Influence of thermal properties on friction performance of carbon composites

Carbon 39 (2001) 1789–1801

Influence of thermal properties on friction performance of carbon composites Christopher Byrne*, Zhiyuan Wang Center for Advanced Friction Studies and Department of Mechanical Engineering and Energy Processes, Southern Illinois University, Carbondale, IL 62901, USA Received 15 May 2000; accepted 31 October 2000

Abstract Three different needled felt C / C composites containing from 5 to 25% fiber oriented normal to the friction surface (z-fiber) were evaluated and tested for friction performance. A laboratory dynamometer was used to simulate cold taxi, hot taxi and normal landing braking events utilizing a single stator and rotor pair. Temperatures measured near the friction surface were lowest for highest thermal diffusivity material demonstrating the effectiveness of z-fiber content at reducing friction surface temperature. Low friction coefficient (|0.12) under cold taxi conditions increased by a factor of two under hot taxi conditions due to water-desorption transitions. Composites with high diffusivity needed greater braking power to experience transition relative to lower diffusivity materials. Higher braking power was needed to produce a transition under higher humidity conditions. Highest z-fiber content discs showed the lowest wear rate which was attributed to higher in-plane shear strength. The wear rate was highest under hot taxi conditions and it was concluded that mechanisms other than oxidation loss were primarily responsible for the wear rate of the C / C composites tested in this study. The C / C composite surfaces had a polished appearance suggesting that the removal of nano-scale particles occurs from the friction process.  2001 Elsevier Science Ltd. All rights reserved. Keywords: A. Carbon / carbon composites; D. Acoustical properties; Frictional properties; Thermal diffusivity

1. Introduction Carbon / carbon composites (C / C) are now widely used as brake materials in the aerospace industry owing to their high strength at elevated temperature, low density, low coefficient of thermal expansion, high heat capacity, chemical inertness, self-lubrication ability, and good wear resistance. Of the various applications such as brakes, clutches, rocket motors, heatshields for re-entry vehicles and refractory components, 63% by volume of the C / C produced in the world are used as aircraft brakes. Compared to metal-based friction materials, C / C has many advantages including decreased brake weight, long disc life, less noise, smoother operation and absorption of more thermal energy per brake unit [1–4]. The friction behavior of carbon / carbon is influenced by many factors, such as composite type, friction parameters *Corresponding author. Tel.: 11-618-453-4560; fax: 11-618453-5260. E-mail address: [email protected] (C. Byrne).

(load, sliding speed, energy), environment (temperature, humidity, atmosphere) and prior history of composite surface [3,5]. Among them, surface temperature is a very important factor. If too high, a brake material may be more susceptible to abrasion, galling or oxidation. Size, shape and thermal diffusivity of the contact regions will influence the magnitude and direction of heat flow at the surface. In particular, increasing through-thickness thermal diffusivity will increase the rate of heat flow from the interface and thus lower the friction surface temperature. While this seems obvious, there has been no proof given in the literature regarding this effect on friction behavior. With increased use of needled felt preforms, and their inherent capability for predetermined architecture, there is an immediate need for understanding the influence of through-thickness properties on friction performance.

2. Background The tribology of conventional solid carbons has been

0008-6223 / 01 / $ – see front matter  2001 Elsevier Science Ltd. All rights reserved. PII: S0008-6223( 00 )00296-7

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studied for a long time. Bragg suggested the high lubrication of graphite is due to the crystal structure that allows easy shearing between adjacent planes [6]. However, this view was questioned after studies of the high wear rate (‘dusting wear’) of carbon commutator brushes at high altitude [7]. Roselman and Tabor thought the high friction behavior during dusting wear is the result of the interaction between the dangling covalent bonds of carbon atoms created by wear [8]. Savage showed that condensable water vapor in the atmosphere resulted in low wear rate and friction coefficient [9,10]. Rowe proposed that lubricating vapor (water or oxygen) intercalates the graphite lattice and reduces shear strength [11]. However, measurements of the lattice parameter of polycrystalline graphite established that interlayer spacing is not affected by water vapor [12]. Furthermore, non-graphitic carbon materials with ‘turbostratic structure’ [13] also exhibit similar low friction behavior but their lattice does not shear as readily as that of graphite [14]. Deacon and Goodman put forth an alternative explanation for the low friction behavior [15]. They suggested that low intercrystallite adhesion in polycrystalline graphite affected the lubricity of graphitic materials. In their view, the edge of a graphite crystal reacts with water and oxygen in the atmosphere and forms a surface containing various oxygenated groups. This reduces the force between graphite crystallites and induces a low friction behavior. Lepage and Zaida suggest that some water molecules dissociate and interact with carbon atoms to form hydrogen and hydroxyl complexes, to which other mobile physisorbed water molecules subsequently attach [16]. Marchon et al. considered that the lubricity imparted by water vapor depends on the availability of adsorption sites on a graphite surface [17]. When sites are available, oxygen feeds the formation of various complexes like semiquinone and lactone on the graphite surface, which can be polar sites for the adsorption of water molecules [18]. Molecules attached to polar surface sites can act as secondary sites for further adsorption via hydrogen bonds, forming two-dimensional clusters on the graphite surface [18]. Diatomic gases such as nitrogen and hydrogen do not impart a lubricative effect [7,19,20]. Although the lubricative effect of water vapor has been known for a long time, the exact friction and wear mechanism is still not entirely clear. Several investigators have considered that the low friction of graphite results from basal plane contact at the sliding surface [20–22]. Porgess and Wilman showed a slight tilt of basal planes to the sliding surface [21]. The crystallites were tilted so that the basal plane normals were slightly against the direction of motion at the sliding surface, and then parallel to the resultant of normal and friction force. Hill suggested that the low friction is due to a film of debris and, once disrupted, the coefficient increases [23]. Longley et al. found film formation to be affected by carbon structure and hardness of the counter-

