1. Introduction
Total hip arthroplasty (THA) is the definitive treatment for arthritis of the hip. It is the second most successful operation after cataract surgery and is performed in vast numbers throughout the world. Pioneering THA designs of the 1960–1970 era incorporated metal-on-polyethylene (MPE) and metal-on-metal (MOM) bearings, the latter using mainly cobalt-chrome alloys (as-cast CoCr). History shows that the 1st generation MOM bearings encountered many failures, considered multifactorial due to limited knowledge in design, instrumentation, and bone-fixation concepts [
1,
2,
3,
4,
5,
6]. In some cases, adverse CoCr wear produced severe biological reactions around the hip joints [
3,
7]. Nevertheless, subsequent studies showed some MOM patients achieved excellent clinical success, sometimes over 20 years [
6,
8,
9,
10,
11].
In hip simulator laboratories, a major body of research identified fluid-film lubrication as one of the goals for achieving success with hip bearings [
12,
13,
14,
15,
16,
17,
18]. Large MOM bearings appeared particularly suited to this concept. In the mid 1990’s, believing that problems with design, fixation and tribology had been addressed, implant companies developed the 2nd generation, large-diameter, MOM bearings. These included the femoral-head resurfacing type (
Figure 1: RSA) and the now standard ball and socket configuration (
Figure 2 and
Figure 3: THA). RSA and THA diameters ranged 36 to 60 mm and were widely accepted in their first ten years of use [
19,
20]. However, subsequent failures in many centers were blamed on production of CoCr debris and release of high concentrations of Co and Cr ions [
21,
22,
23]. Implant retrievals revealed that metal debris frequently blackened the hip capsules and surrounding musculature, the resulting production of pathological fluids, cysts, and tissue necrosis creating the appearance of “pseudotumors” [
23,
24,
25]. This immune tissue response was termed an adverse reaction to metal debris (ARMD) [
25]. Once again, cup design, surgical positioning and fixation were implicated as major risks for adverse wear. The most commonly reported failure mechanism appeared to be mechanical in origin, termed “edge loading” of steeply-inclined cups (
Figure 1C) [
21,
22,
23,
26,
27,
28,
29,
30,
31]. Edge-loading of THA bearings in patients’ hips may occur in several ways. For this report, we defined three likely mechanical risks as; (I) constant rim wear of the cup due to mal-positioned implants, (II) rim wear occurring intermittently at some extreme of patient motion, and (III) rim wear occurring intermittently due to hip-impingement episodes (head destabilized and forced over cup rim).
The second negative clinical experience with MOM bearings provided us with the incentive to re-examine methodology used in the simulator wear studies. Understanding and predicting performance of THA devices remains a complex and demanding task. Over a 5-decade history, hip-simulator laboratories developed pre-clinical wear tests that guided the medical industry [
32,
33,
34,
35,
36,
37,
38,
39,
40,
41]. Beginning in the 1960’s [
42], wear-simulation of THA bearings contemplated integration of hip-joint biomechanics and tribology with many related parameters that included material combinations and implant designs. In addition, there were many patient variables to be considered (
Figure 1), and much needed simplified to formulate feasible research protocols. Thus, simulator guidelines used by various regulatory agencies [
43,
44,
45] represented a necessary simplification. The internationally accepted methodology shows the cup mounted above the femoral head (
Figure 2), thereby simulating the “Anatomic” hip configuration as viewed on patient radiographs (
Figure 1) [
43]. The stipulated test inclination for the cup (
Figure 2: 60° to load axis) represents a 30° angle (
I°) in the simulator’s horizontal plane, this believed to approximate the 45° cup inclination in patients (
Figure 1A).
Wear studies with large-diameter MOM bearings followed standard simulator guidelines (
Figure 2) with some preferring to use a 35° cup inclination [
46,
47,
48,
49,
50,
51,
52,
53,
54,
55]. Resulting wear-rates appeared acceptably low at 0.5 to 0.8 mm
3 per million load cycles (Mc). Three of these non edge-loaded studies also reported on their cup wear-patterns [
46,
47,
48].
With emerging clinical insight regarding the risks of “edge-loading” in steeply inclined cups (
Figure 1C) [
21,
30,
31,
49], simulator laboratories responded with steep-cup wear studies. All used the same simulator guideline, i.e., constant edge-loading of cup (fixed-inclination) that represented mode-I clinical risk. There were no guidelines for selection of test parameters. Cup inclinations were different in each study and MOM diameters varied 38.5 to 48 mm (
Table 1). Even in non edge-loaded tests, it is recognized that some MOM bearings will produce much higher wear-rates for reasons unknown, i.e., “run-away wear” or “break-away wear” (“BAW”) trends [
48,
50,
51]. Such BAW wear phenomenon represents a dramatic contrast to the anticipated classical run-in and steady-state wear phases (
Table 2: “STD”). Thus, the large variation in steep-cup MOM wear-rates (1.9–19.5 mm
3/Mc) remained open to interpretation (
Table 1).
