHES and E. BROSZEIT The effect of dynamic loads in tribometers - analysis and experiments H. HESHMAT Written contributions List of authors List of delegates

Session XVI Wear 2

Session XVII Viscoelasticity

Session XVIII Hard coatings

Session XX Coating evaluation

337

35 1

359

37 1

379

389

399

409

417

429

435

445

453

465 475

489 48 1

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I X

INTRODUCTION

The Sixteenth Leeds-Lyon Symposium on Tribology was held at the Institut National des Sciences AppliquCes de Lyon from the 5th to the 8th of September 1989. It was dedicated to the memory of the late Professor Daniel Berthe who had contributed much to earlier symposia in Lyon. As all other Leeds-Lyon Symposia, it discussed only one topic which this time was "Mechanics of Coatings". The subject was chosen because it seemed timely to bring together men of different disciplines connected with coatings to find ways of extending the industrial use of these coatings particularly where tribology was concerned.

It was indeed necessary to get mechanical engineers, surface and volume physicists, theorists and experimenters, applicators and users together to measure how coatings could mean different things to different men and also to note the different criteria retained for the qualification of coatings in each speciality. It was too much to expect that a single symposium could get these specialists to agree on a given course, it was however reasonable to hope that each would take stock of that difference and in time take steps to reduce it. We hope to have met that goal.

Some papers were invited and the "Call for Papers" was a great success. A Review Board was set up to examine the abstracts. Close to 60 papers were programmed which meant that triple sessions had to be held on one of the three days of the symposium. These papers were dispatched in 20 sessions which discussed theory, experimental data, hard and soft, smooth and rough coatings, solid lubricants and oxides, coating properties and coating property measurements, wear and failure mechanisms of coatings, coatings in bio-mechanics and coating evaluation.

The meeting was attended by more than 150 delegates from 17 countries. This is an all time high for Lyon and of course a great encouragement to all of the organisers. We were, as always on these occasions, delighted to host a large and very active contingent from Leeds headed by Professor Dowson and Dr. Taylor. Imperial College was also very present both inside and outside of the sessions and it was very rewarding to see both familliar and new faces in such large numbers.

After the opening session, on Tuesday, September 5th. delegates were taken to VilliC-Morgon where the conference banquet was held. We were priviledged in having Mrs Daniel Berthe with us. The dinner was prepared by two well known Chefs of the Lyon area, MM. Troump and Marguin and the wines were chosen by a special commission set up by the wine experts of the Laboratoire de MCcanique des Contacts. Some commotion was caused as one of the courses was rather more "Lyonnais" than expected but we hope that those who were surprised when they found out what they had eaten have since forgiven us.

Water jousts were organised on Thursday evening for the delegates at Givors which is 25 kms south of Lyon. Jousters mounted, one at a time, on the pontoon of a small flat bottomed boat. Two boats come side by side and each jouster tries to "duck" his opponent by pushing him overboard with a 5 to 7 meter lance. The fight takes place in a basin which was originally part of the port of Givors on the Rhone and which was walled in relatively recently. This local form of tournament reaches back to roman times and the Givors club has been at it quite a while as it celebrated its 100th anniversary in 1986. The performance was followed by a reception given by the Lord Mayor of Givors. AU of this was accompanied by music which punctuated each ducking or lance breaking with traditional energetic tunes.

Thanks to Andy Olver of Westland Helicopters, the joust spirit was carried over the next day during the Friday evening barbecue organized by the laboratory technicians. Minor changes in the

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X

rules had to be introduced as delegates served as both boats and jousters and brooms were used instead of lances. The Saturday trip to beautiful Auvergne took us to Ambert, la Chaise-Dieu and le Puy and marked the end of the 16th symposium.

None of this could have happened without a heavy commitment of all members of the Laboratory. To all, thanks are due for their energy, their good humour and inventiveness.

We are looking forward to going to Leeds to attend the XVIIth symposium on Vehicle Tribology.

Maurice GODET

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xi

Professor Daniel Berthe

The 16th LeedsLyon Symposium was, dedicated to our friend and colleague Professor Daniel Berthe who died in February after a long illness. All of the members of the Laboratoire de MCcanique des Contacts de 1INSA along with many of our friends from Leeds and elsewhere wanted to honour his memory and chose this way of doing it. It seemed fitting, as he contributed greatly to many of the earlier symposia as one of the co-editors of the Proceedings and as an organizer of the event itself when it was held in Lyon.

He joined the Laboratory in 1968 and worked from the start on roughness effects in hertzian contacts and on many of the problems such as cavitation and fatigue associated with roughness. A man of great intuition and tremendous finesse, he tackled that very difficult subject in ways both orikinal and productive. Discrete, curious of all things scientific, he welcomed discussions with each one of us and helped us all in the formulation of our ideas and in the advance of our research.

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SESSION I CONFERENCE THEMES

Chair man : Professor M. Godet

PAPER

PAPER

PAPER

(i) On the elastic constants of thin solid lubricant films

(ii) Frictional properties of lubricating oxide coatings

(iii) Elastic and viscoelastic analysis of two multiply layered cylinders rolling over each other with coulomb friction

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3

Paper I (i)

On the elastic constants of thin solid lubricant films

M.N. Gardos

The literature was searched to find numerical values of elastic constants for layered hexagonal lubricants (e.g., graphite, MoS2, NbSeg, GaSe and InSe) and for selected hardcoat underlays such as TiN and Tic. The data are critically reviewed for accuracy and usefulness to a tribologist. An attempt is made to correlate the electronic and crystal structure of selected hexagonal solid lubricants with the degree of ionic-to-covalent bonding that both characterizes the crystal systems and determines the magnitude of their elastic constants. The basic intent is to provide the appropriate Young's modulus ( E ) , shear modulus (G) and Poisson's ratio ( u ) data to computer diagnoslicians, who need t o substitute realistic values into their iterated programs predicting the actual concentrated contact stresses associated with hardcoat/softcoat layered systems.

1 INTRODUCTION

Under concentrated contact conditions, e.g., in rolling element bearings o r gears, modifying the bearing surfaces by the addition or hard and sofL lubricant coatings will alter the size and distribution of the contact stresses.

It is well-established that under static loading, the region of maximum Hertzian stresses is below the surface o r an uncoated bcaring material. The effect of high friction/ traction and normal forces at the surface is to increase the magnitude of the maximum shear stress and raise its region of occurrence closer t o the contact surface (Pig. 1). Fedorchenko (1) showed that when an increasing number of hard inclusions are embedded in the surface of a softer matrix, and each progressively larger set carries the same normal load, the peak octahedral stress will a l s o be broughL toward the surface (Fig. 2). As the number and size of the inclusions grow, one recognizes that the upper limit is reached when a uniform, high elastic modulus hardcoat of a given thickness is formed. Superimposing a solid lubricant softcoat on top of an already hardcoated bearing surface will further alter the size and the location of the maximum (modified) HerLzian stress region. A thin coating of material more elastic than its substrate will lower the apparent elastic modulus in the contact zone and lower the maximum stress. Reversibly glassifled lubricating oils in concentrated contacts achieve the same erfect. The obvious advantages of hardcoat/oil and hardcoat/ softcoat combinations have been well-documented in the literature (1 through 6 ) .

A large number of investigators brought forth models which represent, static (2 Lhrough 8 ) and dynamic-sliding (9,lO,ll) conditions. These models predict the depth, magnitude and distribution of stress fields as a function of substrate/coating(s) elastic constants and

coating thicknesses. The constants of particular interest are the Young's modulus ( E l , the Poisson's ratio ( u ) and, occasionally, the shear modulus (G).

By the use of computers, the predictive models can iterate a variety of likely combinations, depending on arbitrarily selected, single and compound values of E, G, and u. However, accurate prognosis of the real contact sLresses associated with a given moving mechanical assembly, e.g., meshing steel gears coated with an 23 urn thick layer of reactively ion-plated TiN and further covered with 400 nm of sputtered MoS2. requires the subsLitution of accurate elastic constants to obtain an exact solution. These data become an important part of predicting the wear life of each assembly component.

A search of the tribology literature for reliable values of E, G, and u for hard and soft solid lubricant films resulted in very few pieces of data. However, recent papers from basic science and optical-electronic-magnetic device coatings literature have established the basis for the estimation of elastic constants for layered hexagonal solid lubricant species such as graphite, MoS2, NbSe2, InSe and GaSe as well as for their hardcoat underlays, e.g., Tic and TiN. The present work is intended as an overview of elastic constant fundamentals and data on selected soft and hard lubricant layers. An emphasis is placed on the highly anisotropic nature of layered-hexagonal solid lubricant materials, especially those in the thin film form, where the basal planes of the crystal 13 tes are more- or-less aligned in the plane of sliding o r rolling.

