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Tribology Letters

, 66:100 | Cite as

Measuring Evolution of Transfer Film–Substrate Interface Using Low Wear Alumina PTFE

  • Jiaxin Ye
  • Wei Sun
  • Yan Zhang
  • Xiaojun Liu
  • Kun Liu
Original Paper

Abstract

Polymeric solid lubricants lay down their own wear debris onto hard metallic counterfaces to form a protective transfer film which reduces friction and wear effectively without lubrication. Adhesive shear strength at the hidden interface between the film and substrate determines the film persistence and correlates with system wear qualitatively. Previous studies showed that an ultralow wear (k ~ 10− 7 mm3/Nm) alumina-PTFE solid lubricant forms an extremely adherent and complete transfer film, and strong chemical bonds between wear debris and counterface perpetuate the film–substrate adhesion very early in the sliding. In this paper, we aimed to test the permanence of such adhesion by removing pre-developed transfer films using sliding rubber contact and measuring the topographical evolution of the interface throughout the course of a standard wear test using the well-studied alumina-PTFE system. The results unexpectedly showed continuous wear of the counterface across the wear track, and counterface wear rate decreased proportionally from 3 × 10− 7 to 3 × 10− 8 mm3/Nm with increased film area fraction and sliding distance. A proposed rule-of-mixtures wear model coincided closely with the experimental results and strongly suggested a coupled mechanism of adhesive and fatigue wear of the counterface. The upper limit of the interfacial counterface fatigue wear rate was predicted to be 3 × 10− 8 mm3/Nm.

Keywords

Alumina-PTFE Transfer film–substrate interface Counterface abrasion Wear 

1 Introduction

When slid against hard metallic counterfaces, polymeric solid lubricants sacrifice debris to the counterface in the form of a transfer film which separates the sliding bodies and reduces friction and wear. Normally, polymers wear much faster than the counterfaces and wear rates of the two are about inversely proportional to their hardness ratio as first reported by Archard [1]. Polytetrafluorethylene (PTFE), for example, has a prohibitively high wear rate against steel and 10–40 wt% micro-sized fillers (e.g., glass fiber, metal oxides) are often used to reduce wear by up to 99.99% [2, 3]. The high content hard fillers tend to accumulate at the sliding surface and are supposed to preferentially support the normal load and reduce wear [4]. However, they also tend to pierce the otherwise lubricious PTFE-rich transfer film and cause increased friction and counterface abrasion [5].

The fact that micro-fillers often reduce polymer wear at the sacrifice of increased counterface abrasion and deteriorated transfer films lead tribologists to seek nanosized fillers as alternatives since they are of the same scale of roughness asperities and much less abrasive to the counterface. Starting from 2000, Li et al. [6], Chen et al. [7], and Sawyer et al. [8] found that certain nanofillers could reduce polymer wear as good as micro-fillers without evident counterface abrasion. Burris and Sawyer [9, 10] and many [11, 12] later demonstrated that nanofillers like alumina could reduce PTFE wear up to 30,000 × with as little as 0.2–2% loadings. The accompanying transfer films in these systems were extremely thin, continuous, complete, and adherent [13, 14]. Krick et al. [15] surveyed alpha-phase alumina fillers from a range of different vendors and found that the most effective wear reducing alumina fillers in PTFE are those which form porous and friable micro-sized agglomerates that get broken down into nanosized fragments at the sliding surface. The accumulation of such nanosized filler fragments mechanically reinforces the composite’s running surface and lead to a dramatic reduction in wear.

The most crucial part of transfer film quality is the adhesion strength. Bahadur and Tabor [16] slid PTFE against pre-developed transfer films of low wear PTFE composites and found the transfer films immediately removed and wear of PTFE unaffected by the existing films. They suggested debris bonded weakly to counterface asperities through mechanical interlocking and transfer films were instantaneously removed and replenished during sliding. Briscoe [17], Bahadur and Gong [18], and Schwartz and Bahadur [19] suggested that transfer film adhesion was of chemical nature and system wear rate decreased with increased adhesion strength.

A famous ultralow wear alumina PTFE solid lubricant provides probably the best opportunity for transfer film studies to date. The system deposits nanoscale debris so strongly adherent to counterfaces that film removal is a rare event and transfer film grows thicker and more uniform continuously during sliding [20]. Film adhesion strength is ~ 15 × greater than in unfilled PTFE [21] and transfer film removal rate during sliding is on the order of 10− 9 mm3/Nm, a value of ~ 1% of the composite’s wear rate and similar to wear rate of commercial wear-resistant coatings. Harris et al. [22] recently proved the combination of ambient air and dry sliding induces degradations of PTFE chains which chemically bond the polymer to the metallic counterface and alumina fillers to form the extremely adherent transfer film and a wear-resistant running film at the bulk’s running surface. This is a significant breakthrough in recent transfer film studies and further confirms the importance of tribochemistry in transfer film adhesion.