face material [24]. Debris would more easily form between two soft graphite surfaces than two non-graphitic carbon surfaces. In disc-on-disc friction tests of C / C composites many factors have been shown to affect the results. The initial surface condition can affect braking behavior. Run-in (surface roughness, Ra 5 2.0 mm) specimens have exhibited much higher friction coefficient and wear rate than polished (Ra 5 1.2 mm) specimens under the same condition [25]. Yen and Ishihara mention two kinds of surface morphology formed after sliding [26]. One is dull-looking gray with a machine-finished appearance. The other is a lustrous black with mirror-like polished appearance under room light. The lustrous surface was covered with a layer of thin debris film of the order of 1 mm thick. During braking the friction surface temperature increases according to the energy absorption rate. In a study of C / C composites using a ring-on-ring drag test, Yen and Ishihara report the friction coefficient rises during two transitions as a result of the friction process [27]. The first, reported at 150–2008C, occurs because desorption of water vapor from the solid carbon surface decreases the lubrication effect. The other one at 650–7008C suggests that the increased friction coefficient involves an oxidation mechanism. Chen and Ju also discussed the first transition [28]. Using drag tests they compared several types of hybrid matrix C / C composites and noted transitions for some of them. Pre-transitional friction coefficients were |0.1–0.2. During transition, the initially formed thin, smooth lubricative film was suddenly disrupted and turned into a thick powdery debris layer which raised the friction coefficient to 0.5–0.9. The powdery debris on certain composites was easily ‘ironed’ into a smooth and tight lubricative film to cause both friction and wear to decline. Another factor influencing friction behavior is the atmosphere. Friction mechanisms of C / C composites in dry nitrogen and dry air have been studied [29]. The friction coefficient in air is much lower than that in dry nitrogen at 150–2008C. An explanation put forth by the investigators is that oxygen can have a limited lubricating effect on C / C composites but is less effective compared to water vapor. Chen et al. studied the effect of humidity on tribological behavior [30]. Relative humidity level had a strong effect on the water-desorption transition. Low humidity and high sliding speed generally gave higher friction coefficient and wear rate. High humidity and low sliding speed gave lower friction coefficient and wear rate. Clark, Tanvir and their co-workers [31,32] studied the contact load effect on friction. They found for bulk carbon / graphite sliding against itself that higher load resulted in a lower friction coefficient. Chen et al. reported the effect of load for C / C [33]. Their results indicated that the friction coefficient and wear rate variation with sliding distance depended on load and composite type. They categorized the worn surface morphology into three types. Type I featured a thin, smooth and bright debris film. Type II was

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3. Experimental

Fig. 1. Schematic diagram of sub-scale brake dynamometer.

defined as a thick, rough and dark powdery debris layer. Type III was characterized as a smooth, dense and bright debris film. A higher load accelerated the transition from type I to type II, but impeded the transition from type II to type III. Other factors that may influence friction behavior are speed, energy condition, matrix, and fiber type [34]. Stanek studied the effect of fiber orientation [35]. Discs with fibers oriented normal to the friction surface gave lower wear rates due to reduced fiber removal. Based on the concept that a smooth debris layer reduces the friction coefficient and wear rate, it was proposed that a large proportion of normal fibers in the wear surface may reduce the amount of debris required to reduce wear rates. His results suggested that fibers oriented normal to the friction surface (z-fibers) are important for high performance brake discs. Under landing braking conditions the friction surface temperature of C / C aircraft brakes can rapidly exceed 10008C and increase oxidation rates [3]. This is one of the main factors often reported to influence wear rates and result in decreased C / C brake life. If temperatures can be lowered by increasing through-thickness thermal diffusivity, it may be possible to extend brake life. Although there are some references regarding friction behavior of C / C composites, there is no open literature addressing friction behavior of high thermal diffusivity carbon / carbon.

