On the tribological behaviour of mechanical seal face materials in dry line contact Part I. Mechanical carbon

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1 Wear 256 (2004) On the tribological behaviour of mechanical seal face materials in dry line contact Part I. Mechanical carbon G.A. Jones Aeronautical and Mechanical Engineering Department, University of Salford, Newton Building, Salford M5 4WT, UK Received 4 September 2002; received in revised form 2 July 2003; accepted 2 July 2003 Abstract Mechanical face seal performance is critically dependant on the tribology of the seal face materials. In this research the tribology of the premier seal face materials under dry running contact has been examined. It is shown that the PV (pressure velocity) capability, of these premier face materials, is dependant on the development of a carbon graphite contact film. Experimentation has shown that these contact films breakdown via a seizure mechanism. The breakdown of the contact film is shown to result in a transition from a very low wear rate regime to a severe wear rate regime. Analysis of the results and the performance profiles produced has allowed the mechanisms of contact film formation and the functionality of the film in solid contact to be identified. A contact model is proposed that describes the low friction characteristic of mechanical carbon by combining the theories of Lancaster and Savage. The model is finally related to the mechanical seal face design Elsevier B.V. All rights reserved. Keywords: Tribology; Seals; Carbon; Graphite; Friction; Wear 1. Introduction Mechanical carbon refers to a series of complex polycrystalline materials. Within this series of materials there are grades that possess the mechanical properties and tribological characteristics that have resulted in their wide use as face materials in mechanical seals. The present leading seal face combination is a high duty mechanical carbon running against a reaction bonded or alpha sintered silicon carbide. The mechanical seal designer when selecting a mechanical carbon grade as a seal face material is primarily concerned in obtaining the best friction and wear behaviour. However, the selection process is very much an empirical process, depending on the data from seal tests in the laboratory and information derived from field applications. It is accepted that under dry running conditions the tribological performance of the leading seal face materials is due, predominantly, to the low frictional character of the mechanical carbon. However, the defining material characteristics, which determine this friction and wear behaviour, is extremely complex and the topic is still the subject of detailed research. Current address: Tribo-Tech Consultancy, Gledco Engineered Materials Ltd, Bankfield Terrace, Leeds, LS4 2JR. Tel.: ; fax: address: gordon@gajones.fsbusiness.co.uk (G.A. Jones). During the period a significant amount of work was carried out on the friction and wear of carbon and graphite and a strong body of knowledge was generated that developed an understanding of specific influential mechanisms. Subsequent work has built on this knowledge and understanding, though many aspects still remain imperfectly understood. From the available body of work three main conclusions can be drawn that describes the basic fundamentals governing the friction and wear of carbon and graphite: 1. The friction and wear of carbon and graphite is extremely sensitive to the environmental conditions within the contact zone. 2. The formation of stable surface films of wear debris, following structural breakdown of the carbon or graphite surface, significantly influences the friction and wear behaviour developed. 3. The wear of carbon and graphite can be by means of several processes but it is invariably driven and dictated by (1) and (2) above. The purpose of this research was to add to the currently available body of work by: (1) contributing to the development of a greater understanding of the tribo-dynamics of the carbon graphite surface films and film forming process, (2) developing an integrated theory and model that combines /$ see front matter 2003 Elsevier B.V. All rights reserved. doi: /s (03)

2 416 G.A. Jones / Wear 256 (2004) Nomenclature a semi-width for a cylinder on a plane A constant F total friction force M molecular weight of the bombarding gas P gas partial pressure Q activation energy of evaporation of a vapour from the surface R universal gas constant T absolute gas temperature T frictional heat generated T s surface temperature W load Greek symbols µ c friction coefficient of contact film µ t total friction coefficient µ v friction coefficient of vapour monolayer the three conclusions listed and (3) relating the developed theory and model to mechanical face seal tribo-performance. The experimental rationale adopted simulates the tribology of the sealing interface without the difficulty of complex design interactions of a mechanical face seal arrangement. In actual operation seal faces are not parallel, they deviate in terms of their macro-geometry in both the radial and tangential directions. The tangential deviation produces a circumferential waviness contact profile at the sealing interface and the contact at the peaks of the waves can, to a first approximation, be likened to a line contact geometry. Accepting this analogy, a pad on cylinder geometry was used. This presented line contact and permitted the tribo-dynamics of the contact film and film forming process to be evaluated. Several commercial mechanical carbons of varying degrees of graphitic order in contact with alpha sintered and/or reaction bonded silicon carbide were examined. Two experimental methods were adopted, one examined the PV capabilities of the mating combinations and the other measured the stability of performance under constant operating conditions. 2. Experiment An Amsler friction and wear testing machine in the pad on cylinder configuration was used, Fig. 1. The pads were manufactured from alpha sintered silicon carbide and reaction bonded silicon carbide and the rotating cylinders were manufactured from various commercial grades of mechanical carbon. Table 1 provides some nominal mechanical and physical property data on the materials used. The mechanical carbon grades selected were derived from different formulations to provided subtle variations in thermal and elastic properties. Samples of material from batches from which the subject materials were taken were examined metallographically to confirm material quality. In order to get a complete picture of the tribological behaviour of these materials under dry running conditions three experiments were carried out Experiment 1 determination of pressure velocity (PV) capability The PV limit was used in this experiment to evaluate the tolerance of the various seal face material combinations to varying degrees of solid contact. A high PV limit demonstrates an ability to operate under conditions of high applied loads and surface velocities and, in the case of lubricated contact, to cope with periods of fluid film starvation. The surface presentation and the surface texture parameters used here were selected in order to permit subsequent analysis by modelling the contact condition to that of a plane on a regular undulating surface. This contact arrangement was also selected so that the development and breakdown of the carbon graphite contact film could be seen and monitored and the impact of film dynamics on the PV limit could be evaluated. Table 2 details the material combinations examined. The experimental procedure followed was that the mechanical carbon cylinders were rotated at a constant surface velocity of 0.5 m/s. The load was increased in increments up to the point of contact instability. The details of the procedure are that the mating surfaces of the cylinder and the pad were prepared to give the following surface parameters: R a ( m) R q ( m) S m ( m) Cylinder Pad < >500 Prior to the test run the pad and cylinder components were ultrasonically cleaned in acetone and dried in an oven at 80 C. A thin wire thermocouple was bonded to the stationary pad approximately 0.25 mm from the mating face using an Araldite adhesive which was left for 24 h to obtain maximum strength. The thermocouple had an accuracy of ±0.1 C. The pad and cylinder were carefully located on the apparatus and the pad was held so that it accurately presented its contacting surface parallel to the longitudinal axis of the cylinder. The rotational speed was set at zero and the machine was switched on. The rotational speed was rapidly ramped up to the rpm that gave a surface velocity of 0.5 m/s then the first experimental load was applied. This start up procedure took on average 15 s. After completing the start up, the load was increased in increments, decided by the performance of the combination. At each stage, the friction and temperature were allowed to stabilise for 30 min and then a period of stable operation of around cycles was permitted before an incremental load adjustment was made. This stepwise increase in load continued until stable operation could not be sustained. The friction coefficient

