Effect of synovial fluid temperature on wear resistance of different polymer acetabular materials

Abstract

In order to investigate the effect of frictional heat on the wear resistance characteristics of polymeric acetabular materials, the tribological tests and wear numerical analysis of three common polymer acetabular materials were carried out under different synovial fluid temperatures. The study results show that XLPE and VE-XLPE exhibit superior wear resistance compared to UHMWPE in high-temperature, heavy load environments. The coefficient of friction of three materials gradually decreases as the temperature of the synovial fluid increases. The wear depth and wear volume of the three materials increased with the increase of the temperature of the synovial fluid, and the forms of wear at 46°C and 55°C were mainly adhesive wear and plastic deformation. The higher temperature of the synovial fluid accelerates the oxidative degradation of the material surface and generates oxidation functional groups, which leads to the breakage of C-C bonds in the surface molecular chains under the sliding shear effect, thus reducing the mechanical properties of the material. Specifically, the surface of the polymer material will soften at a higher ambient temperature, mainly due to the decrease of hardness, and then deteriorate in the friction property, and finally increase the wear rate. Ansys results showed that the volume wear of the three materials increased with the increase of synovial fluid temperature, and the trend could be approximately linear. Numerical calculations predict that VE-XLPE has the highest wear of 0.693 mm³ among the three materials at 37°C, followed by XLPE at 0.568 mm³ and UHMWPE with the lowest wear of 0.478 mm³. At higher synovial fluid temperatures (46°C, 55°C), VE-XLPE still has the largest wear volume among the three materials, while XLPE and UHMWPE have similar wear. The wear cloud pictures showed that the maximum wear volume occurred near the edge of the acetabulum.

Keywords

Polymer acetabular materials hip joint prosthesis synovial fluid temperature frictional heat wear model

Introduction

Arthroplasty is an effective treatment for advanced bone and joint disease, and total hip arthroplasty (THA), for example, has been increasing in demand for the procedure at a high rate in recent years, with a reported 71% increase in initial THA expected in the United States by 2030.^1,2 However, cases of prosthesis failure due to in vivo loosening after THA are common,³ and current studies suggest that wear particles from the relative motion of the femoral head to the acetabulum are the main cause of loosening failure.^4–6

Currently, artificial acetabulums are usually made of ultra-high molecular weight polyethylene (UHMWPE) materials, but with the continuous development of materials in recent years, highly cross-linked polyethylene (XLPE) has been produced and put into clinical application in order to improve the wear resistance of UHMWPE, and at the same time, high cross-linked polyethylene doped with vitamin E (VE-XLPE) has gradually gained widespread attention in order to improve the problem of insufficient oxidation resistance of XLPE.^7,8 Taddei et al.⁹ compared the wear resistance of three polymers, UHMWPE, XLPE, and VE-XLPE, using a hip simulator, and found that VE-XLPE acetabulums exhibited better wear resistance properties only under long-term or more severe test conditions. In a study by Chen et al.,¹⁰ it was found that VE-XLPE has better wear resistance than XLPE and has stronger oxidation resistance properties, which reduces the probability of surface fatigue crack generation. Grupp et al.¹¹ found that the strength of oxidation resistance of the three materials by artificial aging tests was: VE-XLPE > XLPE > UHMWPE.

In order to prolong the service life of hip prostheses and reduce the probability of postoperative revision due to prosthesis failure, it is important to conduct research on improving the wear characteristics of acetabular prostheses. However, the in vivo environment in which the acetabulum serves is extremely complex, and there are many factors affecting the frictional properties of the acetabular prosthesis, such as the weight load of different patients,¹² the surface microstructure of the acetabular prosthesis,¹³ and the tissue composition of the synovial fluid.¹⁴ Among them, the effect of temperature rise of the prosthesis and surrounding synovial fluid¹⁵ on the wear resistance of the prosthesis, especially the polymer acetabulum, due to the frictional heat generated by the relative femoral-acetabular motion, has not received sufficient attention. Temperature rise caused by friction heat may lead to multiple failure modes of prosthesis, such as fretting wear and creep. Few studies have been conducted on the fretting wear properties of polymeric materials, which are prone to wear damage under high loads even at very small displacements because of their flexibility.¹⁶ In a study by shen et al.,¹⁷ it was found that the temperature rise on the surface of PEEK polymers due to frictional heating under fretting wear conditions was very small. A study by Wang et al.¹⁸ found that both frictional corrosion on the surface of the artificial hip joint and fretting corrosion at the head-neck joint released chromium- and phosphate-rich particles, which led to inflammation of the osteoarticular tissue and artificial joint failure. In addition, creep is also a cause of premature failure of UHMWPE acetabular cup liners,¹⁹ and current research on the creep properties of polymer acetabular liners is less, and there are many factors that affect the rate and magnitude of creep strain, like molecular weight, crystallinity, crosslink density, and additive concentration, etc. Takahashi et al.²⁰ experimentally investigated the effect of size and thickness on the effect of size and thickness on the compressive creep strain rate in UHMWPE and VE-XLPE acetabular liners was investigated experimentally, and it was found that the creep strain rate decreased with increasing the inner diameter and thickness.

