- Email: [email protected]
A novel pin-on-apparatus
Wear 254 (2003) 111–119
A novel pin-on-apparatus Naoki Fujisawa a,∗ , Natalie L. James b , Richard N. Tarrant c , David R. McKenzie c , John C. Woodard b , Michael V. Swain a,1 a
Biomaterials Science Research Unit, Faculties of Dentistry and Engineering, University of Sydney, Suite G11, National Innovation Centre, Australian Technology Park, Eveleigh, NSW 1430, Australia b VentrAssist Division, Ventracor Limited, 126 Greville Street, Chatswood, NSW 2067, Australia c Department of Applied and Plasma Physics, School of Physics, University of Sydney, Sydney, NSW 2006, Australia Received 26 March 2002; received in revised form 11 September 2002; accepted 29 October 2002
Abstract A novel pin-on-disk apparatus was developed that provides a repetitive impact loading between periods of sliding through alternate lifting and dropping of a spring-suspended spinning disk, away from, and onto, a spring-supported pin, respectively. The combination of the repetitive impact loading and sliding achieved in the apparatus was able to induce film adhesion failure of a thin film coated disk within 20 min, which, in the absence of the impact loading, would have survived the test due to the adequate sliding wear resistance. The impact/sliding pin-on-disk apparatus developed is therefore a useful means of predicting the sliding wear resistance and film adhesion of a coated system simultaneously. © 2002 Elsevier Science B.V. All rights reserved. Keywords: Pin-on-disk apparatus; Impact/sliding; Film adhesion failure; Thin film
1. Introduction A conventional pin-on-disk apparatus is commonly employed as a means of predicting the tribological performance of a thin film coating in continuous sliding contact. A typical pin-on-disk apparatus provides a normal contact load between a stationary pin and a revolving disk, and measures the resulting frictional force to evaluate the coefficient of friction [1]. However, although a specific time-mean contact load may be set for a given pin/disk pair at a given sliding velocity, the dynamic fluctuation of the contact load can vary depending upon the loading mechanisms used in the test apparatus (e.g. spring, dead-weight or pneumatic loading) [2]. The wear mechanism of a coated system involves a progressive volume loss of the film material due to sliding friction, or abrupt film adhesion failure when the localised strain energy exceeds the adhesion energy at the film/substrate interface (Griffith’s criterion). Considering the energy threshold-sensitive mechanism of film adhesion failure, a slight difference in dynamic loading condition alone could alter the failure mode of the coated system. A repeated contact event under impact loading conditions in addition to ∗ Corresponding author. Tel.: +61-2-9351-1814; fax: +61-2-9351-1815. E-mail address: [email protected] (N. Fujisawa). 1 Co-corresponding author.
the conventional continuous sliding often occurs in a practical wear situation, thus increasing the likelihood of the coated system to meet Griffith’s criteron. To date at least several impact/sliding wear test apparatuses have been developed to study the wear resistance mainly of bulk metals and polymers. These include a wear test rig that directs bullet-shaped projectiles against a vertical metal disk surface which is either stationary or in motion [3], a pin-on-disk-type apparatus with a spatial crank oscillator mechanism to provide an adjustable impact load impulse in synchrony with the rotational disk motion [4], a cam-based device that utilises two independent cam/cam follower/lever systems, with one system generating a horizontal reciprocated motion of an upper specimen by an eccentrically mounted cam, and the other system allowing the upper specimen assembly to drop repetitively onto the lower specimen using a cut-out cam and gravity [5], and so forth. Most recently, a micro-impact testing method was developed [6] with the particular aim of testing thin films for film adhesion failure, whereby an actively oscillating (diamond) test probe contacts the vertical specimen surface repetitively at a much greater frequency (up to 1000 Hz) than occurs in other impact/sliding devices. The specimen position can be moved in the transverse direction during testing but at a much slower sliding speed (∼1 m/s). Film adhesion failure under this micro-impact loading was evaluated as the time
0043-1648/02/$ – see front matter © 2002 Elsevier Science B.V. All rights reserved. PII: S 0 0 4 3 - 1 6 4 8 ( 0 2 ) 0 0 3 0 7 - 1
112
N. Fujisawa et al. / Wear 254 (2003) 111–119
when the probe position (depth), which was monitored continuously during the test, showed a step increase resulting from the localised film loss. The results reported [6] demonstrated clearly that a repetitive high frequency micro-impact could promote film adhesion failure. In this study, a pin-on-disk apparatus was developed that provides a repetitive ‘macro’-impact loading of a much lower frequency (∼0.1 Hz) between periods of continuous sliding, by alternately lifting a spinning disk away from a pin using a solenoid, and dropping it onto the pin under gravity. The pin-on-disk apparatus was designed primarily to simulate a potential transient sliding contact event of mating titanium alloy surfaces that are, in normal operating conditions of the device, fully separated from each other via a highly stable fluid film bearing. A contact between the two surfaces may occur only when the device is exposed to a sudden acceleration or deceleration of an excessive magnitude, although that is a condition highly unlikely for the device to experience.
2. Pin-on-disk apparatus 2.1. Apparatus design Cylindrical pins (6 mm diameter by 20 mm length) and flat circular disks (19 mm diameter by 8 mm thickness) were cut from Ti-6Al-4V ELI (Extra Low Interstitial grade) rods. One end of the pin was machine-finished to a 24.8 mm-radiused tip with a resulting surface roughness average in the range of 0.2 m. The working surface of the disk was polished to a roughness average smaller by an order of magnitude. The indentation hardness and biaxial modulus of the titanium alloy substrate were in the range of 4.5 and 114 GPa, respectively [7]. The schematic drawing of the pin-on-disk wear tester is shown in Fig. 1. A brushless DC-servomotor (3556K024B, Minimotor SA, Croglio, Switzerland) was used to revolve the disk about the disk centre. The rotational speed of the disk was speed-controlled to approximately 2000 rpm by a four-quadrant pulse-width modulation servo amplifier (BLD 5606-SH4P, Minimotor SA, Croglio, Switzerland). A 50 mm-long titanium extension shaft transmitted the motor shaft revolution to the disk via a cylindrical disk holder (19 mm × 8 mm diameter), which was screw-attached to the end of the extension shaft in axial alignment. The disk was attached onto the flat disk-holder face using an adhesive double-sided tape (ScotchTM Brand #4616AFT, 3M Australia Pty. Ltd., Sydney, Australia). The component consisting of the disk and the motor with the disk surface facing down was suspended vertically via an extension spring from a precision single-axis translation stage whose vertical level was adjustable by a built-in micrometer (Melles Griot Inc., Irvine, CA, USA). The spring-suspended disk component was allowed to move only in the vertical direction with the use of a linear motion guide-rail type bearing (9 mm rail
width with 31 mm bearing length, RSR9ZMUU + 167 l, THK Co. Ltd., Tokyo, Japan). The pin, with the radiused pin tip pointing upwards, was held securely in the acrylic pin carriage component. The pin carriage component was free to slide in the vertical direction alone by incorporating two ball-bushing bearings (6 mm i.d. × 19 mm length, LM6-AJ, THK Co. Ltd., Tokyo, Japan) parallel to each other and to the pin axis. The pin carriage component was supported by a compression spring from the component bottom at 65 mm front of the pin axis. The pin carriage component was made of acrylic mainly for the lightness of the moving component. The disk/pin couple was immersed in a blood volume expansion medium (Haemaccel® , Hoechst Marion Roussel Australia Pty. Ltd., Lane Cove, NSW, Australia) in a plastic bath at room temperature to simulate potential applications in a biological environment. The horizontal pin position was located so that the tip of the pin slid on the disk along a circle of approximately 5 mm radius. The vertical position of the single-axis translation stage, from which the disk was spring-suspended, was adjusted using the micrometer such that the time mean of the contact force between the disk and the pin would be approximately 1 N during steady sliding contact. To induce an impact loading, the spring-suspended disk was lifted upwards to a defined height using a linear pull-action solenoid (43 series, 12 V dc, BLP, Suffolk, UK), held elevated for 1 s, and subsequently dropped onto the pin component under gravity. The lifting and dropping of the disk component were repeated at every 10 s by alternately energising and disconnecting the solenoid using a solid-state relay with a rectangular pulse generated by a function generator as the on-off switching signal. 2.2. Friction coefficient measurement The coefficient of friction in the steady sliding condition was obtained as the ratio of the frictional force to the contact force during the three quarters of the 10 s duty cycle period (∼7.5 s) starting at 1.5 s after every impact loading. The friction coefficient during the initial 1.5 s of the duty cycle, in which transient fluctuations of the contact force, frictional force and the motor speed took place, was excluded from consideration. A strain gauge type load cell (1 kg capacity, A.L. Design Inc., Buffalo, NY, USA) was located beneath the compression spring to measure the normal force in the compression spring. The load cell output was amplified using a strain gauge signal-conditioning module (AMA-RM-044, 2.5 kHz frequency response, Applied Measurement Australia Pty. Ltd., Oakleigh, Vic., Australia). The contact force between the pin and disk equates to the vector summation of the upward force in the compression spring, a downward force due to the total weight of the pin component, and an upward buoyancy force associated with the partial submersion of the pin component in the lubricating fluid. Although the buoyancy force is dependent on the submerged volume or
N. Fujisawa et al. / Wear 254 (2003) 111–119
113
Fig. 1. Schematic drawing of the pin-on-disk apparatus (without lubricant bath), where the moving disk and pin components are shaded in different intensities.