3 G.A. Jones / Wear 256 (2004) Fig. 1. (a) Amsler friction and wear testing machine; (b) test specimen arrangement showing pad and cylinder geometry and the jig for holding the test pads such that line contact is presented. was measured continuously through the dynamic frictional torque generated (error ±0.05 N m). The wear rate was measured by the decrease in distance, over time, between the test head of the apparatus and a reference point, using slip gauges and feeler gauges. The frictional heat generated was measured through the thermocouple bonded to the sidewall of the pad. Relative humidity in the laboratory in the vicinity of the experiment was monitored and this was consistently around 55% RH. Two test runs per material combination were carried out. Each test run provided information on PV performance. Each material combination provided a PV profile including a PV

4 418 G.A. Jones / Wear 256 (2004) Table 1 Mechanical and physical property details of the mechanical carbon and the silicon carbide materials, used in this work a Material grade Material description Mechanical and physical properties Density (kg/m 3 ) Hardness Elastic modulus (GPa) Specific heat (J/kg K) Mechanical carbon Grade 001 Carbon + resin impregnation Shore A Grade 002 Carbon/graphite + resin impregnation Shore A Grade 003 Graphite + resin impregnation Shore A Grade 004 Carbon/graphite + resin impregnation Shore A Grade 005 Carbon + resin impregnation Shore A Grade 006 Carbon/graphite + resin impregnation Shore A Grade 007 Carbon/graphite + resin impregnation Shore A Silicon carbide Grade 013 Alpha sintered silicon carbide, >98.5% Hv Grade 014 Reaction bonded silicon carbide Hv Thermal conductivity (W/m K) a Metallographic analysis: percent resin and percent porosity by volume across the range of mechanical carbon materials was estimated to be between 6 and 10 and <3%, respectively. Mean grain size of the alpha sintered silicon carbide was 7.00 m and the percent porosity by volume was estimated to be <3%. Mean grain size of the reaction bonded silicon carbide was 9.00 m, the percent free silicon and percent porosity by volume was estimated to be 10 and <3%, respectively. limit above which stable conditions could not be sustained. Also, information was obtained on friction coefficient, wear rate, frictional heat generated and the role of carbon graphite contact films on the performance profile generated Experiment 2 performance under constant conditions In support of the information generated from the PV testing the stability of performance under constant contact conditions was examined. The purpose was to evaluate the Table 2 Material combinations examined in this work Mechanical Silicon carbide counterface carbon Experiment 1 Grade 001 Grade 013 (polished; R a = 0.05 m) Grade 002 Grade 003 Grade 004 Grade 005 Grade 006 Grade 007 Grade 001 Grade 014 (polished; R a = 0.05 m) Grade 002 Grade 003 Grade 004 Experiment 2 Grade 001 Grade 013 (polished; R a = 0.05 m) Grade 002 Grade 003 Grade 004 Grade 005 Grade 006 Grade 007 Experiment 3 Grade 001 Grade 013 (coarse; R a = 0.15 m) Grade 002 Grade 003 Grade 004 potential for the transition from stable to unstable operating conditions being cyclic and time dependent. The experiment was prepared and set up as per experiment 1. A constant PV value was set. The surface speed was maintained at 0.5 m/s and the load was adjusted to give a PV value of approximately 50% of the measured PV limit for the material combination under examination. The duration of the experiment was timed to give a total sliding distance of 25 km ( cycles). Two test runs per material were carried out. The experiment provided further information on friction coefficient, frictional heat generated, wear rate and contact film dynamics and stability under constant operating conditions. The friction coefficient, frictional heat generated and the wear rate were measured in the same way as described above, with the same degree of accuracy. In addition the wear rate was further assessed by measuring the weight change before and after the test run, to five decimal places and to an accuracy of ±0.01 mg. Table 2 details the material combinations examined in this experiment Experiment 3 the influence of SiC surface texture In this experiment the influence of surface texture of the silicon carbide counterface on the contact film dynamics and its consequent influence on the tribological performance was evaluated. Experiment 1 was repeated, using various mechanical carbon alpha sintered silicon carbide combinations, Table 2. The starting surface texture of the silicon carbide was produced to a coarser finish. This was done by increasing the amplitude parameters, R a and R q and reducing the spatial parameter S m. The mechanical carbon cylinders were left in the as-machined condition. The surface texture on the alpha sintered silicon carbide pads were produced by controlled lapping and polishing to give the following parameters:

5 G.A. Jones / Wear 256 (2004) R a ( m) R q ( m) S m ( m) Cylinder (as turned) Pad At least two test runs per material combination were carried out. Each material combination provided a PV profile and information on the effect of the starting surface texture of the alpha sintered silicon carbide on the tribo-dynamics of the carbon graphite contact film produced. Also, information was obtained on friction coefficient, wear rate, and frictional heat generated. 3. Results 3.1. Bearing surfaces The surface conditions of the mating materials before and after each test run are described in Table 3. In all the cases examined in the three experiments both bearing surfaces exhibited evidence of contact film formation. The pad on cylinder configuration allowed the formation of a carbon graphite contact film on the mechanical carbon cylinder itself to be observed during test. From experiment 1 the point of contact instability was indicated by the disintegration of this film. Visual examination and surface roughness measurements of the contact surface of the cylinder, after test, indicated the formation of a very random coarse surface in which relatively large tracts of material had been removed by ploughing, tearing or cutting. The edges of the cylinder, which constituted the edges of the contact area, also became heavily chipped. Examination of the surface of the pads also revealed the presence of a contact film, which after test was, also, in a fragmented and fractured condition. Repeat runs, terminated at the point of contact film failure, identified blistering of the film as the key failure mechanism. The photomicrographs in Fig. 2 show the surfaces exhibiting blistering of the contact film. Table 3 Details of the wear surfaces before and after each test run in experiments 1 and 3 Mechanical carbon grade R a ( m) Rotating mechanical carbon cylinder Stationary SiC pad (grade 013) Stationary SiC pad (grade 014) Before After Before After Before After Experiment 1 Test #1 Grade Test #2 Repeat Test #3 Grade Test #4 Repeat Test #5 Grade Test #6 Repeat Test #7 Grade Test #8 Repeat Test #9 Grade Test #10 Repeat Test #11 Grade Test #12 Repeat Test #13 Grade Test #14 Repeat Test #15 Grade Test #16 Repeat Test #17 Grade Test #18 Repeat Test #19 Grade Test #20 Repeat Test #21 Grade Test #22 Repeat Experiment 3 Test #30 Grade Test #31 Repeat Test #32 Grade Test #33 Repeat Test #34 Grade Test #35 Repeat Test #36 Grade Test #37 Repeat

6 420 G.A. Jones / Wear 256 (2004) Fig. 2. Photomicrographs of carbon graphite contact film failure: (a) and (b) the blistering of the contact film on the mechanical carbon material; (c) and (d) the blistering of the contact film formed on the silicon carbide counterface.

7 G.A. Jones / Wear 256 (2004) Fig. 2. (Continued ) Experiment 1 determination of the pressure velocity (PV) capability The PV was calculated by relating the contact area to the normal load supported and the surface velocity of the mechanical carbon cylinder. The contact area, for a cylinder on a plane, was calculated from the Hertzian equation for semi-width, a, such that the contact area = 2al, where l is the length of the cylinder, and the contact pressure was derived by dividing the normal load by the calculated contact area. Any warping of the contact area resulting from tangential traction s was neglected. Fig. 3 illustrates the form of the PV profiles produced and Table 4 details the PV limits recorded for the material combinations examined. In all cases the PV limit was shown to be the value at which the frictional character of contact and the wear mechanism transformed to a more severe regime. This transformation was accompanied by gross failure of the contact film. The results produced show the broad range of triboperformance possible from these seal face materials. The results show that the mechanical carbon grade 004 can offer a PV capability that is twice that of grade 005 and progressively better, on average, than that given by the other intermediate grades. These results clearly demonstrate that, under dry running conditions, the PV capability is critically dependant on the stability of the contact film formed. Also, the importance of the counterface in supporting the stability of the contact film was shown. The PV performance of mechanical carbon in contact with alpha sintered silicon carbide was shown to be appreciably better than when in contact with reaction bonded silicon carbide. The results show that the PV capability was significantly reduced and the difference in PV performance from one mechanical carbon grade to another was less. However, the ranking order was shown to remain essentially the same. Table 4 PV limits for the various material combinations examined in experiments 1 and 3 Mechanical carbon cylinder PV limit of silicon carbide pad (MPa m/s) Grade 013 Grade 014 Experiment 1 Test #1 Grade Test #2 Repeat Test #3 Grade Test #4 Repeat Test #5 Grade Test #6 Repeat Test #7 Grade Test #8 Repeat Test #9 Grade Test #10 Repeat Test #11 Grade Test #12 Repeat Test #13 Grade Test #14 Repeat Test #15 Grade Test #16 Repeat Test #17 Grade Test #18 Repeat Test #19 Grade Test #20 Repeat Test #21 Grade Test #22 Repeat Experiment 3 Test #30 Grade Test #31 Repeat Test #32 Grade Test #33 Repeat Test #34 Grade Test #35 Repeat Test #36 Grade Test #37 Repeat 59.55