A related study found that the in vivo hip prosthesis femoral head temperature could reach 43°C after one hour of walking,²¹ and there was a significant increase in the synovial fluid temperature near the hip prosthesis after 1 hour of walking, with a temperature rise of up to 7°C in the ceramic ball on the synovial fluid near the polyethylene hip prosthesis.²² When polymeric acetabular materials are used for a long time in a high temperature environment, they accelerate their own oxidation process,²³ which affects their wear resistance. Saikko et al.²⁴ investigated the effects of lubricant serum type and synovial fluid temperature on the frictional behavior of VE-XLPE under mating with CoCr and found that the average wear coefficient of VE-XLPE decreases significantly as the synovial fluid temperature increases from low to 37°C. In addition, it has been reported that the temperature rise of the synovial fluid near the prosthesis can adversely affect both the wear characteristics and the lubricating properties of the synovial fluid, such as a decrease in the density and viscosity of the synovial fluid with increasing temperature,²⁵ as well as a change in the lubricating properties caused by the denaturation of proteins in the synovial fluid at higher synovial fluid temperatures.^26,27 At the same time, a research result has shown that UHMWPE materials in the synovial fluid at higher temperatures are more susceptible to be induced and undergo oxidative degradation,²⁸ which leads to deterioration of wear resistance. However, there is no systematic evaluation on the changes of wear resistance of the three polymer materials under the influence of different synovial fluid temperatures, and there is a lack of research on the mechanism of friction heat affecting the service life of polymer acetabular prosthesis.

Therefore, in this paper, with reference to the peak contact interface or synovial fluid temperature of joint prostheses obtained in related studies (as shown in Table 1), combined with equivalent load conversion, friction tests were carried out regarding three polymer acetabular prostheses at different synovial fluid temperatures and different contact loads. By analyzing the friction coefficient, wear rate and wear morphology, the mechanism of the frictional thermal effect on the wear resistance of polymer acetabular materials was revealed. In order to further elucidate the thermogenic wear mechanism of the femoral head-acetabular contact interface during the long service life of the joint prosthesis in vivo, a wear model of the acetabular prosthesis material under the influence of frictional heat was constructed using Ansys. The results of this thesis provide important theoretical support for the in vivo failure analysis of joint prostheses and the design of new polymeric acetabular materials. The research route is shown in Figure 1.

Table 1.

Studies on temperature rise due to frictional heat at the hip joint interface.

References	Research materials and areas	Peak temperature
Bergmann²¹	Polyethylene cups	43.1°C
Pritchett²²	Synovial fluid	47°C
Roussignol²³	Polyethylene cups	52.65°C
Lu²⁹	Polyethylene cups	51.3 ± 5.2°C
Bergmann³⁰	Synovial fluid	46°C
Rocchi³¹	CoCrMo femoral head	51°C

Materials and methods

Sample preparation

Ultra-high molecular weight polyethylene, XLPE and VE-XLPE rods (Shanghai Mitsubishi Chemical High-Tech Materials Co., Ltd.) are processed into discs of φ25 mm × 6 mm. The surface of the disc sample was polished by using sandpaper to 2000#, and then polished by using polishing machine (MP-1B, Shaoxing Jingbo Testing Instrument Company) after completion, and the surface roughness was less than 0.8 μm. The upper sample for conducting the friction test is a ZrO₂ ceramic ball (Shanghai Gongtao Ceramics Co., Ltd.) with a diameter of φ6.35 mm and a surface roughness of less than 0.01 μm. The simulated synovial fluid was prepared using neonatal bovine serum (Hangzhou Four Seasons Company) and deionized water in the ratio of 1:3. Based on the ambient temperature in the human body proposed in the ISO 14242 standard, combined with the temperature at which proteins in the synovial fluid tend to precipitate in large quantities,³² and the highest transient peak temperature at the prosthetic interface measured in the relevant literature (as shown in the study by Lu et al. in Table 1, which was taken as 55°C to control for consistent differences in parameters), the synovial fluid temperatures for this study were selected as 37°C, 46°C, and 55°C, respectively.

Tribological test

A reciprocating friction tester (HRT-02A, Jinan Hengxu Testing Machine Technology Co., Ltd.) was used to conduct the ZrO₂ ball-three polymer acetabular replacement material disk friction test (the schematic diagram is shown in Figure 2). The tribological test focuses on the wear of two body of mating parts, especially the three polymer materials used for the disk specimen. The friction and wear testers we use can be modified with fixtures and stroke adjustments to perform fretting tests, but they cannot be used for creep testing. To equivalently convert the contact load at the femoral head-acetabular interface to the load of the ball-disc friction test, a femoral head of φ14 mm was used as the reference object, and 30% of the surface area of the femoral head³³ was used as the actual bearing area of the acetabulum, and the effective contact area between the femoral head and the acetabulum was calculated to be approximately 740 mm². Based on the peak load of 3000 N for gait motion in the ISO 14242 standard, the load per unit area of the acetabulum is calculated to be about 4.05 N/mm², assuming an initial contact area of 1 mm² for the ball-disk sample and choosing a safety factor of 3,³³ according to equation (1).