Needled felt C / C discs 10.2 cm OD and 7.6 cm ID were tested on a sub-scale aircraft brake dynamometer manufactured by Link Engineering Inc. This consists of a motor driven shaft with removable inertia plates to which a rotor disc is attached as indicated in Fig. 1. The stator disc is attached to a pneumatically actuated section that applies the contact force and also has an arm for monitoring brake torque. Friction coefficient was calculated based on an effective swept radius of 4.4 cm. Stops were controlled to maintain constant brake torque. An inertia of 2.64 kg m 2 was used in all tests. For taxi stops an initial velocity of 546 rev. / min gave an energy of 3842 J which was dissipated in 5.5 s. Cold taxi conditions began with initial brake temperature of 308C. Hot taxi conditions were conducted with higher initial temperatures depending upon the heat remaining from previous stops. For normal landing stops an initial velocity of 2944 rev. / min (125.6 kJ) was brought to a stop in 28 s (normal landing conditions are lower energy than service landing where the aircraft is heavier). Thermocouples were placed at various depths from the friction surface at the outside diameter of the stator disc. Preforms of needled felts containing P25 fiber heat treated to 24008C were directly CVI densified. Details of the preforms are shown in Table 1 as provided by the manufacturer. All consisted of a needled chopped fiber but with z-fiber varying from |5% (A-type) to |25% (Dtype). The fiber volume of |25% is typical of a needled felt and results in a composite with properties significantly affected by the matrix. The CVI matrix consisted of laminar phases and isotropic phases, at times forming more than two concentric layers about the fibers. A pair of discs was made from pressed, chopped prepreg (phenolic) fiber tows that were carbonized and then CVI densified. P25 fiber without heat treatment was used and the resulting composites contained a fiber volume of |50%. This results in a material with nominally isotropic in-plane properties and through-thickness properties largely matrix influenced. This disc pair was used as a control

Table 1 Needled felt preform characteristics (as provided by Amoco Polymers) Type

Fiber

Structure

Density (g / cm 3 )

Approx. volume fiber (%)

Heat-treat temp. (8C)

Relative z-fiber

A

P25

Chopped fiber / needle

0.588

0.28

2400

Low: |5% of total

B

P25

Chopped fiber / needle

0.485

0.23

2400

Medium: |15% of total

D

P25

Chopped fiber / needle

0.451

0.22

2400

High: |25% of total

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against which the needled felts were compared and are designated as Y-type. Density and open porosity were measured from wafers of material at multiple depths from the disc surface. Archimedes principle was employed after weighing immersed samples in mineral spirits to give impervious density and open porosity. Room temperature thermal diffusivity was measured by a flash laser technique on small plugs of material from each composite. As an indicator of through-thickness mechanical properties, acoustic velocity was measured. Contact piezoelectric transducers (1 MHz) in a transmission configuration with spike excitation were used to measure time of flight. The values are then combined with specimen thickness to give acoustic velocity. A complete assessment of elastic moduli was not attempted. Comparing the velocities of each material gives a qualitative comparison of throughthickness modulus as influenced by fiber architecture.

4. Results and discussion Typical density profiles through the thickness of C / C will indicate a maximum near the surface and minimum at the center of the piece. This is a result of CVI densification that deposits carbon more rapidly at the skin, often choking off the still porous interior. The needled felt C / C used in this study generally showed this type of profile. However some discs contained high density interior regions, indicating substantial inhomogeneity at the millimetric scale beyond that which creates a density gradient [36]. Table 2 gives the average values obtained for the composites used. The bulk density of the needled felts ranged from 1.61 to 1.75 g / cc and the open porosity ranged from 14.3 to 22.6%. The material with highest z-fiber content (D-type) had the highest density and lowest open porosity. Sections of material polished for microscopic examination revealed characteristics of the carbon matrix and fiber architecture. Images shown in Fig. 2 reveal the presence of z-fibers created by needling. The structure of the low z-fiber C / C is characterized by small and widely spaced z-tows (Fig. 2a). In-plane arrangement of fibers, in this and