the vertical position of the moving pin component, the amplitude of buoyancy fluctuation calculated (<0.01 N) was much smaller than the maximum amplitude of force fluctuation in the compression spring (∼1.6 N). Therefore, the contact force was estimated on the assumption of a constant buoyancy force (∼0.217 N). The frictional force was calculated as the motor torque due to friction alone (obtained as the total motor torque minus the motor torque whilst no contact between the disk and the pin) divided by the radius of the wear track. The measured back-emf voltage waveform of the motor was sinusoidal with negligibly small amplitudes of its harmonics, and the inductance of the motor was negligibly low, and therefore the motor torque may be determined as the back-emf
constant multiplied by the RMS (root mean square) of the fundamental of the line-to-line phase current multiplied by 3 (the number of motor phases). In fact, the actual phase current that was measured by a current transducer (LA25-NP, DC to 150 kHz (−1 dB) bandwidth LEM, Geneva, Switzerland) was a square waveform in each phase of width 120◦ on and 60◦ off with a distinctive ‘W’-shaped waveform on top of each 120◦ on sub-phase. However, the RMS of the fundamental phase current was found to be proportional to that of the quasi-square waveform by a ratio factor (∼0.95) regardless of the amplitude or frequency, as was confirmed by fast Fourier transformation analysis. Therefore, the fundamental RMS could be evaluated simply as the RMS of the filtered current multiplied by the ratio factor. The RMS of the filtered
114
N. Fujisawa et al. / Wear 254 (2003) 111–119
current was obtained through a wide-band RMS-to-DC converter (AD637, Analog Devices Inc., Norwood, MA, USA), which was configured for two pole Sallen-Key filter option with an averaging time constant of 120 ms. The motor shaft speed was also recorded using a frequency-to-voltage converter (Microgenic Systems Version 2/1 TACHO, 100 ms response time, Farnell Components, Chester Hill, Australia). The three signals (contact force, motor torque and motor speed) were recorded at a sampling rate of 100 Hz. 3. Mechanical characterisation of the apparatus 3.1. Impact loading The spinning disk struck the pin at the start of the test. The magnitude of impact was adjusted experimentally by varying the disk height for dropping to a level that was sufficient to cause gross film adhesion failure of a coated disk against an uncoated pin within a 20 min test period. The instantaneous acceleration of the moving pin component at the time of the initial impact loading was estimated using two piezoelectric accelerometers (Type 4375, ≤2% error below 10 kHz to 10% error at 16.5 kHz, Brüel and Kjær, Nærum, Denmark), coupled with a signal-conditioning amplifier (NEXUS(tm) Type 2692, Brüel and Kjær, Nærum, Denmark). One accelerometer was attached onto the back of the acrylic part behind the blind hole where the pin was inserted (defined in Fig. 1 as position 1, and the other on top of the acrylic part beside one of the ball-bushing bearings of the pin component (defined in Fig. 1 as position 2, both using beeswax. Measurements were conducted in the absence of lubricating fluid due to one of the non-water-proof accelerometers placed in position 1. An untreated titanium pin and a non-spinning untreated titanium disk were used to estimate the maximum instantaneous impact force possible for an unworn pin/disk couple. Fig. 2 shows accelerometer signals recorded at the two accelerometer positions using a band-frequency range of 1 Hz to 22.4 kHz of the signal-conditioning amplifier and a sampling rate of 5 MHz. The upper band frequency limit of 22.4 kHz was used because the use of higher upper frequency limits (30 and 100 kHz) did not significantly alter the signal waveform profile. These waveforms appear to have reflected the actual acceleration signals reasonably well since the frequency of the initial peak is 6–7 kHz, which is well below the resonance frequency of the accelerometer (∼55 kHz). As can be seen in the figure, not only the magnitude but also the sign of the initial peak measured away from the pin were different from those beneath the pin. In addition, the initial acceleration peak recorded away from the pin lagged behind that recorded beneath the pin. These differences and the phase lag suggest an instantaneous deflection/deformation of the acrylic part to a considerable extent. Further, the recorded peak value of 5680 m/s2 beneath the
Fig. 2. Accelerometer signals on the acrylic part of the pin carriage component beneath the pin (solid line) and away from the pin beside the ball-bushing bearings (dashed line) as a function of time, t, using a 5 MHz sampling rate. Acceleration was referenced to increase in the direction of gravity.
pin may still underestimate the true instantaneous acceleration experienced by the pin/disk couple because the acrylic section between the pin bottom and the accelerometer could have absorbed part of the instantaneous impact energy due to the low stiffness of the material. The presence of the lubricating fluid would also affect the peak acceleration. Due to the difficulty in obtaining a realistic acceleration peak value, an alternative method was also undertaken to characterise the peak impact force. The method was based on estimating the impact force from the size of contact deformation dimension on the disk caused by the radiused pin in the presence of the lubricating fluid. The polished disk surface after a single impact revealed a microscopically visible localised area of slightly different surface appearance. However, due to the relatively large radius (24.8 mm) of the pin, it was difficult to quantitatively characterise the impacted disk surface. To cause a more definitive plastic deformation on the disk, a ball bearing with a much smaller radius (1 mm) was held on a shallow small blind hole made on the vertex of the pin. The identical pin/disk couples were also loaded statically to various levels of forces using a compression/tensile tester (Model 5565 with Type 2525-805 (5 kN) load cell, Instron Ltd., Buckinghamshire, UK) to obtain the relationship between the static force and the radius of plastic deformation. The levels of maximum forces used were 0, 10, 20, 30 50 and 100 N, and the compression rate used was 50 m/min. Based on the force-deformation radius curve, the radius of the plastic deformation due to the impact loading requires a static force of 20 N, approximately. That is, the impact force is roughly equivalent to a static force in the range of 20 N. The Hertzian indentation theory in the case of an uncoated-Ti6Al4 V pin of a 25 mm tip radius contacting an uncoated Ti6Al4 V disk under static loading predicts plastic deformation to occur for contact forces only at 86 N
N. Fujisawa et al. / Wear 254 (2003) 111–119
Fig. 3. Accelerometer signal on the acrylic part beneath the pin recorded at a 200 kHz sampling rate.