F = K \times P \times S

(1)

Where K is the safety factor, according to existing studies,³⁴ the gait activity load of the hip joint is higher in actual human gait activity, so in some current theoretical test load conversions,^35,36 it often needs to be scaled up so that it is more similar to the actual human hip joint load. P is the unit bearing, S is the contact area, and the obtained ball-disk interface load is about 12.15 N. To simplify the calculation, an approximate value of 10 N is taken. Also referring to the study of Shankar et al.,³⁵ 20 N and 30 N were selected as the other two sets of test loads, respectively, considering the larger body weight loads and a series of daily human movements (running, carrying, going up and down stairs, etc.). The samples were ultrasonically cleaned in anhydrous ethanol for 20 min before the test was carried out, then dried for 20 min using a constant temperature drying oven (DZF-1, Beijing Zhongxing Weiye Century Instruments Co., Ltd.), and finally weighed 5 times using an electronic balance (JM-B2002 type, Zhuji Chaoze Weighing Equipment Co., Ltd.) (the average value was taken after rounding off the maximum and minimum values). Friction test parameters are shown in Table 2.

Table 2.

Ball-disc reciprocating friction test parameters.

Test parameters	Numerical value
Friction sub	Ball-disc
Simulated synovial fluid	25% NBS
Synovial fluid temperature	37 °C/46 °C/55°C
Displacement frequency	1 Hz
Applied load	10 N/20 N/30 N
Cycle time	3600 s

After the friction test, the samples were washed in anhydrous ethanol for 20 min and dried and weighed, and the mass wear of the samples was measured five times (the average value was taken after rounding off the maximum and minimum values), followed by observation of the wear marks by a three-dimensional surface profiler (UP-Lambda, Rtec, USA), calculation of the wear volume by the width and depth of the wear marks, and calculation of the average wear rate by equation (2).

K = \frac{V}{S F}

(2)

where V is the wear volume, K is the average wear rate, S is the total sliding travel, and F is the loading load. Scanning electron microscopy (SEM, SU-8010, Hitachi, Japan) was used to observe the wear morphology of UHMWPE discs, and an energy spectrometer (EDS, SU-8010, Hitachi, Japan) was used to perform elemental analysis of the wear surface.

Wear finite element analysis

Through the finite element software ANSYS Workbench, numerical simulation of wear of polymer acetabulum under the influence of different synovial fluid temperatures without using polymer acetabulum was carried out to obtain the wear amount of one gait and the wear amount of walking for 60 s for three different polymer acetabulums, and the wear amount of walking for 1 year was predicted based on the results. In general the use of XLPE liner and VE-XLPE liner is better for conventional acetabular cup shell design.^37,38 It has also been shown that the performance of the combination of polyethylene liner and titanium acetabular cup meets the requirements of use and that the wear marks of both are mainly generated during the assembly and disassembly of the liner.³⁹ In this regard, the study in this paper ignores the difference in performance between the acetabular cup and polymer liner combination, and the acetabular cup and acetabular liner are studied as a whole. The simplified structural parameters of the hip joint are shown in Figure 3(b). According to the clinical application, the diameter of femoral head is 28 mm and the thickness of acetabular cup is 5 mm. The radial clearance is 0.125 mm. The acetabular anteversion is 45°. The ceramic femoral head, which is widely used in clinical applications, was selected as the object of study for the polymer acetabular (COP) prosthesis mate, and the mechanical property parameters of the material are shown in Table 3. In order to simulate the motion of the ceramic ball in the polymer acetabulum, the equivalent replacement of the experimental model with the finite element model is shown in Figure 3(a), and the friction coefficient and wear coefficient obtained from the ball-disc friction test (to consider a larger motion load, 20 N load) are equivalently substituted into the finite element numerical calculation. The wear model used in this study is the Archard wear model,⁴⁰ whose volume wear can be calculated by equation (3). In order to use the Archard wear model under curved surface conditions, the contact surface detection method was set to “Nodal Normal to Target” and the “Adjust to touch” option for the curved surface was enabled before running the FEM simulation.

d W = \frac{K}{H} p^{m} v^{n}

(3)

Where K is the hardness in MPa, K is the wear coefficient, P is the contact pressure, V is the sliding speed, m is the contact pressure exponent, and n is the sliding speed exponent.

Table 3.

Hip joint material properties.