all C / Cs studied, is random. The high z-fiber material contains large and tightly spaced z-tows (Fig. 2b). These z-tows are not perfectly straight and uniform and when absent result in a composite with 2-D random structure (Fig. 2c). The low fiber volume of the needled felts can result in considerable porosity with large inter-tow voids as shown in Fig. 2c. A typical z-fiber structure is represented in Fig. 2d. These z-fibers serve as a conduit for throughthickness thermal transport that is particularly effective when combined with a directional CVI matrix. For the carbon fiber the conductivity is greatest in the axial direction owing to a preferred orientation of graphene layers. Moreover, the matrix deposited from CVI also has a preferred orientation of graphene layers parallel to the fiber axis. The end result is the formation of a relatively large diameter thermal pathway the size of a z-tow. This arrangement also influences the mechanical properties of the composite and indicates the size scale of inhomogeneities present. In addition to the solid-phase inhomogeneities, the examination reveals that much of the porosity results from large voids as seen in Fig. 2c. The thermal diffusivity of the low z-fiber composite (A) measured an order of magnitude higher for the in-plane vs. the through-thickness directions. Measured values shown in Fig. 3 demonstrate that needled felts with high z-fiber content (D) can have through-thickness diffusivity that is equal to in-plane values. The effect of z-fiber content on through-thickness diffusivity is clearly indicated by the data. In each case through-thickness diffusivity for A-type C / C is lower than that of B-type which is lower than that of D-type. For near surface samples the values ranged from 0.045 cm 2 / s to 0.180 cm 2 / s. In both A- and B-type needled felts the radial diffusivity is near 0.45 cm 2 / s. The high z-fiber C / C has a lower radial diffusivity due to the lower proportion of in-plane fibers. There is a distinct difference in through-thickness diffusivity between near surface and center portions of the needled felts. Microscopy observations indicated a higher z-fiber content in the center of the discs. This results from a needling operation that forms z-fibers with each added felt layer. When needling is performed through the entire stack a high proportion of z-fiber in the original, mid-plane

Table 2 Average bulk density, impervious density and open porosity of C / C discs Sample

Average bulk density (g / cm 3 )

Average impervious density (g / cm 3 )

Average open porosity (%)

A1 A2 B1 B2 D1 D2 Y1 Y2

1.65 1.61 1.66 1.65 1.70 1.75 1.62 1.62

2.07 2.08 2.10 2.07 2.04 2.05 1.74 1.74

20.2 22.6 20.9 20.2 16.8 14.3 7.3 7.0

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Fig. 2. Micrographs of needled felt C / C fiber architectures.

Fig. 3. Thermal diffusivity of C / C composites in through thickness (surface and center) and radial directions.

felt results. The higher z-fiber content then imparts higher through-thickness diffusivity. The control material (Y-type) was found to have the lowest diffusivity of all the materials in this study. The low through-thickness diffusivity is partly due to a lack of z-fiber and partly from the phenolic char matrix component. In addition, the P25 fibers used were not graphitized. That reduces radial and axial fiber diffusivity. This is the main factor in the low radial diffusivity of the control C / C. Acoustic velocity of the needled felt composites through the thickness exhibits a trend similar to that of thermal diffusivity. Shown in Fig. 4 are the average measured values for the C / C composites used in this study. The low z-fiber material has the lowest velocity while the high z-fiber material has the highest velocity. This is a direct consequence of the z-fibers contributing to the stiffness of the composite normal to the friction surface. Fiber and matrix graphene structure in z-tows results in increased through-thickness elastic modulus. Another consequence

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Fig. 4. Through thickness acoustic velocity of C / C discs.

of this structure will be higher shear strength in the friction plane owing to the orientation of graphene layers. While no measurement of shear strength was conducted, the data clearly show how increasing the quantity of z-fibers gives measurably higher through-thickness composite diffusivity and acoustic velocity. Dynamometer testing of discs was performed while measuring temperatures at various depths from the friction surface. In taxi and normal landing stops temperature profiles were substantially affected by through-thickness thermal diffusivity. The dynamic near surface (1 mm) temperature of each material, tested under the same conditions, is plotted in Fig. 5. The initial time lag is mostly a result of the brake apply time which results in an |1 s delay before full torque is achieved. The temperature–time profile of each needled felt C / C is very similar with the clear exception that higher temperatures are attained as thermal diffusivity decreases. Another difference is the time at which maximum temperature is attained. The high diffusivity C / C reaches maximum temperature before the other materials due to the combination of rapid transport from the interface and a diminishing energy dissipation rate. The energy dissipation rate for the

Fig. 5. Near surface temperature during cold taxi stops for C / C composite types.

other materials is the same but the transport of thermal energy is slower so that higher temperatures are reached and the maximum temperature occurs closer to the end of a stop. The control material (Y-type) demonstrated a distinctly different temperature–time profile than the needled felt composites. It reaches a maximum temperature at the end of the stop owing to very low diffusivity. The data demonstrate qualitative trends regarding the effect of thermal diffusivity on brake temperatures. Making quantitative comparisons between materials was not feasible in this study due to several factors. Material inhomogeneity was found to significantly affect the uniformity of temperature profiles. In addition, thermocouple measurements are only representative of a trend since size, location and thermal coupling of the bead to the C / C can influence measurements. With this understanding care is taken when using measured or calculated temperatures. Yet to be perfected is the ability to measure surface temperatures of friction couples. Despite this, analytical and experimental approaches can be used to give an estimate of average surface temperature. A one dimensional model [37] that incorporates temperature-dependent material properties was used to calculate temperatures in the C / C during a braking event. A comparison to measured values from seven depths below the friction interface is shown in Fig. 6. Several features are worth noting. At all times during a stop a negative temperature gradient occurs. The temperature rises first near the friction surface with a measurable time lag deeper in the material. The experimental data qualitatively match the calculated temperatures for the locations closest to the surface. In all cases the measured values are lower. The shapes of the calculated curves have a poor match to the experimental values far from the surface. Several explanations can be offered for the differences seen. These include (a) poorly defined thermal properties, (b) rate of energy input not well defined, (c) non-uniform energy input, and (d) thermal losses (none taken into account) are significant. Despite the

Fig. 6. Experimental (thin) and calculated (bold) temperatures at depths from friction surface ranging from 1.5 to 10.5 mm. Cold taxi condition, D-type C / C.