or beyond. Therefore, in theory, plastic deformation of the uncoated pin/disk couple due to the pin-tip radius was unlikely at the time of the impact loading. This suggests that the localised area of slightly different surface appearance on the impacted disk could have resulted from the asperity contact by the pin that had a surface roughness average 10 times greater than that of the disk. In order to visualise the overall impacting event in a larger time scale, an accelerometer signal on the acrylic part beneath the pin measured for a 50 ms period is shown in Fig. 3. The signal was recorded in the absence of the lubricating fluid using a 1 Hz to 100 kHz band-frequency range of the signal-conditioning amplifier and a 200 kHz sampling rate. In the figure, there is another peak at approximately 20 ms after the initial peak. A few minor acceleration peaks are also apparent, subsequent to the second peak. The number, magnitudes and the timing of these minor peaks, however, varied from time to time. After the two major impacts and a few randomly occurring minor impacts, the pin and disk remained in contact with each other until the disk was lifted again for the next impact event. The velocity of the disk immediately before impacting the pin, vi , was also estimated by time-integrating the instantaneous acceleration of the dropping disk component to the point of impact. The disk component acceleration was measured with the same accelerometer using a 1 Hz to 22.4 kHz band-frequency range and a 200 kHz sampling rate. The vi values obtained were in the range from 0.14 to 0.16 m/s (n = 24 measurements). 3.2. Transient contact force oscillation post impact After striking the pin at the start of the test, the spinning disk oscillated vertically while remaining in contact with the pin, with the oscillatory amplitude attenuating to zero in about 0.7–0.8 s in a quasi-exponential manner. This is evidenced in the typical waveform of the contact force, or
115
Fig. 4. Typical contact force waveform immediately after the first impact, recorded during the wear test of a 2 m-thick carbon-coated titanium pin on an uncoated titanium disk in Haemaccel using a sampling frequency of 100 Hz.
Fn , immediately after impact as shown in Fig. 4. Sliding contact without oscillation followed for about 8.3 s before the disk was lifted to impact the pin again at 10 s intervals for a 20 min period. The oscillatory amplitude attenuation immediately after the impact event is most likely caused by friction in the bearings and possibly by fluid-dynamic stresses acting on the submerged pin-component surfaces. To characterise the dynamic interactions between the disk and pin components during the 0.7 s period of oscillation, the following analytical modelling was conducted. By first neglecting the damping element in the system for simplicity, the equation of motion for the disk component while in contact with the pin component is, M
d2 x + Kx + Fn = 0, dt 2
(1)
where x is the displacement of the disk/pin component, which was set to zero in the steady sliding state at a contact force Fn = 1 N, and was referenced to increase in the direction of gravity, t the time in second, M = 0.771 kg the total mass of the disk component, and K = 491 N/m the extension spring constant. Similarly, for the pin component, m
d2 x + kx + B − Fn = 0, dt 2
(2)
where m = 0.123 kg is the total mass of the pin component, k = 407 N/m the compression spring constant, and B the buoyancy associated with the partial submersion of the pin component in the lubricating medium, and is expressed as a linear function of x as B = b1 x + b2 ,
(3)
where b1 = 2.26 N/m is the rate of buoyancy increase with increasing x, and b2 = 0.217 N the total buoyancy force
116
N. Fujisawa et al. / Wear 254 (2003) 111–119
when Fn = 1 N. By taking the sum of Eqs. (1) and (2), the time-varying Fn is eliminated as (m + M)
d2 x + (k + K + b1 ) x + b2 = 0 dt 2
(4)
Solving the differential equation for x gives x = X cos(2πf0 t + α) + x0 ,
(5)
where X is the oscillatory amplitude in the absence of fric√ tional loss, f0 = (k + K + b1 )/(m + M)/2π (5.05 Hz) is the frequency of oscillation of the joined disk/pin component, x0 = −b2 /(k + K + b1 ) + c is the midpoint of oscillation, and α and c are constants. The midpoint of oscillation was defined to be zero, and therefore c = b2 /(k +K +b1 ) = 0.24 mm. Since substitution of b1 = 0 and b2 = 0 results in negligible difference in both f0 (5.04 Hz) and x0 (0.24 mm), the effect of buoyancy on the frequency is negligibly small. The positive x0 value for b1 = 0 and b2 = 0 implies a downward shift of the midpoint of oscillation in the absence of buoyancy. Furthermore, the resonance frequencies of the individual disk and pin components while having no √ interaction with each other are K/M/2π = 4.02 Hz and √ k/m/2π = 9.16 Hz, respectively, and therefore the frequency f0 of the joined disk/pin component is shown to take a value closer to the resonant frequency of the disk component than to that of the pin component. The oscillatory amplitude X may be estimated by equating the sum of the kinetic and spring energy of the disk component immediately before striking the pin component to the total energy accumulated in the two springs when the joined pin/disk component was at the lower extreme position post impact. That is, 0.5Mvi2 + 0.5K(xi − xd0 )2 = 0.5k(X − xi )2 + 0.5K(X − xd0 )2
(6)
where x = xi is the position of the disk component immediately before impacting the pin, and x = xd0 the midpoint of oscillation of the disk component in the absence of the pin component. It is, however, noted that this equation assumed a negligible change in B and a negligible frictional loss in the bearings during the transition between the two points in x. Since vi in this equation is known from the previous section, and xi and xd 0 are measurable, X is calculated to be in the range between 4.9 and 5.4 mm depending on the value of vi used (0.14–0.16 m/s). To include the effect of damping on the system, Eq. (4) may be rewritten as d2 x dx + 2(m + M)γ dt dt 2 +(k + K + b1 ) x + b2 = 0
(m +M)
Solving the differential equation for x yields −γ t 2 2 cos (2πf0 ) − γ t + α + x0 x = Xe where γ is the damping factor.
(7)
(8)
Fig. 5. Vertical displacement of the joined pin/disk component with time post impact (dots) and predicted vertical displacement waveform using Eq. (8) and a damping factor of 3.2 (solid line).
Fig. 5 shows an instantaneous displacement waveform post impact while the disk was not spun, which was obtained as an instantaneous contact force signal divided by the compression spring constant, k. On the graph, the time-displacement curve predicted by Eq. (8) with γ = 3.2 and X = 5.2 mm is superimposed as the dashed line. The damping factor was used as it provided the best fit of the predicted displacement to thefirst two cycles of the oscillation. The angular frequency (2πf0 )2 − γ 2 was calculated to be 31.57 rad/s for γ = 3.2. Since the angular frequency for γ = 0 is 31.73 rad/s, the effect of damping on the frequency is in fact slight. Although the predicted frequency of oscillation is slightly less than the experimental one, the predicted displacement curve approximated the experimental curve reasonably well until the third positive peak of oscillation. Beyond this point in time, the oscillatory amplitude attenuated at a greater rate and eventually to a steady state. This indicates an increase of the viscous damping factor towards the complete decay of the oscillation, presumably because of the enhanced frictional loss in the bearings at lower velocities of oscillation. 3.3. Transient disk speed fluctuation post impact After being impacted by an uncoated pin, the rotational speed of a carbon-coated disk was shown to slow down initially, then to complete one cycle of oscillation before levelling out 1.5 s after the impact (not shown). The carbon film was deposited on the disk as multiple carbon film layers of a 48 nm thickness to an approximate full thickness of 2 m using cathodic arc and plasma immersion ion implantation. The resulting indentation hardness and biaxial modulus were in the range of 10 and 100 GPa, respectively [7]. For this frictional couple, the oscillation amplitudes were in the range of 3% of the baseline level (2000 rpm).
