Wear, 49 (1978) 67 - 78 0 Elsevier Sequoia S.A., Lausanne - Printed in the Netherlands

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A METALLOGRAPHIC EXAMINATION OF ROLLERS SUBJECTED TO WEAR UNDER ROLLING-SLIDING CONDITIONS

D. M. FEGREDO* and C. PRITCHARD

Engineering Physics Section and Tribology Section, The Railway Technical Centre, London Road, Wilmorton, Derby, DE2 8UP (Gt. Britain)

(Received August 19, 1977)

Summary

Wear rate uersus maximum contact pressure p. curves were obtained at 3%,6% and 9% creep for rollers of a fracture tough rail steel (0.40 C-l.46 Mn) that were tested in a twin-disc wear and lubrication machine. Subsurface cracks were observed to form below and above the mild-severe wear transi- tion; these increased in depth and profusion with increasing creep and increasing po. Cracks close to the surface are almost parallel with it and even- tually break through to form flake-like particles of wear debris. Undeformed inclusions or inclusions with a low index of deformation are considered to be crack initiators, and AlsO particles were mainly associated with the cracks although more malleable MnS was also found. Tangential contact forces obtained from positive creep and an adequate coefficient of friction are necessary for crack formation. Microhardness tests disclosed that the wear rate appears to correlate with hardness at the extreme surface.

1. Introduction

Work has recently been published [l] on wear obtained under rolling- sliding conditions on a laboratory-scale machine in simulation of the contact conditions between the flange of a railway wheel and the side of a rail. The study concentrated on the transitions observed between mild and severe wear modes. Mild wear was characterized by a shiny worn surface with black oxidative wear debris, severe wear by a matt grey surface with flake metallic debris. The wear transitions were connected with the theoretical shakedown limits calculated by Johnson [ 21 and it was suggested that they relate to the shear strength of the steel. When the ratio T/iV of tangential and normal forces between the contacting surfaces is low (T/N < 0.32), as when the sur-

-

*Permanent address: Engineering Physics Section, Physical Metallurgy Research Laboratories, Department of Energy, Mines and Resources, Ottawa, Canada.

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faces are lubricated, the maximum shear stress occurs below the surface and it was suggested that failure arises from cracks initiating at this depth. Under unlubricated conditions T/N exceeds 0.32 and the maximum shear stress occurs at the surface. Cracks are then thought to propagate from the surface and to result in a flaking-type wear.

It was postulated that fatigue [ 1, 31 was important in severe wear by cyclically strain-hardening the material and that inclusions or surface cracks caused metallic wear flakes. However, the formation of cracks at inclusions was not demonstrated experimentally and neither was microhardness work reported. The work briefly described below was therefore done to investigate these aspects.

2. Experimental methods

The same laboratory Amsler lubrication and wear testing machine employed by Beagley [ 1] was used. The test specimens were two rollers of 35 - 40 mm diameter and normally 10 mm width which were rolled at about 200 rev min-l against one another under load and were constrained by gear- ing so that one rotated 1.104 times faster than the other. Some sliding thus occurred at the contact, and the ratio of sliding speed to rolling speed (called creep) was predetermined by carefully machining the specimens to the required diameters. Beagley did most of his experiments on the most commonly used British rail steel, to B.R. specification BS 11. The present experiments were done on specimens of a fracture tough rail steel of composition 0.40 C, 1.46 Mn, 0.24 Si, 0.027 S, 0.041 P and less than 0.01 Ni, Cr and MO. Similar experimental techniques to Beagley’s were adopted: (a) using a wire brush to remove debris from the lower roller to simulate side-cutting/flange wear (the wear rate is substantially reduced if no brush is used although the brush itself, acting on the roller, does not add to the wear rate); (b) employing a number of short runs of about 5 min duration to determine wear rates in the regime of severe wear; (c) reducing the width of the rollers below 10 mm to obtain a maximum contact pressure p. greater than 880 MN mW2. Any surface broadening caused by high contact stress was removed, when necessary, between runs by very careful machining. p. is taken as [l, 41 equal to

o 418 EW, +R2) ‘I2

* 1 -x5,- t

and creep is defined as ( Vf - V&/V, where W’ is the load per unit width of the two rollers in contact, RI, R2 are their respective radii, E is the Young’s modulus of elasticity, and Vf and V, are roller peripheral velocities. I’, pertains to the lower (faster) roller.