There is little doubt strong adhesion is crucial for high-quality transfer film formation, yet it remains unclear how long such adherence could last. It is almost intuitive to assume friction and wear only occur or initiate at the polymer-on-polymer interface between the bulk’s running surface and the transfer film top surface once a persistent film is formed and low wear reached. In other words, transfer film adhesion is permanent unless wear of film reaches the interface or delamination occurs. In this study, we aim to test this hypothesis by removing pre-developed transfer films and measuring the evolution of the film–substrate interface throughout the course of a wear test. The goal is to shed more light on the relations between transfer film adhesion, counterface abrasion, and overall wear reduction. A well-studied alumina-PTFE nanocomposite is chosen for its ultralow wear rate and extremely persistent transfer film as previously documented in detail [5, 9, 10, 13, 20, 21, 22, 23, 24, 25, 26, 27, 28, 29, 30, 31, 32]. Ultralow wear is defined in this study as having wear rates below 10− 6 mm3/Nm, a definition for wear-resistant PTFE composites suggested by Harris et al. [22].

2 Materials and Methods

2.1 Composite and Counterface Preparation

The polymeric solid lubricant used in this study was a 5 wt% α-Al2O3 PTFE nanocomposite. The composition and preparation method have been described by the same authors [31, 32] and are comparable to those used in many studies [9, 22, 25, 26, 27, 28, 29, 33]. The PTFE powder was a DuPont Teflon™ 7C molding resin with a reported average particle size of 40 µm. Alpha-phase alumina nanoparticles were obtained from Shanghai Maikun Chemical Co., Ltd (99.99%, product link: https://goo.gl/RR5QfV) with a reported average particle size of 30 nm measured with scanning electron microscopy (SEM) as illustrated in Fig. 1a. These nanoparticles tend to form micro-sized agglomerates likely due to their manufacturing process and high surface-to-volume ratio. Krick et al. did a thorough study on wear reducing alpha-phase alumina nanoparticles in PTFE and found a particular effective alumina filler with vendor-specified particle size of 27–43 nm (Nanostructured and Amorphous Materials, Stock #1015WW) formed similar micro-agglomerates which were porous and nanostructured under SEM inspections [15]. Based on this finding, it is probably inaccurate to describe the alumina filler used in this study as nanosized. However, the same authors suggested that, unlike traditional micro-sized dense alumina particles, the micro-agglomerates as shown in Fig. 1a are easy to get broken up into nanosized filler fragments at the sliding surface and can reduce PTFE wear significantly through a coupled mechanical and tribochemical mechanism [15].

Fig. 1

a SEM image of the alumina particles shows individual particle has a diameter of 30 nm but fillers tend to form micro-sized agglomerates which has a porous and nanostructured morphology (image courtesy of Shanghai Maikun Chemical). b ‘Stripe test’ used in this study to acquire the history of transfer film development with incrementally decreased linear reciprocating length and increased sliding distance. c and d Optical images of the same area showing the clean counterface, formed transfer film and counterface after transfer film was erased. e Line scans showing counterface abrasion caused by 200 k cycles of sliding against the composite pin. Each line is the average of 15 scans separated by 0.1 mm spacing (total width of 1.5 mm). Two profiles were matched at regions outside the wear track. (Color figure online)

The two powders were first hand-mixed in a polyethylene container with a 5:95 wt% ratio. The mixture was added into anhydrous ethanol with a 1:2 volume ratio and dispersed using a 500 W ultrasonic horn with a 25.8-mm-diameter titanium tip; full power was pulsed on for 5 s and off for 3 s over 3 min to improve dispersion state. The mixture was drained through a lab filter paper with 8 µm pore size (grade 40, diameter 90 mm, Whatman® by Sigma–Aldrich) for 20 min to remove the bulk ethanol and vacuum dried at 90 °C for 30 min. The dry powder was cold pressed using a Φ10 × 20 mm cylindrical mold and heated to 365 at 120 °C per hour, held for 3 h, and cooled at the same rate. The sintered samples were machined into 6.4 × 6.4 × 15 mm pins prior to the wear test.

Counterface used in this study was a 120 × 30 × 3 mm grade 304 stainless steel plate pre-ground and polished to ~ 50 nm Ra using a commercial automatic fine-grinding machine (Fangda FD-24LP, China). The finished surface had no obvious directionality to the roughness. Prior to the wear test, counterfaces were ultrasonicated with acetone for 10 min, cleansed in distilled water, and air dried for 30 min.