8 422 G.A. Jones / Wear 256 (2004) Frictional Torque (N.m) PV Value (MPa.m/s) Fig. 3. PV profile range for the mechanical carbon grades in contact with alpha sintered silicon carbide and reaction bonded silicon carbide, in grey. The combinations examined typically produced profiles that fell within the boundaries shown. The friction coefficient was calculated from the measured frictional torque to an accuracy of ±0.01. Fig. 4 shows the form of the friction coefficient/pv profiles produced. In all cases the friction coefficient reduced with increasing PV and this was common and consistent across the range of combinations examined. A typical profile was characterised by the two transition points. The first transition was qualitative and signified full coverage of the contact surfaces with a contact film. The second transition was a classic transition from mild to severe wear. The friction coefficients recorded for those mechanical carbon grades in contact with reaction bonded silicon carbide were consistently lower, for the same contact conditions, than those recorded for contacts with alpha sintered silicon carbide, Fig Friction Coefficient First Transition: Full coverage of the contact surfaces with a contact film. Second Transition: Contact film failure PV Value (MPa.m/s) Fig. 4. Friction coefficient profile range for the mechanical carbon grades in contact with alpha sintered silicon carbide and reaction bonded silicon carbide, in grey. The combinations examined typically produced profiles that fell within the boundaries shown.

9 G.A. Jones / Wear 256 (2004) The maximum temperature at which contact film disintegration took place for each material combination was examined. The results obtained demonstrated that the temperature at film failure is approximately the same for all mechanical carbon grades examined in contact with either polished alpha sintered silicon carbide or polished reaction bonded silicon carbide. The temperatures recorded were predominantly within 10 C of a mean temperature of 73 C. These trends demonstrate that contact film disintegration is very closely linked to the interface temperature. Fig. 5. Photomicrographs showing details of contact film development. Photomicrographs (a) and (b) show the structure of the film developed on the mechanical carbon; (c) (e) show the structure of the film developed on the silicon carbide counterface; (f) and (g) compliment the patch build up hypothesis.

10 424 G.A. Jones / Wear 256 (2004) Fig. 5. (Continued ).

11 G.A. Jones / Wear 256 (2004) Due to the nature of the experiment the wear rate was difficult to measure and quantifiable data was not collected. The experimental observations of the wear process, however, are detailed. In all cases wear was not detected until the second transition point. At this point the wear rate transformed from an apparent zero wear rate to approximately 15 mm/h and this was roughly the same for all combinations examined. Microscopic examination of the contact surfaces before the second transition revealed that wear debris was in fact generated and entered into the process of contact film formation. In the cases examined a significant amount of the wear particles produced remained within the contact zone and only a relatively few were expelled in the form of loose wear debris. Those particles that remained in the contact zone were comminuted by repeated deformation and fracture. When the particles had been fragmented to some critical size, they became agglomerated at suitable sites on the wear surfaces, due to adhesion forces resulting from the relative surface energies of the contact materials, to form the stable contact films observed. The photomicrographs in Fig. 5 show the stages in the development of this film on both surfaces. Further microscopic examination of the contact surfaces revealed that during sliding there are two competing processes taking place. Selected repeat runs were interrupted at low PV values and examination of the contact surfaces revealed the presence of film disruption, areas of film regeneration and the presence of minor blistering of the contact film. This identifies a balance between the breakdown of the film or layers of film, which cause wear, and agglomeration, or pseudo sintering, between the particles within the film or layers of film, which lead to consolidation. The second transition is the point at which the balance between these two competing processes is lost in favour of film breakdown. The photomicrographs in Fig. 6 show the general contact surface condition and the minor blisters on both contact surfaces Experiment 2 performance under constant conditions Fig. 7 shows the form of the profile produced for the friction coefficient (±0.01) as a function of distance travelled. Table 5 details the maximum, minimum and mean friction coefficients and the typical temperatures generated on the stationary pad at 0.25 mm from the contact surface. Also included in the table is the overall wear of the mechanical carbon cylinder. These results show a consistency of operation. In all the cases examined the contact film built up over a short running in period at the start of the test run. After this period the test continued uninterrupted and without change in the operating mode. There was no evidence of periodic film disintegration and re-establishment on a major scale, as alluded to by other workers. There was evidence, however, of the balance between film/layer breakdown and agglomeration, as mentioned in experiment 1. The wear rate was again difficult to measure because of the low rate of loose particle generation but a wear rate calculation was made from the change in weight projected over the duration of the test run. The wear rate results produced were scattered across two orders of magnitude with no discernable correlation. It is proposed here that the steady state wear rate of mechanical carbon is a function of the dynamics of the contact film disintegration regeneration process and the consequent rate of loose particle expulsion from the contact zone. This is a complex relationship, which is the subject of further work by this author Experiment 3 the influence of SiC surface texture Increasing the surface roughness parameter R a and reducing the spatial parameter S m of the starting surface texture of the alpha sintered silicon carbide resulted in an approximate 30 40% reduction in the PV limit compared with that obtained in experiment 1 for polished alpha sintered silicon carbide. Also, the frictional heat developed at the PV limit was reduced by 10 30%. There was again evidence of contact film formation and there was evidence of the film/layer breakdown and consolidation process. The PV limit was indicated by gross failure of the contact film, which took place at a relatively common temperature and temperature rise ( T 57 C). The change in surface texture had little effect on the friction coefficient. The wear rate was again very difficult to measure Examination of the contact films formed SEM, TEM and optical microscopy of selected contact films, taken from both contact surfaces, showed that they were made up from fine debris. This debris was shown to fill the surface irregularities on both the mechanical carbon and the counterface and become consolidated into larger particles and structures by an agglomeration/pseudo-sintering process. Optical microscopy and SEM, Figs. 2, 5 and 6, showed that adjacent layers overlap each other in a manner similar to the tiles on a roof. TEM, Fig. 5, showed that the layers were made up of grains of around 1.0 m in diameter and these were, in turn, made up of subgrains of about 0.1 min diameter. These observations suggest that the build up of the contact film, on both surfaces, occurs in patches. These are developed from the breakdown of the randomly orientated polycrystalline grains of the mechanical carbon, to a size of approximately 1.0 m in diameter. These patches form separately, grow, impinge and overlap until they produce a continuous film. The thickness of the film is determined to a very large extent by the surface irregularities on both the contacting surfaces. In this work film thickness was shown to be in a range of 1 5 m. Microprobe analysis of grades 001, 002, and 003, and their associate contact films, from contact with alpha sintered silicon carbide, detected, in all cases, a significant amount of O 2 in the contact film compared to that in the base material.

