- Email: [email protected]
Aspects of fretting wear of sprayed cermet coatings
Wear, 106 (1985)
63
63 - 76
ASPECTS OF FRETTING
WEAR OF SPRAYED
CERMET COATINGS*
T. C. CHIVERS Centraf Electricity Generating GLl3 9PB (Gt. Brituin)
Board
Berkeley
Nuclear
Laboratories,
Berkeley,
Glos.
summary Two experimental fretting programmes which investigated aspects of fretting wear of sprayed cermet coatings are reviewed. These programmes were conducted in support of components used in the advanced gas-cooled reactor. It is speculated that the results from these programmes are compatible with a simple two-stage wear model. This model assumes that an initial wear process occurs which is dominated by an interlocking and removal of asperities. Such a phase wilI be dependent on the superficial contact areas and possibly the interfacial load, but the latter aspect is not considered. This initial wear is of very short duration and is followed by a mild, oxidative, wear mode. Coatings data are also compared with those for structural steels, In short-term low temperature tests it appears that structural steels have comparable performance with the cermet coatings but it is argued that this is an artefact of the wear process. However, at high temperatures (600 “C) wear of stainless steel could not be determined, the specimens showing a net weight gain. It is concluded that for in-reactor fretting applications cermet coatings will have advantages over structural steels at low temperatures. Even in high temperature regions some operation at low temperatures is experienced and consequently cermet coatings may be useful here also.
1. Introduction Within nuclear reactor systems bearing surfaces are required to accommodate, for example, differential thermal expansions or some mechanism function. In some plant areas conventional lubricants cannot be employed and many tribologically benevolent materials are excluded because of the requirements for circuit compatibility. There is also a conflict between satisfying the structural aspects of material selection and the bearing aspects. One way of compromising these requirements is to provide a surface *Paper presented at the Institution of Mechanical Engineers’ Fretting Wear Seminar, Nottingham, Gt. Britain, April 2 - 3, 1985. 0043-1648/85/$3.30
@ Elsevier Sequoia/Printed in The Netherlands
64
treatment or coating which gives a favourable performance under unlubricated conditions. One treatment adopted for use in the British advanced gas-cooled reactor (AGR) was the sprayed cermet coating. Experimental work had shown that in large amplitude sliding wear (several millimetre stroke) cermet coatings had improved wear, friction and static adhesion characteristics compared with stainless steels. Wear rates were observed to be about an order of magnitude lower on the cermet coatings than on the substrate materials. Mean frictional characteristics for the two material categories were not significantly different, but the statistical spread of the data was lower with the cermet coating. Whilst low carbon steels and austenitic steels exhibit static adhesion at temperatures greater than 300 “C! and 550 “C respectively, this phenomenon does not become apparent until temperatures exceed 650 “C for the cermets. For such reasons sprayed cermet coatings were adopted on some bearing surfaces in the AGRs where temperatures up to 700 “C are encountered. The use of high temperature and pressure CO2 as the heat transfer medium in the AGR gives a compact design with a high energy density, but with high noise and flow energy levels. As a result vibration of some components can occur with consequential fretting wear. In some instances the potential for fretting wear was not apparent until late in the AGR development programme. Under these conditions design freedoms were very limited and modifications to the bearing surfaces using cermet coatings had a number of attractions. Apart from the tribological benefits discussed above, there were production advantages. For example, little, if any, preand post-deposit machining is required, there is good dimensional control and there are no metallurgical changes in the substrate material. These advantages are offset by having only a thin coating (100 - 200 pm) available for wear, no metallurgical bonding, no accredited non-destructive examination system and, at the time of making decisions, an absence of fretting wear data. In this paper certain aspects of fretting wear of cermet coatings are discussed. Two programmes of work concerned with specific components will be considered in particular. In one programme very limited data were obtained, and analysis of the other is not yet complete. 2. Sliding wear processes Studies of sliding wear under unlubricated conditions have shown the specific wear rate to vary with a large number of variables. One such variable is the slip amplitude, and indeed this is the variable generally used to distinguish between fretting and sliding wear. Many workers have produced data showing wear rate to reduce as amplitude reduces. Whilst the underlying trend is well established it is not easy to distinguish between the roles of various parameters in modifying the wear processes. Transitions from initial severe wear modes to more benevolent mild wear regimes commonly occur, and the role of amplitude in modifying the transition is not determined.
65
Nevertheless both cermet coatings and structural steels demonstrate a reduction in wear rate as slip amplitude reduces, and both material types exhibit transitions in wear modes as contact conditions and operating parameters change. These transitions are important considerations in the discussions which follow. Throughout this paper reference will be made to the specific wear rate which is defined as
where WV is the wear volume lost, L is the interfacial load carried by the wearing interface and S is the distance slid. Its use follows from the adhesive wear laws formulated by Archard [l] based on various concepts, e.g. that the real area of contact is proportional to the load and the wear volume is proportional to the sliding distance. In the situations to be discussed adhesive wear may not necessarily be present and wear transitions will lead to different wear modes. Under these conditions eqn. (1) will be used on the basis of the data obtained at the end of the test. However, other definitions will also be employed on the basis of the assumption that for any particular wear regime the wear volume is linear with sliding distance, i.e.
It should be noted that under some conditions the wear volume loss is dominated by that which occurs in the initial “running-in” or severe wear mode. Subsequent volume loss in the post-transition or mild wear regime is very small and may be indistinguishable from the scatter in the data. In these circumstances WV is essentially independent of the sliding distance S, and with constant load the wear rate is proportional to the reciprocal of the sliding distance and a plot of log K uersus log S will have a slope of -1.