A brief overview of micromechanical techniques employed to measure elastic constants i s also given. Nearly all of these measurement techniques have been used for metallic layers and hardcoats only. The application of these techniques are suggested

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4

0.2 0.1 0 -0.1 -0.2 -0.3 Y

ROLLING DIRECTION -

0.2 0.1 0 -0.1 -0.2 -0.3 Y

UNDER STATIC LOAD

0.2 0.1 0 -0.1 -0.2 -0.3 Y

OCTAHEDRAL STRESS

Fig. 1 Concentrated contact stress contours under various friction/traction conditions; shaded areas represent the region of maximum stresses.

for hardcoat/softcoat combinations, where the soft lubricant layers consist of sputtered films or layers transferred from s e 1 f - lu b r i c a t i ng c omp o s i t e s .

Pig. 2 Distribution and depth of maximum subsurface stresses, as influenced by a progressively large number of hard surface inclusions from (a) to (c), under the same total load; from (1).

2 THEORY AND PHACTrCE OF ELASTIC CONSTANT MEASUREMENTS

2.1 A Brief Overview of Elasticity Fundamentals

As described in Belenkii, et al's excellent treatise (131, the relationship between the stress and strain tensors in the weak (elastic) deformation of a solid is known as Hooke's Law:

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5

where dik and ulm are the synunetric, second rank stress and strain tensors, respectively. The fourth-rank tensor Ciklm is known as the elasticity tensor and its components are called the elastic moduli o r elastic constants.

A completely isotropic solid has two elastic constants (E and G), cubic crystals have three constants (C11, C12, C44) and crystals with hexagonal synunetry typical of layered hexagonal materials such as graphite and MoS2 are described by five elastic constants.

The elastic constants for the hexagonal synunetry are:

1.

3.

5.

cxxxx = cyyyy (C11 = C22) 2. czzzz (C33)

4 * cxxyy (C12) cxxzz = cyyzz (c13 = c23)

cxzxz = cyzyz (C55 = C44)

where the following compact notation system is used: xx-1, yy-2, 22-3, yz-4, xz-5, xy-6. The same notation is applicable to cubic crystals, although the crystallographic axes do not subtend the same angle. The coordinate (xy) axes of a cube are Cartesian, while the crystallographic xy axes of a hexagonal crystal form 120 angles. The xz and yz angles are, of course, 90' in both cases.

A hexagonal crystal is, therefore, characterized by Cll, C12, C13, C33 and C44:

-~ C11 and C12 represent the binding interaction between atoms within layers, whose elastic properties in a synunetry plane are isotropic; one can best imagine these constants as those involved in idealized tensile testing of a single, hexagonal layer. They determine E and u in the synunetry (basal) plane (i.e., E l l ) .

C13, along with the C33 and C44, determines interlayer binding and, to a large measure, u on application of pressure along the synunetry axis. It is difficult to measure, because the sample has to be polished at an angle to the planes of the layers: there is, therefore, a wide variation in the C13 values generated by different investigators. In many cases the value of C13 was not determined directly, but estimated from measurements and on the basis of certain assumptions.

s.33 governs E in the direction perpendicular to the basal planes. This elastic constant represents the interaction between layers under compression (i.e., El).

g44 represents the stresses in the basal plane which are generated as a result of shear between neighboring planes being tangentially displaced relative to each other (i.e., C44 = G). Its value can vary widely, depending on the number of basal plane defects, o r incorrect sequencing of the layers relative to each other.

It is shown in (13) that

c13 (2) u = --

c11 f c12

where u represents the change in the dimensions of a hexagonal crystallite in the plane of its layer on application of pressure along the synunetry axis. One can also imagine u of a hexagonal crystallite as the thinning of the platelet as it is being elongated by tensile forces within the basal plane. This interesting symmetry becomes obvious when we note that u may be generally defined for a uniaxial stress state as

where cX is the strain in the direction of the applied stress and cY is the strain in the orthogonal direction. Theoretical thermo- dynamic considerations dictate that in the absence of phase changes 0 < u < 0.5 (11). Since it is a positive number, the above equation indicates that an applied negative (i.e., compressive) stress results in orthogonal strains which are positive (i.e., tensile).

Due to the highly anisotropic nature of hexagonal crystals, the well-known relationship connecting E with G:

E G =

2(1 - u) (4)

which is applicable to isotropic (e.g., cubic) materials only, cannot be used with layered hexagonal materials.

As also discussed in (131, relationships have been established between the experimental values of the elastic constants and atomic- level parameters of model calculations which determine the bonding-controlled interactions between atoms. The elastic properties and phonon spectra of the crystals are often calculated using model force constants governing the displacement of a given atom from its equilibrium position and the forces exerted on this atom by all other atoms of the crystal lattice. The simple assumption is frequently invoked that the interactions between atoms exist along the line connecting them and depend only on the interatomic distance. Since the interaction forces between atoms decrease quite rapidly with distance, lattice dynamics in a specific crystal structure can often (but not always) be confined to nearest neighbors. For example, the consideration of only the central forces does not apply t o graphite. It is, however, generally valid or a large number of layered transition metal dichalcogenides and analog compounds (13) . 2.2 Elastic Constants for Hexagonal Crystals

Table I contains a compilation of elastic constant values for layered crystals, collected by Belenkii, et a1 (13) from a large variety of literature sources. The data were obtained by

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(x 1011 dynes/cm2)

106 t- 2+ 144 t- 20ff

3.9 f 0.4

3.56 -. >1.8

Table I. Elastic constants of layered crystals (in units of 1011 dynes/cm2), from (13).

Reference

(16) (17)

(18) (19) (20)

Hexagonal Crystals Rhombohedra 1 Cry s t a 1 s

Elastic Constant Graphite GaS TiSe2 TaSe2 NbSe2 SnSe2 GaSe InSe

15.7

3.3

1.5

3.6

0.8

10.3

2.9

1.2

3.4

0.9

7.3

2.7

3.0

3.6

1.2

12

4.2

-

3.9

1.4

22.9

10.7

5.4

1.9

19.4

9.1

_ _ _ _

4.2

1.8

10.3

-__-

2.8

1.8

c11

c12

c13

c33

c4 4

106

18

1.5

3.7

0.018- 0.035

Table 11. Selected elastic constants of graphite.

investigations of compressibility, neutron scattering, calculations of specific heat, light scattering and ultrasound propagation. Although the accuracy of certain values is open to question due to the difficulties associated with the measurement techniques, the quality of the test specimens and disagreement among investigators (see thorough discussion in Ref. 13), the following important observations can be made from the data in Table I:

Elastic Constant

c11

1. A strong anisotropy of the elastic properties of layer crystals is indicated by the fact that Cll, C12 >> C13, C33, C44. A large difference exists between E within the basal plane (Ell) and the elastic modulus normal to it (EL)

2. The strongest elastic anisotropy is exhibited by graphite. The anisotropies of the constants of the other compounds are considerably less. As discussed later in this paper, the degree of reduction depends on the ionicity of bonding and the resultant crystal structure changes. The covalent or ionic bonding character plays an especially important role in determining the anisotropy of the elastic constants. For example, the size of C44 (i.e., G) tends to be higher in more ionic compounds, also yielding some clues as to the anticipated magnitude of the critical resolved shear stress. None of these aspects are immediately obvious from the data, as assembled and discussed in (13).

c3 3

0.42 ? 0.2 0.405* 0.292**

c4 4

+ Experimental data. ++ Theoretical calculations. * Canadian natural graphite. ** Pile graphite, with smaller crystalline

size than Canadian graphite above.

do not represent all the reliable data that can be found in the literature.

Conspicuously absent from Table I are the elastic constants of MoS2, a solid lubricant very important to tribologists. Those data were taken from (21). along with additional information about NbSe2 (Table 111).

It is of interest to compare elastic constants calculated or measured for single crystals of graphite and MoS2 by basic researchers with data generated during quests to determine G and E of these materials for tribological applications.

3. The wide variation in the C44 of graphite stems from the quality of the specimens. As shown in (151, a reduction in the number of basal dislocations increases C44 significantly. The data in Table I1 indicate that not only the quality of the crystal but also the crystallite size influence C44: a smaller crystallite size appears to produce a smaller shear modulus. It is further shown in Table 11 that in spite of the fact that the moduli of graphite have been investigated the most extensively for decades, the values in Table I still

Martin et a1 (22) evaluated G of graphite and Nos2 powders in an epoxy resin binder. Thin beams of the respective composites were cast and the moduli determined by the vibrating reed technique. Test beams with lubricant percentages higher than 32% by volume could not be made, because the liquid epoxy used at such ratios was not sufficient to hold all the solid lubricant particles together, or to produce a

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Table 111'. Elastic constants of 2H-MoS2 and 2H--NbSe2 (in units of 1.011 dynes/cm2), from (17).