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differences the results reasonably represent the nature of thermal profiles in C / C brake applications. The end result is that substantial thermal stresses can be generated and surface temperatures may be quite high as a result of thermal transport characteristics of the friction material. The dynamometer results from simulated taxi stops also give an indication of the influence of material variables. Friction coefficient can vary from one stop to the next or within a single stop. The friction can also vary with test material. In this study dynamometer stops were performed repeatedly with time in-between to achieve repeatable recycle temperatures. When discs are freshly machined several stops are generally needed to break-in the friction surfaces. Then friction coefficient, as averaged over the duration of a stop, may (1) remain stable, (2) gradually rise with repeated stops as surfaces evolve, or (3) change significantly from one stop to the next. In the cold taxi stops all materials followed scenario 2 with a very gradual increase in friction over 100 stops. Then for hot taxi condition friction jumped by a factor of 2 or more after several stops warmed the discs. A summary of the results is given in Table 3. When operating under cold taxi conditions the C / C discs demonstrated a low friction coefficient ( m ) which remained stable after break-in. An exception occurred in one series of stops for D-type material when relative humidity dropped from 64 to 52%. During that time friction increased by 30–100% [36]. This is a result of the influence of humidity level on the occurrence of moisture desorption friction transitions. There is a correlation between z-fiber content and m. Greater amounts gave higher friction. In addition, the higher z-fiber content materials had the shortest break-in period. When operating in hot taxi conditions a friction transition was observed to occur which resulted in a sudden increase in friction after a number of warm-up stops. The materials with lower through-thickness thermal diffusivity went through the transition in the fewest stops owing to higher temperature gradients and higher maximum surface

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temperatures. The transitions observed often occurred during braking such that a rapid increase in m resulted in contact force reduction in an attempt to maintain constant brake torque. When a sudden and substantial increase occurs, system response is not fast enough and momentary increase in energy dissipation rate results. In addition, a sudden increase in m brings to question the appropriateness of a stop average value since important information may be overlooked. Reporting of two values for m along with a metric of change would give a more thorough description. In the results given here, however, brevity is preferred. Results from simulated normal landing stops demonstrate a different trend in friction versus material type. In this case all stops had temperatures high enough such that braking was entirely above the moisture desorption friction transition. On average the lower z-fiber corresponded to higher stop average m for the 100 consecutive stops performed. Type A C / C had m ranging from 0.4 to 0.5, with significant variation from stop to stop. Type B ranged from 0.3 to 0.5 and type D from 0.2 to 0.4. No cause for the significant variations could be established. For all tests some stops displayed significant vibrations but no cause was determined. For types A and B C / C thermally induced red emission was produced at the friction interface. Type D C / C did not give noticeable emissions which is consistent with the lower temperatures measured. As with taxi conditions, type A gave the highest near surface temperatures and type D gave the lowest. Temperatures measured near (|1 mm) the friction surface were well above 5008C for all but D-type C / C which typically had a high of |3008C. Average surface temperatures were higher in all cases by 20%, or more, compared to the temperatures measured. The control material (Y-type) gave m results that were similar to that of the needled felts under taxi conditions. Specifically, the cold taxi stops exhibited only a very small increase in m after a substantial run-in period as did the low z-fiber (A) material. For hot taxi conditions the

Table 3 Summary of stop average friction coefficient results from taxi and normal landing tests Disc

Test

m low

m high

Comments

A-type B-type D-type Y-type A-type B-type D-type Y-type A-type B-type D-type Y-type

Cold taxi Cold taxi Cold taxi Cold taxi Hot taxi Hot taxi Hot taxi Hot taxi Landing Landing Landing Landing

0.09 0.095 0.10 0.08 0.10 0.10 0.15 0.10 0.52 0.51 0.40 0.27

0.10 0.12 0.16 0.10 0.30 0.30 0.32 0.25 0.36 0.28 0.18 0.13

Gradual rise after 20 stop run-in period Gradual rise after 6 stop run-in period Gradual rise after 2 stop run-in period Gradual rise after 25 stop run-in period Friction jumps 33 after 4 stops, becomes variable Friction jumps 33 after 16 stops, becomes variable, drops Friction jumps 2.13 after 40 stops, drops after 65 stops Friction jumps 2.53 after 12 stops, keeps rising Friction typically varies by 20% over 100 stops Friction varies by 40% over 100 stops Friction varies, but overall declines over 100 stops Friction varies by 50% over 50 stops