3. Case studies 3.1. Circulator components Large demands are made on AGR gas circulators. For each reactor some 3700 kg of CO2 coolant is pumped each second with a pressure rise of 3 bar. Centrifugal flow gas circulators are employed, and in most stations the circulators and their drive systems are contained within the reactor pressure circuit. The circulator can be isolated from the reactor circuit by closing an isolation dome and an inner sealing ring. During the extensive proving trials conducted on the circulators these dome and sealing ring components suffered fretting damage. Details of the mechanical, vibrational and life assessment aspects are described by Chivers et al. [2]. Here only the fretting palliative and component substantiation work will be described. Given the apparent improved wear performance of cermets it
66
was decided to coat the mild steel contacting surfaces with the Union Carbide Ltd. detonation-gun-deposited cermet LW5 (mixed tungsten and chrome carbides with a nickel binder). A programme of work was initiated to quantify the improvement in performance that might be obtained from the design modifications, and aspects of the experimental work will now be discussed. 3.1.1. Specific wear rate data Two laboratories were involved in acquiring data, i.e. the Wolfson Institute of Interfacial Technology, Nottingham, and the National Nuclear Corporation (NNC), Whetstone. Table 1 outlines the test programme and compares the test parameters with those for normal operation, whilst Table 2 summarizes the results obtained. Direct comparisons of the specific wear rates obtained in the two laboratories are not meaningful in that different assessment methods have been used. In the Wolfson tests weight loss measurements were made following ultrasonic cleaning of the specimens. The metal density was then used to calculate the volume loss and hence the wear rate. In the lower amplitude tests at NNC volume loss only from below the original surface was measured. A simple assessment would be that corrections to the data to move them to a common base would result in an increase in the specific wear rate determined from the Wolfson tests relative to those from NNC. However, this is an oversimplification. Limited profile traces of the worn surface areas in the Wolfson tests showed no evidence that the material was TABLE
1
In-reactor
conditions
Parameter
and LW5 test programme Reactor
Wolfson
\GR gas
Environment
280
(“C)
280
280
(bar)
40
4
Movement peak to peak (pm)
25 at normal operation
50
Frequency
20
100
Nominal area
contact
- 50
25 50
Large annulus on annulus
Annulus
> 0.065
160 mm*
160 mm* 4.6
mz
Nominal contact pressure (MN rne2)
1.1 - 2.3
10.3
Duration
30 years
150
Acceleration over reactor conditions
(Whetstone)
AGR gas
Gas pressure
(Hz)
NNC
AGR gas
Temperature
Geometry
test
=30
on flat
- 300 h
Flat on flat
800 h =4
test
67 TABLE 2 Results from the LW5 and mild steel test programmes Laboratory
Specimen
K (m3 N’
m-l)
Test duration
K VL a (m3 N-l m-l)
(h) Mild steel versus mild steel Wolfson Annulus Flat NNC (Whetstone) Top Bottom Top Bottom L W5 versus L W5 Wolfson
NNC (Whetstone)
Annulus Flat Annulus Flat Top Bottom Top Bottom
5.3 x lo-‘71 2.6 x 10-17j
150
800
2.1 3.2 0.6 0.6
x x x x
lo-‘7) lo-l7 \ lo-l7 lo-l7
2.6 1.1 1.6 7.1
x x x x
lo-l6 10-16 lo-l6 10-1s
8.7 3.9 1.7 1.9
x x x x
lo-l7 10-17 10-17 10-16
150 300 800
aKVL is based on the volume loss below the original surface datum and not the total volume loss. raised above the original surface. The tendency was to produce a smoother but lower surface. This contrasted with the lower amplitude NNC tests which showed pitting on both surfaces but with an elevated rim surrounding the pit. The net result was a growth in the volume of material. There are anomalies in the 150 and 300 h tests conducted on LW5 at. the Wolfson Institute in that the weight loss from the longer test was significantly lower than that in the other test. Whilst statistical variation could contribute to this there may well be other explanations as will be discussed subsequently. Bearing in mind these, and other observations made above, it would appear that the hard-coating specific wear rate is not significantly different from that for mild steel. 3.1.2. Wear morphology Wear rate is not the only criterion on which to judge relative performance. In many applications the way in which wear proceeds and its distribution are also of relevance. For the mild steel-mild steel pairing metallographic examination showed evidence of plastic deformation in the substrate and material transfer. This resulted in a layered and oxidized mound of transferred material on one face and a corresponding pit in the other. Characteristic height and depth dimensions for these features were typically 50 - 75 pm on the shorter duration tests and up to 200 pm in the longer tests.
68
The tests with LW5 against itself showed loss of peaks in the contact areas and resultant infilling of the valleys, the net result being a general smoothing of the surface. In the longer term, smaller amplitude, tests some pitting was observed but now in both surfaces and to a depth of some 50 pm. In these tests, however, the rims of the pits were raised above the level of the original surface, the built-up edge being a compact oxide. 3.2. Flow control The gas flow rate through any fuel channel of the AGR can be varied throughout the fuel life. Flow control systems are installed at the exit to the channel; here temperatures are in the region of 600 “C and mass flow rates are some 10 kg s-l. The mating surfaces are coated with the Union Carbide Ltd. detonation-gun-deposited cermet LClB (chromium carbides with an Ni-Cr binder) to provide protection in normal operation. The high velocity gas past the components can result in some vibrational excitation and consequential fretting wear. Studies have been undertaken to investigate the performance of this LClB coating in a fretting wear mode and under various test conditions. These experiments were conducted in the NNC Risley laboratories. Also examined was the behaviour of the substrate material (316 austenitic stainless steel) to determine the likely behaviour should the coating be worn through. 3.2.1. Specific wear rate data The parameters employed in the test programme are summarized in Table 3. The specific wear rates for the LClB are presented as a function of sliding distance in Fig. 1. These data have been obtained using surface profilometry techniques employing a ball-ended stylus. The use of this stylus offers some operational advantages, but for rough surfaces, as is the case with as-sprayed cermets, it leads to an overestimate of wear volumes and hence wear rates [3]. This aspect is currently being assessed with a view to establishing a “correction” factor. Clearly scatter exists in the data but it is not possible to discern effects attributable to either gas composition or pressure, or temperature for the COz-based environments, but it must be noted that only a very limited TABLE 3 Experimental parameters used in the LClB test programme Environments
Air at 20 “C and 600 “C CO* at 300 “C and 600 “C COz and CO at 600 “C t COz and CO at 600 “C
Interfacial load (N) Slip amplitude (pm) Frequency (Hz)
900 s 10 50 - 80
LClB versus LClB 3 16 stainless steel uersus 316 stainless steel
69 16
14
:
7
\
10‘
0’ .