Elastic Constant

c11

c12

c13

c33

c44

Layered Crystal

MoS2

23.8

-5.4

2.3

5.2(5.2)**

1.89(1.90)**

NbSe2

10.6/17.1*

1.4/ 7/9*

3.1/-0.2*

5.4(5.5)**

1.95(1.98)**

* Results of separate estimation methods ** Values in parentheses are rigid layer

constants.

mixture wet enough to be cast into specimens The authors in (20) recommended that G for higher percentages of lubricant may be found by extrapolation. Although extrapolation to 100 percent lubricant content involves an error of an unknown magnitude, such curve fitting was attempted here by the use of a desktop computer in Fig. 3. Substituting 100 in place of x in the polynomials, G M ~ s ~ - 2.82 x lo6 psi - 19.44 GPa = 1.94 x 10l1 dynes/cm2, and Ggrapp*te - 2.79 x lo6 psi 19.24 GPa - 1.92 x 10 dynes/ cm2. These data are in good agreement with C44 (MoS2) in Table 111, but off by a substantial amount of the various C44 (graphite) values in Tables 1 and 11.

30

20

10

0 "' 0 10 20 30 40 50 60 70 80 90 100

MoS2 GRAPHITE

Yo LUBRICANT BY VOLUME y = 1.8498 + 1.0863e-2x + 2.5284e-3x-2 R*2 = 0.997 y = 1.8004 + 4.4180e-2x + 2.1705e-3~"2 R"2 = 0.999

Fig. 3 Shear modulus of MoS2 and graphite, as extrapolated from data in (22).

In their exceptionally thorough work, Griffin and Ward (23) determined G M ~ s ~ by (a) extruding polystyrene - MoS2 powder compositions to achieve a highly oriented lay of the pigment particles by the laminar flow of the shear process, and (b) assessing the degree of orientation (and thus G) by ultrasonic pulse propagation.

The best G M ~ s ~ values were determined as:

x - direction (Gx) - 17.94 GPa = 1.79 x 1011 dynes/cm2

y -- direction (Gy) - 31.67 GPa = 3.17 x 1011 dynes / cm2

* z - direction (Gz) 12.94 GPa 5 1.29 x 10l1 dynes/cm2

If the system were comprised of perfectly aligned hexagonal crystallites in the direction of extrusion, then Gx - Gy > GZ. data indicated that Gx < Gy, there appeared to be some peculiarity of MoS2 particle fracture and shear anisotropy in the extrusion direction of the xy plane and perpendicular to it. As such, it is not unreasonable to average G,, Gy and G,. The average value is 20.83 GPa = 2.08 x 1011 dynes/cm2 and it is also in good agreement with C44 in Table 111, and with G#,s2 from Martin et al's work in (20).

Since the

Unfortunately, Martin and coworkers calculated EMoS2 from G M ~ s ~ , by using Eq. (4). The appli cabil i ty of thj s formula to hexagonal crystall ites, even jf they were more or less random in a cast epoxy beam, is highly questionable. As previously mentioned, it does not apply at all t o individual packets of hexagonal platelets, or an ensemble of these particles whose basal planes are generally parallel with the bearing substrate, forming a complete layer. Such alignment has been achieved f o r MoS2 by special sputtering techniques (24, 25, 26) and by run-in of either sputtered (27) or polymer-bonded (28) films. Monolithic, polycrystalline graphite compacts have also undergone basal plane alignment on their surfaces during sliding (29). If the degree of alignment can be determined by X-ray diffraction or, at the very least, by SEM photomicrography (e.g., see Fig. 4 for a schematic and Fig. 5 for the actual appearance of sputtered and run-in MoS2 films), then the best value for E must clearly lie somewhere between Ell (i.e., C11) and El (i.e., C33). Everything depends on the extent of the alignment. Judging from Fig. 5, the bending of the platelets t o a partial lay of the basal planes normal t o the applied load calls f o r some scaling to establish a realistic value for a compressive modulus. In the case of perfect alignment, El should be used.

Crystal 1 i te a1 ignment-induced modulus changes are equally significant to those who use graphite fiber reinforcements in structural and self-lubricating composites. The compound moduli of these composites can be measured by convent) onal tensile, compression or flexural test methods. However, to correlate theory with practice, the modulus of a fiber or a laminating plate, as well as that of the matrix, must be known (28, 29). Therefore, any of the previously described, alignment-induced modulus changes importanE to the tribologist are equally important to the composites technologist. The data in Fig. 6 from (30) exemplify the anisotropy of moduli that exist in the various forms of graphite.

The 150 x lo6 psi - 1034 GPa = 103.4 x 10l1 dynes/cm2 in Fig. 6 for a single crystal under

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8

MODULUS (106 PSI)

(103 PSI) STRENGTH

TYPE I EDGE PLANES EXPOSED

SINGLE BULK

CRYSTAL GRAPHITE "A" " C 150 5 1.5

3000 4 1.5

TYPE II BASAL PLANES EXPOSED

~

4

10 - LOW FRICTION ~

150 120

50 500

-&=z---z -r/ q / p r t / ' ' I - ,,p$;FF-$Z; -- TRANSITION REGION

INTERFACE //// / / / l////,l,//

BURNISHED TYPE I

,

u

Eig. 4 Schematic r e p r e s e n t a t i o n of MoS2 f i l m s s p u t t e r e d i n d i f f e r e n t c r y s t a l l i t e o r i e n t a t i o n s , and t h e e f f e c t s of burn ish ing on Type I f i l m s ( c o u r t e s y of Dr. P. D . F l e i s c h a u e r , The Aerospace Corpora t ion) .

C GaSe InSe MoS2 NbSe2

0.01* 0 . 0 9 0.30 0 .13 O.ll**/ 0.26***

Fig . 5 SEM photomicrograph of v a r i o u s l y run- i n , Type I s p u t t e r e d MoS2 f i l m depic ted schemat ica l ly i n F ig . 4 [ c o u r t e s y of D r . P. D . F l e i s c h a u e r , The Aerospace Corp. , a l s o see (27) I .

b a s a l t e n s i o n i s very c l o s e t o t h e experiment- a l l y measured C GPa = 3 .45 x lo i i dynes/cm2 is a l s o c l o s e t o t h e C33 ( E l ) values i n Tables I and 11. a l s o provides c l u e s a s t o t h e modulus d i f f e r - ences of p r e - g r a p h i t i c carbon versus h i g h l y ordered p y r o l i t i c ( p o l y c r y s t a l l i n e ) g r a p h i t e , both i n t h e random-bulk and t h e a l igned-bulk ( p l a t e o r f i b e r ) form. I t is of i n t e r e s t t o n o t e t h a t t h e i n d i c a t e d alignment of c r y s t a l - l i tes i n F i g . 6 resemble more t h e u l t r a h i g h modulus, pi tch-based f i b e r s . I n PAN-based f i b e r s , t h e b a s a l p lanes of c r y s t a l l i t e s a r e arranged around t h e per iphery i n an "onionskin" f a s h i o n , where t h e middle is more r a d i a l l y ordered , l i k e spokes i n a wheel (33 , 3 4 ) .

( E l l ) ; the 5 x lo6 p s i = 34.47

Fig . 6

The e x t e n t and type of a l ignment i n g r a p h i t e f i b e r s prepared from d i f f e r e n t

F ig . 6 C r y s t a l s t r u c t u r e , e l a s t i c and s t r e n g t h p r o p e r t i e s of var ious forms of g r a p h i t e , from ( 3 2 ) .

p r e c u r s o r s l e a d t o n o n l i n e a r e l a s t i c e f f e c t s . For example, t h e r e i s a dramat ic i n c r e a s e i n modulus a t h i g h e r t e n s i l e s t r a i n s , due t o s t ra in- induced r e o r i e n t a t i o n of the b a s a l p l a n e s p a r a l l e l t o t h e d i r e c t i o n of t h e t e n s i l e f o r c e along t h e f i b e r a x i s ( 3 5 ) . The magnitude of i n c r e a s e was found l a r g e r f o r pi tch-based f i b e r s , which have a much l a r g e r modulus and h i g h e r degree of p r e e r r e d c r y s t a l l i t e o r i e n t a t i o n from t h e o n s e t .

The t e n s i l e force- induced al ignment i n g r a p h i t e f i b e r s and t h e shear-induced al ignment of s p u t t e r e d MoS2 f i l m s b r i n g up t h e q u e s t i o n of u s i n g c o r r e c t va lues of u f o r layered s o l i d l u b r i c a n t s . Mart in e t a1 ( 2 2 ) may have com- pounded t h e i r e r r o r of t r y i n g t o employ Eq. (4) f o r c o n v e r t i n g GM0s2 t o E#,s2 by assuming a U M ~ S - 0.30. If one presumes t h a t t h e only u whicz has meaning t o a t r i b o l o g i s t i s t h e one descr ibed by Eq. (21, then i t can be seen from t h e d a t a i n Table I V , t h a t Mart in e t a1 were o f f by a f a c t o r of two t o t h r e e . Genera l ly , t h e l a r g e r t h e e l a s t i c a n i s o t r o p y i n t h e c r y s t a l , t h e smaller i s t h e va lue of u . A l o t more

Table I V . P o i s s o n ' s Rat io f o r S e l e c t e d , Layered S o l i d L u b r i c a n t s , as c a l c u l a t e d by Eq . ( 2 ) , us ing d a t a from Tables I , I1 and 111.