Components	Materials	Density ρ(kg/m³)	Elastic modulus E (GPa)	Poisson’s ratio υ
Femoral head	ZrO₂	6050	210	0.3
	UHWMPE	935	0.8	0.4
Acetabular	XLPE	935	1.0	0.4
Acetabular	VE-XLPE	963	0.9	0.4

To improve the computational efficiency, the automatic meshing method was used for the acetabulum and the multi-area meshing method was used for the femoral head, with the mesh size set to 2 mm, and the number of mesh elements obtained was 27,548 and the number of nodes was 43,601. The contact subsets were set up using the CONTA174 3D contact unit, and the contact type between the acetabulum and femoral head was set to friction, and the friction coefficient was averaged from the reciprocal sliding friction test data under a 20 N load. Referring to the gait load in ISO14242 standard, the displacement constraint and load constraint were applied to the acetabular and femoral head models as shown in the curve in Figure 3(c). We refer to the ISO 14,242 standard and use dynamic forces as the dynamic load input for the finite element simulation. The time of a complete gait is set to 1 s, and a complete gait cycle is divided into 5 time nodes in detail, the maximum peak load is 3000 N. The inclination angle of the acetabulum to the horizontal plane was set to 45° to simulate the human standing posture, and the back of the acetabulum was also set to full constraint to simulate the close contact between it and the pelvis.

Results and Discussion

Analysis of friction coefficient results

Figure 4 shows the variation curves of friction coefficients of ZrO₂ ball-UHMWPE discs at different temperatures of the synovial fluid and under different loads. It is found that the friction coefficient of ZrO₂ ball-UHMWPE discs gradually decreases with the increase of the synovial fluid temperature under 10N and 30N load, which is due to the protein precipitated by the heat of the synovial fluid adhering to the surface of UHMWPE discs to form a protective film to achieve the friction reduction effect, and the higher the temperature, the more precipitated. With the increase of applied load, the friction coefficients at different synovial fluid temperatures are reduced, which is related to the characteristics of UHMWPE plastic material, with the increase of load, the instantaneous contact area increases, leading to the reduction of friction coefficient during the friction process.

Figure 5 shows the variation curves of friction coefficients of ZrO₂ ball-XLPE discs at different synovial fluid temperatures and under different loads. It is found that the friction coefficient of ZrO₂ ball-XLPE discs gradually decreases with the increase of the synovial fluid temperature under 20 N and 30 N load, which is due to the protein precipitated by the heat of the synovial fluid adhering to the XLPE surface to form a protective film to achieve the friction reduction effect, but the friction coefficients at 46°C and 55°C gradually converge with the change of time and remain around 0.1. In addition, it was noticed that the friction coefficient was gradually becoming larger at 10 N and 37°C. This may be due to the fact that the protein components precipitated by the synovial fluid as the third body intensified the friction of the contact surface, and there were fewer protein molecules in the synovial fluid at low temperature, and the protein protective film had not been formed yet, while the contact area was small at low load, and the protein molecules were more easily involved in the whole friction process in the form of third body particles.

Figure 6 shows the variation curves of friction coefficient of ZrO₂ ball-VE-XLPE disc at different synovial fluid temperatures and different loads. It is found that the friction coefficient of ZrO₂ ball-VE-XLPE disc gradually decreases with the increase of the temperature of the synovial fluid under three loads, and the greater the load, the same trend of friction coefficient changes at 46°C and 55°C and the values are extremely close, indicating that VE-XLPE shows more excellent wear resistance under severe environment (high temperature and heavy load). However, at 20 N and 37°C, the friction coefficient is gradually becoming larger, guessing that there are fewer protein molecules in the synovial fluid at low temperature, and the protein protective film has not yet been formed, and the VE-XLPE abrasive chips mixed with protein as the third body intensifies the friction.

Figure 7 shows the change curve of the average coefficient of friction, and the results can be found that the average coefficient of friction under the three polymer material mating pairs changes as the temperature of the synovial fluid increases, among which, the average coefficient of friction under the three loads of XLPE material decreases with the increase of the temperature of the synovial fluid, and similar conclusions can be obtained for VE-XLPE and UHMWPE materials under the loads of 20 N and 30 N. In this regard, it is important to consider that the various components of synovial fluid have an important influence on joint wear,^41–43 for example, there are many reports of improved lubrication of the hip friction interface due to the large aggregation of proteins in synovial fluid caused by increased temperature.⁴⁴ This is because when the temperature increases, the protein molecules precipitated in the synovial fluid will aggregate under the action of sliding shear and change from fluid-like to gel-like,⁴⁵ forming a protective film with the substrate material and acting as a friction-reducing agent.