The severity of edge-loading will depend to what degree the normal wear-pattern becomes truncated by the cup rim. This can only be determined if size and location of wear patterns are known. In an earlier study, we developed an algorithm that predicted wear-pattern areas and the degree of edge-loading likely to be present in a patient’s MOM hip joint [
57]. As a simulator test method (
Figure 4), when the cup-profile angle (
P), and its included angle (
Ac) are known, the margin-of-safety (MOS) protecting the cup from edge-loading can be calculated using,
The rim angle (P) for the sub-hemispherical profile of large diameter cup is given by,
The wear area (CAP = 442 mm
2) measured in 60 mm cups [
48] subtended a wear-pattern angle (
Ac) estimated (using standard spherical equations) to be 45.6°. The profile angle in this large, sub-hemispherical cup was given by
P = 11.7°. Therefore, the 45° test inclination used (
Figure 4) would produce a margin-of-safety of MOS = 10.5° (Equation (1)), i.e., no edge-loading. This Anatomic-cup study [
48] also revealed that measured head wear-areas averaged 3.78-fold greater than in cups (
Table 2). This is a result of the head’s ±23° orbit under the simulator’s load-axis (Z). The wear-pattern formed on the head (
Figure 4) subtends the angle
Ah =
Ac + 46° = 91.6°. Using standard spherical equations, it can be shown that this larger wear-pattern has an estimated area of 1712 mm
2, i.e., 3.87× times larger than in cups. This calculation validated our previous experimental wear-pattern ratio [
48].
Notable in the prior Anatomic-cup study was that wear-patterns on heads represented 29% to 34% of the nominal surface area but cups only 7% to 13% (
Table 2). As visualized on our large plastic models, head and cup wear-patterns presented dramatically different areas (
Figure 5). It was therefore notable that implant retrieval studies with MOM and ceramic-on-metal (COM) bearings revealed that wear-patterns in vivo extended over 50–60% of nominal cup surfaces [
58,
59]. With a goal to creating clinically-relevant cup wear-patterns in edge-loading simulations, we decided that a novel test strategy would be to reverse the standard Anatomic-cup configuration. Simply mounting the cups in the inverted position would change the distributed wear-pattern from heads (
Figure 5) to cups. This new strategy also introduced a constantly varying cup inclination, which would simulate for the first time the mode-II clinical risk of edge-loading [
60].
Our trial MOM simulator study with Inverted-cups (1 Mc duration) validated the algorithm’s predictions, showing no edge-loading evident with cup inclinations peaking at 40° and 50° (
Figure 6A) [
60]. The follow-up wear study with 70° peak cup inclination produced edge-loading over 5 Mc duration. Wear-pattern analysis showed edge wear was achieved and validated over 5 Mc duration (
Figure 6B) [
60].
Compared to the 60 mm Anatomic-cup test [
48], edge-loading in Inverted-cups produced approximately 3-fold higher wear-rates (
Table 3) [
60]. This was an encouraging result but the 5 Mc wear study suffered from two limitations. Gravimetric wear assessment of the bearings was compromised due to transfer of proteinaceous debris from the lubricant [
61,
62]. This contamination appeared during “run-in” and “steady-state” wear phases. In addition, one test-station experienced an simulator-cam failure, resulting in initial loss of one MOM bearing. These limitations reduced the effectiveness of the Inverted-cup model and also made wear predictions uncertain for 60 mm MOM. Therefore, the purpose of the current study was to extend the Inverted-cup test from 5-million [
60] to 10-million cycles. This longer duration would aid definition of steady-state wear and thereby provide better estimates for run-in wear-rates. The hypotheses were that, (1) wear trends in 60 mm heads and cups would be linear from 5 to 10-million cycles duration, (2) re-assessment of run-in wear phases would reveal transition to steady-state wear occurred prior to 0.75 Mc duration, (3) 60 mm MOM wear-rates would average 1.7 to 2 mm
3/Mc overall [
60], and (4) cup wear magnitude would represent >70% of total MOM wear [
48].
2. Materials and Methods
The four 60 mm MOM bearings used in the 10-million cycle study were the originals donated for the 5 Mc study [
60] by DJO-Global (Austin, TX, USA). The cups were mounted at a 47° face angle in Polyacetal adaptors (
Figure 7A) with locking rings and anti-rotation pegs to guard against high torques associated with large CoCr bearings. The orbital hip simulator (Shore Western, Monrovia, CA, USA) was identical to that used in the prior study and test methods duplicated that work [
60]. Bearing wear was assessed by gravimetric techniques at each 0.5 Mc interval.