The program adopted was to use one pair of rollers per creep value to obtain p. uersus wear rate curves, by increasing p. for each of three differ- ent creep values: 3%, 6% and 9%. Having established the mild-severe wear

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transitions, four separate pairs of rollers were then tested at different pc values for each creep value, two pairs below and two above the mild-severe transition. Wear rates are reported as weight loss in grams of the rollers per metre rolled of the lower roller per 10 mm width of contact. In general, the lower roller wore faster than the upper. For this reason, it was thought that metallographic examination of the former would be more likely to disclose subsurface cracks and crack nucleators, and lower rollers were usually sectioned, nickel-plated and prepared by standard techniques for optical examination. The polished and etched sections looked at were perpendicular to the roller axis and in the centre of the width. Several upper rollers were also examined. Microhardness tests were mainly done on lower rollers at a 200 g load in a Leitz microhardness tester. The tip of the nearest indentation to the nickel-steel interface for the different specimens was never greater than 10 pm in distance from the interface.

3. Experimental results

Figure 1 illustrates p. uersus wear rate curves for three separate roller pairs for 3%, 6% and 9% creep. These rollers were 10 mm wide for 6% and 9% creep but were progressively narrowed for 3% creep in order to extend the contact pressures above 880 MN m-‘. Mild-severe wear transitions were obtained at approximately 870,735 and 620 MN me2 for 3%, 6% and 9% creep respectively, the values of p. at the transition decreasing with increasing creep. The separate points plotted in Fig. 1 represent the average wear rate (for a number of runs) of the further experiments, on distinct roller pairs, at selected p. values above and below the transitions. They are a good fit to the curves except at maximum contact pressures greater than 880 MN me2 for 6% and 9% creep, i.e. for roller widths less than 10 mm. (Points

500 700 900 1100

Maximum Contact Pretsur~ (p,,) MN/m’

Fig. I. Maximum contact pressure po vs. wear rate for - - -0, 3%, -.-A, 6% and -O, 9% creep,

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X and Y are referred to in Section 4.) Black oxidative and grey metallic flake debris were obtained below and above the wear transitions respectively. Fric- tion coefficients were fairly steady for any one creep value but did increase slightly with increasing creep. The ranges of friction values obtained were 0.52 - 0.56, 0.58 - 0.64 and 0.58 - 0.65 for 3%, 6% and 9% creep respectively.

Since wear rate obviously depends upon p. and the amount of creep, two experiments were done at a high p. of about 1660 MN mm 2 for (i) 0% creep (free rolling condition) and (ii) 9% creep with the rollers lubricated by oleic acid (coefficient of friction p = 0.04). The tangential force T generated was approximately 0 kgf in (i) and approximately 6 kgf (normal load 170 kgf) in (ii) and essentially no wear was obtained if wear is defined as “the progressive loss of substance from the surface of a body brought about by mechanical action” [ 51. However, surface mushrooming occurred and Fig. 2 shows this effect after the final two runs which totalled 6 min for the free rolling 0% creep test.

Fig. 2. Roller surface spreading effect in a free rolling (0% creep) test at pg = 1660 MN mP2.

Cracks were found below the extreme surface which increased in length and profusion with increasing p. and increasing creep. The depths at which visible cracks appeared increased and the surface topography grew rougher with increasing creep. The cracks were generally associated with approxima- tely spherical inclusions. Figures 3(a), 3(b) and 3(c) show cracks in rollers subjected to (a) 3% creep at p. = 435 MN rnM2 (mild wear), (b) 3% creep at p. = 870 MN mP2 (severe wear) and (c) 9% creep at p. = 886 MN m-’ (severe wear); Figure 3(d) depicts inclusions dotted along the crack trace of a potential wear flake. Subsurface cracks can thus form above and below the mild-severe wear transition and, in general, A1203 particles with and with- out attached MnS have been identified in the associated inclusions by electron microprobe analysis. Flattened MnS inclusions were also occasional- ly found in the crack path. Figure 3(e) shows no crack formation below the surface of the 9% creep specimen lubricated with oleic acid at p. x 1660 MN rnh2, although a number of potentially dangerous inclusions are visible.

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The etched microstructure of Fig. 4(a) illustrates the inter nse surf; ace and subsurface shear occurring in the specimen of Fig. 3(c). This type of ’ sur- face shear occurs to varying degrees in the different rollers deper rding up Ion test conditions. The extreme surface of the lower roller in the lubricat *ed

(a)

(b)

Fig. 3. (For caption seep. 73.)

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Cd)

Fig. 3. (For caption see opposite.)

13

(e)

Fig. 3. Arrows indicate typical subsurface cracks in rollers subjected to: (a) 3% creep at p. = 435 MN m-2 (320x);(b) 3%creepatpo = 1370MNmP2 (400x);(c)9%creepat po = 886 MN me2 (400x). (d) Inclusions dotted along the crack trace of a potential wear flake (3% creep at po = 1500 MN me2, 400x). (e) The untracked subsurface region of a specimen lubricated with oleic acid and subjected to 9% creep at po = 1660 MN mu2 (160x

test (Fig. 4(b)) shows evidence of shear but even this is absent in the free rolling experiment (0% creep at about 1660 MN mh2) although the layers beneath show compressed grains similar to those in Fig. 4(b).