12 426 G.A. Jones / Wear 256 (2004) Fig. 6. Photomicrographs of the minor blisters that generate on the contact surfaces during stable operation. Photomicrographs (a) and (b) show those formed on the mechanical carbon; (c) (e) show those formed on the silicon carbide surface.

13 G.A. Jones / Wear 256 (2004) Fig. 6. (Continued ). Table 5 Details from experiment 2 a Test number Mechanical carbon grade Normal load (N) Friction coefficient Typical T ( C) Wear rate (m 3 /m) Maximum Minimum Mean 23 Grade Grade Grade Grade Grade Grade Grade a Mechanical carbon in contact with alpha sintered silicon carbide (grade 013) under constant operating conditions.

14 428 G.A. Jones / Wear 256 (2004) Friction Coefficient Distance Travelled (Cycles x 10 3 ) Fig. 7. Friction coefficient as a function of distance travelled. All combinations examined in experiment 2 produced profiles that fell within the range shown. 4. Discussion analysis of results The results of this work have demonstrated the fundamental importance of stable contact films to the performance of mechanical carbon in line contact with silicon carbide under dry running conditions. The work has identified a mechanism of formation, a mechanism of failure and the tribo-dynamics under stable operating conditions. The critical features have been identified as: 1. The breakdown of the surface structure of the mechanical carbon into units of wear debris of around 1.0 m in diameter. 2. The consolidation and agglomeration of the generated wear debris on to both the contacting surfaces. 3. The development of the adhering wear debris into larger layer type structures, which overlap to form a film, which covers large areas of the contact surfaces. 4. The development, during stable operation, of a balance between film breakdown and film consolidation. 5. The identification of film failure by a seizure and blistering type mechanism, which occurs at a common temperature within the contact zone. 6. The increased presence of O 2 in the contact film and therefore the potential of condensable vapours in the film/layer forming process. The experiments have shown that these characteristics combine to generate a third phase between the contacting surfaces which has associated with it a low resistance to tangential sliding and it is the stability of this third phase that dictates the tribological performance. In this work a number of different mechanical carbon grades have been examined. These were made up from different formulations and had different bulk mechanical properties. Irrespective of this difference they showed significant similarity in their tribological performance. The exception to this was the PV limit, which was effectively the only defining characteristic. It has been shown that similarities existed across the range of combinations examined in terms of the friction coefficient/pv profiles (with the exception of the PV limit), the temperature at the point of contact film failure, and the wear character. Also it was shown that the mechanical carbon materials examined responded in the same manner to a change in the surface texture of the alpha sintered silicon carbide and a similar response was also shown when reaction bonded silicon carbide was used in place of the alpha sintered silicon carbide counterface. These similarities in performance can be related to the macro-similarities in the nature and tribo-characteristics of the contact film formed. The mechanical carbon grades examined in this work, though different in formulation and mechanical properties, are constructed from the same basic graphite crystallite unit. The contact films formed are, therefore, composites of this same base unit and it is this common structural feature that provides the macro-similarity between the resultant contact films. The defining PV limit, however, cannot be described or explained at this macro-level. The PV limit defines the specific bearing power above which the contact film becomes unstable and film failure results. To understand this and identify the fundamental principles involved, the influence of the graphite crystallite size and orientation and the associated