N. Fujisawa et al. / Wear 254 (2003) 111–119
117
However, when the carbon coating delaminated from the disk, the oscillation amplitude increased to >5%, presumably due to an increase of friction. The transient fluctuation of the disk speed could have influenced the tribological performance of the pin/disk specimens. 3.4. Repeatability of test results The frequency of disk revolution used in this study is in the range of 33 Hz, which is significantly greater than the frequency of the impact loading cycle of 0.1 Hz. Because of the large difference between the two frequencies together with fluctuations of the disk speed, locations of impacts on the disk and their sequence are determined randomly in this test system. This means that there is a chance for a local site of a coated disk surface to be impacted more often than elsewhere, thus presenting slight variability of the tribological performance of the coating. A greater degree of variability in the tribological performance of the disk coating may be expected if the pin loses its original surface profile more rapidly as the impact pressure decreases progressively with impact number or time. 4. Friction and wear results Fig. 6 shows the contact force (Fn ), frictional force, or Ft , and the coefficient of friction of the uncoated pin/carbon coated disk couple, tested for a 20 min period. Since the mechanical time constant of the DC servomotor whilst not loaded is 27 ms according to the motor specifications (i.e. frequency response of about 6 Hz), and the amplitude spectra of torque signals beyond 3 Hz were insignificant relative to those below 3 Hz, the three signals sampled at 100 Hz were low-pass-filtered by n-spikes filter with n = 8 three times, which provided an effective cut-off frequency of 3.3 Hz. As seen in the top panel of Fig. 6, the contact force fluctuated between 1.0 and 1.1 N in this experiment (referred later as exp. 1), for two reasons. The Fn fluctuation in the majority of cases resulted from the vertical positions of the moving disk and pin components varying slightly with time, presumably because of the inconsistent frictional effect of the bearings. Furthermore, where there was a large deviation (∼0.1 N) of Fn from Fn = 1 N, the vertical position of the single-axis translation stage, from which the disk was spring-suspended, was adjusted manually using the micrometer to achieve Fn = 1 N. The adjustment was done, however, only when the disk was lifted away from the pin so as not to impose any transient changes of the contact force during continuous sliding. Despite the fluctuation in contact force, the waveform profile of the frictional force was consistent with that of the friction coefficient, confirming negligible effects of the Fn fluctuation on the wear mode. The coefficient of friction curve of this experiment remained relatively flat and smooth until around 450 s, beyond which it began to fluctuate with greater amplitude up
Fig. 6. Contact force, frictional force and coefficient of friction with time for a uncoated pin/carbon-coated disk couple (top three panels), and friction coefficient waveform of an uncoated pin/disk couple as a control (bottom panel).
to 700 s. This was followed by a significant increase both in fluctuation amplitude and in the base line level (0.1 → 0.3–0.4). The control uncoated pin/disk couple, on the other hand, showed a higher time-mean coefficient (∼0.4) throughout the test period with much greater amplitude of fluctuation due to scuffing of the titanium-on-titanium sliding contact. Considering the higher friction and tendency to scuffing of the titanium-on-titanium contact in the control experiment, the increased amplitude of friction fluctuation of the coated disk after 450 s was likely to be due to an increase in the number and/or the total area of localised regions where coating had been removed, causing a more direct titanium-on-titanium contact and unstable disk revolution. The subsequent increase both in the fluctuation amplitude and in the base line level of the friction coefficient to levels comparable to those of the titanium-on-titanium contact indicates gross coating removal along the entire sliding path.
118
N. Fujisawa et al. / Wear 254 (2003) 111–119
Fig. 7. Coefficient of friction vs. time for two identical uncoated pin/coated disk pairs tested under the repetitive impact loading cycle (top panel), and for two other identical pairs subjected to one or two impact loadings at the beginning of the test (bottom panel), where markers represent time-mean values over 7.5 s between two sequential impact loadings, or over every 10 s in the case of continuous sliding without repetitive impact loading.
Another identical pin/disk pair was tested in the same manner as in exp. 1 (exp. 2), and two other identical pairs in the same way but with only one impact loading cycle at the start of the test (exp. 3 and 4). However, due to an error in the solenoid switching operation, the disk in exp. 4 was unintentionally dropped one more time at approx. 20 s but from a much lower height. The resulting friction curves of the four experiments (exp. 1–4) are shown in Fig. 7. The coated disks subjected to the repetitive impact loading were exposed over substantial areas as shown in the upper panel of Fig. 8 (shown only for exp. 1). The surface profile of the disk measured across the wear track radially using surface profilometry (SJ-400 Surface Roughness Tester, Mitsutoyo Corporation, Kawasaki, Kanagawa, Japan) is shown in the lower panel of Fig. 8. Based on the step change of surface profile at either end of the wear track width, the film appears to have delaminated rather than worn progressively. Also, the wear depths greater than the coating thickness (∼2 m) by up to 5 m indicate removal of the underlying titanium substrate. The counter body or the pin also wore substantially to have a wear scare diameter identical to the width of the wear track on the disk. On the other hand, the coating exposed to a single impact loading cycle was lost significantly in one case (exp. 3) and remained intact with minimal wear also on the mating pin surface in the other (exp. 4). Microscopic inspection of the coating disk of exp. 4 revealed a light sliding wear trace on the coating all along the sliding path except at four discrete small regions, where the coating was locally lifted and bulged up from the surface (Fig. 9). The surface profilometry across the light sliding wear trace also did not show any measurable loss of the film. Since the number of
Fig. 8. A photograph of the carbon-coated disk from exp. 2 is shown in the upper panel, where the brighter regions correspond to exposed titanium substrate surfaces. The lower panel shows the surface profile measured radially across the wear track of the same disk along the dashed line indicated in the upper panel. The radial distance was referenced to decrease towards the disk centre.
coating-bulged regions was four, which was greater than the number of disk drop (1 or 2), not only the initial prime impact but also the following few minor impacts in the impact loading cycle were sufficient to cause localised coating lift-off. The disk on which there was good coating retention, except at the impact-loaded points of exp. 4 suggests that the same coating in the absence of the repetitive impact loading would remain intact for a much longer period of time. Based
Fig. 9. A scanning electron micrograph of the carbon coating that was locally bulged up from the titanium alloy substrate due to the initial prime impact.
N. Fujisawa et al. / Wear 254 (2003) 111–119
on the film’s sliding wear resistance and impact-induced local bulging from exp. 4, the gross film adhesion failure in the remaining three experiments is most likely to have been initiated by the local film bulging that progressed along the sliding path with increasing number of the impact loading cycles. The coating appeared to have been removed at a region on the sliding path where the delaminated/lifted coating was no longer able to withstand the frictional force by the sliding pin. Finally, the coating removal process is likely to have propagated all the way along the sliding path. The timing of gross film adhesion failure was much earlier for the coating that underwent a single impact loading cycle at the start of the test in exp. 3 (approx. 150 s for exp. 3 as opposed to approx. 1000 s for exp. 1 and 2). Although the exact reason for this is unknown, there could have been a greater residual stress in this particular film, which possibly resulted in a local bulging of the coating of a greater magnitude.
5. Conclusions A repetitive impact loading between periods of continuous sliding was shown effective in a novel pin-on-disk apparatus for testing simultaneously both the sliding wear resistance and film adhesion of a coated system. The repetitive impact loading generated by the apparatus was able to promote adhesion failure of a thin film coating applied on a titanium alloy disk. This coating would have survived the test within the 20 min test period in the absence of the repetitive impact loading due to its low friction coefficient and adequate sliding wear resistance. The mechanisms of film
119
adhesion failure involved initial local debonding of the film at discrete regions along the sliding path, where the repetitive pin/disk contact occurred, and subsequent removal of the bulged coating by the frictional force exerted by the sliding pin.
Acknowledgements This study was supported by the Australian Research Council under the Strategic Partnerships with Industry, Research and Training (SPIRT) Scheme. References [1] ASTM Standards G99-95a(200)e1, Standard Test Method for Wear Testing with a Pin-on-Disk Apparatus, ASTM International, West Conshohocken, PA, 2000, Electronic Version. [2] E.-S. Yoon, H. Kong, O.-K. Kwon, J.-E. Oh, Evaluation of frictional characteristics for a pin-on-disk apparatus with different dynamic parameters, Wear 203–204 (1997) 341–349. [3] R.G. Bayer, P.A. Engel, J.L. Sirico, Impact wear testing machine, Wear 19 (1972) 343–354. [4] S.L. Rice, Reciprocating impact wear testing apparatus, Wear 45 (1977) 85–95. [5] R.J. Pick, K. Brown, A. Plumtree, Techniques in the study of impact and sliding wear of zircaloy-4, Wear 52 (1979) 381–392. [6] B.D. Beake, S.R. Goodes, J.F. Smith, Micro-impact testing: a new technique for investigating thin film toughness, adhesion, erosive wear resistance, and dynamic hardness, Surf. Eng. 17 (2001) 187–192. [7] N. Fujisawa, M.V. Swain, N.L. James, R.N. Tarrant, J.C. Woodard, D.R. McKenzie, Nano-indentation studies of brittle thin films on a titanium alloy substrate, J. Mater. Res. 17 (2002) 861–870.