The average coefficient of friction under ZrO₂ ball-VE-XLPE disc and ZrO₂ ball-XLPE disc pairings were similar, which is similar to the results of Affatato et al.,^46,47 adding VE was not effective in improving the tribological properties of UHMWPE, but the average coefficient of friction values under ZrO₂ ball-UHMWPE disc pairings were the smallest, which may be due to the UHMWPE is more likely to adsorb the protein molecules precipitated from the synovial fluid. Under a large load of 30 N, the smallest average coefficient of friction can be found for all three materials, with the largest average coefficient of friction for UHMWPE, because when the applied load becomes large, the elastic force exceeds the adhesion force and the contact bonds on the surface break, leading to a decrease in the coefficient of friction.⁴⁸ In addition, the proteins precipitated from the synovial fluid at higher temperatures, besides improving the wear at the joint contact interface, sometimes act as a third body to aggravate friction (especially at low loads). Meanwhile, under light load, UHMWPE exhibits more wear resistance compared to VE-XLPE and XLPE, but at higher synovial fluid temperatures (46°C, 55°C), the wear characteristics of VE-XLPE and XLPE vary less by synovial fluid temperature, and the friction coefficients are almost the same, exhibiting a stronger resistance to aging. The mechanism of the effect of synovial fluid temperature on the friction process of three polymeric acetabular materials is shown in Figure 8. When the temperature of the synovial fluid increases due to friction heat, the protein accumulation and adhesion as well as hyaluronic acid (HA) precipitation appear in the friction area of the sample surface, and the friction coefficient is easily affected by the synovial fluid state. The frictional heat generated during the friction process will further aggravate the softening effect of the polymer material, and the wear and oxidation of the surface will increase at the higher temperature of the synovial fluid. As the possible mechanism of chemical degradation reaction shown in Figure 8, (a) part of the molecular chain of the surface layer breaks (like C-C bond and C-H bond) under the cyclic frictional shear action and thermal oxidation of the higher temperature synovial fluid, and the free radical chain formed on the surface of the polymer material combines with oxygen in the synovial fluid to form peroxide radicals. At higher temperatures in synovial fluid, peroxide radicals continue to react to form ketone and alcohol, which negatively affect the mechanical properties of the material.⁴⁹ With the continuous cyclic shearing action and the increase of oxidation, it also makes some of the flexible molecular chains on the surface of the material untwist and even break away from the surface of the substrate, which eventually increases the wear.

Analysis of wear morphology and wear mechanism results

Taking the friction specimen under 20 N load as an example, Figure 9 shows the three-dimensional morphology of the three polymer material mating pairs after the friction test at different synovial fluid temperatures (37°C, 46°C and 55°C, respectively). With the increase of the synovial fluid temperature, the wear depth under all three mating pairs increased, among which the wear depth of VE-XLPE could reach 60.1 μm at a sliding fluid temperature of 55°C, and the wear depth of XLPE was also 53.2 μm. This illustrates that the higher temperature of the synovial fluid affects the surface layer of the three polymer disc materials, and the adsorption and lubrication of the synovial fluid on the friction interface is reduced, thus intensifying frictional wear,⁵⁰ while accelerating the aging and oxidative degradation of the polymer material surface and reducing the mechanical properties of the material, as evidenced by the deepening of the wear depth with the increase of the synovial fluid temperature, which also indirectly affects the wear resistance of the material surface and leads to an increase in the wear rate. Observing Figure 9(a)-(i), it is found that the wear depth in the Z-direction of each material increases with the increase of the sliding fluid temperature, which is because the higher sliding fluid temperature affects the mechanical properties of the surface layer of the material, thus intensifying the surface wear. Among them, VE-XLPE has the largest wear depth and XLPE has the smallest wear depth, while the wear resistance of UHMWPE is less affected by the synovial fluid temperature, which is similar to the findings of Harsha et al.,⁵¹ who concluded that the frictional wear of XLPE may be more severe under unlubricated, highly abrasive conditions compared to UHMWPE, and therefore the corrosive wear mechanism under higher temperature synovial fluid lubrication should be further explored.

Figure 10 shows the 3D morphological data after the friction test. Observing Figure 10(a), it is found that under the influence of synovial fluid temperature, the average wear rate values under ZrO₂ ball-UHMWPE disc mating are smaller compared to VE-XLPE and XLPE, and the maximum wear rate under ZrO₂ ball-XLPE disc mating occurs at 55°C and 30 N (at the maximum load and maximum synovial fluid temperature conditions), which is due to the fact that under the effect of higher temperature synovial fluid and transient frictional heat, the wear zone easily reaches the thermal deformation temperature of the polymer material, thus accelerating the rate of oxidative degradation reaction of the surface layer. Comparing the wear depth volume data illustrated in Figure 10(b) and the wear volume data in Figure 10(c), it can be found that there is little difference in the volume wear volume of the three polymers at light load (10 N), and combined with the three-dimensional morphology after the friction wear test illustrated in Figure 9(a)–(i), the volume wear volume of VE-XLPE is the largest at 20 N load for all three synovial fluid temperatures, 37°C, 0.15 mm³ at 46°C, 0.123 mm³ at 46°C and 0.118 mm³ at 55°C, respectively. The volume wear of XLPE and UHMWPE are similar, which is similar to the findings of Affatato et al.,⁴⁶ who found that the relationship between the magnitude of mass wear after accelerated aging is VE-XLPE > UHMWPE > XLPE, and the effect of adding vitamin E on improving the wear characteristics of XLPE is not significant, and the results of Su et al.⁵² also mentioned that the wear rate of VE-XLPE is much higher than that of XLPE acetabular.