15 G.A. Jones / Wear 256 (2004) surface reactivity need to be evaluated. The following model is proposed An integrated model for dry running contact The contact has to satisfy a mechanical condition and a thermal condition. However, the contact environment, the surface reactivity of mechanical carbon and the properties of the contact film formed will greatly influence how these conditions are satisfied. Savage [1] and Savage and Schaefer [2] suggested that the low friction of graphite was associated with low surface forces due to the surface adsorption of water or other condensable vapours. This definitive work was further supported by subsequent studies [3 5] in this field. Jenkins [6], Porgess and Wilman [7], Midgley and Teer [8] and Quinn [9], showed that worn graphite surfaces exhibit a preferred orientation of basal planes. Workers [7 9] further demonstrated that in fact the basal planes were tilted at some small angle with their normals pointing against the direction of sliding. Roselman and Tabor [10] identified the basal planes as low energy surfaces. Also, assuming that the contacts during sliding are mainly between basal plane surfaces, they suggested that shear occurred primarily at these interfaces, rather than beneath the surface, leading to low friction and wear. Deacon and Goodman [11] and Campbell and Kozak [12] highlighted the significance of edge site interactions on the friction and wear behaviour developed. Lancaster [13] and Clarke and Lancaster [14] examined the contact films formed by various carbon brush materials and by carbon materials used for dry bearings. They related the low friction of these materials to the low strength of the contact films and suggested that the film acts as a thin solid lubricant. It was also suggested that the low strength of the contact film was a consequence of weak intercrystalline bonding. The results produced here lend further support to the Lancaster/Clarke film theory but also the results suggest that an integrated model for dry running contact should include a component relating to the adsorption of condensable vapours. A contact model that combines both these concepts is proposed. In this work it has been shown, in agreement with [15 18], that during sliding the contact film, on a macroscopic scale, is in a continuous state of disintegration and regeneration and it is shown that total film failure occurs when the disintegration part of this cycle dominates. The failure of the film is shown to have a blistering component to the process and this is in agreement with the findings of Arnell et al. [3], Midgley and Teer [4] and Swinnerton and Turner [19]. In addition to this, measurements of the frictional heat generated close to the contact zone imply that contact film failure occurs at a common interface temperature, and this is in agreement with Lancaster [20]. These factors together describe a seizure mechanism in which the physisorbed components are removed from the contact surfaces and the resulting intimate contact between these surfaces generates high adhesive forces and consequent contact film failure. The breakdown of the randomly orientated polycrystalline grains at the surface of the mechanical carbon generates wear debris that is highly reactive (orders of magnitude more reactive than activated charcoal [16]) and adsorbs polar gas (H 2 O, O 2,CO 2,NH 3,CH 3 OH, etc.) from the surrounding environment. The adsorption generally takes place at the edges of the crystallites within the debris where there are unsatisfied bonding forces. The saturation of these active edge sites promotes the creation of a contact film of low shear strength, which is covered by a transient monolayer of adsorbed vapour molecules. This lowers the free energy of the contacting surfaces and generates an interface that has a low resistance to tangential sliding [21]. Increasing the severity of the contact by increasing the PV parameter causes the interface temperature to increase. At the PV limit the interface temperature is such that water vapour, etc. is no longer physisorbed on to the contact surface and adhesive crystallite edge interactions can now take place and the film is destroyed. Microprobe analysis identified a higher level of O 2 in the contact films than that found in the bulk material. This supports the adsorption theory and aligns with the findings of Buckley and Johnson [22] who suggested a carbon oxygen bond and noted a 1000 increase in wear rate when the air pressure surrounding the interface was reduced to N/m 2. They also noted the disappearance of the carbon graphite contact film from the mating surfaces. The disintegration/regeneration process associated with contact film stability is also a symptom of the above. Under stable operating conditions with both bearing surface covered with a smooth contact film contact between the two surfaces will take place at discrete microscopic locations, which are continually changing. The flash temperatures at these locations will be such that the physisorbed vapour will be driven off, adhesive crystallite edge interactions will take place and the film will be disrupted. Subsequent localised contact relief will then allow the film to regenerate and repair in the denuded areas while film disruption takes place elsewhere within the contact zone. If the PV parameter is now increased more contact spots are created, more film disruption takes place and the general interface temperature increases. Film regeneration, however, keeps up with the demand until, at the PV limit, it is unable to do so and the film is completely destroyed. The effects described above can be visualised by relating the rate at which vapour molecules bombard the surface to the rate at which they are evaporated. The rate of bombardment is given by [1] n b = ( cm 2 s 1 )p (MT) 1/2 (1) where p is the gas partial pressure, M the molecular weight of the bombarding gas and T the absolute gas temperature.

16 430 G.A. Jones / Wear 256 (2004) The rate of evaporation from the surface is given by [1] n e = A e Q/RT s (2) where A is a constant, Q the activation energy of evaporation of a vapour from the surface, T s the surface temperature and R the universal gas constant. At the flash temperatures generated at the local contact points, the rate of evaporation will be significantly greater than the rate of bombardment and the film will be disrupted. On a larger scale, when the general surface temperature reaches a critical value the rate of evaporation across the whole contact surface will be greater than the rate of bombardment and total film disintegration then takes place. Relating this rationale to the friction condition, the load W is supported partly (fw) by the monolayer of vapour having a friction coefficient of µ v and partly ((1 f)w) by the contact film having a friction coefficient µ c. The total friction force F (µ t W) can be expressed as F = µ v fw + µ c (1 f)w (3) This expression, to a first approximation, represents the state of equilibrium between film disintegration and film regeneration and it allows the description of the tribo-performance in terms of the friction coefficient. At the PV limit f tends towards 0 and the friction condition is then dominated by adhesive crystallite edge interactions and this leads to contact film failure. On the wear rate of mechanical carbon under stable operating conditions, this work as shown that this is quite benign. The debris from localised contact film failure is available for film regeneration and it only constitutes wear in terms of material removal if it is expelled from the contact zone. This can be related to mechanical seals where there are a number of adjacent contact zones within the sealing interface and there are only two boundaries from which debris can be expelled as wear. The contact geometry in mechanical seals facilitates the reuse of debris, from localised contact film failure, in the film regeneration process and under stable operating conditions benign wear rates can be expected. The mode by which the contact film fails and generates reusable debris is very complex. Surface fatigue is considered to be a contributor [14,15,17], so too must Hertzian contact stresses [17] and elastic compressive and thermal strain recovery. The above model combines both the Lancaster and Savage theories on the frictional behaviour of carbon graphite materials. The model proposes that mechanical carbon materials depend for tribo-performance on the generation of a contact film of low shear strength on one or both contacting surfaces and the generated contact film requires the presence, within the contact zone, of water or other vapour molecules for the development of its structure and properties. It has not been possible to develop a full mathematical model of the contact condition because further more detailed information about the crystallite structure and the properties of the contact film formed would be required. Also, the precise mode of contact film rupture needs further clarification. Nevertheless this model still offers a very useful tool Use of the model This work has shown that the film forming characteristics of the mechanical carbon grades examined, the surface texture of the counterface and the type of silicon carbide used influence the tribo-performance produced. The usefulness and functionality of the model is demonstrated in explanation of these effects. With respect to the film forming characteristics of the mechanical carbon grades, wear debris composed of small graphite crystallites will have a higher concentration of edge sites and hence adsorption sites. Materials manufactured from formulations containing a high proportion of fine non-graphitic fillers will develop wear debris and contact films with a high concentration of small crystallites and consequently have a high demand for condensable vapours. With this higher proportion of small crystallites there is a greater probability of adhesive crystallite edge interactions. Both film disintegration and regeneration occur more rapidly and the interface temperature is higher for a given contact condition. The balance between the rate of molecular bombardment and evaporation is harder to maintain and as a consequence f tends towards 0 at a lower PV value for these types of formulations. Grades 001 and 005 are materials with a very high concentration of fine non-graphitic fillers and small graphite crystallite components and grade 003 is a material with a higher graphitic character composed of larger graphite crystallites. The remaining grades have compositions between these two extremes. Grades 001 and 005 failed at a much lower PV than grade 003. Grades 002 and 004 produced the highest PV capability. These materials contained, within their composition, components that supported the film disintegration and regeneration process, to a greater degree. Grades 006 and 007 have a basic carbon graphite composition with an even distribution of graphite crystallite sizes and these gave PV limits that were within the mid-range. The reduction in the PV limit, for a given mechanical carbon grade, observed by increasing the coarseness of the contacting surface of the alpha sintered silicon carbide can be evaluated in the same way as the above. The increased textural coarseness increased the volume of the surface interstices that need to be filled with carbon graphite wear debris before a smooth surface is formed. This results in an increase in the gross film thickness. The contact film, in relation to the bulk silicon carbide, is in effect a thermal barrier. The structure of the contact film, being derived from the breakdown of randomly orientated polycrystalline grains at the surface of the mechanical carbon, will be highly disordered in terms of its graphitic nature. This will give a relatively high thermal resistance particularly when compared to that of silicon carbide. Increasing the film thickness, by increasing the textural coarseness, increases the efficiency of this