The Inverted-cup method provided variable edge-loading conditions. The four cups set with inclination
L = 47° in the home-cam position (
Figure 7B) alternated between “shallow” (
I = 24°) and “steep” (
I = 70°) inclinations during each 1-s orbit of the simulator cam (
C = ±23°). As in prior studies, the bovine-serum lubricant (Hyclone, Ogden, UT, USA) was diluted 30% to a concentration of 20 mg/mL in each 450 mL test chamber [
48,
60]. New lubricant was installed after each cleaning cycle. Weight-loss changes due to wear were measured at 0.5-million cycle intervals to 10 Mc duration. Implant cleaning procedures were rigorous in anticipation of protein contamination from lubricant precipitation. When weight-gain trends were identified, those bearings were taken back for inspection and additional cleaning. Head and cup weight-loss trends were compared by linear-regression techniques and box-plot analyses from 1 to 10 Mc duration. In the 5 Mc Inverted-cup study, MOM-1 was omitted from consideration [
60] and remaining three MOM bearings used for analysis. The present study continued in this way but re-introduced MOM-1 for comparison. Bearing wear areas were identified visually and by light microscopy, stained red for photography and taped to minimize cup reflections (
Figure 6) [
48,
60].
In prior studies without edge-loading (
Figure 6A), it was possible to track the size of head and cup wear-patterns (CAP mm
2;
A°) and the margin of safety (MOS°). In the prior Anatomic-cup study, wear-patterns on heads averaged 1668 mm
2 in area [
48]. The trial Inverted study (1 Mc, 17° inclination) produced virtually identical cup wear-patterns (average area 1663 mm
2) [
60]. Using standard equations for spherical geometry, the corresponding control angle (
A°) for cup wear-patterns in this study was calculated as
A = 90.2°. The corresponding measured MOS-angles 15.4° and 5.3° for 17° and 27°-mounted cups demonstrated the 10° shift in wear-patterns [
60].
In contrast, an edge-loading test has no margin-of-safety and the standard wear-pattern (
A°) will be truncated by the cup rim. For this test, we tracked estimated non-wear (
N°) and truncated wear-zone (
B°) angles predicated on simulator cam-rotation being ±23° (
Table 4). At the cam’s home position (
Figure 8), the truncated wear-pattern was estimated as
B = 82.4°, this representing 7.8° of edge-loading. As the cam rotates down to its minimum position, cup inclination will reduce to 24° thereby providing a 15.2° margin-of-safety. As the cam rotates up to maximum position, cup inclination increases to 70° thereby indicating
E = 30.8° as the maximum edge-loading.
With non-wear angle (
N°) estimated to vary 63° to 109° (
Table 4), the home position (
Figure 8: datum Q) was used as a control to track measurements comparing the estimated values
N = 86.2° and
B = 82.4°. At 8.5, 9.5 and 10 Mc durations, the corresponding of B-chords and N-chords of cup wear-patterns (
Figure 5) were measured and converted to angular data for analysis.
It was necessary to predict the degree of edge-loading represented in each wear study. Here the critical inclination (
Ic) is defined as that cup position where MOS = zero, i.e., cup wear = pattern is juxtaposed to the cup rim, and given by
The severity of edge-loading (EL) may be defined as,
For the 60 mm cup in Anatomic test mode, the critical inclination was calculated as
Ic = 61.5° (
Table 5). In the Inverted test, the larger wear–pattern lowers the critical cup inclination to
Ic = 39.1°. Estimated risks of edge-wear were, (i) 9% edge-loading at home-cam position, (ii) 34% edge-loading at high-cam position, these contrasting with (iii) margin-of-safety MOS = 15.2° at the low-cam position (
Figure 8).
For comparison of published “steep-cup” data (
Table 1) it was necessary to determine the wear-pattern (A) and profile (P) angles applicable to each cup design. It is known that the initial contact areas enlarge as bearing wear progresses during run-in phase [
46,
47,
48]. When the contact-stresses are sufficiently reduced to allow optimal tribological conditions, the resulting equilibrium producing lower wear-rates in the “steady-state” phase. The steady-state wear-patterns measured in MOM studies with no edge-loading typically ranged 380 to 480 mm
2 in area (
Table 6). These averaged approximately 4-fold larger than contact-areas predicted using Hertzian equations for elastic deformation of spheres (
Figure 9A). The wear-pattern angles (
Ac) were calculated (using standard spherical equations) for large diameter cups (38–60 mm) and plotted (
Figure 9B). This treatment revealed a linear relationship, with angle-
Ac inversely proportional to cup diameter (D mm). Omitting two outliers (#1, 8), the wear-pattern angle (
Ac) was defined by,
3. Results
As plotted over 10-million cycles duration (
Figure 10), it was apparent overall that, (i) heads and cups demonstrated run-in wear and both showed steady-state wear phases 1 to 10 Mc duration, (ii) cups (#2–4) produced more wear than heads (#1–4), and (iii) total wear in cup-1 was approximately double that of the other cups. Cup wear-patterns stained for photography at 10 Mc duration demonstrated the truncated (
B) and large non-wear areas (
N) typical of edge-loaded cups (
Figure 11). Their area assessments were represented on average by angles
N = 86.3° and
B = 82.5°.