Figure 11 shows the electron microscopic micrographs of the wear morphology of UHMWPE, XLPE, and VE-XLPE after friction tests under 20 N load and different temperature synovial fluid lubrication. As shown in Figure 11(a)–(c), UHMWPE has a small amount of abrasion marks on the friction surface at 37°C synovial fluid temperature, and when the temperature of the synovial fluid increases, a small amount of debris is observed on the disk surface. The temperature rise of the synovial fluid leads to denaturation of the proteins in the synovial fluid, which can affect its lubricating properties.²⁶ Observation of Figure 11(f) revealed that a thin mucous film with cracks could be clearly observed on the worn surface of XLPE. The proteins precipitated by the heat of the synovial fluid may participate in the friction as a third body, which is similar to the findings of Sava et al. Bovine serum leads to a high friction coefficient because the inability of the protein molecules to bind to the UHMWPE surface exacerbates the friction,⁵³ which validates the scenario that the friction coefficient of VE-XLPE discs becomes progressively larger at 37°C. Comparing Figure 11(c) and (i), it is found that the surface peeling of UHMWPE and VE-XLPE at 55°C synovial fluid temperature is obvious, and the mechanical properties of the surface polymer are degraded due to the oxidative degradation of the surface polymer material, and the wear resistance of the polymer material is deteriorated, the surface peeling of the sample is intensified and the plastic deformation of the surface is deepened, which verifies the previous analysis of the 3D morphological results. On the other hand, protein precipitation may also form a protective film on the surface to reduce friction. High-viscosity protein films and protein debris were found at the abrasion marks, suggesting that protein precipitation from the synovial fluid will aggregate due to sliding shear, and form a gel film with hyaluronic acid under pressure and thermal effects, which overall acts as a boundary lubricant and helps protect the substrate material, and macroscopically shows a decrease in friction coefficient. SEM image at 46°C synovial fluid temperature as shown in Figure 11(h), which is consistent with the previous results of smaller friction coefficients for ZrO₂ ball-VE-XLPE discs with a protective film near the wear marks, indicating that the protein does react with the substrate material⁵⁴ to form a thin deposit on the surface layer. Results like those of Hussain et al. also confirm⁴² that a protective film of serum forms on the friction surface under human serum lubrication, which reduces the wear rate.

As shown in Figure 12, the results of the surface element test of three polymer specimens after friction under 20 N load and different synovial temperatures, the surface of the samples shown in Figure 12(a) and (b) have obvious blocky red O element accumulation. It has been claimed that such flake particles form a protective layer of deposited material on the friction surface as a transfer film, thus reducing the wear rate of UHMWPE,⁵⁵ and similar lubricating films can be observed in samples of synovial fluid in humans.⁵⁶ This explains the lower friction coefficient of UHMWPE in the previous analysis of friction results at the three synovial fluid temperatures because the proteins precipitated at higher synovial fluid temperatures formed a certain protective film to play a friction-reducing effect, and there is also speculation that⁵⁷ the change in the viscoelasticity of UHMWPE affected by the synovial fluid temperature may have an important effect on its friction behavior. It is also in Figure 12(d) and (i) can be found the accumulation of red lumpy O elements on the surface of XLPE and VE-XLPE, which also proves the process of protein precipitation, but in probably due to the lower concentration mixed with abrasive chips as a third body exacerbating the friction, which also corresponds to the result of the increased friction coefficient shown earlier in Figure 5(b) and Figure 6(b). The comprehensive analysis shows that the O element increases with increasing synovial fluid temperature, and it can also be observed that the O element percentage increases significantly for UHMWPE at 46°C synovial fluid temperature and VE-XLPE at 46°C synovial fluid temperature as shown in Table 4, which is consistent with the report that protein precipitates from synovial fluid in large amounts when the synovial fluid temperature is higher than 45°C.³²

Table 4.

EDS test results on disk surface after 20 N load friction test.

Element	UHMWPE			XLPE			VE-XLPE
	Propotion (%)			Propotion (%)			Propotion (%)
	37°C	46°C	55°C	37°C	46°C	55°C	37°C	46°C	55°C
C	96	94	96	95	95	83	97	84	95
N	1	4	1	1	1	3	1	3	1
O	1	2	1	1	1	9	1	8	1
Zr	3	3	3	3	3	5	2	5	3

As shown in Figure 12(a), (d), (h), and (i) for the EDS element distribution, it can be observed that the yellow C element appears as a delamination cut scene, which from a microscopic point of view is due to the sliding shear action causing the C element tearing on the surface layer of the material. It indicates that the higher temperature sliding fluid and frictional shear action together intensify the aging and oxidation process of the friction surface, with the appearance of broken chains and free radicals, thus reducing the mechanical properties of the material, and macroscopically showing an increase in the wear rate. The phenomenon is easily observed for UHMWPE and XLPE in synovial fluid at 37°C, while VE-XLPE is observed in synovial fluid at 46°C and 55°C, which to some extent reflects the stronger aging resistance of VE-XLPE compared to UHMWPE and XLPE.⁵⁸ In addition, it is not difficult to find from Table 4 that the Zr element also changes with the increase of the sliding fluid temperature, such as the Zr element has a significant increase in the elemental distribution of the wear zone at the synovial fluid temperature of XLPE at 55°C and VE-XLPE at 46°C. This indicates that material transfer occurs during the friction test, the wear is in the form of adhesive wear, and the higher synovial fluid temperature causes protein denaturation leading to increased adhesive wear,⁵⁹ which is more obvious during the friction at 46°C synovial fluid.