E
7
z
“E E a
a
lo-
16 t
f
Y ”
iri u i!d u-l 10‘
lo-
18
L
10'
I
I
10' SLIDING
J
104 DISTANCE
(
105
m)
Fig. 1. Sliding fretting wear data for LClB US. LClB (load, 900 N; amplitude, 10 pm) for various environments and temperatures: 0, air at 20 “C; n, air at 600 “C; 0, CO2 at 300 “C; 0, CO2 at 600 “C; fl, CO2 plus CO at 600 “C.
number of tests were conducted at 300 “C. The line passing through K = 2 X 10-l* m3 N’ m-l at S = lo5 m has a slope of -1 which corresponds to zero wear. This line is not a calculated best fit to the data, but neither is it entirely unrepresentative. This suggests that following an initial high wear rate mode material removal rates drop to very low levels. Wear rate data obtained in air at room temperature separate from the CO2 results and would appear to be saturating at about 10-l’ m3 N-l m-l. Only one test was conducted at 600 “C in air; this datum may demonstrate an enhanced wear rate, but it could conceivably be within the general scatter of the data. Specific wear data for the stainless steel on stainless steel combination could not be determined as all specimens showed a volume gain rather than a loss. The mechanism appeared to be that of forming a transformed metal and/or oxide mound having an increased volume over that of the base metal. One can conclude therefore that any real metal loss is low.
70
4. Discussion 4.1. Hard coatings The information which is now becoming available on the fretting wear of sprayed cermet coatings shows wear rate reducing with reducing slip amplitude, as has been widely reported for steels. It appears, however, that under some conditions the reduction in wear rate with steels is much greater, such that the cermet coatings lose their clear advantage demonstrated at large slip amplitudes, but this may be an oversimplification. Skinner [4] studied the wear behaviour of a sprayed cermet coating in a crossed cylinder geometry where a pin was loaded against a continuously rotating shaft. He identified three sequential wear regimes for the coatings as follows: (i) a short-lived initial wear, probably resulting from the mechanical interlocking of asperities, (ii) an adhesive wear mode involving the homogenization of surface layers and (iii) mild wear in which oxidative wear provides a major contribution. Skinner also comments that for the pin, which is in continuous contact, the second (adhesive wear) regime was of very short duration. In other studies on the longer term fretting wear of low carbon steels it has been observed that mild wear can result, with any severe or adhesive wear mode making very little contribution to the generation of wear volume (see for example ref. 5). Given this information and that continuous contact exists it will be speculated that for the fretting wear of cermet coatings two wear regimes dominate. These regimes are (i) and (iii) above, i.e. that probably resulting from asperity interaction, followed by a mild wear process. A corollary of this model is that the first component will be influenced by the superficial area of contact, i.e. the area over which asperities can interact during relative movement. This will frequently be less than the potential contact area associated with the test geometry and is recognizable from the damage pattern on the surface (examples are shown in Fig. 2). The second component will, conventionally, be associated with the real areas of contact. For the fretting tests conducted here notionally flat-on-flat geometries have been employed. In reality alignment problems have resulted in very variable contact conditions and hence variable notional contact areas. Let us consider first the LClB data where each point shown represents a discrete experiment. The data at 1.1 X lo2 m sliding distance are apparently anomalous in that the volume removed is much greater than that lost at much longer sliding distances. An explanation may be scatter in the data but specimen examination suggests a greater than average superficial contact area. Indeed, with a prejudicial eye one can detect a correlation between higher than average wear rates and larger contact areas and vice versa. It will also be assumed that the first wear regime will not be environment depen-. dent, although it may be load dependent. This will not concern us here as the load remains constant. Also, because of the large number of data we do not need to consider contact areas in detail and hence we can consider the “mean” line through the results.
(b)
(a)
Fig. 2. Macroscopic views of LW5 specimens from Wolfson tests (the nominal contact conditions were the same for both tests). (a) Macroscopic view of the annulus after cleaning (the large nominal contact area in this test should be noted). (b) The same view as in (a) but this test had a very low nominal contact area.
For such a model our wear volume equation WV= K&A
becomes
+K,(S-SI)L
=KSL
(3)
where the subscript I refers to initial conditions and m refers to mild wear conditions. It is assumed that S, is very small and hence eqn. (3) reduces to
W,=K$&A =KSL
+K,SL (4)
Here K, is used as a wear volume per unit notional area occurring over the sliding distance Si. As stated earlier, any load dependence is not addressed here. As the curve through the COz data is so close to a slope of -1 (Fig. 1) it is not sensible to use those results to evaluate the parameters in eqn. (4). If, however, the notional curve through the air data is employed then K,S,A is calculated as 2 X lo-l3 m3 and K,ai,as 2 X 10-l’ m3 N-i m-l. If this value of K&A is considered to be independent of environment then K mco, is calculated as 2 X lo-l9 m3 N-l m-l which gives a good correlation within the data. It should be noted that as in Skinner’s work the wear in air is greater than that in CO* environments. Let us now consider the LW5 data, again assuming that the model discussed above is applicable. Here very little information is available and cross correlation between laboratories is complicated by changes in slip amplitude, interfacial load and the material loss assessments in the two test series. Hence only the Wolfson data are considered further. For these two
results it is necessary to make an assessment of the relevant contact areas. The shorter term experiment had a very much greater notional contact area than the longer running experiment and the area ratio has been independently assessed as 1:14 (see Fig. 2). Simultaneous equations can now be set up as % = and
K,SIA,+KmSL (5)
WV,, =KISIAb+K,SL where the subscripts a and b refer to the two different tests. For the purpose of this exercise it is sufficient to use relative contact areas, and if the experimental values for wear volume etc. are substituted then Km is evaluated as 5 X lo-‘* m3 N-’ m-i and K,SIA,as 1 X lo-l3 m3. The latter value is close to that quoted above, but there is no real relevance to this similarity. Had the lower area been used as the reference area then KISIAbwould have been an order of magnitude lower. The “predictions” of wear volume as a function of sliding distance on the basis of the proposed model for these data are shown in Fig. 3 together with the datum points. Also included is a spread based on the variation in the different wear volumes associated with the two specimens in a pairing. It is assumed that Sr is small and, on the scale shown, indistinguishable from zero. The NNG data have not been assessed in a similar manner because of the difference in volume measurement and the difficulty in estimating contact areas.