Microhardness uersus distance from the nickel-steel interface is plotted in Fig. 5 for selected lower rollers, each of which had run several thousand revolutions at their particular contact pressure (after many thousands of running-in cycles at lower contact pressures). One upper roller (T4) is included as a comparison. The results show that after unlubricated rolling with positive creep, such that there is a substantial frictional force (in these experiments T/N = 0.52 - 0.65), the microhardness is greatest at the extreme surface. The slope of the curves decreases with increasing distance from the surface and progresses asymptotically to the as-received hardness value of HVzcc, = 230. The curves for rollers B8, B9 and B15 which sustained increas- ing normal pressure p. at the same 3% creep rate are of similar shape but are displaced vertically, as are the curves for Bll and B18 obtained at 9% creep. The two curves obtained when T/N was small were quite different. That obtained with roller B13, run at 9% creep but which was lubricated with oleic acid, coincides with that obtained with roller B14, run at 0% creep, and shows a maximum hardness in the interior.

The results imply some correlation between the extreme surface micro- hardness of the worn rollers and their wear rate. The roller B4 had a lower surface hardness than would be expected if this were the case, but this may

Fig. 4, Etched microstructures showing: (a) surface and subsurface shear deformation occurring in a roller subjected to 9% creep at pg = 886 MN rn”“% (160x); (b) general absence of shear deformation in a specimen ~~br~~ated with oleic acid and subjected to 9% creep at p0 = 1660 MN me2 ff60K).

be because of the nature of the extreme surface where there is the highest co~centr~t~o~ of partially cracked regions. T4, which roBed against BC, apparently developed a greater surface microhardness although its wear rate was about 47% of the total. Efforts were naturally made to select sound areas for the indentations but there is always the possibility of cracks exist- ing beneath an app~e~t~y good poXished surface.

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Micro

hardr

H”Z0

Distance of indentation tip horn Ni-steel junction pm

Ii141 0 166Ol 0

‘Oleic acid lubrication

Fig. 5. Microhardness (HVz& us. distance from the Ni-steel interface; B and T indicate lower and upper rollers respectively.

4. Discussion

The sharp rise in microhardness with decreasing depth (Fig. 5) corre- lates with the increasing shear deformation made visible in the etched micro- structures of the surface layers, e.g. see Fig. 4(a). Both indicate that the tangential force, which is responsible together with creep for causing the wear, has an influence on the crystal structure up to a depth of about 200 (urn, depending upon the test conditions. Subsurface cracks were observ- ed only in material which had experienced obvious shear and it appears therefore that in this type of wear the existence of cracks, severe metallic wear and increased surface hardness are all associated with considerable plastic deformation in shear of the subsurface material. The cracks propagate in the direction of metal flow and when they are close to the surface are almost parallel to it. In consequence, when a crack finally breaks through at the surface, the metallic debris is released in the form of a flake. Flaking wear was illustrated by Beagley [l] (his Fig. 11) and the overall picture is similar to that being developed in detail by Suh and his coworkers [6, 71 although they investigated sliding systems which do not involve the rolling contact studied in the present experiments.

The periodic form of the applied stress suggests that the cracks seen in Figs. 3(a), 3(b) and 3(c) are due to fatigue and the photographs indicate that they initiate at hard relatively undeformed inclusions in the highly sheared zone below the surface. Suh and his coworkers have discussed the role of inclusions in nucleating voids and give evidence that metals without inclu- sions exhibit lower wear rates [7]. In the present work the initiating inclu- sions appear to be oxides with and without sulphides; however, Marich and Curcio [ 81 concluded in a very recent report that silicates were the primary initiators of various defects in-their rails. A somewhat analogous situation exists in hot rolling during which compressive and shear stresses are applied;

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sharp cavities are produced around non-deformable inclusions or inclusions with a low index of deformation [9,10]. The intense flow of metal in one direction produces the same effect in the present experiments and the cavities act as further stress concentrators to generate fatigue cracks. There is thus the double effect of higher intrinsic stress concentration with increasing crack length and the increasing stress severity with decreasing depth that combines to produce severe wear.

Beagley suggested that when friction was high (T/N > 0.32) cracks would initiate at the surface. However, Suh et al. [6 J give theoretical argu- ments why cracks should not be nucleated at the surface because of the largely compressive stresses close under the contact and because they postulate the existence of a soft metal layer at the surface. In the present specimens nearly all the traces observed of cracks which intersected the sur- face were found to be along a direction which had faced away from the on- coming counterface. This is consistent with their having initially been subsur- face cracks, having propagated in the flow direction of the plastically deformed metal and eventually having broken through by fracture initiated at their tips as the extreme surface wore down to them.