A novel pin-on-apparatus Naoki Fujisawa a,∗ , Natalie L. James b , Richard N. Tarrant c , David R. McKenzie c , John C. Woodard b , Michael V. Swain a,1 a
Biomaterials Science Research Unit, Faculties of Dentistry and Engineering, University of Sydney, Suite G11, National Innovation Centre, Australian Technology Park, Eveleigh, NSW 1430, Australia b VentrAssist Division, Ventracor Limited, 126 Greville Street, Chatswood, NSW 2067, Australia c Department of Applied and Plasma Physics, School of Physics, University of Sydney, Sydney, NSW 2006, Australia Received 26 March 2002; received in revised form 11 September 2002; accepted 29 October 2002
Abstract A novel pin-on-disk apparatus was developed that provides a repetitive impact loading between periods of sliding through alternate lifting and dropping of a spring-suspended spinning disk, away from, and onto, a spring-supported pin, respectively. The combination of the repetitive impact loading and sliding achieved in the apparatus was able to induce film adhesion failure of a thin film coated disk within 20 min, which, in the absence of the impact loading, would have survived the test due to the adequate sliding wear resistance. The impact/sliding pin-on-disk apparatus developed is therefore a useful means of predicting the sliding wear resistance and film adhesion of a coated system simultaneously. © 2002 Elsevier Science B.V. All rights reserved. Keywords: Pin-on-disk apparatus; Impact/sliding; Film adhesion failure; Thin film
1. Introduction A conventional pin-on-disk apparatus is commonly employed as a means of predicting the tribological performance of a thin film coating in continuous sliding contact. A typical pin-on-disk apparatus provides a normal contact load between a stationary pin and a revolving disk, and measures the resulting frictional force to evaluate the coefficient of friction [1]. However, although a specific time-mean contact load may be set for a given pin/disk pair at a given sliding velocity, the dynamic fluctuation of the contact load can vary depending upon the loading mechanisms used in the test apparatus (e.g. spring, dead-weight or pneumatic loading) [2]. The wear mechanism of a coated system involves a progressive volume loss of the film material due to sliding friction, or abrupt film adhesion failure when the localised strain energy exceeds the adhesion energy at the film/substrate interface (Griffith’s criterion). Considering the energy threshold-sensitive mechanism of film adhesion failure, a slight difference in dynamic loading condition alone could alter the failure mode of the coated system. A repeated contact event under impact loading conditions in addition to ∗ Corresponding author. Tel.: +61-2-9351-1814; fax: +61-2-9351-1815. E-mail address: [email protected] (N. Fujisawa). 1 Co-corresponding author.
the conventional continuous sliding often occurs in a practical wear situation, thus increasing the likelihood of the coated system to meet Griffith’s criteron. To date at least several impact/sliding wear test apparatuses have been developed to study the wear resistance mainly of bulk metals and polymers. These include a wear test rig that directs bullet-shaped projectiles against a vertical metal disk surface which is either stationary or in motion [3], a pin-on-disk-type apparatus with a spatial crank oscillator mechanism to provide an adjustable impact load impulse in synchrony with the rotational disk motion [4], a cam-based device that utilises two independent cam/cam follower/lever systems, with one system generating a horizontal reciprocated motion of an upper specimen by an eccentrically mounted cam, and the other system allowing the upper specimen assembly to drop repetitively onto the lower specimen using a cut-out cam and gravity [5], and so forth. Most recently, a micro-impact testing method was developed [6] with the particular aim of testing thin films for film adhesion failure, whereby an actively oscillating (diamond) test probe contacts the vertical specimen surface repetitively at a much greater frequency (up to 1000 Hz) than occurs in other impact/sliding devices. The specimen position can be moved in the transverse direction during testing but at a much slower sliding speed (∼1 m/s). Film adhesion failure under this micro-impact loading was evaluated as the time
0043-1648/02/$ – see front matter © 2002 Elsevier Science B.V. All rights reserved. PII: S 0 0 4 3 - 1 6 4 8 ( 0 2 ) 0 0 3 0 7 - 1
112
N. Fujisawa et al. / Wear 254 (2003) 111–119
when the probe position (depth), which was monitored continuously during the test, showed a step increase resulting from the localised film loss. The results reported [6] demonstrated clearly that a repetitive high frequency micro-impact could promote film adhesion failure. In this study, a pin-on-disk apparatus was developed that provides a repetitive ‘macro’-impact loading of a much lower frequency (∼0.1 Hz) between periods of continuous sliding, by alternately lifting a spinning disk away from a pin using a solenoid, and dropping it onto the pin under gravity. The pin-on-disk apparatus was designed primarily to simulate a potential transient sliding contact event of mating titanium alloy surfaces that are, in normal operating conditions of the device, fully separated from each other via a highly stable fluid film bearing. A contact between the two surfaces may occur only when the device is exposed to a sudden acceleration or deceleration of an excessive magnitude, although that is a condition highly unlikely for the device to experience.
2. Pin-on-disk apparatus 2.1. Apparatus design Cylindrical pins (6 mm diameter by 20 mm length) and flat circular disks (19 mm diameter by 8 mm thickness) were cut from Ti-6Al-4V ELI (Extra Low Interstitial grade) rods. One end of the pin was machine-finished to a 24.8 mm-radiused tip with a resulting surface roughness average in the range of 0.2 m. The working surface of the disk was polished to a roughness average smaller by an order of magnitude. The indentation hardness and biaxial modulus of the titanium alloy substrate were in the range of 4.5 and 114 GPa, respectively [7]. The schematic drawing of the pin-on-disk wear tester is shown in Fig. 1. A brushless DC-servomotor (3556K024B, Minimotor SA, Croglio, Switzerland) was used to revolve the disk about the disk centre. The rotational speed of the disk was speed-controlled to approximately 2000 rpm by a four-quadrant pulse-width modulation servo amplifier (BLD 5606-SH4P, Minimotor SA, Croglio, Switzerland). A 50 mm-long titanium extension shaft transmitted the motor shaft revolution to the disk via a cylindrical disk holder (19 mm × 8 mm diameter), which was screw-attached to the end of the extension shaft in axial alignment. The disk was attached onto the flat disk-holder face using an adhesive double-sided tape (ScotchTM Brand #4616AFT, 3M Australia Pty. Ltd., Sydney, Australia). The component consisting of the disk and the motor with the disk surface facing down was suspended vertically via an extension spring from a precision single-axis translation stage whose vertical level was adjustable by a built-in micrometer (Melles Griot Inc., Irvine, CA, USA). The spring-suspended disk component was allowed to move only in the vertical direction with the use of a linear motion guide-rail type bearing (9 mm rail
width with 31 mm bearing length, RSR9ZMUU + 167 l, THK Co. Ltd., Tokyo, Japan). The pin, with the radiused pin tip pointing upwards, was held securely in the acrylic pin carriage component. The pin carriage component was free to slide in the vertical direction alone by incorporating two ball-bushing bearings (6 mm i.d. × 19 mm length, LM6-AJ, THK Co. Ltd., Tokyo, Japan) parallel to each other and to the pin axis. The pin carriage component was supported by a compression spring from the component bottom at 65 mm front of the pin axis. The pin carriage component was made of acrylic mainly for the lightness of the moving component. The disk/pin couple was immersed in a blood volume expansion medium (Haemaccel® , Hoechst Marion Roussel Australia Pty. Ltd., Lane Cove, NSW, Australia) in a plastic bath at room temperature to simulate potential applications in a biological environment. The horizontal pin position was located so that the tip of the pin slid on the disk along a circle of approximately 5 mm radius. The vertical position of the single-axis translation stage, from which the disk was spring-suspended, was adjusted using the micrometer such that the time mean of the contact force between the disk and the pin would be approximately 1 N during steady sliding contact. To induce an impact loading, the spring-suspended disk was lifted upwards to a defined height using a linear pull-action solenoid (43 series, 12 V dc, BLP, Suffolk, UK), held elevated for 1 s, and subsequently dropped onto the pin component under gravity. The lifting and dropping of the disk component were repeated at every 10 s by alternately energising and disconnecting the solenoid using a solid-state relay with a rectangular pulse generated by a function generator as the on-off switching signal. 2.2. Friction coefficient measurement The coefficient of friction in the steady sliding condition was obtained as the ratio of the frictional force to the contact force during the three quarters of the 10 s duty cycle period (∼7.5 s) starting at 1.5 s after every impact loading. The friction coefficient during the initial 1.5 s of the duty cycle, in which transient fluctuations of the contact force, frictional force and the motor speed took place, was excluded from consideration. A strain gauge type load cell (1 kg capacity, A.L. Design Inc., Buffalo, NY, USA) was located beneath the compression spring to measure the normal force in the compression spring. The load cell output was amplified using a strain gauge signal-conditioning module (AMA-RM-044, 2.5 kHz frequency response, Applied Measurement Australia Pty. Ltd., Oakleigh, Vic., Australia). The contact force between the pin and disk equates to the vector summation of the upward force in the compression spring, a downward force due to the total weight of the pin component, and an upward buoyancy force associated with the partial submersion of the pin component in the lubricating fluid. Although the buoyancy force is dependent on the submerged volume or
N. Fujisawa et al. / Wear 254 (2003) 111–119
113
Fig. 1. Schematic drawing of the pin-on-disk apparatus (without lubricant bath), where the moving disk and pin components are shaded in different intensities.