17 G.A. Jones / Wear 256 (2004) barrier and limits the heat sink provided by the counterface. The result of this is that the balance between the rates of molecular bombardment and evaporation is harder to satisfy and as a consequence f tends towards 0 at a lower PV value and the film disintegrates. A similar effect to this was observed when various mechanical carbon grades were run against reaction bonded silicon carbide. The elastic and thermal properties of reaction bonded silicon carbide and alpha sintered silicon carbide are very similar and the difference in the PV limits observed would not be expected. A potential explanation is suggested by microscopic examination of the material microstructure, wear surfaces and reference to the model and analysis above. The structure of reaction bonded silicon carbide consists of silicon carbide grains in a network of approximately 10% free silicon. The structure of alpha sintered silicon carbide contains no second separate phase. During the action of surface preparation and the action of sliding there are significant differences in the rates of material removal from the silicon carbide and free silicon phases in the reaction bonded silicon carbide resulting from differences in their hardness and wear resistance. This causes the free silicon to lie below the hard silicon carbide grains, producing a micro-surface texture not produced on the alpha sintered silicon carbide. This micro-surface texture provides surface interstices that need to be filled with carbon graphite wear debris producing a thicker contact film and again f tends towards 0 at a lower PV value. Arnell et al. [3] and Midgley and Teer [4] produced evidence of periodic film disintegration and re-establishment on a major scale and Teo and Lafdi [18] have further examined these friction transitions. This situation can be related to a process of over filming. With highly graphitic mechanical carbons there is the potential for the formation of relatively thick heavy surface film deposits. Williams et al. [17] confirmed this in their study of fine textured electro-graphite s and epoxy impregnated baked carbons. To overcome this problem many mechanical carbon manufacturers incorporate abrasive particles into the material formulation. The purpose of this is reported to be to prevent very smooth surfaces forming resulting in high real area of contact. The model proposed here offers a slightly different explanation. The model suggests that it is desirable to generate a thin contact film and promote the minor film disintegration and regeneration process. The role of the abrasive component is that it encourages both of these conditions. It is widely recognised that an abrasive particle size beyond 0.3 m will produce a large increase in wear rate of both the mechanical carbon and its associate counterface. This is due, primarily, to excessive film removal. All the mechanical carbons in this work contained within their formulation, to a greater or lesser degree, an abrasive component. Examination of the photomicrographs in Figs. 2, 5 and 6 shows minor abrasive wear on the surface of the films formed thus supporting the notion of micro-abrasion to control contact film thickness and promote the minor film disintegration and regeneration process Relevance of the contact model to mechanical face seal design As shown, the performance of a mechanical face seal under dry running conditions idealised to line contact, utilising a mechanical carbon as one of the seal faces, depends critically on the development of a stable contact film. If this does not form then the seal will function inefficiently and the service life will be short. The model developed here enables the mechanisms taking place at the sealing interface and the requirements for stable operation to be visualised and evaluated. The central objective is to maintain a dynamic balance between contact film disintegration and regeneration. The model permits a qualitative assessment of whether the duty conditions and seal face design will permit this. In gas turbine engine mainshaft bearing compartments, contacting face seals (and carbon ring seals) are subjected to high speeds and high temperatures. Notwithstanding the requirement to prevent the mechanical carbon from oxidising the stability of the contact film is critical to the performance of the seal. The model here provides the seal material designer with both a diagnostic tool and developmental tool. The model enables an evaluation of the likely impact of the contacting condition and environmental aspects on the tribo-dynamics of the contact film and film forming process. It provides the format for seal face and seal design performance mapping. It provides a framework for the development of seal face materials and seal face design. The model also offers a route(s) for the development of mechanical carbons for seal face applications in vacuum and inert gas atmospheres. There is also use for this model in lubricated mechanical face seals. In a liquid or gas lubricated mechanical seal there is always a degree of mechanical contact either at stop/starts and during general operation due to the close dimensional tolerances associated with thin liquid/gas films. During these periods of partial lubrication the contacting faces need to be able to cope with the tribo-stresses presented. In the case of mechanical carbon, it needs to generate a stable contact film and the model can again be used to predict the likely outcome. Importantly the model proposed here enables the key features of mechanical face seal design to be related to the critical tribological characteristics of the leading seal face materials. This is an important connection. The future development of mechanical seal face material technology hinges on a greater understanding and subsequent development of the tribo-mechanisms that promote low frictional forces within the sealing interface. 5. Conclusions Clearly, the results and analysis carried out in this research have added to the knowledge and understanding of the tribological behaviour of mechanical carbon and the leading seal face materials. The work has developed analysis and