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transition occurred early as with A-type material. Under landing conditions the friction coefficient of Y-type was lower than that of the needled felts, ranging from 0.17 to 0.27 over consecutive stops. This was slightly lower than that for hot taxi conditions. It is important to note that some taxi stops demonstrated a rapid increase in m when passing a moisture desorption transition. For the landing stops no such jump was discernible but this does not mean desorption did not occur. As will be shown later, the sudden increase in m may occur during the early part of the stop depending upon the braking power, humidity level, or thermal transport characteristics. Under landing conditions desorption would occur very quickly during the initial load application. Examination of worn surfaces by optical and electron microscopy was performed. Whole discs were mounted for observation thus avoiding damage incurred by sample preparation. Examination by polarized light microscopy revealed that the needled felts developed very little friction film. The film is comprised of finely divided carbon packed to substantial thickness, of the order of 1 mm, that prevents the examination of the bulk material. Fiber and matrix structures were clearly observable as shown in Fig. 7a. Film and debris is present filling large and small voids. However, in most regions the microstructures of individual carbon components (fibers and matrix) are directly exposed to the friction surface. The layered CVI matrix structure around a z-fiber tow is clearly revealed. Regions where

in-plane fibers were exposed also had a polished appearance such that carbon structure was discernible. The high polish of the composite surface suggests that friction processes remove very small particles (.1 mm) from the surface. If only large particles were removed, surfaces would not be as highly polished. Microscopy indicated mostly particles of 100 nm, or less, combined with some larger debris. X-ray diffraction of wear debris showed the original crystal structure was mostly destroyed by friction leaving an amorphous powder. This occurred for both taxi and landing stops [36]. The surfaces of the control material (Y-type) did develop a significant friction film under taxi conditions as seen in Fig. 7b. The films are comprised of very small particles compacted together with some larger pieces of debris. This is typical of friction films on C / C composites as found on other materials tested in our facility. The film was absent from Y-type discs after landing stops. The cause of film formation and removal, and the role in the friction process is still poorly understood. Different materials may generate a film as shown here but how they influence performance is still vague. Taxi conditions did not produce evidence of oxidation loss. Landing stops did produce oxidation at the inner and outer non-friction surfaces. Micrographs reveal that severe pitting of the CVI matrix occurred in A-type C / C as shown in Fig. 8. Oxidation was most pronounced at the outside radii relative to inside radii, presumably due to the

Fig. 7. Polarized light micrograph of D-type needled felt C / C after cold taxi stops, (a) showing polished surface. Electron micrograph of Y-type C / C after hot taxi stops, (b) showing friction film and underlying composite.

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Fig. 8. Electron micrographs of oxidation loss near outside radius (a), and inside radius (b).

ready supply of air. Under the same conditions the D-type C / C showed very little sign of oxidation loss due to lower temperatures resulting from the increased thermal transport from the friction surfaces. The Y-type C / C also demonstrated severe oxidation on exposed surfaces. No evidence of oxidation loss from the friction surface was detected by microscope investigations. In these oxygen-starved regions very fine, loose powder was found in unfilled voids. No oxidation pitting of the CVI carbon or phenolic char was found. Sharp, reactive edges did not exhibit rounding over. Thus, when oxidation was severe at air-exposed areas no such damage was seen at the friction surface. This does not rule out the possibility of chemisorption of oxygenated species after the stop finished and discs slightly separated. Such adsorption and minor oxidation loss, including surface activation, may have significant influence on friction behavior. Friction surfaces of needled felt C / C after landing stops exhibited a failure mode unique to a material with substantial lamellar structure. The micrographs in Fig. 9 show a region where fibers are oriented perpendicular to the sliding surface. Some very thin film is revealed in Fig. 9a and appears as narrow bands oriented perpendicular to the sliding direction. This was only discernible in the regions surrounding the z-fiber. Fibers were often debonded from the matrix as in Fig. 9b. Also revealed is matrix cracking from near surface stresses induced by sliding. This failure follows the weakly bonded regions between graphene layers in the CVI carbon. CVI matrix growth occurs radially outward from the fibers causing the pattern seen. Different spacings of these cracks are indicated by arrows

1, 2 and 3. These cracks are believed to result from shear stresses developed during sliding that cause a rotation of the surface region to the level that mode II shear failures between graphene layers occur. This is analogous to the bending of a stack of cards forming a stair-step morphology at the edge. In the friction scenario this stair-step surface is worn flat during sliding. When stresses are removed at the end of braking, the material should rebound to give a saw-tooth surface morphology. An additional effect, shown in the upper left section of Fig. 9b (arrow 3), is the removal of a section of matrix as though several graphite domains have been removed. This was viewed in a few regions near to, but not in contact with, a z-fiber. The matrix cracking reported here was found at the leading and trailing quadrants (relative to the sliding path) of the fibers. Apart from the cracking described above, the surface morphology of the matrix and fiber was found to be different. At the micrometre level the CVI carbon has a smooth surface suggesting that either sub-micron particles are worn from the surface or debris has filled in surface roughness. The fiber ends show considerable roughness. This indicates that the surfaces are not being filled in by debris significantly. It also demonstrates that there are different wear mechanisms involved in each carbon phase of a C / C. The fibers have small crystalline domains less than 1 mm in width and several mm in length. Intergranular fractures may then produce submicron debris. The matrix has a crystal structure of much greater size and intergranular weakness does not contribute to the production of the fine debris. Wear of that phase occurs as discussed above