2.2 Wear and Friction Testing

Wear test was performed in ambient air condition (20 °C, 40% RH) using a custom-built linear reciprocating tribometer as previously described in detail [31]. The sliding conditions were selected to best match those of comparable studies [5, 13, 15, 22, 25, 26, 27, 28, 29, 33] with a normal load of 250 N, contact pressure of 6.5 MPa and sliding speed of 50 mm/s. A ‘stripe test’ was used following Harris et al. [22, 28, 30] by decreasing the reciprocating length from 90 to 20 with 10-mm intervals intermittently at predetermined sliding cycles of 1, 2, 5, 10, 20, 50, 100, and 200 k as illustrated in Fig. 1. Such test configuration preserved areas of early stage transfer films and provided the opportunities for further analyses of transfer film development. An unfilled PTFE pin was selected as the control sample and slid against the same steel counterface under identical test conditions (FN, P, V) except with a fixed stroke length of 25 mm. The control group test was stopped at 10 k sliding cycles (500 m) due to excessive wear of the pin.

At each test interruption, mass loss of the composite pin was measured with a mass balance (± 10 µg) and volume loss was calculated using mass loss data and predetermined sample density. In situ wear rate of the composite pin at any given test interval was calculated as the slope of the linear regression line of the measured wear data, one point before and one point after, and had the form of
$$k=\frac{{\Delta {V_{{\text{loss}}}}}}{{\Delta ({F_{\text{N}}} \cdot d)}}=\frac{{\Delta {m_{{\text{loss}}}}}}{{\rho \cdot {F_{\text{N}}} \cdot \Delta d}}\left[ {\frac{{{\text{m}}{{\text{m}}^{\text{3}}}}}{{{\text{Nm}}}}} \right]$$
(1)
where FN is the normal load, d the sliding distance, mloss the mass loss of the pin, and ρ is the sample density. The linear regression analysis uncertainty of in situ wear rate was determined using Monte-Carlo method described by Burris and Sawyer [34] and Ye et al. [23]. Friction coefficient was calculated as the ratio of average friction force and normal force during sliding. Uncertainty in friction coefficient was calculated following methods described by Schmitz et al. [35].

2.3 Transfer Film Characterization

The difficulty of transfer film morphology characterization lies in the separation of the film from counterface. Scanning electron microscopy (SEM) was mostly used in literature as it is sensitive to surface chemistry despite its insensitivity to surface topography. Contact profilometry and white-light interferometry are good at topography characterization but poor at polymer film detection. Optical microscopy was used in this study as it provides the best contrast between transfer film and counterface as illustrated in Fig. 2. The composite used in this study has been known to produce a thin and transparent transfer film that manifests colorful interference fringes under a broadband-light-source equipped optical microscope [20, 21, 23, 25, 26]. Transfer film thickness could be determined from the fringe pattern which is basically a height contour map of the film. The film thickness (e) for a certain color fringe is a function of the interference order, k:

Fig. 2

A semi-quantitative way of transfer film thickness assessment based on white-light interference pattern: a the principle and pattern of white-light interference near the edge of a thin and transparent film, b film thickness estimation of the kth order interference fringe, and c an example of transfer film thickening process observed and quantified using this method. (Color figure online)

$$e=~\frac{{\left( {2k - 1} \right)}}{{4n}}\lambda ,\quad k=1,2,3 \ldots$$
(2)
Here, n is the diffraction index of the film; \(\lambda\) is the wavelength of the fringe color (e.g., \({\lambda }_{\text{red}}=650 \text{nm}\)). The diffraction index of PTFE (n = 1.36) was used here as the transfer film is mainly composed of PTFE [20, 25, 28]. Due to the broadband nature of the halogen light source used in this study, the interference fringe color oscillates from violet to red from the edge to center of the film as the film thickness increases. This is best illustrated in Fig. 2a, b. Near the edge of the film, transfer film appears dark due to the half wave loss when the light was reflected off of the top surface that caused the 0th order dark fringe. Transfer film’s peak thickness, tp, defined here as the maximum film thickness in the image view-field is calculated using the highest order interference fringe discernable in the film area. A simple calculation suggests two full spectrums of colorful fringes should be succinctly visible before neighboring fringes overlap. In practice, a highest 4th order red fringe was visible in all cases, likely because red light is the strongest component in the halogen light spectrum. The lower and upper limit of transfer film thickness detection in this study is therefore:
$${e_{(1{\text{th}}\;{\text{violet}})}}=\frac{{\left( {2 \times 1 - 1} \right)}}{{4 \times 1.36}}~ \times ~{\lambda _{{\text{violet}}}}=\frac{{\left( {2 \times 1 - 1} \right)}}{{4 \times 1.36}} \times 400\;{\text{nm}}=74\;{\text{nm}}$$
(3)
$${e_{(4{\text{th}}\;{\text{red}})}}=\frac{{\left( {2 \times 4 - 1} \right)}}{{4 \times 1.36}} \times {\lambda _{{\text{red}}}}=\frac{{\left( {2 \times 4 - 1} \right)}}{{4 \times 1.36}} \times 650\;{\text{nm}}=836\;{\text{nm}}~$$
(4)