** Using average va lues of c l o s e s t C 1 1 and C22 from Tables I and 111, and C13 = 3 . 1 x 1011 dynes/cm2.

*** Using f i r s t column of v a l u e s from Table 111.

I

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needs to be said about the Poisson effect in coatings comprised of these crystallites, but not without experimental evidence.

The magnitude of v has a great deal to do with interfacial stresses transmitted to a coating/substrate region by a concentrated contact load. Lancaster and Wade (36), using Matthewson's analysis from (lo), showed that these stresses increase with increasing v and with reduced h/a ratio (h = film thickness; a = Hertzian half-width of contact, i.e., contact radius). In case of resin bonded solid lubricants or transfer films from polymeric self-lubricating composites, the v of the polymers is equally significant. The data in (36) and (37) indicate that the v of several polymers is both strain and strain rate dependent. At higher strains, polymers such as polypropylene, polyethylene and nylon exhibited progressively higher Poisson's ratios. At high rates of loading polymers and their composites become more incompressible, as the time of loading approaches the molecular relaxation time. A progressively increasing rate of loading increases the effective value of u to 0.5. That, in turn, leads to higher interfacial stresses. The respective elastic constants of both the pigment and the binder enter the theory and practice of determining the viscoelastic behavior of polymer layers with particulate inclusions (38).

2.3 Elastic Constants for Hard Coatings

Recent reviews of modern measurement methods for the mechanical properties of thin films deal virtually with hardcoats and metallic layers only e.g., see (39). These coatings are analyzed either in a free-standing mode (where the substrate has been etched away), or still firmly attached to it. The former methods include measurement of the internal-stress- induced curling of microbeams (401, deflecting these thin beams electrostatically and by mechanical vibration (411, by deflecting with a nanoindenter (42), or simply by tensile testing of the free standing films themselves (43). With the coating still attached to the substrate, the techniques include ultramicroindentation (44, 45, 46), the use of small vibrating reeds (46, 47, 48), and Bri1louj.n scattering (49). The authors in (49) aptly state that the techniques which separate the coating from its underlay are fraught with error. The elastic properties of films may change when removed from the substrate.

Therefore, the best values for TiN and Tic, deposited onto various steels, were taken from (46) and (48). In the earlier work, E T ~ N for reactively sputtered, stoichiometric compositions was found to be 640 GPa, with v = 0.25; ET~C = 460 GPa and v = 0.17. conference, our Swiss colleagues are reporting E T ~ N ranging from 384 to 446 GPa, depending on the type of the steel substrate (48). They also report a higher than previous E T ~ C of 555 GPa. The literature values compared with their own data indicated the following: (a) E tends to increase with the year of measurement, (b) Ecoating < Ebulkr and (c) the magnitude of E is both substrate type and coating microstructure

At this

dependent. These findings agree with those of Petersen and Guarnieri (41) in that 400 to 800 nm thick sputtered and chemically vapor deposited hardcoats such as Si02, Si3N4, Sic and chromium generally exhibited E different from those of the equivalent bulk materials. In most cases, the values were considerably lower than the bulk. The same thing was observed for W, Cu and A 1 thin films ( 5 0 ) .

Since hard, polycrystalline coatings are deposited with one or another low index crystallographic plane more or less preferentially aligned parallel with the substrate plane (51), one would expect El (the modulus of interest in a Hertzian contact) to depend on the specific coating texture. As shown by Perry for TiN, it is indeed the case (51). E for the (110) and (111) plane of TiN is lower and that for the (100) plane is higher than the bulk value. It is due to the fact that the magnitudes of C11, C12 and C44 are different; there are also residual stresses causing plastic deformation in a polycrystal-- line, textured film and the resultant altera- tion of the apparent moduli. For the bulk, Perry used E T ~ N = 640 GPa and V T ~ N = 0.2.

3 DISCUSSION

The anisotropic nature of the various layered crystal structures and their elastic constants is controlled by the magnitude of the atomic interaction within a layer (intralayer bonding) versus bonding between layers (interlayer bonding). It is the electronic structure of a material which determines the crystal structure. Only certain crystal structures exhibit the requisitely strong intralayer and weak interlayer bonding required for a platelet-like solid lubricant.

Inasmuch as the crystal- and electronic- structure related behavior of layered transition metal dichalcogenides and the GaSe/ InSe compounds have begun to be treated with some understanding (3, 13, 52 through 571, one needs to point out here only that the magnitude of the elastic constants can serve as a guide with respect. to the usefulness of these compounds a:; solid lubricants.

For example, the nature of bonding in NbSe2 is -25% more ionic than in MoS2 (52), and InSe is -15% more ionic than in GaSe (54). The data in Tables I, I11 and IV indicate that the constants of the more ionic materials (a) are less anisotropic, (b) exhibit greater interlayer attraction i.e., C44(InSe) > C44(GaSe); C44(NbSe2) > C44(MoS2); note that lower C44(G) values portend lower critical resolved shear stresses, (c) are more incompressible i.e., u(1nSe) > u(GaSe); v(NbSe2) > v(MoS2), therefore their sputtered layers under Hertzian loads would transmit more interfacial stress (which tends to enhance delamination), and (d) are probably weaker in ternis of basal plane load carrying capacity, due to the respectively lower C11 (intralayer bonding) values. On the whole, GaSe and InSe appear to be promising as solid lubricants [see (56)], complementing graphite and MoS2 for lower load, ultralow friction applications.

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I n some c o n t r a s t with t h e s e p r e d i c t i o n s i s t h e behavior of g r a p h i t e . While t h e h igh C 1 1 ( E l l ) values i n d i c a t e very high load- c a r r y i n g capac i ty [ n o t e t h a t t h e a r i t h m e t i c average of t h e t h e o r e t i c a l and experimental va lues i n Table I1 i s 1250 GPa, which is t h e E of diamond], t h e u l t r a - l o w C44 ( G ) va lues b e l i e t h e high shear s t r e n g t h of g r a p h i t e i n vacuum. I n t h e absence of any i n t e r c a l a t e d mois ture o r o t h e r i n t e r c a l a n t s , one would expect G t o be cons iderably h i g h e r . While on t h e fundamental b a s i s it makes sense t h a t b a s a l p lane d e f e c t s (vacancies , s t e p s , k i n k s , s t a c k i n g f a u l t s ) would reduce 71 - bonding i n t e r a c t i o n between p lanes and t h u s cause t h e p r e v i o u s l y d iscussed reduct ion i n C44 (and t h e s h e a r s t r e n g t h ) , they should be a b l e t o do s o only i n t h e presence of some i n t e r c a l a n t . I n t h e c a s e of MoS2 and NbSe2, t h e e f f e c t s of s u l f u r o r selenium vacancies i n t h e b a s a l p lane should be e x a c t l y t h e o p p o s i t e . The presence of po in t -defec t - caused dangl ing bonds t h e r e would i n c r e a s e i n t e r l a y e r bonding [ o r bonding t o a s u b s t r a t e , see (57)1, and t h u s i n c r e a s e C44 (G). Moisture would have t h e same e f f e c t , b u t no t f o r t h e same reason. Hydrogen-bond-induced a t t r a c t i o n of oxidized MoS2 edge (o r d e f e c t i v e b a s a l p l a n e ) s i tes bonded with water molecules would equal ly change ( i n c r e a s e ) t h e apparent C44 ( G ) of a water-vapor- s a t u r a t e d sample. These a r e t h e cavea ts one must keep i n mind when t r y i n g t o use e l a s t i c cons tan t va lues a s guides f o r s o l i d l u b r i c a n t s e l e c t i o n . A g r e a t d e a l depends on t h e q u a l i t y of t h e sample, a s w e l l a s t h e atmosphere and t h e technique of e l a s t i c cons tan ts de te rmina t ion .