When negligible tangential force was applied to rollers B13 and B14, maximum hardness was found about 250 ,um below the surface. If the posi- tion zI of maximum shear stress below the surface of two cylinders in c contact is taken as [ 4 3

then .a1 is about 200 pm for the 9% creep {lubricated) and 0% creep tests. This value compares reasonably well with the depth of maximum hardness at 150 pm observed for B13 and B14 although it must be remembered that truly elastic conditions did not apply, The contact pressure p. of about 1660 MN m-a was enough to mushroom the specimens plastically (Fig. 2). No subsurface cracks were observed in these specimens and wear was minimal. However, Beagley [ 11 observed that severe wear in similar condi- tions occurred after several hours of running and he postulated that this arose from cracks which nucleated at the depth at which the maximum shear stress occurred.

He also speculated why, when normal and tangential stresses were kept constant, wear increased more rapidly than linearly with increasing creep. Similar results can be abstracted from the wear curves in Fig. 1 which show, for example, that normal and tangential forces of 200 and 125 kgf respectively per unit width of contact (point X in Fig. 1, roller B18 in Fig. 5) generated far more wear at 9% creep than did the higher forces of 585 and 305 kgf respectively at 3% creep (point Y, roller B15). Beagley suggested that higher surface temperatures generated at the higher slip speeds {about 3.5 cm s-l at 9% creep) may be a factor, but the microhardness measure- ments in Fig. 5 show changes due to increasing creep which are clearly evident some distance below the surface. Comparing rollers run at similar p.

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values (e.g. B9 with B4 and with B18, and B8 with Bll) shows that running at a higher creep rate developed a greater hardness which was evident to a greater depth. Overall specimen temperatures generated in these tests were small; the rollers were never too hot to hold. In similar experiments to study the effect of humidity on friction [ll] the surface temperature in dry air increased only to about 35 “C in several minutes’ sliding at about 3% creep and a pressure p. of 430 MN m- ‘. The rapid increase in wear is thus associated simply with the greater strain hardening made evident by the microhardness measurements. This in turn implies a greater accumulated shear strain. As creep increases the slip speed increases, but perhaps more significantly each point on one surface interacts while in the contact area with a longer length of the counterface. (If the length b of the contact area is assumed to be constant at any given load, an asperity on the fast roller spends a time b/V, in the contact region. The counterface is sliding past at a speed V, - V,, so the asperity interacts with a length on the counterface equal to (Vf - V,) b/V,, i.e. a length which is proportional to the creep rate (Vf - Vs)/ V, .) It is evident that the increment of subsurface damage (initiation and propagation of cracks considered together) caused in one pass through the contact area owes more to the interactions between the surfaces as they slide this small length than to the overall forces experienced whilst traversing the contact length b.

5. Conclusions

Severe wear in rolling-sliding is associated with plastic deformation in shear in the subsurface. The wear rate appears to correlate with the micro- hardness generated at the extreme surface. Subsurface cracks, commonly initiated at hard undeformed inclusions, propagate almost parallel to the sur- face and eventually break through to form flake-like wear debris.

Acknowledgments

Thanks are due to Mr. N. Green for his assistance, to Mr. R. Hill for machining the specimens, to Dr. R. Packwood and Mrs. V. Moore for electron microprobe analysis and to the British Railways Board and the Dep~ment of Energy, Mines and Resources, Canada, for permission to publish this paper.

References

1 T. M. Beagley, Wear, 36 (1976) 317. 2 K. L. Johnson, Wear, 9 (1966) 4. 3 Y. P. Chin, T. E. Tallian and J. I. MeCool, Wear, 17 (1971) 433. 4 S. Timoshenko and J. N. Goodier, Theory of Elasticity, McGraw-Hill, New York, 1951,

p. 382.

5 J. Halling (ed.), Principles of Tribology, MacMillan, London, 1975, p. 94. 6 N. P. Suh, S. Jahanmir, J. Fleming, E. P. Abrahamson, N. Saka and J. P. Teixeira,

Second Progress Rep. on the Delamination Theory of Wear to Advanced Research Project Agency D.O.D., M.I.T., Cambridge, Mass., 197 5.

7 S. Jahanmir, N. P. Suh and E. P. Abrahamson, Wear, 28 (1974) 235. 8 S. Marich and P. Curcio, Development of high-strength alloyed rail steels suitable for

heavy duty applications, BHP Melbourne Research Laboratories Rep. MRL/083/76/ 015, presented at the ASTM Symp. on Rail Steels, Denver, Colo., U.S.A., November 1976, to be published.

9 R. Kiesling and H. Nordberg, Production and application of clean steels, IS1 Publ., 1972, p. 179.

10 R. Tricot, ISIPubl., 1972, p. 199. 11 T. M. Beagley, I.J. McEwen and C. Pritchard, Wear, 33 (1975) 141.