This is well illustrated in Fig. 2b where the maximum film thickness reaches the measuring limit and higher orders of interference fringe became indiscernible. Figure 3c shows an example of the thinnest transfer film discernable using this method (tp ~ 74 nm) and a film thickening process observed and quantified using this method.

Fig. 3

a Polymer wear volume plotted against sliding distance. Low wear sliding was reached after the first 180 m and differential wear rate of the composite stabilized at 7 × 10− 7 mm3/Nm at 12 km of sliding distance. b Average friction coefficients were calculated for each set of sliding cycles with the same reciprocating stroke length indicated with gray bars. Error bars correspond to the standard deviations of the friction coefficient per set of cycles

Whereas the white regions between the island-like domains in Fig. 2 appear to be film-free, there likely exist ultra-thin transfers of aligned polymer chains in these regions beyond the detection limit of the current microscopy. Makinson and Tabor first reported that PTFE always draws fibrous transfers of about 10–40 nm thickness during sliding [36]. Harris et al. [22] provided a physical framework for PTFE chain scission and transfer which suggests a single PTFE chain can get broken down into a few nanometers long filaments and transferred to the sliding counterface from pure van der Walls attractions. However, such ultra-thin films are weakly correlated with wear mode of the polymer [20, 37, 38]. A typical example is the unfilled PTFE which is a standard high wear polymer and generates a thin and continuous film underneath the usually reported thick and patchy transfer film [37]. Based on this reasoning and the accumulation of hard fillers at the bulk’s running surface [15], we hypothesize that such film neither provides wear protection for the bulk composite nor the counterface. For this reason, only optically detectable and likely debris-formed regions of transfer film as shown in Fig. 2 are included in transfer film analysis in this study. A pixel intensity thresholding technique following Bhimaraj et al. [39] and Ye et al. [21] using Photoshop™ was used to convert the image into a binary map (e.g., black and white image insets in Fig. 1d) from which the transfer film area fraction, X, defined as the ratio between film-covered area and total surface area, was calculated.

2.4 Transfer Film Removal and Surface Profilometry Measurements

The primary aim of this study is to measure the evolution of transfer film–substrate interface throughout the course of a wear test. To expose the hidden interface, transfer films at different stages of development were slid against a 6-mm diameter fluoroelastomer rubber ball (E = 4.5–10.5 MPa) across the wear track width as shown in Fig. 1b. Experiments were conducted using a custom-built microtribometer equipped with in situ contact imaging capability (see Appendix for details). The sliding conditions were 3 N normal load, 2.65 MPa contact pressure, 10 mm/s sliding speed, and 10 mm reciprocating length. This method is based on a previous finding that these films adhere extremely to rubber even at very low contact pressures (e.g., 0.7 MPa) and shear stresses (e.g., 0.6 MPa), and can be easily removed by sliding against rubber, the details of which could be found in Ref. [21] and the Appendix. All microtribometry tests were run for 200–1000 cycles to ensure complete removal of the transfer film and an example result was shown in Fig. 1b. Exposed counterface areas were cleaned with a lint-free wipe.

Surface profilometry measurements were conducted using a 3D laser scanning confocal microscope with automated image stitching capability (Keyence VK-X100). Original and transfer film-erased counterface topography of 3 × 12 mm areas were measured at eight predetermined locations in the center of each stripe segment. Laser-engraved fiducial marks were used to align the two images in the microscope analysis software (VK-Analyzer™) and line scans of the original and film-erased counterface were plotted and compared as shown in Fig. 1e. Each line in Fig. 1e is the average of 15 scans separated by 0.1-mm spacing. Regions outside the wear track were used to match the two profiles and the average height of the original bare counterface was set as the zero height. The gray cross-sectional area, Α, was used to calculate the abrasion volume of counterface, Vsteel, by sweeping it through the wear track length. Average counterface abrasion depth, δ, was calculated as the quotient of the abraded area and wear track width (6.4 mm). Based on the stripe test configuration, equivalent Archard wear rate of the counterface could be computed and had the form of
$$k=\frac{{\Delta {V_{{\text{steel}}}}}}{{{F_{\text{N}}} \cdot \Delta d}}=\frac{{\Delta A \cdot l}}{{{F_{\text{N}}} \cdot 2l \cdot \Delta N}}\left[ {\frac{{{\text{m}}{{\text{m}}^{\text{3}}}}}{{{\text{Nm}}}}} \right]$$
(5)
where FN is the normal load; l is the stripe segment length (10 mm), and N is the total sliding cycles. In situ wear rate of the counterface was calculated as the slope of the wear measurements at the point of interest, one point earlier and one point after. Wear rate uncertainty was determined using Monte-Carlo method following Burris and Sawyer [34].