the vertical position of the moving pin component, the amplitude of buoyancy fluctuation calculated (<0.01 N) was much smaller than the maximum amplitude of force fluctuation in the compression spring (∼1.6 N). Therefore, the contact force was estimated on the assumption of a constant buoyancy force (∼0.217 N). The frictional force was calculated as the motor torque due to friction alone (obtained as the total motor torque minus the motor torque whilst no contact between the disk and the pin) divided by the radius of the wear track. The measured back-emf voltage waveform of the motor was sinusoidal with negligibly small amplitudes of its harmonics, and the inductance of the motor was negligibly low, and therefore the motor torque may be determined as the back-emf
constant multiplied by the RMS (root mean square) of the fundamental of the line-to-line phase current multiplied by 3 (the number of motor phases). In fact, the actual phase current that was measured by a current transducer (LA25-NP, DC to 150 kHz (−1 dB) bandwidth LEM, Geneva, Switzerland) was a square waveform in each phase of width 120◦ on and 60◦ off with a distinctive ‘W’-shaped waveform on top of each 120◦ on sub-phase. However, the RMS of the fundamental phase current was found to be proportional to that of the quasi-square waveform by a ratio factor (∼0.95) regardless of the amplitude or frequency, as was confirmed by fast Fourier transformation analysis. Therefore, the fundamental RMS could be evaluated simply as the RMS of the filtered current multiplied by the ratio factor. The RMS of the filtered
114
N. Fujisawa et al. / Wear 254 (2003) 111–119
current was obtained through a wide-band RMS-to-DC converter (AD637, Analog Devices Inc., Norwood, MA, USA), which was configured for two pole Sallen-Key filter option with an averaging time constant of 120 ms. The motor shaft speed was also recorded using a frequency-to-voltage converter (Microgenic Systems Version 2/1 TACHO, 100 ms response time, Farnell Components, Chester Hill, Australia). The three signals (contact force, motor torque and motor speed) were recorded at a sampling rate of 100 Hz. 3. Mechanical characterisation of the apparatus 3.1. Impact loading The spinning disk struck the pin at the start of the test. The magnitude of impact was adjusted experimentally by varying the disk height for dropping to a level that was sufficient to cause gross film adhesion failure of a coated disk against an uncoated pin within a 20 min test period. The instantaneous acceleration of the moving pin component at the time of the initial impact loading was estimated using two piezoelectric accelerometers (Type 4375, ≤2% error below 10 kHz to 10% error at 16.5 kHz, Brüel and Kjær, Nærum, Denmark), coupled with a signal-conditioning amplifier (NEXUS(tm) Type 2692, Brüel and Kjær, Nærum, Denmark). One accelerometer was attached onto the back of the acrylic part behind the blind hole where the pin was inserted (defined in Fig. 1 as position 1, and the other on top of the acrylic part beside one of the ball-bushing bearings of the pin component (defined in Fig. 1 as position 2, both using beeswax. Measurements were conducted in the absence of lubricating fluid due to one of the non-water-proof accelerometers placed in position 1. An untreated titanium pin and a non-spinning untreated titanium disk were used to estimate the maximum instantaneous impact force possible for an unworn pin/disk couple. Fig. 2 shows accelerometer signals recorded at the two accelerometer positions using a band-frequency range of 1 Hz to 22.4 kHz of the signal-conditioning amplifier and a sampling rate of 5 MHz. The upper band frequency limit of 22.4 kHz was used because the use of higher upper frequency limits (30 and 100 kHz) did not significantly alter the signal waveform profile. These waveforms appear to have reflected the actual acceleration signals reasonably well since the frequency of the initial peak is 6–7 kHz, which is well below the resonance frequency of the accelerometer (∼55 kHz). As can be seen in the figure, not only the magnitude but also the sign of the initial peak measured away from the pin were different from those beneath the pin. In addition, the initial acceleration peak recorded away from the pin lagged behind that recorded beneath the pin. These differences and the phase lag suggest an instantaneous deflection/deformation of the acrylic part to a considerable extent. Further, the recorded peak value of 5680 m/s2 beneath the
Fig. 2. Accelerometer signals on the acrylic part of the pin carriage component beneath the pin (solid line) and away from the pin beside the ball-bushing bearings (dashed line) as a function of time, t, using a 5 MHz sampling rate. Acceleration was referenced to increase in the direction of gravity.
pin may still underestimate the true instantaneous acceleration experienced by the pin/disk couple because the acrylic section between the pin bottom and the accelerometer could have absorbed part of the instantaneous impact energy due to the low stiffness of the material. The presence of the lubricating fluid would also affect the peak acceleration. Due to the difficulty in obtaining a realistic acceleration peak value, an alternative method was also undertaken to characterise the peak impact force. The method was based on estimating the impact force from the size of contact deformation dimension on the disk caused by the radiused pin in the presence of the lubricating fluid. The polished disk surface after a single impact revealed a microscopically visible localised area of slightly different surface appearance. However, due to the relatively large radius (24.8 mm) of the pin, it was difficult to quantitatively characterise the impacted disk surface. To cause a more definitive plastic deformation on the disk, a ball bearing with a much smaller radius (1 mm) was held on a shallow small blind hole made on the vertex of the pin. The identical pin/disk couples were also loaded statically to various levels of forces using a compression/tensile tester (Model 5565 with Type 2525-805 (5 kN) load cell, Instron Ltd., Buckinghamshire, UK) to obtain the relationship between the static force and the radius of plastic deformation. The levels of maximum forces used were 0, 10, 20, 30 50 and 100 N, and the compression rate used was 50 m/min. Based on the force-deformation radius curve, the radius of the plastic deformation due to the impact loading requires a static force of 20 N, approximately. That is, the impact force is roughly equivalent to a static force in the range of 20 N. The Hertzian indentation theory in the case of an uncoated-Ti6Al4 V pin of a 25 mm tip radius contacting an uncoated Ti6Al4 V disk under static loading predicts plastic deformation to occur for contact forces only at 86 N
N. Fujisawa et al. / Wear 254 (2003) 111–119
Fig. 3. Accelerometer signal on the acrylic part beneath the pin recorded at a 200 kHz sampling rate.