Wear finite element results analysis

The results of three acetabular wear values at different synovial fluid temperatures were obtained by simulating 1 complete gait of normal walking and walking for 1 min by ANSYS Workbench, and VWEAR wear clouds were output using the contnmisc189 command. According to the real-time variation curves of finite element wear of the three acetabular materials under one complete gait as illustrated in Figure 13(a)–(c), it can be found that the contact area between the femoral head and the acetabulum is larger at the beginning of the motion, and with the change of gait load, the relative forces and displacements of the contacting subsets occur and frictional wear begins to occur. According to the wear cloud maps in Figure 13(a)–(c), it can be seen that the maximum wear volumes of all three acetabulars occur near the edges, which is due to the concentration of edge loading, which intensifies the wear of the three materials. Analysis of the volumetric wear change curves reveals that the wear volumes all increase with the increase in temperature, and the maximum wear volume occurs at about 0.62 s. The higher the temperature variation from 37°C to 55°C, the greater the wear volume variation. Among the three acetabular materials, VE-XLPE showed the largest FEM wear results and UHMWPE showed the smallest FEM wear, with the volume loss of 2.25 × 10⁻¹² mm³ for VE-XLPE and 1.56 × 10⁻¹² mm³ for UHMWPE at 37°C. The volume wear of UHMWPE was 1.97 × 10⁻¹² mm³. The volume loss of VE-XLPE at 55°C was 3.49 × 10⁻¹² mm³, and the volume wear of UHMWPE was 2.95 × 10⁻¹² mm³, while the wear of XLPE was closer to that of UHMWPE.

The numerical simulations of acetabular wear under normal walking for 60 s are shown in Figure 14, and it was found that for all three polymer materials, the volume wear increased with increasing temperature, which could be approximated as a linear relationship. Among them, similar to the results of 1s gait motion, VE-XLPE has the largest volume wear and XLPE and UHMWPE have similar wear finite element results. After walking for about 60 s, the maximum wear amount of VE-XLPE at 37°C can reach 2.46 × 10⁻⁹ mm³, the maximum wear amount generated at 46°C is about 2.95 × 10⁻⁹ mm³, and the maximum wear amount at 55°C is 3.82 × 10⁻⁹ mm³. It can be seen that the influence of temperature parameter on the wear amount of polymer acetabular is obvious, the higher the temperature, the greater the wear amount of UHMWPE, the wear amount generated at 55°C is about 1.5 times the wear amount at 37°C.

In order to simplify the calculation, by fitting the wear data results of 60 gait cycles under three materials and three synovial fluid temperatures, the change relationship of the wear volume function with time as the independent variable under different temperatures was obtained as shown in equation (4), and the volume wear prediction volume of one million cycles (one million cycles) was obtained using this function model, and the prediction results are shown in Table 5. It can be found that the volume wear under each material pairing increases with the increase of synovial fluid temperature, among which, the volume loss of VE-XLPE is more obvious, and its wear at 55°C reaches 1.076 mm³/mc, while the volume wear of UHMWPE and XLPE is closer. It was also found that the predicted volumetric wear for 1 million gait cycles at all three temperatures was relatively small, because the clinically applied ceramic-headed-polyethylene acetabular mating actually produces much less wear compared to metal-based femoral head mating,⁶⁰ for example, one study predicted that the wear of CoCrMo-UHMWPE acetabulars could reach 30.94 mm³ per year,⁶¹ even up to 54.1 ± 13.9 mm³ when cross-shear effects are taken into account.⁶² In contrast, the results of Shankar et al.⁶³ found that the finite element simulation of the ZrO₂ femoral head with the UHMWPE acetabulum under frictional mating was about 0.1 mm³ per 1 million gait cycles, which is closer to the volume wear of the UHMWPE acetabulum predicted by the present method under one year of normal gait walking. Also in the study by Zietz et al.,⁶⁴ the wear per million cycles of the XLPE acetabulum at 37°C synovial fluid temperature obtained by the hip simulator was about 3.35 mm³, which is about 6 times higher than the 0.568 mm³ volumetric wear at 37°C that we obtained by finite element calculations. Compared with existing studies, the predicted values of 1 million volume wears obtained by numerical simulation are to some extent able to explain the frictional heat-induced wear at the acetabular interface of the artificial hip joint. However, the data used for the finite element numerical calculations were taken from the ball-disc friction test, which is different from the actual friction data at the interface of the artificial hip joint.

V (x) = p_{1} x^{2} + p_{2} x + p_{3}

(4)

Where V is the volume wear volume, x is the number of gait cycles, p₁, p₂, p₃ are the coefficients respectively, and their specific values are shown in Table 5.

Table 5.

Formula coefficients and volume wear after 1 million gait cycles.