3 Results

3.1 Wear and Frictional Behavior of the Composite

During sliding, the composite material used in this study is known to undergo a high wear run-in period before transitioning into low wear steady-state sliding. After the first 180 m, the composite pin lost 0.37 mm3 of material, whereas it barely lost any material between 180 and 340 m as illustrated in Fig. 3a. This abrupt transition is related to the formation of a protective transfer film during the run-in and was first reported in detail by Ye et al. for this material [20]. Previous measurements found that the exact low wear transition point occurs at about 10–20 m of sliding distance under the same sliding speed and contact pressure [20]. For this reason, in situ wear rate at D = 180 m was computed as the slope of the wear data between 180 and 340 m using the pre-described method and had a value of 1.43 ± 6.35 × 10− 7 mm3/Nm as shown in Fig. 5a. This is the only case where the linear regression analysis window is not centered at the point of interest since it is inappropriate to use run-in data to calculate wear rate after the transition point. Wear of the composite increased gradually after the transition point and stabilized at a wear rate of 7 × 10− 7 mm3/Nm as shown in Fig. 5a, a value slightly higher than reported wear rates of the system (k ~ 1–3 × 10− 7 mm3/Nm) with fixed stroke length experiments [5, 9, 10, 13, 25, 26].

Fig. 4

a Representative 2D profilometry measurements of transfer film–substrate interface across wear track width after 1, 10, 100, and 200 k sliding cycles. Each line represents the average of 15 scans separated by 0.1 mm (total width of 1.5 mm) using 3D profilometry data. Original counterface profiles outside the contact region were used as height reference and were set as zero height. The abraded volumes of the counterface were shown in gray color. b Average interface roughness, Ra, and c abrasion depth, δ, plotted against sliding cycles. Error bars represent one standard deviation from 15 measurements

Fig. 5

a Wear rates of the composite and the counterface plotted against accumulated sliding cycles in log–log scale with associated uncertainties. The control group data of unfilled PTFE are also shown. The control group test was forced to stop at 10 k cycles due to excessive wear of the pin and single point wear rates of the pin and the counterface were reported. b Transfer film area fraction against sliding cycles. Error bars represent one standard deviation from five measurements. c Representative optical images of transfer film morphology at 1, 10, 100, 200 k cycles. Area fraction was calculated using a pixel thresholding technique as described by Bhimaraj et al. [39] and Ye et al. [21]. Images correspond to gray data labels in b

Friction coefficient remained stable throughout the test despite the changes in stroke length and reciprocating frequency. The average friction coefficient was 0.189 ± 0.011 with a 95% confidence interval and the friction coefficient decreased slightly between 1000 and 2000 sliding cycles. This is in accordance with a previous measurement reporting the material had the lowest friction right after the transition point [20].

3.2 Wear and Topographical Evolution of the Film–Substrate Interface

The original lapped counterface had a mean roughness of Ra = 50 nm and Rsk = 0.05. As illustrated in Fig. 4a, after the first 1000 sliding cycles, visible scratches started to scatter across the film–substrate interface resulting in an increased roughness of Ra = 150 nm and decreased skewness of Rsk = − 2.5. The average abrasion depth was δ = 0.1 µm. At 10 k cycles, the interface was ubiquitously abraded across the track width and average abrasion depth increased to δ = 0.3 µm. Roughness increased to Ra = 300 nm and skewness increased to Rsk = − 1.2. Counterface abrasion increased with increased sliding cycles, and both roughness and abrasion depth increased with increased sliding cycles as shown in Fig. 4b, c, respectively. At 200 k cycles, the original steel surface was worn down 0.8 µm on average by the sliding pin and had a roughness of Ra = 445 nm and Rsk = − 1.0.