or beyond. Therefore, in theory, plastic deformation of the uncoated pin/disk couple due to the pin-tip radius was unlikely at the time of the impact loading. This suggests that the localised area of slightly different surface appearance on the impacted disk could have resulted from the asperity contact by the pin that had a surface roughness average 10 times greater than that of the disk. In order to visualise the overall impacting event in a larger time scale, an accelerometer signal on the acrylic part beneath the pin measured for a 50 ms period is shown in Fig. 3. The signal was recorded in the absence of the lubricating fluid using a 1 Hz to 100 kHz band-frequency range of the signal-conditioning amplifier and a 200 kHz sampling rate. In the figure, there is another peak at approximately 20 ms after the initial peak. A few minor acceleration peaks are also apparent, subsequent to the second peak. The number, magnitudes and the timing of these minor peaks, however, varied from time to time. After the two major impacts and a few randomly occurring minor impacts, the pin and disk remained in contact with each other until the disk was lifted again for the next impact event. The velocity of the disk immediately before impacting the pin, vi , was also estimated by time-integrating the instantaneous acceleration of the dropping disk component to the point of impact. The disk component acceleration was measured with the same accelerometer using a 1 Hz to 22.4 kHz band-frequency range and a 200 kHz sampling rate. The vi values obtained were in the range from 0.14 to 0.16 m/s (n = 24 measurements). 3.2. Transient contact force oscillation post impact After striking the pin at the start of the test, the spinning disk oscillated vertically while remaining in contact with the pin, with the oscillatory amplitude attenuating to zero in about 0.7–0.8 s in a quasi-exponential manner. This is evidenced in the typical waveform of the contact force, or
115
Fig. 4. Typical contact force waveform immediately after the first impact, recorded during the wear test of a 2 m-thick carbon-coated titanium pin on an uncoated titanium disk in Haemaccel using a sampling frequency of 100 Hz.
Fn , immediately after impact as shown in Fig. 4. Sliding contact without oscillation followed for about 8.3 s before the disk was lifted to impact the pin again at 10 s intervals for a 20 min period. The oscillatory amplitude attenuation immediately after the impact event is most likely caused by friction in the bearings and possibly by fluid-dynamic stresses acting on the submerged pin-component surfaces. To characterise the dynamic interactions between the disk and pin components during the 0.7 s period of oscillation, the following analytical modelling was conducted. By first neglecting the damping element in the system for simplicity, the equation of motion for the disk component while in contact with the pin component is, M
d2 x + Kx + Fn = 0, dt 2
(1)
where x is the displacement of the disk/pin component, which was set to zero in the steady sliding state at a contact force Fn = 1 N, and was referenced to increase in the direction of gravity, t the time in second, M = 0.771 kg the total mass of the disk component, and K = 491 N/m the extension spring constant. Similarly, for the pin component, m
d2 x + kx + B − Fn = 0, dt 2
(2)
where m = 0.123 kg is the total mass of the pin component, k = 407 N/m the compression spring constant, and B the buoyancy associated with the partial submersion of the pin component in the lubricating medium, and is expressed as a linear function of x as B = b1 x + b2 ,
(3)
where b1 = 2.26 N/m is the rate of buoyancy increase with increasing x, and b2 = 0.217 N the total buoyancy force
116
N. Fujisawa et al. / Wear 254 (2003) 111–119
when Fn = 1 N. By taking the sum of Eqs. (1) and (2), the time-varying Fn is eliminated as (m + M)
d2 x + (k + K + b1 ) x + b2 = 0 dt 2
(4)
Solving the differential equation for x gives x = X cos(2πf0 t + α) + x0 ,
(5)
where X is the oscillatory amplitude in the absence of fric√ tional loss, f0 = (k + K + b1 )/(m + M)/2π (5.05 Hz) is the frequency of oscillation of the joined disk/pin component, x0 = −b2 /(k + K + b1 ) + c is the midpoint of oscillation, and α and c are constants. The midpoint of oscillation was defined to be zero, and therefore c = b2 /(k +K +b1 ) = 0.24 mm. Since substitution of b1 = 0 and b2 = 0 results in negligible difference in both f0 (5.04 Hz) and x0 (0.24 mm), the effect of buoyancy on the frequency is negligibly small. The positive x0 value for b1 = 0 and b2 = 0 implies a downward shift of the midpoint of oscillation in the absence of buoyancy. Furthermore, the resonance frequencies of the individual disk and pin components while having no √ interaction with each other are K/M/2π = 4.02 Hz and √ k/m/2π = 9.16 Hz, respectively, and therefore the frequency f0 of the joined disk/pin component is shown to take a value closer to the resonant frequency of the disk component than to that of the pin component. The oscillatory amplitude X may be estimated by equating the sum of the kinetic and spring energy of the disk component immediately before striking the pin component to the total energy accumulated in the two springs when the joined pin/disk component was at the lower extreme position post impact. That is, 0.5Mvi2 + 0.5K(xi − xd0 )2 = 0.5k(X − xi )2 + 0.5K(X − xd0 )2
(6)
where x = xi is the position of the disk component immediately before impacting the pin, and x = xd0 the midpoint of oscillation of the disk component in the absence of the pin component. It is, however, noted that this equation assumed a negligible change in B and a negligible frictional loss in the bearings during the transition between the two points in x. Since vi in this equation is known from the previous section, and xi and xd 0 are measurable, X is calculated to be in the range between 4.9 and 5.4 mm depending on the value of vi used (0.14–0.16 m/s). To include the effect of damping on the system, Eq. (4) may be rewritten as d2 x dx + 2(m + M)γ dt dt 2 +(k + K + b1 ) x + b2 = 0
(m +M)
Solving the differential equation for x yields −γ t 2 2 cos (2πf0 ) − γ t + α + x0 x = Xe where γ is the damping factor.
(7)
(8)
Fig. 5. Vertical displacement of the joined pin/disk component with time post impact (dots) and predicted vertical displacement waveform using Eq. (8) and a damping factor of 3.2 (solid line).
Fig. 5 shows an instantaneous displacement waveform post impact while the disk was not spun, which was obtained as an instantaneous contact force signal divided by the compression spring constant, k. On the graph, the time-displacement curve predicted by Eq. (8) with γ = 3.2 and X = 5.2 mm is superimposed as the dashed line. The damping factor was used as it provided the best fit of the predicted displacement to thefirst two cycles of the oscillation. The angular frequency (2πf0 )2 − γ 2 was calculated to be 31.57 rad/s for γ = 3.2. Since the angular frequency for γ = 0 is 31.73 rad/s, the effect of damping on the frequency is in fact slight. Although the predicted frequency of oscillation is slightly less than the experimental one, the predicted displacement curve approximated the experimental curve reasonably well until the third positive peak of oscillation. Beyond this point in time, the oscillatory amplitude attenuated at a greater rate and eventually to a steady state. This indicates an increase of the viscous damping factor towards the complete decay of the oscillation, presumably because of the enhanced frictional loss in the bearings at lower velocities of oscillation. 3.3. Transient disk speed fluctuation post impact After being impacted by an uncoated pin, the rotational speed of a carbon-coated disk was shown to slow down initially, then to complete one cycle of oscillation before levelling out 1.5 s after the impact (not shown). The carbon film was deposited on the disk as multiple carbon film layers of a 48 nm thickness to an approximate full thickness of 2 m using cathodic arc and plasma immersion ion implantation. The resulting indentation hardness and biaxial modulus were in the range of 10 and 100 GPa, respectively [7]. For this frictional couple, the oscillation amplitudes were in the range of 3% of the baseline level (2000 rpm).