Sliding
Fig. 3. Speculated
fretting
wear models.
distance
5
73 TABLE 4 Values for mild wear rates for cermet
Environment
coatings
on the basis of the wear model
LW5
LClB
LClB
LClB
CO2 plus CO
Air
CO2
280
CO2 and CO, plus co 300 - 600
Wear mode
Fretting
Room temperature Fretting
K, (calculated) (x10” m3 N-l m-l)
0.5
2
Temperature
(“C)
presented
Fretting 0.02
30 - 100 Continuous slide (ref. 4) 30
Table 4 summarizes the wear rates derived. Care is required in interpreting this information, particularly as there are very little data for LW5. The mild wear rates evaluated for both cermet coatings in COz environments differ by an order of magnitude. This result contrasts with work at large slip amplitude which shows similar wear rates for the two materials. The difference identified here may be associated with amplitude effects, and this view may be supported by the NNC data which showed volume growth at a lower amplitude than that employed in the Wolfson tests. The area aspect also needs to be addressed when examining the LClB data. It has already been observed that a wide variation in superficial contact area arose as a consequence of different alignments. If a variation of an order of magnitude is assumed then this would result in a factor of 3 variation from the mean in the volume loss associated with the first wear stage. This deviation would be reflected in specific wear rate calculations which, in addition to other statistical changes, would result in a large amount of scatter in the data, but a scatter reducing with increasing sliding distance. The broad trend in data is in accord with such a view. 4.2. Structural steels Very few data on the structural steels were obtained in these test series. Only the mild steel tests yielded specific wear data and these were compatible with other fretting data on low carbon steels. The overall view formed is that at these low slip amplitudes strong sliding distance effects are not seen. Whilst such a view cannot be formulated from the results obtained here, it will be assumed that any initial wear period is of very limited extent such that linear extrapolation-interpolation of data is possible. The Wolfson mild steel data are included in Fig. 2 on this basis. For the high temperature stainless steel no wear data were obtained; in all cases volume gains were recorded. Whilst the profilometry technique employed has been argued to be in error the overall evidence is for volume growth. Metallographic sections show oxide and/or metallic mounds on the surface. Outside the notional contact areas only very thin oxides were seen.
Hence any real metal loss was low and although surface roughening occurred the fretting wear performance of the stainless steel under the conditions tested was very good. 4.3. Cermet coa tin@ or base steels? A superficial comparison of the mild steels and LW5 data shows no difference in performance. However, the largely u~ubs~ntia~d arguments advanced suggest that this is an oversimplification associated with the very early stages of wear. The models advanced are presented in Fig. 3 and argue for lower wear rates at long sliding distances for cermet coatings. There is operating experience which broadly supports the view expressed here. Low [5] has examined LW5 and low carbon steel surfaces which have operated under similar contact conditions. His interpretation of the results is that the wear rate of LW5 is an order of m~itude better than that of the low carbon steel. However, at the high temperature the measured wear rate performance of LClB was inferior to that of the stainless steel. This may be partly attributable to the fact that the longest stainless steel test ran for a sliding distance an order of magnitude lower than that for the longest LClB test. To polarize the situation one could advocate the use of coatings for improved fretting wear resistance at the lower temperatures, and this will now be considered further. Temperature has been demonstrated to play a significant role in wear transitions in the fretting wear of structural steels (see for example refs. 6 and 7). At the lower temperatures wear rate can increase by orders of magnitude over that at high temperatures, but it is strongly influenced by parameters such as slip amplitude. As the chromium content (this is not necessarily the only significant compositional parameter) increases so does the transition ~mperature. Hence if a structural steel is to be subjected to fretting wear at temperatures below the transition temperature appropriate to its contact and excitational parameters high wear rates may be encountered. This is very important for many in-reactor components as they may “see” proving tests conducted at low tempe~tures, resulting in excessive wear. Perhaps of more importance is the fact that reactor commissioning takes place over a period of several months at gradually increasing temperatures. If 0.2 years is spent below the transition temperature then 40% of a 30 year wear life can be consumed in commissioning if there is a change of two orders of magnitude in wear rate with temperature. If one now examines the LClB data and interprets them in the light of the model advanced then one observes that following the initial wear the low temperature (in air) wear rate is two orders of magnitude greater than that at high temperature. However, this low temperature wear rate is comparable with the high temperature structural steel data. Hence, any operational problems arising from low temperature operation should benefit from the use of sprayed cermet coatings. Another aspect of relevance in comparing performance is surface morphology, Whilst pitting was seen in some of the LW5 tests this was of
a different nature and scale to that observed on the mild steel specimens. The pit-pit configuration compared with the pit-mound geometry seen with the low carbon steel could be of benefit if large operational movements are not to be inhibited. One must, however, be careful in making this last observation as pit-pit damage has been seen also in the fretting of steels. 4.4. Future work In this paper data from two experimental programmes have been presented and interpreted in a very speculative manner. The analysis of the LClB is incomplete and future work can investigate such aspects as superficial contact area. However, the LW5 work, which was completed some years ago, demonstrates the futility of such restricted experimental programmes. Whilst there is a need to conduct short-term experiments on cermet coatings to refine or reject the proposed model, it is also apparent that any test programme needs to include time as a primary variable. The role of interfacial load and slip amplitudes in influencing wear of these coatings also needs to be investigated. It is also important to establish a statistical assessment of the spread in the data.
5. Closure In view of the very speculative nature of the model proposed and the inferences which ensue it is considered inappropriate to draw conclusions, hence this closure. Fretting wear data for sprayed cermet coatings obtained in two experimental programmes have been reviewed and compared with a twostage wear model. It is surmised that an initial wear process results essentially from an interlocking and removal of asperities, and this is followed by a mild, low, wear rate. The initial wear mode results in an apparently high wear rate which is dependent on superficial contact area and this dominates the volume removed for considerable sliding distances. At low temperatures such behaviour provides an explanation for apparently similar behaviour between structural steels and cermet coatings at short sliding distances. At high temperatures (600 “C) the fretting wear performance of structural steels appears to be as good as or better than that of the cermet coatings. However, if components are also required to operate for a period of time below their wear transition temperature then cermets should demonstrate an improved performance overall.
Acknowledgments The author gratefully acknowledges the efforts of his colleagues in the National Nuclear Corporation at Whetstone and Risley and at the Wolfson
76
Institute of Interfacial Technology, Nottingham University, the data presented here. This paper is published by permission Electricity Generating Board.
in producing of the Central
References 1 J. F. Archard, J. Appl. Phys., 24 (1953) 981 - 988. 2 T. C. Chivers, S. C. Gordelier, J. Roy and M. Wharton, Proc. Conf. on Vibration Nuclear Plant, Keswick, 1978, British Nuclear Energy Society, London, 1978. 3 T. C. Chivers and S. J. Radcliffe, Wear, 57 (1979) 313 - 321. 4 J. Skinner, J. Lubr. Technol., 103 (2) (1981) 228 - 235. 5 M. B. J. Low, Wear, 106 (1985) 315 - 335. 6 R. S. Millman, A. Sleightholme and A. A. Parry, Wear, 106 (1985) 77 - 95. 7 D. H. Jones, A. Y. Nehru and J. Skinner, Wear, 106 (1985) 139 - 162.