Wear rates of the composite pin and the film–substrate interface were plotted against sliding cycles in Fig. 5a. While the pin’s wear rate increased slowly and gradually during the test, wear rate of the counterface decreased continuously from 3.0 × 10− 7 mm3/Nm at 1 k cycles to 2.6 × 10− 8 mm3/Nm at 200 k cycles. The monotonic trend of counterface wear rate inversely resembled the evolution of transfer film area fraction as shown in Fig. 5b. At 1 k cycles, the transfer film was composed of sparsely distributed 10–30 µm diameter islands of adherent transfer with a total film area fraction of 33%. At 10 k cycles, the islands grew and merged with each other resulting in an increased area fraction of 71%. Further sliding continued to increase film coverage which saturated at ~ 80% as best illustrated in Fig. 5c. Single point wear rates of an unfilled PTFE pin and the corresponding steel counterface were 4.94 ± 0.24 × 10− 4 and 7.47 ± 6.40 × 10− 9 mm3/Nm, respectively, as shown in Fig. 5a. Notice the counterface wear rate against the composite is only 3 × higher than against pure PTFE during steady-state sliding.

4 Discussion

The slightly higher than expected wear rate of the bulk composite (7 × 10− 7 mm3/Nm) requires a few comments. The same composite pin used in this study was tested by the authors under the same sliding conditions but with a fixed stroke length (e.g., [31, 32]) and had wear rates of 1.5–3 × 10− 7 mm3/Nm, values typical to ultralow wear alumina PTFE as reported in many comparable studies [9, 22, 25, 26, 27, 28, 29, 33]. In two separate works, Harris et al. and their group reported a 0.8 × 10− 7 mm3/Nm wear rate [29] and a 4 × 10− 7 mm3/Nm wear rate [22] for the same alumina PTFE composite both run for a million cycles under identical load, speed, and roughness conditions. The first study used a fixed 25.4-mm stroke length while the second study used a similar ‘stripe’ test configuration and the stroke length decreased intermittently from 88.9 to 27.9 mm. This wear rate discrepancy was never explained but was confirmed by data in this work and two previous publications [31, 32]. We believe stroke length and reciprocating frequency have direct effects on the wear of alumina PTFE system which are beyond the scope of this work.

In the work by Harris et al. [22], the composite wear rate reached a lowest value of 2 × 10− 7 mm3/Nm at 100 k cycles and increased to 4 × 10− 7 mm3/Nm at 1 M cycles. In this study, the composite reached a lowest wear rate of 1.43 × 10− 7 mm3/Nm at 1 k cycles and the wear rate increased to 7 × 10− 7 mm3/Nm at 200 k cycles. Confounding factors for the wear rate discrepancy between this work and Harris et al.’s include the differences in alumina filler, dispersion and sample preparation, etc. However, transfer film produced in this study were visually indistinguishable from those reported in [20, 21, 23, 31, 32] and extremely persistent during sliding. Contact profilometry was used to acquire the topography of the transfer film at 200 k cycles which is compared with the original surface and the film-erased topography in Fig. 6. Notice the transfer film has more volume above the original counterface than below it. This has some interesting implications for transfer film thickness evaluation in tribology literature. However, this is also beyond the scope of this work and will require a standalone study.

Fig. 6

Comparison of the transfer film, original counterface, and film-erased counterface topography at 200 k cycles. All scans were perpendicular to the wear track’s sliding direction and each line represents the average of 15 scans separated by 0.1 mm (total width of 1.5 mm). Transfer film topography was measured with a Taylor Hobson Talysurf-6 contact profilometer (2 µm tip radius, 1 mN contact force). (Color figure online)

To the authors’ best knowledge, this study is the first to measure the topographical evolution of the transfer film–substrate interface in a tribological polymer system. The primary aim of this study is to test the hypothesis that strong adhesion between transfer film and substrate protects the interface from wear and eliminates counterface abrasion locally. For this reason, a well-studied alumina PTFE system was chosen for its extremely low wear and persistent transfer film. Blanchet et al. [40] suggested that low wear PTFE composite transfer films developed by gradually filling in remaining areas of exposed counterface. Ye et al. [20] used in situ optical microscopy to show that transfer film formed by filling in such ‘free-space’ and thickening simultaneously. Optical microscopy measurements in this study confirmed a gradual filling-in process of transfer film development as illustrated in Fig. 5c. However, one evidence against the permanent adhesion theory is the 2D line scans in Fig. 4 showing abrasion of the counterface ubiquitously across the surface in spite of the high transfer film area fraction. A quick way to test the hypothesis is to compare the interface profiles between two different sliding cycles and calculate the range of incremental wear as illustrated in Fig. 7. Here, X1 and X2 are transfer film area fractions at sliding cycle N1 and N2 (N1 < N2), respectively. Mean counterface adhesive wear rate at film uncovered regions is denoted by ka and fatigue wear rate at the film–substrate interface by kf. Adhesive wear of the counterface causes increase in relative depth of film uncovered regions as illustrated in Fig. 7 (d2 > d1). The interface profile at cycle N1 is subtracted from the profile at N2 cycle, and Δ denotes regions where profile 2 is lower than profile 1. Accepting the permanent adhesion theory means kf = 0, counterface wear only occurs at film-free regions, and the subtraction data should be negative at such locations. More importantly, the area fraction of regions with negative differences over the whole track width, \(\sum \varDelta \%\), should be no greater than the film-free region percentage 1 − X1, assuming X1 ≤ X2. The result of a quick check using average 2D line scans from 10 and 100 k sliding cycles is shown in Fig. 8. Within the wear track, the area fraction of further counterface abrasion is 0.74 ± 0.25 with 90% confidence using the student t-statistics. This is much higher than the 0.29 ± 0.14 area fraction of film-free region at 10 k cycles, which suggests that film–substrate interface fatigue wear contributes significantly to the total counterface wear.