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Fig. 9. Surface condition near z-fibers after normal landing, B-specimen. Sliding direction from upper right to lower left.

through the removal of submicron particles by intragranular bond breakage leaving a surface of very low roughness. No difference in film morphology was found in z-fiber regions compared to in-plane fiber regions. A series of experiments was conducted to detail variables determining the onset of the friction transition historically associated with moisture desorption. Stops were performed with different initial velocities and the same brake torque to give different energy dissipation rates and surface temperatures. Three stops of each condition were performed to assure reproducibility. A sufficient time delay between stops assured similar initial temperature and time for adsorption. Relative humidity (RH) of the chamber was controlled. Low z-fiber material was found to go through a transition at 50% RH and initial speed of 546 rev. / min as shown in Fig. 10. Friction is initially low and then rapidly increases from |0.1 to |0.2 at 2 s into the stop. Also shown is the calculated surface temperature for the material. When the transition occurs surface temperatures of 1078C are indicated which is significantly less than the 1508C reported in the literature [29,30]. In addition, m begins to drop near the end of the stop when surface temperatures fall below 1078C suggesting adsorption of moisture occurred. Oscillations in the m values result from torque variations which are most pronounced when contact loads are reduced. In order to maintain constant brake torque, loads of |1000 lb are used when m is 0.1. The dynamometer controller reduces this to |500 lb when m rises to 0.2. Oscillations are friction system-dependent and

are likely the result of slight disc misalignment, an effect lessened at the higher loads. Temperature measurements at various depths during braking indicated temperatures were lower than that calculated. The experimental data in Fig. 11 shows temperature and depth for the time of the transition in the stop of Fig. 10. A polynomial curve fit to the data is used to estimate surface temperature. This method indicates surface temperatures of |758C at the onset of the friction transition. At this temperature moisture desorption will not be complete, and transition should not occur according to literature reports. These data suggest a more complicated mechanism than simple desorption is responsible for the change in friction of carbon. Of course, this temperature

Fig. 10. Calculated surface temperature and friction coefficient during cold taxi stop of A-type material, 50% RH. Temperature at transition is |1078C.

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Fig. 11. Surface temperature extrapolated from experimental data at time of transition for stop shown in Fig. 10. Surface temperature of |758C is suggested. Fig. 13. Influence of composite type (different thermal diffusivity) on occurrence of transition under same stop conditions.

measurement approach contains uncertainties brought about by material inhomogeneity and spatially non-uniform heat input. The calculated estimate also is limited by these factors. To demonstrate energy dissipation effects on transition onset, a typical result for two stops is shown in Fig. 12. For the D-type material initial velocity of 1092 rev. / min resulted in a transition during the stop. This is compared to a stop with lower initial velocity and same RH (50%). Differences in estimated maximum surface temperatures for each stop were clearly demonstrated (not shown). The low velocity disc did not pass through transition. In addition, it was possible to shift the time at which the transition occurred by slightly increasing or decreasing the initial velocity in the range of 900–1200 rev. / min. The transitions were repeatable over three consecutive stops or when returning to that speed after stops at higher initial velocities. When significantly higher velocity was used (.1200 rev. / min), the transition was masked by the initial brake apply giving a high value of m over the entire stop. Friction is again observed to decrease near the end of the

higher energy stop of Fig. 12, presumably due to decreasing surface temperatures and consequent adsorption. Since the transition is thermally driven, materials of different thermal diffusivity should experience the transition under different stop conditions. Fig. 13 shows the results from A-type and D-type C / C discs operating under the same condition (546 rev. / min, 39% RH). The lower z-fiber material showed a transition when the higher z-fiber material did not. This is a direct result of the higher through-thickness thermal diffusivity of D-type material and the demonstrated influence over friction surface temperatures. Humidity level was also found to change the onset of transition for all materials tested. This is shown in Fig. 14 where a stop performed at 95% RH is compared to one at 40% RH, each with the same initial velocity of 360 rev. / min. Transition occurred at the lower humidity and

Fig. 12. Influence of initial stop velocity (energy dissipation rate) on occurrence of transition for D-type C / C, 50% RH.

Fig. 14. Influence of relative humidity on occurrence of transition for Y-type C / C.