One must be e q u a l l y c a r e f u l about t h e t es t technique used f o r h a r d c o a t s , i . e . , us ing t h e r i g h t technique f o r t h e r i g h t specimen. A s pointed out i n ( 4 2 ) , t h e f ree- -s tanding microbeam va lues measure E of s i n g l e c r y s t a l f i l m s o r b a r s c u t from a boule f o r a p a r t i c u l a r c r y s t a l l o g r a p h i c d i r e c t i o n , on a given h a b i t p l a n e . The nanoindenter , on t h e o t h e r hand, measures an average va lue . Thus, t h e two techniques cannot be compared d i r e c t l y . On tex tured T i N and T i C ( o r diamond) h a r d c o a t s , where t h e p o l y c r y s t a l l i t e s a r e t u r b o s t r a t i c a l l y a l igned u s u a l l y wi th a g iven , low index p lane ly ing p r e f e r e n t i a l l y p a r a l l e l wi th t h e s u b s t r a t e , t h e two methods should a l s o y i e l d d i f f e r e n t resul ts . The number of d e f e c t s wi th in t h e cubic c r y s t a l s t r u c t u r e e q u a l l y inf luences t h e magnitude of t h e e l a s t i c c o n s t a n t s . Dis loca t ion and p o i n t d e f e c t s caused by c u t t i n g , g r inding and p o l i s h i n g of a sample can i n c r e a s e t h e moduli. I n c o n t r a s t , h igh temperature annea l ing t h e vacancy complexes, which a c t a s c e n t e r s of d i s s i p a t i o n of t h e mechanical v i b r a t i o n energy, w i l l reduce t h e magnitude of t h e c o n s t a n t s ( 5 8 ) .

T r i b o l o g i s t s would p r e f e r modulus measurement techniques whose specimens resemble t c i b o c o n t a c t s . I n d e n t a t i o n methods, bu t without t h e u s e of sharp ( e . g . , Vickers) i n d e n t e r s , might o f f e r t h e b e s t avenue of approach. Researchers from B a t t e l l e Columbus Labora tor ies s p u t t e r e d a manganin p r e s s u r e t ransducer on t h e bottom, c y l i n d r i c a l r o l l e r of a d u a l - d i s c r i g des ign; t h e t o p , crowned r o l l e r was coated wi th MoS2 or Tic of var ious t h i c k n e s s e s ( 3 ) . Upon r o l l i n g one specimen

a g a i n s t t h e o t h e r under l o a d , they were a b l e t o show t h e g e n e r a l magnitude of stress changes and, i n p a r t i c u l a r , t h e r e d u c t i o n i n stresses a s a f u n c t i o n of s o l i d l u b r i c a n t l a y e r s . These and o t h e r , b a l l - t y p e i n d e n t a t i o n methods have a l r e a d y been shown capable of measuring t h e e l a s t i c response of t h i n polymer l a y e r s on hard s u b s t r a t e s (38 , 59 , 6 0 ) . Ref in ing both t h e a p p l i c a b l e t e s t appara tus and t h e a s s o c i a t e d t h e o r e t i c a l c a l c u l a t i o n s a r e suggested f o r achiev ing a more d e f i n i t i v e c o r r e l a t i o n between t h e theory and p r a c t i c e of mul t icoa ted , concent ra ted c o n t a c t s .

4 CONCLUSIONS

An e x t e n s i v e l i t e ra ture survey was conducted t o f i n d va lues of e l a s t i c moduli f o r layered hexagonal l u b r i c a n t s and f o r s e l e c t e d hardcoat under lays . The d a t a are c r i t i c a l l y reviewed i n terms of accuracy and u s e f u l n e s s t o a t r i b o l o g i s t .

An emphasis was p laced on t h e h i g h l y a n i s o t r o p i c behavior of layered-hexagonal s o l i d l u b r i c a n t materials, e s p e c i a l l y those i n t h i n f i l m form, where t h e b a s a l p l a n e s of t h e c r y s t a l l i t e s a r e more-or-- less a l i g n e d i n t h e p lane of s l i d i n g o r r o l l i n g . It i s shown t h a t i n l i e u of experimental va lues der ived by micromechanical techniques , t h e l i t e ra ture d a t a may be used w i t h t h e fo l lowing cavea ts : ( a ) f o r h i g h l y a l i g n e d ( run- in) c o a t i n g s t h e e l a s t i c c o n s t a n t s a p p l i c a b l e t o s i n g l e c r y s t a l s may a l s o be a p p l i c a b l e t o t h e c o a t i n g s , ( b ) depending on t h e degree of hexagonal c r y s t a l l i t e s i z e and al ignment i n t h e d e p o s i t e d l a y e r s ( t o be determined exper imenta l ly o r a t l e a s t e s t i m a t e d ) , e i t h e r t h e EL o r compound Ell/EL v a l u e s should be used i n t h e stress c a l c u l a t i o n s , and ( c ) an i n c r e a s e i n t h e i o n i c i t y of i n t r a - and i n t e r l a y e r bonding i n c r e a s e s G and u. There a r e l a r g e r c o a t i n g / s u b s t r a t e i n t e r f a c e stresses wi th h igher u, a t l e a s t i n t h e s t a t i c c o n t a c t mode.

With r e s p e c t t o h a r d c o a t s , t h e i r e l a s t i c c o n s t a n t s are dependent on t h e d e p o s i t i o n and pos t - t rea tment c o n d i t i o n s , which c o n t r o l t h e chemis t ry , t e x t u r i n g and d e f e c t s of t h e l a y e r s . More o f t e n than n o t t h e r e s p e c t i v e moduli are lower than t h e publ i shed bulk v a l u e s .

A g r e a t d e a l of work remains i n developing improved, micromechanical i n d e n t a t i o n techniques which can c o r r e l a t e concent ra ted c o n t a c t theory and p r a c t i c e of s o l i d l u b r i c a t e d bear ing s u r f a c e s .

5 ACKNOWLEDGEMENTS

This work was performed under t h e auspices of t h e " D e t e r m i n a t ion o f T r i bo 1 o g i c a 1 Fund amen t a 1 s of S o l i d Lubr ica ted Ceramics" program, DARPA Order No. 5177, AF'WAL Cont rac t No. P33615-85-C- 5087, wi th B . D . McConnell a c t i n g a s t h e AFWAL P r o j e c t Engineer ,

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Ibid, "Part 1L: Turbine Bearings and the Associated Solid Lubricants Research, Vol. 1 : Summary".

J. F. Dill, et at, "Rolling Contact Fatigue Evaluation of Hardcoated Bearing Steels", Proc. Third. Int. Solid Lubr. Conf., ASLE SP-14, p. 230 (1984).

M. N. Gardos, "Physical and Chemical Stabilization of Steel Bearing Surfaces with Titanium Nitride and Titanium Carbide Hard Coatings", Proc. Industry-Academia- Government Workshop, March 12-14, 1984, Vanderbilt U., Nashville, TN, Bureau of Mines, U.S. Dept. of the Interior, 1985.

M. N. Gardos, "The Tribooxidative Behavior of Rutile-Forming Substrates", in New Materials Approaches to Tribology: Theory and Applications. (Ed.'s L.E. Pope, et al), Mat. Res. SOC. Symp. Proc. Vol. 140, p. 325 (1989).

D. Barovich, et al, "Stresscs on a Thin Strip o r Slab with Different Elastic Properties from that of the Substrate due to Elliptically Distributed Load", Int. J. Engng. Sci., 2, p. 253 (1964).

P. K. Gupta and J. A. Walovit, "Contact Stresses Between an Elastic Cylinder and a 1;ayered Elastic Solid", Trans. ASME, J. Lubr. Tech., Ser. F., 96, p. 250 (1974).

R. L. Mehan, et al, "Properties of a Compliant Ceramic Layer", J. Mat. Sci., 16, p. 1131 (1981).

M. J. Matthewson, "Axi-Symmetric Contact on Thin Compliant Coatings", J. Mech. Phys. Solids, 29, p. 89 (1981).

Y. P. Chin and M. J. Hartnett, "A Numerical Solution for Layered Solid Contact Problems with Application to Bearings", Trans. ASME, J. Lubr. Tech., 105, p. 585 (1983).

R. Solecki and Y. Ohgushi, "Contact Stresses Between Layered Elastic Cylinders", Trans. ASME, J. Tribology, 106, p. 396 (1984).

G. L. Belenkii, E. Yu. Salaev, and R. A. Suleimanov, "Deformation Effects in I.ayer Crystals", Sov. Phys. Usp. 31, p. 434 (1988).

L. D. Landau, A. I. Akheiezer and E . M. LiCshj.tz, General. Physics-Mechanics and Molecular Physics, Pergamon Press, London, 1967.

15. E. J. Seldin and C. W. Nezbeda, "Elastic Constants and Electron-Microscope Observations of Neutron-Irradiated Compression-Annealed Pyrolitic and Single-Crystal Graphite", J. Appl. Phys., 41, pp. 3389 (1970).

16. E. S . Seldin, in Proc. 9thSennial Conf. on Carbon, Chestnut- Itill., MA, 1969 (Defense Ceram. Ino. Center, Columbus, OH), p. 59 (1969).

17. R. Nicklow, N. Wakabayashi, and I f . G. Smith, "Lattice Dynamics of Pyrolytic Graphite", Phys. Rev. B., 5 , p. 4951 (1972).

18. G. Dolling and B. N. Brockhouse, "Lattice Vibrations in Pyrolytic Graphite", Phys. Rev., 128, p. 1120 (1962).

19. C. Bownian and J. A. Krumhansl, "The Low-Temperature Specific Heat of Graphite", J. Phys. Chem. Solids, 6, p. 367 (1958).