Fig. 7

Analysis of counterface wear mechanisms and a simple way to test the existence of fatigue counterface wear at the transfer film–substrate interface. Interface topography after N1 cycles was subtracted from topography after N2 cycles (N1 < N2). Areas of negative difference correspond to incremental wear of the counterface. Assuming counterface wear occurs at film-free regions only, area fraction of incremental wear should be no greater than the area fraction of film-free regions at N1 cycles. The validity of this method is also insensitive to the difference of film distribution between the two topographies

Fig. 8

a Comparison between transfer film–substrate interface profiles at 10 and 100 k cycles. Each line is the average of 15 scans and the two profiles are aligned at regions outside the wear track. b Average height difference profile showing continuous counterface wear from 10 to 100 k cycles. Mean deviation between the two profiles outside wear track were considered and a student t test (p = 0.1) is used to calculate the percentage of regions with statistically negative difference over the whole track width

Another strong evidence suggesting the existence of counterface fatigue wear in film-covered regions is a unique pitted morphology of the film-erased counterface as best illustrated in Fig. 9. After 1 k cycles, only minor scratch marks along the sliding direction were visible at the counterface. After 50 k cycles, rows of wear pits along the sliding direction with average depth and length of 1 and 200 µm were discernable. After 200 k cycles, average pit depth remained unchanged while average pit length increased to ~ 700 µm. We believe that failure of the film–substrate interface scavenges material from the counterface bit by bit which causes the increase of pit length. The fact that average wear pit depth did not change between 50 and 200 k cycles in spite of the increased average abrasion depth from 0.48 to 0.82 µm suggests that wear occurred universally across the surface regardless of the transfer film coverage conditions.

Fig. 9

Representative transfer film-erased counterface topography and 2D line scans at a 1 k, b 50 k, and c 200 k sliding cycles showing characteristic wear pits caused by extended sliding against the composite pin. Line scans correspond to red-dashed lines in the images. Pit length is defined as the size of the wear pit along the sliding direction. All images have the same sliding direction and scale bar indicated

The above analysis strongly suggests a combined route of counterface wear: a high adhesive wear rate at film-free regions and a low fatigue wear rate at film-covered regions. Early works by Rabinowicz [41] and Greenwood and Tabor [42] first confirmed the wear of hard metallic counterface when slid against soft polymers under dry sliding conditions due to strong adhesive junctions at asperity level. Failure of a tribological coating–substrate interface under cyclic shear stress is not uncommon and often leads to delamination event without appreciable wear of the coating itself [43, 44, 45]. The transfer film in this study has been proved extremely wear resistant (k ~ 10− 9 mm3/Nm) in the native contact [21]. And yet, results suggest that film–substrate interface might still be the weak link in the tribo-system. To gain some basic understanding of the relative scale of each wear component, we can use a simple rule-of-mixtures model to describe the system in which the total wear rate of the counterface, ksteel, could be written as
$${k_{{\text{steel}}}}=\left( {1 - {\rm X}} \right) \cdot {k_{\text{a}}}+X \cdot {k_{\text{f}}}=\left( {{k_{\text{f}}} - {k_{\text{a}}}} \right) \cdot X+{k_{\text{a}}}$$
(6)

assuming uniform pressure across the contact. Here, X is the transfer film area fraction, ka the counterface adhesive wear rate at film-free regions, and kf is the counterface fatigue wear rate at film-covered regions as illustrated in Fig. 7. From Eq. 6, the plot of ksteel versus X should be a straight line connecting (0, ka) and (1, kf). Wear rates of the film–substrate interface were plotted against transfer film area fractions in Fig. 10. The best-fit using the least square method was shown as the red solid line and the two have a strong linear correlation (R2 = 0.92). The rule-of-mixtures model was plotted as the black dashed line in which the adhesive wear rate constant, ka = 5 × 10− 7 mm3/Nm, was determined using the y-axis intercept of the red solid line. Interestingly, this value is only one order of magnitude lower than wear rates of soft radioactive metals (Cu, Zn, Ag) against PTFE (~ 3 × 10− 6 mm3/Nm) as measured by Rabinowicz and Shooter [46], whereas they used a much higher contact pressure (~ 100 MPa) than in this study. The fatigue wear rate constant, kf = 2.6 × 10− 8 mm3/Nm, was determined using the minimum value of counterface wear rate measured. This is our best guess of kf, since ka is 10 × greater than kf which makes it extremely unreliable to extrapolate kf from the slope constant in Eq. 6. The gray area in Fig. 10 corresponds to the mean variation between the data and the model.