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was ‘eliminated’ by raising the humidity. Increasing the initial velocity at high humidity condition forced the transition to occur. The effect was reversible. The data suggests that more thermal energy (higher temperature) is required to drive off adsorbed moisture under high humidity conditions. The occurrence of a desorption transition has been shown to depend on material property, energy dissipation and humidity level. Thus a critical condition of humidity and brake power must be met for a given C / C to experience transition. The temperatures at which this occurs is still speculative, as we show values may be considerably lower than that reported. In fact, the exact mechanism responsible for the lubricating property of solid carbons is still in question. Clearly, the presence of adsorbed species is required for low friction. Exactly how these adsorbates function in friction phenomena is uncertain and is presently being investigated. The studies using stops of increasing initial velocity demonstrated that stop average friction would rise to a maximum value when transition occurred, or at velocities slightly higher. For Y-type material friction decreased at higher energies. It is presumed that the significant friction film developed on the surface was disrupted by the desorption transitions causing an extreme increase in m. At higher energies m decreased as a more stable surface condition developed. Visual examination confirmed the changes in surface conditions. The needled felts did not demonstrate a maximum near the transition and did not develop a significant friction film. Stops at very high velocities (.3000 rev. / min) caused near surface measured temperatures to exceed 7008C but no other transition was observed. This is in contrast to literature reports of oxygen desorption transitions giving friction increases [26,27]. As

previously mentioned, oxidation was observed under these conditions but only on surfaces directly exposed to air. Wear of the materials tested was evaluated by weighing discs before and after friction testing. The measured wear was not corrected for oxidation loss on exposed surfaces or the containment of wear debris in voids open to the friction surface. For the multiple simulated stops the results are given in Fig. 15 in terms of weight loss per revolution. Clear differences in wear rates are found for the different materials. Higher z-fiber content materials exhibited greater wear resistance. Hot taxi stops resulted in greater wear than normal landing stops despite the greater speeds and temperatures of the latter. The control material, which developed a significant friction film, had the greatest wear rate of all materials in the hot taxi condition. These results suggest that the presence of z-fibers will reduce wear. This is consistent with the concept of increased interface shear strength from a high proportion of graphene layers oriented normal to the sliding direction. The effect of structure orientation on through-thickness mechanical properties was demonstrated by increased acoustic velocity. The effect of reduced surface temperatures on wear rates is not clear but is thought to be a minor contributor to the wear process. This conclusion is based on the fact that taxi conditions often resulted in greater wear than normal landing conditions. This suggests that oxidation loss from high energy stops plays little role in the wear of C / C materials in a disc-on-disc configuration. If oxidation were to weaken the material at or near the friction interface, greater loss should have been detected. In addition, a rougher surface would also be anticipated. Neither occurred, so we conclude that oxygen diffusion into the friction interface during braking is very limited. After braking is finished, oxygen will be able to reach the

Fig. 15. Wear rates for C / C composites under all stop conditions. Needled felt wear resistance increases with z-fiber content.

C. Byrne, Z. Wang / Carbon 39 (2001) 1789 – 1801

interface but temperatures are reduced such that oxidation loss is minimal and a chemisorbed layer is the most significant result.

5. Summary and conclusions Three different needled felt C / C composites containing from 5 to 25% z-fiber and a two-dimensional C / C composite were evaluated and tested for friction performance. It has been found that adding z-fibers increases through-thickness thermal diffusivity and acoustic velocity of discs, and has a distinct effect on friction performance. A laboratory sub-scale dynamometer was used to simulate cold taxi, hot taxi and normal landing braking events utilizing a single stator and rotor pair. Temperatures measured near the friction surface were lowest for highest thermal diffusivity material demonstrating the effectiveness of z-fiber in reducing friction surface temperature. All materials had a low friction coefficient (50.1–0.15) under cold taxi conditions. Hot taxi conditions caused a twofold, or greater, increase which was attributed to a waterdesorption transition. Composites with high diffusivity needed greater braking power to experience transition relative to lower diffusivity materials. It was found that the braking power must be increased to get this transition under higher humidity condition. Highest z-fiber content discs showed the lowest wear rate. This was attributed to higher in-plane shear strength. The wear rate was highest under hot taxi conditions. It was concluded that mechanisms other than oxidation loss were primarily responsible for the wear rate of the C / C composites tested in this study. In addition, no high-temperature, oxygen desorption transition was detected as reported in the literature. Oxidation at the disc outer edge was most severe for the lower diffusivity materials. The C / C composite surfaces had a polished appearance, suggesting that the removal of nanoscale particles occurs from the friction process.

Acknowledgements The authors wish to acknowledge the financial support of the Center for Advanced Friction Studies, an NSF SIUCRC, directed by Dr. Maurice Wright. We are most grateful to Dr. Dan Hecht of Amoco Polymers for supplying needled felt preforms, Aircraft Braking Systems Corporation for CVI densification, Dr. David Marx for use of the 1-D thermal profile software, and laboratory assistance by students Marie-Christine Caron and Jeremy Scott.

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