20. K. Komatsu, "Particle-Size Effects of the Specific Heat of Graphite at Low Temperatures", J. Phys. Chem. Solids, 6, p. 380 (1958).

21. J. L. Feldman, "Elastic Constants of 2II--MoS2 and 211-NbSeg Extracted from Measured Dispersion Curves and Linear Compressibilities", J. Phys. Chem. Solids, 37, p. 1141 (1976).

22. J. T. Martin, C. 11. Balster and F. Abdulhadj., "Measurement of Shear Modulus for Solid Lubricants: Wear Life Coefficients for MoS2 in MoS2 Resins", Lubr. Eng., 28, p. 43 (1972).

23. G. J. L. GriCfjn and D. S. Ward, "Elastic Properties of Oriented Molybdenum Disulfide-Polystyrene Composites", ASLE Proc. Second Int. Conf. on Solid Lubrication, ASLE SP-6, p. 169 (1978).

24. P. D. F'leischauer, "Effects of Crystallite Orientation on the Environmental Stability and Lubrication Properties of Sputtered MoS2 Thin Films", ASLE Trans., 27, p. 82 (1984).

25. P. D. Fleischauer and R. Bauer, "Chemical and Structural Effects on the Lubrication Properties of Sputtered MoS2 Films", STLE Tribology Trans., 31, p. 239 (1988).

26. P. A. Bertrand, "Orientation of RP-sputtered-deposited MoS2 Films", J. Mater. Res., 4, p. 180 (1989).

27. M. R. Hilton and P. D. Pleischauer, "Structural. Studies of Sput Ler- depos ited MoS2 Solid Lubricant films", in New Materials Approaches to Tribolopy: Theory and Applications (Ed.'s L. E. Pope, et al), Mat. Res. SOC. Symp. Proc. Vol. 140, p. 227 (1989).

28. M. N. Gardos, "The Synergistic Effects of Graphite on the Friction and Wear of MoS2

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29.

30.

31.

32.

33.

34.

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42.

Films in Air", STLE Tribology Trans., 31, p. 214 (1987).

Anon, "X-ray Investigation of Structural Changes in Graphite Antifriction Materials During Friction", Izv. A.N. SSSK, Mekh. Mash., 4, pp. 179-184 (1963); Risley-.Trans.-1850-(9091.9F).

R. M. Jones and 11. S. Morgan, "Analysis of Nonlinear Stress-Strain Behavior of Fiber-Reinforced Composite Materials", AIAA J., 15, p. 1669 (1977).

B. Prasad and G. Iiermann, "Response of a Laminated Beam t o a Moving Load", ATAA J., 15, p. 142 (1977).

W. I%. Chambers, "Low Cost, High-Performance Carbon Fibers", Mech. Eng., 97, p. 37 (1975).

M. N. Gardos, et al, "Solid Lubricated Turbine Bearings: Part I - Preparation of 316C Lubricative Composites and Separators", Proc. 3rd Tnt. Conf. on Solid Lubrication, ASLIS SP-14, p. 248 (1984).

N. Ohmae, et al, "Atomic Configuration of Carbon Fibers Studied by Field Ion Microscopy", STLE Tribology Trans., 31, p. 481 (1988).

C. P. Beetz, Jr. and G. W. Budd, "Strain Modulation Measurements of Stiffening Effects in Carbon Fibers", Rev. Sci. Instrum., 54, p. 1222 (1983).

.I. K. Lancaster and D. J. Wade, "The Influence of Reversing Loads on the Performance of Self-Lubricating, Dry Bearings", Proc. 3rd Int. Conf. on Solid Lubrication, ASLIS SP-14, p. 296 (1984).

I. Krause, A . J. Segreto, and 1. Przirembel, "Poisson's Ratio for Viscoelastic Materials", Mat. Sci. Kng., 1, p. 239 (1966).

J. 11. M. van der Ljnder, P.E. Wierenga and E. P. lIonig, "Viscoelastic Behavior of Polymer Layers with Inclusions", J. Appl. Phys., 62, p. 1613 (1987).

D. A. Hardwick, "The Mechanical Properties of Thin Films: A Review", Thin Solid Films, 154, p. 109 (1987).

M. Mehregany, R. T. Howe, and S. D. Senturia, "Novel Microstructures for the In-Situ Measurement of Mechanical Properties of Thin Films", J. Appl. Phys., 62, p. 3579 (1987).

K. E. Petersen and C. R. Guarnieri, "Young's Modulus Measurements of Thin Films Using Micromechanics", J. Appl. Phys. 50, p. 6761 (1979).

T. P. Weihs, S. Hong, J. C. Bravman, and W. D. Nix, "Measuring the Strength and Stiffness of Thin Film Materials by Mechanically Deflecting Cantilever Microbeams", in Thin Films: Stresses and

Mechanical Properties. (Ed's. J. C. Bravman, et al), Mat. Res. SOC. Proc. Vol. 130, p. 87 (1989).

43. ibid, R . W. Hoffman, "Nanomechanics of Thin Films: Emphasis : Tensile Properties", p. 295.

44. Y. Tsukamoto, H. Yamaguchi, and M. Yanagisawa, "Measurements of Ultra- microindentation Hardness, Young's Modulus and Internal Stress", Thin Solid Films, 154, p. 171 (1987).

45. S. Hoshino, K. Fuji, N. Shohata, 1. Yaniaguchi, Y. Tsukamoto, and M. Yanagisawa, "Mechanical Properties of Diamondlike Carbon Films", J. Appl. Phys., 65, p. 1918 (1989).

46. Is. TGrgk, A. J. Perry, L. Chollet, and W. D. Sproul., "Young's Modulus of TiN, Tic, ZrN and IlfN", Thin Solid Films, 153, pp. 37-43 (1987).

47. G. L. Miller, M. Soni, and R. L. Fenstermacher, "A Technique f o r Investigating the Properties of Surfaces, Thin Films, and Interfaces by Means of a Mechanical Marginal Oscillator", J. Appl. Phys., 53, p. 979 (1982).

48. L. Chollet and C. Biselli, "Young's Modulus of TiN and Tic Coatings", this conference.

49. R. Bhadra, M. Grimsditch, and I. K. Schuller, "Elastic Constants of Metal- Insulator Superlattices", Appl. Phys. Lett. 54, pp. 1409-1441 (1989).

50. C. T. Rosenmayer, F. R. Brotzen and R. J. Gale, "Mechanical. Testing of Thin Films", in Thin Films: Stresses and Mechanical Properties, Mat. Res. SOC. Symp. Proc. Vol. 130, pp. 77-86 (1989).

51. A. J. Perry, "The Relationship Between Residual. Stress, X-ray ELastic Constants and Lattice Parameters jn TiN Films Made by Physical Vapor Deposj.tion", Thin Solid Films, 170, pp. 63-10 (1989).

52. A . Madhukar, "Structural Classification of Layered Dichalcogenides of Group IVB, VR and VIB Transition Metals", Solid State Commun., 16, p. 383 (1975).

53. R. Takagi, "Layer-Shaped Structures and Friction Characteristics of MoS2 Family" J. Jap. SOC. Prec. Eng., p. 104 (Nov. 1980).

54. A . Nakanishi and T. Matsubara, "Note on Ionicity of Layered Compounds Gas, GaSe and InSe", J. Phys. SOC. Jap., 51, p. 1339 (1982).

55. W. E. Jamison, "Intercalated Dichalcogenide Solid Lubricants", Proc. 3rd Int. Conf. on Solid Lubrication, ASLE SP-14, p. 73 (1984).

56. M. N. Gardos, "An Analysis of the Ga/In/WSe2 Lubricant Compact", ASLE Trans., 28, p. 231 (1985).

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57. P. D. Fleischauer, "Fundamental Aspects of the Electronic Structure, Materials Properties and Lubrication Performance of Sputtered MoS2 Films," Thin Solid Films, 154, p . 309 (1987).

58. P. A. Maksymyuk, et al, "Changes in the Elastic and Inelastic Properties of Annealed Indium Antimonide Crystals", Sov. Phys. Solid State, 30, p. 1656 (1988).

59. P. E. Wierenga and A. J. J. Franken, "Ultramicroindentation Apparatus for the Mechanical Characterization of Thin Films", J. Appl. Phys., 55, p, 4244 (1984).

60. A . Tonk, J. Sabot, and J. M. Georges, "Microdisplacements Between Two Elastic Bodies Separated by a Thin Film of Polystyrene", paper presented at the ASME/ASLE Joint Lubr. Conf., Hartford, CN, Oct. 18-20, 1983, ASME Paper No. 83-Lub-11.

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Paper I (ii)

Frictional properties of lubricating oxide coatings

M.B. Peterson, S.J. Calabrese, S.Z. Li and X.X. Jiang

A literature search was conducted to identify the properties of surface oxide films that make them effective in reducing friction, wear, and surface damage in sliding contacts. it was concluded that the formation of double oxides of rhenium, molybdenum, and boron would be most effective. A number of nickel, copper, cobalt, rhenium alloys were prepared and their friction temperature properties were compared with those of the oxide films which might be produced on the surface.