in
63
63 - 76
ASPECTS OF FRETTING
WEAR OF SPRAYED
CERMET COATINGS*
T. C. CHIVERS Centraf Electricity Generating GLl3 9PB (Gt. Brituin)
Board
Berkeley
Nuclear
Laboratories,
Berkeley,
Glos.
summary Two experimental fretting programmes which investigated aspects of fretting wear of sprayed cermet coatings are reviewed. These programmes were conducted in support of components used in the advanced gas-cooled reactor. It is speculated that the results from these programmes are compatible with a simple two-stage wear model. This model assumes that an initial wear process occurs which is dominated by an interlocking and removal of asperities. Such a phase wilI be dependent on the superficial contact areas and possibly the interfacial load, but the latter aspect is not considered. This initial wear is of very short duration and is followed by a mild, oxidative, wear mode. Coatings data are also compared with those for structural steels, In short-term low temperature tests it appears that structural steels have comparable performance with the cermet coatings but it is argued that this is an artefact of the wear process. However, at high temperatures (600 “C) wear of stainless steel could not be determined, the specimens showing a net weight gain. It is concluded that for in-reactor fretting applications cermet coatings will have advantages over structural steels at low temperatures. Even in high temperature regions some operation at low temperatures is experienced and consequently cermet coatings may be useful here also.
1. Introduction Within nuclear reactor systems bearing surfaces are required to accommodate, for example, differential thermal expansions or some mechanism function. In some plant areas conventional lubricants cannot be employed and many tribologically benevolent materials are excluded because of the requirements for circuit compatibility. There is also a conflict between satisfying the structural aspects of material selection and the bearing aspects. One way of compromising these requirements is to provide a surface *Paper presented at the Institution of Mechanical Engineers’ Fretting Wear Seminar, Nottingham, Gt. Britain, April 2 - 3, 1985. 0043-1648/85/$3.30
@ Elsevier Sequoia/Printed in The Netherlands
64
treatment or coating which gives a favourable performance under unlubricated conditions. One treatment adopted for use in the British advanced gas-cooled reactor (AGR) was the sprayed cermet coating. Experimental work had shown that in large amplitude sliding wear (several millimetre stroke) cermet coatings had improved wear, friction and static adhesion characteristics compared with stainless steels. Wear rates were observed to be about an order of magnitude lower on the cermet coatings than on the substrate materials. Mean frictional characteristics for the two material categories were not significantly different, but the statistical spread of the data was lower with the cermet coating. Whilst low carbon steels and austenitic steels exhibit static adhesion at temperatures greater than 300 “C! and 550 “C respectively, this phenomenon does not become apparent until temperatures exceed 650 “C for the cermets. For such reasons sprayed cermet coatings were adopted on some bearing surfaces in the AGRs where temperatures up to 700 “C are encountered. The use of high temperature and pressure CO2 as the heat transfer medium in the AGR gives a compact design with a high energy density, but with high noise and flow energy levels. As a result vibration of some components can occur with consequential fretting wear. In some instances the potential for fretting wear was not apparent until late in the AGR development programme. Under these conditions design freedoms were very limited and modifications to the bearing surfaces using cermet coatings had a number of attractions. Apart from the tribological benefits discussed above, there were production advantages. For example, little, if any, preand post-deposit machining is required, there is good dimensional control and there are no metallurgical changes in the substrate material. These advantages are offset by having only a thin coating (100 - 200 pm) available for wear, no metallurgical bonding, no accredited non-destructive examination system and, at the time of making decisions, an absence of fretting wear data. In this paper certain aspects of fretting wear of cermet coatings are discussed. Two programmes of work concerned with specific components will be considered in particular. In one programme very limited data were obtained, and analysis of the other is not yet complete. 2. Sliding wear processes Studies of sliding wear under unlubricated conditions have shown the specific wear rate to vary with a large number of variables. One such variable is the slip amplitude, and indeed this is the variable generally used to distinguish between fretting and sliding wear. Many workers have produced data showing wear rate to reduce as amplitude reduces. Whilst the underlying trend is well established it is not easy to distinguish between the roles of various parameters in modifying the wear processes. Transitions from initial severe wear modes to more benevolent mild wear regimes commonly occur, and the role of amplitude in modifying the transition is not determined.
65
Nevertheless both cermet coatings and structural steels demonstrate a reduction in wear rate as slip amplitude reduces, and both material types exhibit transitions in wear modes as contact conditions and operating parameters change. These transitions are important considerations in the discussions which follow. Throughout this paper reference will be made to the specific wear rate which is defined as
where WV is the wear volume lost, L is the interfacial load carried by the wearing interface and S is the distance slid. Its use follows from the adhesive wear laws formulated by Archard [l] based on various concepts, e.g. that the real area of contact is proportional to the load and the wear volume is proportional to the sliding distance. In the situations to be discussed adhesive wear may not necessarily be present and wear transitions will lead to different wear modes. Under these conditions eqn. (1) will be used on the basis of the data obtained at the end of the test. However, other definitions will also be employed on the basis of the assumption that for any particular wear regime the wear volume is linear with sliding distance, i.e.
It should be noted that under some conditions the wear volume loss is dominated by that which occurs in the initial “running-in” or severe wear mode. Subsequent volume loss in the post-transition or mild wear regime is very small and may be indistinguishable from the scatter in the data. In these circumstances WV is essentially independent of the sliding distance S, and with constant load the wear rate is proportional to the reciprocal of the sliding distance and a plot of log K uersus log S will have a slope of -1.