18 432 G.A. Jones / Wear 256 (2004) tools that could be of use to both the seal designer and the materials technologist. This work has shown that a critical factor governing the tribo-performance of the leading seal face materials, under dry running conditions, is the ability of the mechanical carbon to form a contact film on both the silicon carbide counterface and on itself. The carbon graphite contact film has been shown to be a third phase of low shear strength that modifies the friction condition. The stability of this film is shown to be a function of the balance between film disintegration and regeneration and analysis has related this state to the balance between the rate at which the contacting surfaces are bombarded with vapour molecules, from the surrounding environment, and the rate at which these molecules are evaporated off. The failure of the contact film was shown to be complex and generated a blistering type of process. Adhesive crystalline edge site interactions, surface fatigue and elastic compressive and thermal strain recovery are potential contributory mechanisms. The contact model presented to describe the tribo-mechanisms observed, combines both the Lancaster and Savage theories of the friction of carbon and graphite. The model was used to rationalise the contact film dynamics and the wider experimental observations. The importance of the physical properties and the contact surface texture of the silicon carbide are shown. The model is related to mechanical face seal design and development providing a link between the critical mechanical face seal design criteria and the tribological characteristics of the seal face materials. Acknowledgements The author wishes to express his appreciation to Gledco Engineered Materials Ltd., US Graphite Inc., and the University of Salford for their support of this research. The advice and excellent technical assistance of Professor R.D. Arnell and Mr. G. France, University of Salford, are gratefully acknowledged. References [1] R.H. Savage, Graphite lubrication, J. Appl. Phys. 19 (1948) [2] R.H. Savage, D.L. Schaefer, Vapour lubrication of graphite sliding contacts, J. Appl. Phys. 27 (1956) [3] R.D. Arnell, J.W. Midgley, D.G. Teer, Frictional characteristics of pyrolytic carbon, Proc. Inst. Mech. Engrs. 179 (3j) (1968) [4] J.W. Midgley, D.G. Teer, An investigation of the mechanism of friction and wear of carbon, J. Bas. Eng. 62 (15) (1973) 1 6. [5] R.I. Langley, J.W. Midgley, A. Strang, D.G. Teer, Mechanism of the frictional behaviour of high, low and non-graphitic carbon, Proc. Inst. Mech. Engrs. E (1964) [6] R.O. Jenkins, Electron diffraction experiments with graphite and carbon surfaces, Philos. Mag. 17 (1934) [7] P.V.K. Porgess, H. Wilman, Surface re-orientation, friction and wear in the uni-directional abrasion of graphite, Proc. Phys. Soc. 76 (1960) [8] J.W. Midgley, D.G. Teer, Surface orientation and friction of graphite, graphitic carbon and non-graphitic carbon, Nature 189 (1961) [9] T.F.J. Quinn, A topographical and crystallographic study of the surfaces of rubbed electrographite, Br. J. Appl. Phys. 14 (1963) [10] I.C. Roselman, D. Tabor, The friction of carbon fibres, J. Phys. D 9 (1976) [11] R.F. Deacon, J.F. Goodman, Lubrication of lamellar solids, Proc. R. Soc. London Ser. A 243 (1958) [12] W.E. Campbell, R. Kozak, Studies in boundary lubrication. III. The wear of carbon brushes in dry atmospheres, ASME Trans. 70 (1948) [13] J.K. Lancaster, Dry bearings: a survey of materials and factors affecting their performance, Tribology 12 (1973) [14] W.T. Clarke, J.K. Lancaster, Breakdown and surface fatigue of carbon during repeated sliding, Wear 6 (1963) [15] P.J. Blau, R.L. Martin, Friction and wear of carbon-graphite materials against metal and ceramic counterfaces, Tribol. Int. 27 (6) (1994) [16] B.S. Nau, Mechanical seal face materials, Proc. Inst. Mech. Engrs. 211 (1997) [17] J.A. Williams, J.H. Morris, A. Ball, The effect of transfer layers on the surface contact and wear of carbon-graphite materials, Tribol. Int. 30 (9) (1997) [18] K.M. Teo, K. Lafdi, Friction and wear transitions in carbons (temperature and stress analysis), Tribol. Trans. 44 (4) (2001) [19] B.R.G. Swinnerton, M.J.B. Turner, Blistering of graphite films in sliding contact, Wear 9 (1966) [20] J.K. Lancaster, Transition in the friction and wear of carbons and graphite s sliding against themselves, ASLE Trans. 18 (1975) [21] H. Zaidi, R.D. Paulmier, Influence of absorbed gases on the surface energy of graphite: consequences on the friction behaviour, Thin Solid Films 264 (1995) [22] D.H. Buckley, R.L. Johnson, Mechanism of lubrication for solid carbon materials in vacuum to 10 9 millimetres of mercury, ASLE Trans. 7 (1964)