Fig. 10

Counterface wear rate plotted against transfer film area fraction. The best-fit trendline using the least square method is shown as the red solid line. A rule-of-mixtures model of counterface wear rate is shown as the black-dashed line. The gray area represents the mean variation of the wear data from the model

Fig. 11

Microtribometer used to remove transfer film in this study. Transfer film was removed using adhesive rubber contact as first discovered in Ref. [21]. A Φ6-mm rubber ball was slid against transfer film located on a linear reciprocating stage under the sliding condition of 3 N normal load, 2.65 MPa contact pressure, 10 mm/s sliding speed, and 10 mm stroke length. The contact area was located under a Nikon optical microscope to record the evolution of transfer film removal within a fixed 1 × 1 mm area. Raw images were converted into black (transfer film) and white (film-free area) using a pixel intensity thresholding technique. Film area fraction was calculated as the percentage of black pixels within the whole image view-field and indicated within the binary images

In Fig. 10, there is an increased deviation of the data from the model at the bottom (X > 0.75). This, we believe, could result from two possible factors. First, as transfer film became more continuous, the average size of real microscopic contact area likely increased (Fig. 5c). From basic contact mechanics, the maximum shear stress at the film–substrate interface decreased and caused decreased fatigue wear rate at the interface. Second, for the model, we assume full contact and macroscopically uniform contact pressure at both film-covered and film-free regions such that X and 1 − X also represent the fraction of the total normal load supported on the film-covered and the bare regions, respectively. However, as the transfer film grew more complete and uniform, the residual handful film-free regions might lose contact with the bulk’s running surface during sliding due to their smaller size and thickened transfer film islands. This could also produce lower counterface wear rates than the model predicts.

There are a few important implications of the results in this study. First, contrary to the traditional hypothesis of permanent transfer film adhesion and counterface protection, wear of the counterface persists even in areas covered with a persistent transfer film, at least for the alumina PTFE system. As migration or delamination of the transfer film islands were rare events for this system [20], we believe fatigue failures of the film–substrate interface mostly occur within a close range (1–100 nm) to the interface underneath the apparently persistent transfer film. The mechanism of such fatigue wear might involve complex micro-flows of energy and matter near the interface and require further studies. Second, transfer film coverage correlates more strongly with wear of the counterface than bulk polymer. Previous studies have suggested that polymer wear is also strongly governed by the structure and composition of the running film at the pin’s sliding surface [26]. Third, the counterface wear data shed some light into tribological designs where wear of the hard-metallic component than the solid lubricants is of greater concern. In such cases, interfacial fatigue wear rate sets the lower limit of the metal wear rate and could be a strong function of transfer film coverage, distribution, and adhesion strength.

5 Conclusion

A well-studied alumina-PTFE solid lubricant was known to produce extremely adherent and complete transfer films when slid against hard, metallic counterfaces. After removing the transfer film and exposing the substrate underneath, it was found that wear of the counterface occurred continuously throughout the course of the wear test even after a persistent transfer film was formed on top of the counterface. Counterface wear rate decreased from 3 × 10− 7 mm3/Nm at 1 k cycles to 3 × 10− 8 mm3/Nm at 200 k cycles and had a strong linear relation with transfer film area fraction (R2 = 0.91). A rule-of-mixtures model was proposed to predict wear rate of the counterface as a linear function of film area fraction. The model agreed with experimental results closely and strongly suggested the coexistence of adhesive and fatigue wear mechanisms of the counterface under dry sliding condition.

Notes

Acknowledgements

The authors gratefully acknowledge financial support from the National Natural Science Foundation of China (51505117 and 11472096), the Natural Science Foundation of Anhui Province (1608085QE98) and the Fundamental Research Funds for the Central Universities.

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Copyright information

© Springer Science+Business Media, LLC, part of Springer Nature 2018

Authors and Affiliations

  • Jiaxin Ye
    • 1
  • Wei Sun
    • 1
  • Yan Zhang
    • 1
  • Xiaojun Liu
    • 1
  • Kun Liu
    • 1
  1. 1.Institute of Tribology, School of Mechanical EngineeringHefei University of TechnologyHefeiChina

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