3. Case studies 3.1. Circulator components Large demands are made on AGR gas circulators. For each reactor some 3700 kg of CO2 coolant is pumped each second with a pressure rise of 3 bar. Centrifugal flow gas circulators are employed, and in most stations the circulators and their drive systems are contained within the reactor pressure circuit. The circulator can be isolated from the reactor circuit by closing an isolation dome and an inner sealing ring. During the extensive proving trials conducted on the circulators these dome and sealing ring components suffered fretting damage. Details of the mechanical, vibrational and life assessment aspects are described by Chivers et al. [2]. Here only the fretting palliative and component substantiation work will be described. Given the apparent improved wear performance of cermets it
66
was decided to coat the mild steel contacting surfaces with the Union Carbide Ltd. detonation-gun-deposited cermet LW5 (mixed tungsten and chrome carbides with a nickel binder). A programme of work was initiated to quantify the improvement in performance that might be obtained from the design modifications, and aspects of the experimental work will now be discussed. 3.1.1. Specific wear rate data Two laboratories were involved in acquiring data, i.e. the Wolfson Institute of Interfacial Technology, Nottingham, and the National Nuclear Corporation (NNC), Whetstone. Table 1 outlines the test programme and compares the test parameters with those for normal operation, whilst Table 2 summarizes the results obtained. Direct comparisons of the specific wear rates obtained in the two laboratories are not meaningful in that different assessment methods have been used. In the Wolfson tests weight loss measurements were made following ultrasonic cleaning of the specimens. The metal density was then used to calculate the volume loss and hence the wear rate. In the lower amplitude tests at NNC volume loss only from below the original surface was measured. A simple assessment would be that corrections to the data to move them to a common base would result in an increase in the specific wear rate determined from the Wolfson tests relative to those from NNC. However, this is an oversimplification. Limited profile traces of the worn surface areas in the Wolfson tests showed no evidence that the material was TABLE
1
In-reactor
conditions
Parameter
and LW5 test programme Reactor
Wolfson
\GR gas
Environment
280
(“C)
280
280
(bar)
40
4
Movement peak to peak (pm)
25 at normal operation
50
Frequency
20
100
Nominal area
contact
- 50
25 50
Large annulus on annulus
Annulus
> 0.065
160 mm*
160 mm* 4.6
mz
Nominal contact pressure (MN rne2)
1.1 - 2.3
10.3
Duration
30 years
150
Acceleration over reactor conditions
(Whetstone)
AGR gas
Gas pressure
(Hz)
NNC
AGR gas
Temperature
Geometry
test
=30
on flat
- 300 h
Flat on flat
800 h =4
test
67 TABLE 2 Results from the LW5 and mild steel test programmes Laboratory
Specimen
K (m3 N’
m-l)
Test duration
K VL a (m3 N-l m-l)
(h) Mild steel versus mild steel Wolfson Annulus Flat NNC (Whetstone) Top Bottom Top Bottom L W5 versus L W5 Wolfson
NNC (Whetstone)
Annulus Flat Annulus Flat Top Bottom Top Bottom
5.3 x lo-‘71 2.6 x 10-17j
150
800
2.1 3.2 0.6 0.6
x x x x
lo-‘7) lo-l7 \ lo-l7 lo-l7
2.6 1.1 1.6 7.1
x x x x
lo-l6 10-16 lo-l6 10-1s
8.7 3.9 1.7 1.9
x x x x
lo-l7 10-17 10-17 10-16
150 300 800
aKVL is based on the volume loss below the original surface datum and not the total volume loss. raised above the original surface. The tendency was to produce a smoother but lower surface. This contrasted with the lower amplitude NNC tests which showed pitting on both surfaces but with an elevated rim surrounding the pit. The net result was a growth in the volume of material. There are anomalies in the 150 and 300 h tests conducted on LW5 at. the Wolfson Institute in that the weight loss from the longer test was significantly lower than that in the other test. Whilst statistical variation could contribute to this there may well be other explanations as will be discussed subsequently. Bearing in mind these, and other observations made above, it would appear that the hard-coating specific wear rate is not significantly different from that for mild steel. 3.1.2. Wear morphology Wear rate is not the only criterion on which to judge relative performance. In many applications the way in which wear proceeds and its distribution are also of relevance. For the mild steel-mild steel pairing metallographic examination showed evidence of plastic deformation in the substrate and material transfer. This resulted in a layered and oxidized mound of transferred material on one face and a corresponding pit in the other. Characteristic height and depth dimensions for these features were typically 50 - 75 pm on the shorter duration tests and up to 200 pm in the longer tests.
68
The tests with LW5 against itself showed loss of peaks in the contact areas and resultant infilling of the valleys, the net result being a general smoothing of the surface. In the longer term, smaller amplitude, tests some pitting was observed but now in both surfaces and to a depth of some 50 pm. In these tests, however, the rims of the pits were raised above the level of the original surface, the built-up edge being a compact oxide. 3.2. Flow control The gas flow rate through any fuel channel of the AGR can be varied throughout the fuel life. Flow control systems are installed at the exit to the channel; here temperatures are in the region of 600 “C and mass flow rates are some 10 kg s-l. The mating surfaces are coated with the Union Carbide Ltd. detonation-gun-deposited cermet LClB (chromium carbides with an Ni-Cr binder) to provide protection in normal operation. The high velocity gas past the components can result in some vibrational excitation and consequential fretting wear. Studies have been undertaken to investigate the performance of this LClB coating in a fretting wear mode and under various test conditions. These experiments were conducted in the NNC Risley laboratories. Also examined was the behaviour of the substrate material (316 austenitic stainless steel) to determine the likely behaviour should the coating be worn through. 3.2.1. Specific wear rate data The parameters employed in the test programme are summarized in Table 3. The specific wear rates for the LClB are presented as a function of sliding distance in Fig. 1. These data have been obtained using surface profilometry techniques employing a ball-ended stylus. The use of this stylus offers some operational advantages, but for rough surfaces, as is the case with as-sprayed cermets, it leads to an overestimate of wear volumes and hence wear rates [3]. This aspect is currently being assessed with a view to establishing a “correction” factor. Clearly scatter exists in the data but it is not possible to discern effects attributable to either gas composition or pressure, or temperature for the COz-based environments, but it must be noted that only a very limited TABLE 3 Experimental parameters used in the LClB test programme Environments
Air at 20 “C and 600 “C CO* at 300 “C and 600 “C COz and CO at 600 “C t COz and CO at 600 “C
Interfacial load (N) Slip amplitude (pm) Frequency (Hz)
900 s 10 50 - 80
LClB versus LClB 3 16 stainless steel uersus 316 stainless steel
69 16
14
:
7
\
10‘
0’ .
E
7
z
“E E a
a
lo-
16 t
f
Y ”
iri u i!d u-l 10‘
lo-
18
L
10'
I
I
10' SLIDING
J
104 DISTANCE
(
105
m)
Fig. 1. Sliding fretting wear data for LClB US. LClB (load, 900 N; amplitude, 10 pm) for various environments and temperatures: 0, air at 20 “C; n, air at 600 “C; 0, CO2 at 300 “C; 0, CO2 at 600 “C; fl, CO2 plus CO at 600 “C.
number of tests were conducted at 300 “C. The line passing through K = 2 X 10-l* m3 N’ m-l at S = lo5 m has a slope of -1 which corresponds to zero wear. This line is not a calculated best fit to the data, but neither is it entirely unrepresentative. This suggests that following an initial high wear rate mode material removal rates drop to very low levels. Wear rate data obtained in air at room temperature separate from the CO2 results and would appear to be saturating at about 10-l’ m3 N-l m-l. Only one test was conducted at 600 “C in air; this datum may demonstrate an enhanced wear rate, but it could conceivably be within the general scatter of the data. Specific wear data for the stainless steel on stainless steel combination could not be determined as all specimens showed a volume gain rather than a loss. The mechanism appeared to be that of forming a transformed metal and/or oxide mound having an increased volume over that of the base metal. One can conclude therefore that any real metal loss is low.