Based on these results

1. INTRODUCTION

In their book (1) published in 1907 Archbutt and Deeley stated that "the friction between most so called unlubricated metallic surfaces is, therefore, not a case of true friction between pure metals but between surfaces contaminated by chemically formed films such as oxide and sulfides. In other words the surfaces are partially lubricated. Under moderate pressures such films prevent actual adhesion of metal to metal". Since that time approximately 120 papers have been written concerning the role that the oxide played in tribological processes. These reports fall into four categories: pretreatments, boundary lubrication, high temperature solid lubrication with oxides, and dry sliding. Pretreatrnents are concerned with subjects such as protecting aluminum surfaces by anodizing and the oxidation of ferrous materials to assist in metal working. In boundary lubrication, oxygen present in the lubricant acts competitively with other additives in forming surface films which prevent wear under mild operating conditions and failure under extreme pressure conditions. or polymer films resulting from oxidation of the lubricant. It is also known that metal oxides react with lubricant additives to form metal soaps which are extremely effective lubricants. Metal oxides have been investigated as potential high temperature solid lubricants for temperatures above 350C where many conventional solid lubricants become ineffective. However the largest number of studies have been concerned with metals in dry sliding. Here the oxide determines whether seizing or effective sliding will result. From these studies it is clear certain oxides and conditions favor "mild wear" while others favor "severe wear" (2). The question can be raised whether this accumulated knowledge can be used to develop improved alloys (or lubricants)

Such films may be metallic oxides

'Tribology Group, National Institute of Standards and Technology, Gaithersburg, MD. 2Rensselaer Polytechnic Institute, Troy, NY. 31nstitute of Metal Research, Shenyang, China

which are more effective than those currently in use. Toward this objective, the literature has been reviewed to describe the necessary attributes of an effective oxide film and then to develop and evaluate alloys which might form such oxides under tribological conditions.

2. LITERATURE REVIEW-EFFECTIVE OXIDES

Several excellent reviews have been conducted on the subject of oxidative wear and the role of the oxide. Ludema ( 3 ) reviewed the scuffing literature with particular attention to the role of the oxide. He defines scuffing as a roughening of the surface by plastic flow and suggests that to be effective an oxide should:

(1) enhance the adsorption of selective

(2) be soft and ductile, ( 3 ) wear at a rate less than the

oxidation rate, ( 4 ) not be abrasive as debris, and (5) flake off to remove stresses.

species from the lubricant,

Quinn ( 4 , 5 ) gives a complete review of oxidation wear which was developed for low alloy steels but should be generally applicable for materials for which the wear process is similar. Basically, after run in, plateaus which were formed during run in, oxidize at the interface temperature and grow during the sliding process. thickness ( 1 ~ 3 pn) this oxide film breaks up by a fatigue process to form wear debris. All aspects of the wear process are analyzed and theoretical estimates made of the wear rates. He points out that the diffusion rates of oxygen and iron determine how long it will take to reach the critical film thickness. Thus, control of diffusion rates should control wear. He also points out that stronger oxide films would be beneficial. Batchelors ( 6 ) review was primarily concerned with the factors which affect the growth of oxide films. He points out that there are a number of factors other than diffusion rates which affect the growth (therefore the wear) of oxide films. Lattj-ce misfit causes severe strains in the lattice and form unstable oxides. An oxide in the vitreous

When reaching a critical

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state oxidizes at a lower rate than an oxide in the polycrystalline state since oxygen diffuses faster along the crystalline boundaries. Other factors of importance are the structure of the substrate, mechanical activation, liquid effects, and the mechanical and adhesive properties.

written with great clarity describes the various oxidation/wear models. H e suggests that theories should be confined to the major mechanisms which have been defined for specific materials without trying to work out the completely general problem. He also suggests to have a more stable oxide that:

Sullivan (7,8) in a comprehensive review

0

0

0

0

Of the number

Molar volume of the oxide should be greater than the molar volume of the metal so the film is in compression on the surface. The differential thermal expansion between the metal and the oxide should be small. Both the oxide and the metal should be capable of some plastic flow to relieve the stresses in film and interface. Oxides formed by double diffusion, e . g . , Fe30,, are more stable.

120 papers published between 1929-1984 a (25) merely noted that the oxide was

important in controlling the friction and wear of metals in sliding contact. The effect would be beneficial if it reduced adhesion but harmful if it increased corrosive wear. Such effects were noted in both lubricated and dry sliding.

2.1 Models

A large amount of work has been carried out to understand the process that leads to the formation and wear of oxide layers. Several independent processes have been identified which may or may not be mutually exclusive, These may be identified as sliding on a built up oxide, on an agglomerated debris layer of oxide or on a soft lubricating layer.

The built up oxide layers have received the most attention. Here the oxide grows based on oxidation kinetics and is removed by some wear or failure process (9-21). Four different behaviors have been identified as shown in Figure 1. In the graph labeled FO-FW an oxide is built up on the surface of the asperity in one pass and immediately removed on succeeding passes. For convenience this is called fast oxidation-fast wear. This might apply to the early stages of wear where an insufficient amount of oxide is produced or for a metal which produces a friable oxide layer which cannot withstand the frictional stresses. A second behavior, slow oxidation-fast wear (SO- FW), mostly due to Quinn who worked with low alloy s tee ls , shows that an oxide grows slowly as sliding proceeds until a critical film thickness is reached where the films break up and are removed as flake debris. The controlling factor is the rate of removal and the critical film thickness rather than the oxidation rate as in FO-FW. A third process might be called FO-SW. In this situation an oxide film is rapidly formed and is slowly worn away. When it is completely gone it is replaced by the original FO process. Such a

1 I

I

Tin* Tiwe

SO-SU

_ _ - hn h

L

L

Figure 1. Dynamic models of the formation and removal of oxide films during the wear process.

process has not been found for oxides but may be true for self lubricating composites or other cases where an extremely effective lubricant is used. The final model, SO-SW, refers to a process where the oxide grows slowly and wears away slowly. The resulting film thickness is the difference between the rate formation and the rate of wear. The rate of wear is a function of film thickness. If the film gets too thick the wear rate increases; if it becomes too thin, the rate of oxidation increases. is maintained. This mode is thought to occur if the film failure mode is wear (abrasion) rather than fatigue or fracture. It would also seem to apply when soft lubricating oxide films are formed.

It is clear that these models are not unique for a given materials combination. Changes in operating conditions and atmospheres can cause a change from one model to another.

yield minimum wear if the oxidation process could be controlled so that only the amount needed was produced. To achieve this end it is necessary to identify oxide films which wear slowly. Efforts should be directed to alloys in which this model describes events since it is inherently the lowest wear situation.

low amplitude fretting tests and examined the surfaces after the test. They found that the oxide films were not intact but consisted of l o o s e debris. In fact they identified periods of low friction (f = 0.05) which they attributed to rolling on oxide debris. It was thought that this condition might only apply to small amplitude slip however in 1958 Tamai (23) noted the same effect in reciprocating sliding. Furthermore Inabuchi (24,25) showed that a transition from severe wear to mild wear occurs if sufficient Fe,O, powder is added to the wear track. The powder forms a compacted agglomeration in the damage grooves and will persist for long periods of time. It is interesting to note that they found that 1 pm particles gave the best results since this is the usual thickness of wear debris. In their experiments Stott (26) in studying the friction time behavior of steels at high temperature found friction was controlled by sliding of compacted debris on compacted debris. Two

In this way a stable film

It would be expected that SO-SW would

In 1956 Halliday and Hirst (22) conducted

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forms of oxide production were described: oxide formed on the surface and immediately removed (FO-FW) or metal wear debris was first produced then oxidized. This latter suggestion was first proposed by Tomlinson ( 2 7 ) and has been thought to be important since iron has often been identified as a component in wear debris.

The soft lubricating film is really only a special case of the previous two. It arose from studies of the lubricating characteristics of oxides as lubricants ( 2 8 ) and studies of the sliding characteristics of metals at high temperatures ( 2 9 ) . Basically it was found that certain oxides as powders were effective lubricants and metals which formed these oxides gave the best sliding characteristics at high temperatures either self mated or in combination with ceramics ( 3 0 ) . This was not a particularly revolutionary idea since theories of boundary lubrication have been based on the formation of low shear strength films and materials in common usage (e.g., molybdenum tool steels) at high temperatures contained the ingredients to form soft oxides. Rabinowicz investigated this process in more detail ( 3 1 ) and Kruez ( 3 2 ) considered the formation of borate films in boundary lubrication. No general theory has been developed for the formation or removal of such films; however a deformation mechanism seems consistent with the fact that they are in the plastic state during the shear process.