Acetabular materials	Synovial fluid temperature (°C)	P₁	P₂	P₃	Wear volume (mm³/mc)
UHMWPE	37	4.776 × 10−13	−6.82 × 10⁻¹⁴	−4.932 × 10⁻¹³	0.478
	46	6.069 × 10−13	−8.574 × 10⁻¹⁴	−6.032 × 10⁻¹³	0.607
	55	9.097 × 10−13	−1.929 × 10⁻¹³	−4.28 × 10⁻¹³	0.910
XLPE	37	5.682 × 10−13	−9.157 × 10⁻¹⁴	−5.729 × 10⁻¹³	0.568
	46	6.185 × 10⁻¹³	−7.94 × 10⁻¹⁴	−7.579 × 10⁻¹³	0.619
	55	9.338 × 10⁻¹³	−1.17 × 10⁻¹³	−1.02 × 10⁻¹²	0.934
VE-XLPE	37	6.929 × 10⁻¹³	−1.28 × 10⁻¹³	−4.996 × 10⁻¹³	0.693
	46	8.325 × 10⁻¹³	−1.076 × 10⁻¹³	−9.909 × 10⁻¹³	0.832
	55	1.076 × 10⁻¹²	−1.353 × 10⁻¹³	−1.25 × 10⁻¹²	1.076

The finite element wear models of three polymer acetabulars established based on friction test data show that the models have a good prediction effect to a certain extent by predicting the volume wear for one year of walking. The numerical results show that frictional heat has a significant effect on the wear of polymer acetabulars, and for each material, the maximum wear region occurs near the edges, and the volume wear tends to stabilize with increasing wear time, with an approximately linear increase. Figure 15 shows the wear mechanism of the polymer acetabulum under the influence of frictional heat. As the service time of the polymer acetabulum increases, the prolonged accumulation of frictional heat will intensify the wear of the contact surface. This also explains that when a person walks fast or exercises vigorously, the frictional wear of the femoral head relative to the acetabulum will be more obvious and generate more frictional heat. The increase in temperature of the prosthesis and the surrounding synovial fluid caused by frictional heat will seriously affect the physicochemical properties of the prosthetic material itself, mainly by accelerating the oxidative degradation of the surface polymer material, resulting in a decrease in its mechanical properties and a deterioration in the wear resistance of the polymer material, which will aggravate the frictional wear of the acetabular material. Deeper optimization of the wear model is also needed, for example, by integrating many aspects of human musculoskeletal multibody modeling, synovial lubrication modeling, tribo-corrosion, and contact mechanics.⁶⁵

Conclusion

In this paper, the tribological behavior of three polymer materials under the influence of synovial fluid temperature was investigated experimentally and finite element analysis was conducted to reveal the influence of frictional heat-induced synovial fluid temperature rise on the wear resistance of polymer acetabular materials, which provides important theoretical support for the design and selection of acetabular materials.

a. The coefficient of friction of all three materials gradually decreases with the increase of the temperature of the synovial fluid, and the higher temperature of the synovial fluid will precipitate a large amount of protein, which will adhere to the contact surface to form a protective film and play the effect of friction reduction. Under light load and low temperature, UHMWPE is stronger than VE-XLPE and XLPE for protein molecule adsorption, but under more severe service environment (heavy load and high temperature) XLPE and VE-XLPE show stronger aging and abrasion resistance.

b. The analysis of the wear morphology results found that the wear depth and wear volume of the three materials increased with the increase of the temperature of the synovial fluid, and the temperature increase accelerated the aging and oxidative degradation of the material surface. The forms of wear at 46°C and 55°C were adhesive wear and plastic deformation. The EDS results revealed a cut-like pattern of C elements under sliding shear, as well as a buildup of Zr and O elements, further confirming the presence of protein films and oxidative degradation of the material surface.

c. The finite element analysis shows that the volume wear volume of the three materials increases with the increase of the synovial fluid temperature, and the trend can be approximated as a linear relationship, and the magnitude of volume loss is VE-XLPE > XLPE > UHMWPE, and the results of the wear cloud diagram show that the maximum wear volume occurs at the edge position of the acetabulum. The volume wear of three different acetabular materials under 1 million gait cycles at three synovial fluid temperatures was obtained by numerical simulation. The volume wear of XLPE acetabulum at 37°C, 46°C and 55°C was 0.568 mm³, 0.619 mm³ and 0.934 mm³, respectively. The volume wear of VE-XLPE acetabulum at 37°C, 46°C and 55°C was 0.693 mm³, 0.832 mm³ and 1.076 mm³, respectively. The volume wear of UHMWPE acetabulum at 37°C, 46°C and 55°C was 0.478 mm³, 0.607 mm³ and 0.910 mm³, respectively. The volume wear predicted by the mathematical method used after completion of 1 million gait cycles is low compared to reality, but can be used as a comparative assessment of the service life of different acetabular prostheses.

Footnotes

Declaration of conflicting interests

The authors declared no potential conflicts of interest with respect to the research, authorship, and/or publication of this article.

Funding

The author(s) disclosed receipt of the following financial support for the research, authorship, and/or publication of this article: This paper was supported by the National Natural Science Foundation of China (51705223).

ORCID iD

Songquan Wang

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