70
4. Discussion 4.1. Hard coatings The information which is now becoming available on the fretting wear of sprayed cermet coatings shows wear rate reducing with reducing slip amplitude, as has been widely reported for steels. It appears, however, that under some conditions the reduction in wear rate with steels is much greater, such that the cermet coatings lose their clear advantage demonstrated at large slip amplitudes, but this may be an oversimplification. Skinner [4] studied the wear behaviour of a sprayed cermet coating in a crossed cylinder geometry where a pin was loaded against a continuously rotating shaft. He identified three sequential wear regimes for the coatings as follows: (i) a short-lived initial wear, probably resulting from the mechanical interlocking of asperities, (ii) an adhesive wear mode involving the homogenization of surface layers and (iii) mild wear in which oxidative wear provides a major contribution. Skinner also comments that for the pin, which is in continuous contact, the second (adhesive wear) regime was of very short duration. In other studies on the longer term fretting wear of low carbon steels it has been observed that mild wear can result, with any severe or adhesive wear mode making very little contribution to the generation of wear volume (see for example ref. 5). Given this information and that continuous contact exists it will be speculated that for the fretting wear of cermet coatings two wear regimes dominate. These regimes are (i) and (iii) above, i.e. that probably resulting from asperity interaction, followed by a mild wear process. A corollary of this model is that the first component will be influenced by the superficial area of contact, i.e. the area over which asperities can interact during relative movement. This will frequently be less than the potential contact area associated with the test geometry and is recognizable from the damage pattern on the surface (examples are shown in Fig. 2). The second component will, conventionally, be associated with the real areas of contact. For the fretting tests conducted here notionally flat-on-flat geometries have been employed. In reality alignment problems have resulted in very variable contact conditions and hence variable notional contact areas. Let us consider first the LClB data where each point shown represents a discrete experiment. The data at 1.1 X lo2 m sliding distance are apparently anomalous in that the volume removed is much greater than that lost at much longer sliding distances. An explanation may be scatter in the data but specimen examination suggests a greater than average superficial contact area. Indeed, with a prejudicial eye one can detect a correlation between higher than average wear rates and larger contact areas and vice versa. It will also be assumed that the first wear regime will not be environment depen-. dent, although it may be load dependent. This will not concern us here as the load remains constant. Also, because of the large number of data we do not need to consider contact areas in detail and hence we can consider the “mean” line through the results.
(b)
(a)
Fig. 2. Macroscopic views of LW5 specimens from Wolfson tests (the nominal contact conditions were the same for both tests). (a) Macroscopic view of the annulus after cleaning (the large nominal contact area in this test should be noted). (b) The same view as in (a) but this test had a very low nominal contact area.
For such a model our wear volume equation WV= K&A
becomes
+K,(S-SI)L
=KSL
(3)
where the subscript I refers to initial conditions and m refers to mild wear conditions. It is assumed that S, is very small and hence eqn. (3) reduces to
W,=K$&A =KSL
+K,SL (4)
Here K, is used as a wear volume per unit notional area occurring over the sliding distance Si. As stated earlier, any load dependence is not addressed here. As the curve through the COz data is so close to a slope of -1 (Fig. 1) it is not sensible to use those results to evaluate the parameters in eqn. (4). If, however, the notional curve through the air data is employed then K,S,A is calculated as 2 X lo-l3 m3 and K,ai,as 2 X 10-l’ m3 N-i m-l. If this value of K&A is considered to be independent of environment then K mco, is calculated as 2 X lo-l9 m3 N-l m-l which gives a good correlation within the data. It should be noted that as in Skinner’s work the wear in air is greater than that in CO* environments. Let us now consider the LW5 data, again assuming that the model discussed above is applicable. Here very little information is available and cross correlation between laboratories is complicated by changes in slip amplitude, interfacial load and the material loss assessments in the two test series. Hence only the Wolfson data are considered further. For these two
results it is necessary to make an assessment of the relevant contact areas. The shorter term experiment had a very much greater notional contact area than the longer running experiment and the area ratio has been independently assessed as 1:14 (see Fig. 2). Simultaneous equations can now be set up as % = and
K,SIA,+KmSL (5)
WV,, =KISIAb+K,SL where the subscripts a and b refer to the two different tests. For the purpose of this exercise it is sufficient to use relative contact areas, and if the experimental values for wear volume etc. are substituted then Km is evaluated as 5 X lo-‘* m3 N-’ m-i and K,SIA,as 1 X lo-l3 m3. The latter value is close to that quoted above, but there is no real relevance to this similarity. Had the lower area been used as the reference area then KISIAbwould have been an order of magnitude lower. The “predictions” of wear volume as a function of sliding distance on the basis of the proposed model for these data are shown in Fig. 3 together with the datum points. Also included is a spread based on the variation in the different wear volumes associated with the two specimens in a pairing. It is assumed that Sr is small and, on the scale shown, indistinguishable from zero. The NNG data have not been assessed in a similar manner because of the difference in volume measurement and the difficulty in estimating contact areas.