In previously discussed papers several different approaches to oxidation kinetics and film removal mechanisms have been explored. most cases a linear or parabolic rate has been evaluated though many others have been explored. More recently dynamic kinetics have been explored where the rate can change with time or with operating conditions. However the major question is whether the oxidation process is the same under tribological conditions. Quinn finds that the diffusion coefficients are much faster but the whole question of reactions at high stresses and temperature remains unresolved.

Different wear and failure mechanisms have been identified. Wear of the film can occur by abrasion, fatigue, and deformation. The ultimate failure modes may be due to these wear processes or due to fatigue fracture ( 3 3 ) or blistering of the film. The loss of adhesion or blistering is due to differential stresses between the film and the substrate ( 7 ) .

should be given to certain aspects of the sliding process. For example, Moore ( 3 4 , 3 5 ) showed that when sliding steel against copper, that copper oxide was formed and then folded into the surface with subsequent sliding. Thus the surface is a composite of metal and oxide. Rigney ( 3 6 , 3 7 ) has shown that the wear layer is a mechanically mixed layer of metal and oxide. This effect can occur by either deformation or transfer. Kerridge ( 3 8 ) showed that transfer is the first step in the wear process. Hayler ( 3 9 ) has shown with several sliding systems that two conditions exist depending upon the temperature: oxide sliding against oxide, or metal against metal. condition is not due to the breakdown of the oxide, but rather to the transfer of metal on the surface of the oxide. Molgaard ( 4 0 ) discusses the role of oxide transfer as part of

In

This review has shown that more attention

The metal on metal

the wearing process. the general lack of information on film rheology or "what slides against what. " Under certain circumstances metal may also be sliding against oxide. Several investigators have shown that under such circumstances friction can be quite low ( 4 1 - 4 3 ) . For example, Cu on Fe,O, gives 0 . 2 5 and Nickel on A1,0,, 0 . 5 5 in vacuum,

Film rheology pertains to more than just defining the sliding interface. involves the flow mechanics of the film. Some work has begun in this area. Godet and coworkers ( 4 4 ) have addressed the problem with their so-called "third body approach. " Here particle mechanics are applied to film behavior. Mazuyer ( 4 5 ) has studied thin film rheology using bubble techniques, drawing dramatic conclusions at the affect of film thickness on deformation mechanisms. Other approaches such as continuum mechanics, micromechanics, molecular dynamics, or flow mechanics have been suggested and are under investigation in the Office of Naval Research Tribology Program. Such approaches attempt to predict friction and wear based on film behavior rather than determining experimentally how friction and wear affect the film.

These points emphasize

It also

2 . 2 Variables

One of the first investigations of surface films in the U.S. was conducted by Campbell ( 4 6 ) in 1 9 3 9 . He used preformed films and determined that friction dropped significantly when the film thickness reached 0 . 3 pm. Rabinowicz (31) found that a thickness of 0.01 pm was effective in reducing friction at high temperatures. others the maximum film thickness should be less than 3 pm to avoid failure. This suggests film values should be maintained at about 0.1 pm; however this value will depend on load and the hardness of the substrate material.

occurring oxide film on copper broke down between 10 g's and 100 g's load. He also found that rougher surfaces gave a lower value. It was suggested that to prevent breakdown the oxide and the substrate should be of equal hardness. Cocks ( 4 8 , 4 9 ) found, using electrical resistance measurements that the oxide did not breakdown from normal loading but that a lateral force was needed which was less than the sliding friction force. He also noted load transitions due to frictional heating. Hirst ( 5 0 , 5 1 ) measured the load capacity of oxide films and found that films preformed by heating at high temperatures gave little surface protection. The highest load capacity was with films formed slowly at low temperatures.

rate of 6 0 / 4 0 brass sliding against tool steel over a range of loads, velocities, and temperatures. He found that the film was protective at light loads and low speeds where sufficient time was available between passes to form the film and at high temperatures where the oxide formed a film faster than new surface was exposed. Higher temperatures were needed at higher loads and lower speed to give a protective film. This demonstrated the importance of frictional heating. Lancaster found that the mean surface temperature

From the work of Quinn and

Whitehead ( 4 7 ) found that the naturally

Lancaster ( 5 2 , 5 3 ) investigated the wear

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controlled film behavior not the flash temperature. The high temperature transition occurred when the mean temperature was 300C for brass on steel.

Earls and coworkers (54-57) studied the wear behavior of steel surfaces in great detail. They found that there are a range of loads and velocities where friction and wear are unstable. In this unstable region the oxide film is suddenly removed from the pin and partially from the track. Friction and thus temperature increase; new oxide is formed, and friction and wear return to their lower values. The oxide is removed at a critical thickness which decreases with increasing load. Film failure is by cracking or blistering when adhesion is low. The behavior of the films is controlled by the mean surface temperature which is related to W1I2V (W = load, V = velocity). Atmosphere particularly moisture plays a very strong role in the wear rates (53,58-62). Increasing the concentration of reactants increases the rate of wear and the transition values.

2.3 Properties

Early work (63) with oxide films suggested that oxide layers were easily broken on soft metals and that heating softens the underlying metal which aided in this disruption (64). Amorphous films were found to be more protective than crystalline films. Welch (65- 67) conducted extensive work on the role of frictional heating in the wear of steels. He concluded that the observed transitions were not only influenced by the formation and removal of oxide films but by structural changes and hardening caused by martenistic transformations, This hardening enhanced the protective properties of the oxide film.

of a large number of inorganic compounds at various loads between high pressure anvils. From these data friction coefficients can.be derived and plotted against the hardness of the oxide. These data are shown in Figure 2 for a normal stress of 100 kg/mm2 (142,000 psi). It can be seen that friction increases almost directly with hardness (Mohs) up to a value of 0.20 to 0.25 where it then remains relatively constant. This transition occurs at a Mohs hardness of 3.5 which is equivalent to about 150 kg/mm2. are lower and the transition moves to higher hardness. suggesting that shear only occurred at low hardness values and surface slip occurred with the harder oxides. It is not clear whether surface slip meant metal/oxide slip or oxide/oxide slip. Since in many cases the oxide was firmly attached to the anvils it might be assumed that oxide/oxide slip occurred; this however is not a certainty. He also noted substantial changes in composition and structure under pressure. Crystal structure changes occurred which often yielded free metal at the metal oxide interface.

Peterson (28) evaluated powdered oxides as lubricants at various temperatures. Certain oxides (CuO, MOO,, COO, Cu,O) and double oxides (molybdates and tungstates) were effective lubricants at 700C while other powdered oxides were not. Those that were effective had low melting points. This point is illustrated in

Bridgeman (68) measured the shear strength

At high pressures friction values

Bridgeman explained these results by

I . H , e ! I II.PM 2 . MOO

4 . CUlO 5 . ZnO 6. 0aO

8 . cdo

12. A 1 1 0 1 13. CaO

I S . C r I O s 16. Fez03 11. NIO I S . A s m s

3* n- 14. T I02

7 . nm,

I 9. smo, 19. 5r0 10. YO,

1 2 3 4 5 6 7 8 9 Hardness (nohi)

Figure 2. Effect of hardness on the friction coefficient of different oxide powders used as lubricants ( 8 8 ) .

Figure 3 where friction coefficient is plotted against melting point for a series of double oxides. It can be seen that as the melting point approaches the test temperature, friction decreases to a value of about 0.15 and there after remains constant. This is essentially the Bridgeman curve where melting point is a reasonable measure of the oxide hardness. Thus these oxides are seen to be effective lubricants when their melting temperature is within 200C of the test temperature.

Metal oxides were also evaluated as metal working lubricants at high temperatures (69). Low friction was found when the oxide was soft and ductile or when a friable layer developed. CuO and ZnO were found to be effective. The ZnO behaved as a loose powder rather than a sheared film.

Unfortunately the oxides which form in sliding situations are usually a very complex mixture. For example in Lancaster's work (53) the most probable film was 46 percent Fe,O,, 21.5 percent CuzO, 22.5 percent ZnO and 11 percent WO,. Other investigators using steel found various percentages of Fe,O, , Fe,O, and FeO depending upon the surface temperatures. Working with aluminum alloys of silicon and copper Razavizadeh (70,71) found that the film was composed of compacted debris whose thickness increased with load. Additions of silicon and copper to the alloy extended the mild wear region by improving the stability of

1 Copper 2 Rhenate 2 P o t d l l l m 2 Holybddte

4 Sodtun Tungitdte 5 Wlybdenm Oxide 6 Cobalt Rhenati 7 P O t a i i l U m Molybdatn 8 Nickel Molybdite

3 copper Rhemtc

9 copper TYngltate 10 Lead Molybdate I 1 Lead lungstate I2 CalcIm Halybdate 13 Calc lm Tunpitate I4 Nickel lungitate 15 Tungsten Oxide

Figure 3. Effect of melting temperatures on the coefficient

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