N. Fujisawa et al. / Wear 254 (2003) 111–119
117
However, when the carbon coating delaminated from the disk, the oscillation amplitude increased to >5%, presumably due to an increase of friction. The transient fluctuation of the disk speed could have influenced the tribological performance of the pin/disk specimens. 3.4. Repeatability of test results The frequency of disk revolution used in this study is in the range of 33 Hz, which is significantly greater than the frequency of the impact loading cycle of 0.1 Hz. Because of the large difference between the two frequencies together with fluctuations of the disk speed, locations of impacts on the disk and their sequence are determined randomly in this test system. This means that there is a chance for a local site of a coated disk surface to be impacted more often than elsewhere, thus presenting slight variability of the tribological performance of the coating. A greater degree of variability in the tribological performance of the disk coating may be expected if the pin loses its original surface profile more rapidly as the impact pressure decreases progressively with impact number or time. 4. Friction and wear results Fig. 6 shows the contact force (Fn ), frictional force, or Ft , and the coefficient of friction of the uncoated pin/carbon coated disk couple, tested for a 20 min period. Since the mechanical time constant of the DC servomotor whilst not loaded is 27 ms according to the motor specifications (i.e. frequency response of about 6 Hz), and the amplitude spectra of torque signals beyond 3 Hz were insignificant relative to those below 3 Hz, the three signals sampled at 100 Hz were low-pass-filtered by n-spikes filter with n = 8 three times, which provided an effective cut-off frequency of 3.3 Hz. As seen in the top panel of Fig. 6, the contact force fluctuated between 1.0 and 1.1 N in this experiment (referred later as exp. 1), for two reasons. The Fn fluctuation in the majority of cases resulted from the vertical positions of the moving disk and pin components varying slightly with time, presumably because of the inconsistent frictional effect of the bearings. Furthermore, where there was a large deviation (∼0.1 N) of Fn from Fn = 1 N, the vertical position of the single-axis translation stage, from which the disk was spring-suspended, was adjusted manually using the micrometer to achieve Fn = 1 N. The adjustment was done, however, only when the disk was lifted away from the pin so as not to impose any transient changes of the contact force during continuous sliding. Despite the fluctuation in contact force, the waveform profile of the frictional force was consistent with that of the friction coefficient, confirming negligible effects of the Fn fluctuation on the wear mode. The coefficient of friction curve of this experiment remained relatively flat and smooth until around 450 s, beyond which it began to fluctuate with greater amplitude up
Fig. 6. Contact force, frictional force and coefficient of friction with time for a uncoated pin/carbon-coated disk couple (top three panels), and friction coefficient waveform of an uncoated pin/disk couple as a control (bottom panel).
to 700 s. This was followed by a significant increase both in fluctuation amplitude and in the base line level (0.1 → 0.3–0.4). The control uncoated pin/disk couple, on the other hand, showed a higher time-mean coefficient (∼0.4) throughout the test period with much greater amplitude of fluctuation due to scuffing of the titanium-on-titanium sliding contact. Considering the higher friction and tendency to scuffing of the titanium-on-titanium contact in the control experiment, the increased amplitude of friction fluctuation of the coated disk after 450 s was likely to be due to an increase in the number and/or the total area of localised regions where coating had been removed, causing a more direct titanium-on-titanium contact and unstable disk revolution. The subsequent increase both in the fluctuation amplitude and in the base line level of the friction coefficient to levels comparable to those of the titanium-on-titanium contact indicates gross coating removal along the entire sliding path.
118
N. Fujisawa et al. / Wear 254 (2003) 111–119
Fig. 7. Coefficient of friction vs. time for two identical uncoated pin/coated disk pairs tested under the repetitive impact loading cycle (top panel), and for two other identical pairs subjected to one or two impact loadings at the beginning of the test (bottom panel), where markers represent time-mean values over 7.5 s between two sequential impact loadings, or over every 10 s in the case of continuous sliding without repetitive impact loading.
Another identical pin/disk pair was tested in the same manner as in exp. 1 (exp. 2), and two other identical pairs in the same way but with only one impact loading cycle at the start of the test (exp. 3 and 4). However, due to an error in the solenoid switching operation, the disk in exp. 4 was unintentionally dropped one more time at approx. 20 s but from a much lower height. The resulting friction curves of the four experiments (exp. 1–4) are shown in Fig. 7. The coated disks subjected to the repetitive impact loading were exposed over substantial areas as shown in the upper panel of Fig. 8 (shown only for exp. 1). The surface profile of the disk measured across the wear track radially using surface profilometry (SJ-400 Surface Roughness Tester, Mitsutoyo Corporation, Kawasaki, Kanagawa, Japan) is shown in the lower panel of Fig. 8. Based on the step change of surface profile at either end of the wear track width, the film appears to have delaminated rather than worn progressively. Also, the wear depths greater than the coating thickness (∼2 m) by up to 5 m indicate removal of the underlying titanium substrate. The counter body or the pin also wore substantially to have a wear scare diameter identical to the width of the wear track on the disk. On the other hand, the coating exposed to a single impact loading cycle was lost significantly in one case (exp. 3) and remained intact with minimal wear also on the mating pin surface in the other (exp. 4). Microscopic inspection of the coating disk of exp. 4 revealed a light sliding wear trace on the coating all along the sliding path except at four discrete small regions, where the coating was locally lifted and bulged up from the surface (Fig. 9). The surface profilometry across the light sliding wear trace also did not show any measurable loss of the film. Since the number of
Fig. 8. A photograph of the carbon-coated disk from exp. 2 is shown in the upper panel, where the brighter regions correspond to exposed titanium substrate surfaces. The lower panel shows the surface profile measured radially across the wear track of the same disk along the dashed line indicated in the upper panel. The radial distance was referenced to decrease towards the disk centre.
coating-bulged regions was four, which was greater than the number of disk drop (1 or 2), not only the initial prime impact but also the following few minor impacts in the impact loading cycle were sufficient to cause localised coating lift-off. The disk on which there was good coating retention, except at the impact-loaded points of exp. 4 suggests that the same coating in the absence of the repetitive impact loading would remain intact for a much longer period of time. Based
Fig. 9. A scanning electron micrograph of the carbon coating that was locally bulged up from the titanium alloy substrate due to the initial prime impact.
N. Fujisawa et al. / Wear 254 (2003) 111–119
on the film’s sliding wear resistance and impact-induced local bulging from exp. 4, the gross film adhesion failure in the remaining three experiments is most likely to have been initiated by the local film bulging that progressed along the sliding path with increasing number of the impact loading cycles. The coating appeared to have been removed at a region on the sliding path where the delaminated/lifted coating was no longer able to withstand the frictional force by the sliding pin. Finally, the coating removal process is likely to have propagated all the way along the sliding path. The timing of gross film adhesion failure was much earlier for the coating that underwent a single impact loading cycle at the start of the test in exp. 3 (approx. 150 s for exp. 3 as opposed to approx. 1000 s for exp. 1 and 2). Although the exact reason for this is unknown, there could have been a greater residual stress in this particular film, which possibly resulted in a local bulging of the coating of a greater magnitude.
5. Conclusions A repetitive impact loading between periods of continuous sliding was shown effective in a novel pin-on-disk apparatus for testing simultaneously both the sliding wear resistance and film adhesion of a coated system. The repetitive impact loading generated by the apparatus was able to promote adhesion failure of a thin film coating applied on a titanium alloy disk. This coating would have survived the test within the 20 min test period in the absence of the repetitive impact loading due to its low friction coefficient and adequate sliding wear resistance. The mechanisms of film
119
adhesion failure involved initial local debonding of the film at discrete regions along the sliding path, where the repetitive pin/disk contact occurred, and subsequent removal of the bulged coating by the frictional force exerted by the sliding pin.
Acknowledgements This study was supported by the Australian Research Council under the Strategic Partnerships with Industry, Research and Training (SPIRT) Scheme. References [1] ASTM Standards G99-95a(200)e1, Standard Test Method for Wear Testing with a Pin-on-Disk Apparatus, ASTM International, West Conshohocken, PA, 2000, Electronic Version. [2] E.-S. Yoon, H. Kong, O.-K. Kwon, J.-E. Oh, Evaluation of frictional characteristics for a pin-on-disk apparatus with different dynamic parameters, Wear 203–204 (1997) 341–349. [3] R.G. Bayer, P.A. Engel, J.L. Sirico, Impact wear testing machine, Wear 19 (1972) 343–354. [4] S.L. Rice, Reciprocating impact wear testing apparatus, Wear 45 (1977) 85–95. [5] R.J. Pick, K. Brown, A. Plumtree, Techniques in the study of impact and sliding wear of zircaloy-4, Wear 52 (1979) 381–392. [6] B.D. Beake, S.R. Goodes, J.F. Smith, Micro-impact testing: a new technique for investigating thin film toughness, adhesion, erosive wear resistance, and dynamic hardness, Surf. Eng. 17 (2001) 187–192. [7] N. Fujisawa, M.V. Swain, N.L. James, R.N. Tarrant, J.C. Woodard, D.R. McKenzie, Nano-indentation studies of brittle thin films on a titanium alloy substrate, J. Mater. Res. 17 (2002) 861–870.