Sliding
Fig. 3. Speculated
fretting
wear models.
distance
5
73 TABLE 4 Values for mild wear rates for cermet
Environment
coatings
on the basis of the wear model
LW5
LClB
LClB
LClB
CO2 plus CO
Air
CO2
280
CO2 and CO, plus co 300 - 600
Wear mode
Fretting
Room temperature Fretting
K, (calculated) (x10” m3 N-l m-l)
0.5
2
Temperature
(“C)
presented
Fretting 0.02
30 - 100 Continuous slide (ref. 4) 30
Table 4 summarizes the wear rates derived. Care is required in interpreting this information, particularly as there are very little data for LW5. The mild wear rates evaluated for both cermet coatings in COz environments differ by an order of magnitude. This result contrasts with work at large slip amplitude which shows similar wear rates for the two materials. The difference identified here may be associated with amplitude effects, and this view may be supported by the NNC data which showed volume growth at a lower amplitude than that employed in the Wolfson tests. The area aspect also needs to be addressed when examining the LClB data. It has already been observed that a wide variation in superficial contact area arose as a consequence of different alignments. If a variation of an order of magnitude is assumed then this would result in a factor of 3 variation from the mean in the volume loss associated with the first wear stage. This deviation would be reflected in specific wear rate calculations which, in addition to other statistical changes, would result in a large amount of scatter in the data, but a scatter reducing with increasing sliding distance. The broad trend in data is in accord with such a view. 4.2. Structural steels Very few data on the structural steels were obtained in these test series. Only the mild steel tests yielded specific wear data and these were compatible with other fretting data on low carbon steels. The overall view formed is that at these low slip amplitudes strong sliding distance effects are not seen. Whilst such a view cannot be formulated from the results obtained here, it will be assumed that any initial wear period is of very limited extent such that linear extrapolation-interpolation of data is possible. The Wolfson mild steel data are included in Fig. 2 on this basis. For the high temperature stainless steel no wear data were obtained; in all cases volume gains were recorded. Whilst the profilometry technique employed has been argued to be in error the overall evidence is for volume growth. Metallographic sections show oxide and/or metallic mounds on the surface. Outside the notional contact areas only very thin oxides were seen.
Hence any real metal loss was low and although surface roughening occurred the fretting wear performance of the stainless steel under the conditions tested was very good. 4.3. Cermet coa tin@ or base steels? A superficial comparison of the mild steels and LW5 data shows no difference in performance. However, the largely u~ubs~ntia~d arguments advanced suggest that this is an oversimplification associated with the very early stages of wear. The models advanced are presented in Fig. 3 and argue for lower wear rates at long sliding distances for cermet coatings. There is operating experience which broadly supports the view expressed here. Low [5] has examined LW5 and low carbon steel surfaces which have operated under similar contact conditions. His interpretation of the results is that the wear rate of LW5 is an order of m~itude better than that of the low carbon steel. However, at the high temperature the measured wear rate performance of LClB was inferior to that of the stainless steel. This may be partly attributable to the fact that the longest stainless steel test ran for a sliding distance an order of magnitude lower than that for the longest LClB test. To polarize the situation one could advocate the use of coatings for improved fretting wear resistance at the lower temperatures, and this will now be considered further. Temperature has been demonstrated to play a significant role in wear transitions in the fretting wear of structural steels (see for example refs. 6 and 7). At the lower temperatures wear rate can increase by orders of magnitude over that at high temperatures, but it is strongly influenced by parameters such as slip amplitude. As the chromium content (this is not necessarily the only significant compositional parameter) increases so does the transition ~mperature. Hence if a structural steel is to be subjected to fretting wear at temperatures below the transition temperature appropriate to its contact and excitational parameters high wear rates may be encountered. This is very important for many in-reactor components as they may “see” proving tests conducted at low tempe~tures, resulting in excessive wear. Perhaps of more importance is the fact that reactor commissioning takes place over a period of several months at gradually increasing temperatures. If 0.2 years is spent below the transition temperature then 40% of a 30 year wear life can be consumed in commissioning if there is a change of two orders of magnitude in wear rate with temperature. If one now examines the LClB data and interprets them in the light of the model advanced then one observes that following the initial wear the low temperature (in air) wear rate is two orders of magnitude greater than that at high temperature. However, this low temperature wear rate is comparable with the high temperature structural steel data. Hence, any operational problems arising from low temperature operation should benefit from the use of sprayed cermet coatings. Another aspect of relevance in comparing performance is surface morphology, Whilst pitting was seen in some of the LW5 tests this was of
a different nature and scale to that observed on the mild steel specimens. The pit-pit configuration compared with the pit-mound geometry seen with the low carbon steel could be of benefit if large operational movements are not to be inhibited. One must, however, be careful in making this last observation as pit-pit damage has been seen also in the fretting of steels. 4.4. Future work In this paper data from two experimental programmes have been presented and interpreted in a very speculative manner. The analysis of the LClB is incomplete and future work can investigate such aspects as superficial contact area. However, the LW5 work, which was completed some years ago, demonstrates the futility of such restricted experimental programmes. Whilst there is a need to conduct short-term experiments on cermet coatings to refine or reject the proposed model, it is also apparent that any test programme needs to include time as a primary variable. The role of interfacial load and slip amplitudes in influencing wear of these coatings also needs to be investigated. It is also important to establish a statistical assessment of the spread in the data.
5. Closure In view of the very speculative nature of the model proposed and the inferences which ensue it is considered inappropriate to draw conclusions, hence this closure. Fretting wear data for sprayed cermet coatings obtained in two experimental programmes have been reviewed and compared with a twostage wear model. It is surmised that an initial wear process results essentially from an interlocking and removal of asperities, and this is followed by a mild, low, wear rate. The initial wear mode results in an apparently high wear rate which is dependent on superficial contact area and this dominates the volume removed for considerable sliding distances. At low temperatures such behaviour provides an explanation for apparently similar behaviour between structural steels and cermet coatings at short sliding distances. At high temperatures (600 “C) the fretting wear performance of structural steels appears to be as good as or better than that of the cermet coatings. However, if components are also required to operate for a period of time below their wear transition temperature then cermets should demonstrate an improved performance overall.
Acknowledgments The author gratefully acknowledges the efforts of his colleagues in the National Nuclear Corporation at Whetstone and Risley and at the Wolfson
76
Institute of Interfacial Technology, Nottingham University, the data presented here. This paper is published by permission Electricity Generating Board.
in producing of the Central
References 1 J. F. Archard, J. Appl. Phys., 24 (1953) 981 - 988. 2 T. C. Chivers, S. C. Gordelier, J. Roy and M. Wharton, Proc. Conf. on Vibration Nuclear Plant, Keswick, 1978, British Nuclear Energy Society, London, 1978. 3 T. C. Chivers and S. J. Radcliffe, Wear, 57 (1979) 313 - 321. 4 J. Skinner, J. Lubr. Technol., 103 (2) (1981) 228 - 235. 5 M. B. J. Low, Wear, 106 (1985) 315 - 335. 6 R. S. Millman, A. Sleightholme and A. A. Parry, Wear, 106 (1985) 77 - 95. 7 D. H. Jones, A. Y. Nehru and J. Skinner, Wear, 106 (1985) 139 - 162.
in