All cups showed an initial weight-gain artifact due to adherence of protein contaminants (
Figure 10: Flag-1). Cups #2–4 behaved similarly as a group, revealing a 2nd weight-gain artifact at 3 Mc (Flag-3). Their wear trends appeared to plateau at 5 Mc (Flag-4) and again at 8 Mc (Flag-5). Even with cyclical weight-gain artifacts evident, the overall wear in cups trended linearly to provide an average wear-rate of 9.9 mg/Mc ± 10% over 1 to 10 Mc (
Figure 12A: minimum R = 0.98). Extrapolating steady-state trends back to the
Y-axis, revealed cups #2–4 transitioned into steady-state phase by 1.5 Mc (
Figure 12B). Omitting the weight-gain artifacts at 0.25 Mc, cup run-in wear-rates were estimated to be 13.4 to 16.3 mg/Mc, averaging 14.9 mg/Mc. This run-in estimation was 1.5-times greater than their ensuing steady-state trend. Cup-1 in this study showed anomalous behavior. By 1 Mc it produced a run-in wear magnitude of 58 mg (
Figure 10: Flag-7). By 2 Mc duration, it appeared to be transitioning into a steady-state phase (Flag-8) only to accelerate with another run-in wear phase to 5 Mc (Flag-9). The final steady-state wear phase 5 to 10 Mc appeared reasonably linear (Flag-10) 9.2 mg/Mc) and a good match to that of cups #2–4 (Flag-6).
Heads #1–4 did not show any weight-gain artifacts as measured initially in cups. Their steady-state wear trending was negligible to 4 Mc (
Figure 10: Flag-4) and they only began to demonstrate some mild wear beyond 5 Mc. It was interesting that head-1 with the lowest wear-rate was paired with the highest wearing cup-1. Regression analysis for steady-state wear in heads #1–4 averaged 2.3 mg/Mc ± 5% (
Figure 13A: 1–10 Mc), indicating good precision with high regression coefficients (minimum R = 0.96). Extrapolating steady-state trends back to the
Y-axis revealed heads transitioned into steady-state phase by 0.75 Mc (
Figure 13B). These data provided run-in wear estimates ranging 7 to 20 mg/Mc, with average 19.1 mg/Mc. This run-in wear estimation for heads was 8.3-times greater than the ensuing steady-state trend.
At 10 Mc duration, the cumulative volume of wear in three MOM bearings (#2–4) represented 17.1 mm
3 with cup wear providing 75% of total (
Figure 14). The corresponding wear volume in MOM-1 was 27.9 mm
3 (
Figure 12), with cup wear providing 93% of total.
4. Discussion
This simulator study examined the wear response of 60 mm metal-on-metal (MOM) bearings run to 10-million cycles duration (10 Mc) with cups mounted under transient edge-loading conditions. This may be the first simulator study using an algorithm to predict severity of edge-loading in steeply-inclined cups. The 60 mm cup diameter represented the largest in wear studies to date and featured the largest margin of safety (MOS). The standard Anatomic-cup test, defined in international simulator guidelines, was reversed in this study so that hip motion could be input via the acetabular cup. This created the larger cup wear-patterns considered more clinically-relevant for simulation of edge-loading. This novel method also provided the first demonstration of intermittent edge-loading in steeply-inclined cups (mode-II clinical risk) rather than the standard fixed edge-loading condition (mode-I clinical risk). With cup inclination varying continuously (24° to 70°) in each cycle period of 1-s, the overall wear-rate in three MOM bearings averaged 1.7 mm3 per million cycles while one MOM, considered an outlier, was almost 60% higher at 2.7 mm3 per million cycles. The Inverted-cup model demonstrated satisfactory run-in and steady-state wear phases to 10 Mc duration in both heads and cups, thereby satisfying hypothesis #1.
There are three major considerations in laboratory wear simulations, namely (1) experimental design, (2) credibility of wear trends, and (3) clinical relevance. The most disturbing result from this review of both standard and steep-cup MOM wear studies was that laboratory wear-rates varied from less than 1 to over 19 mm
3/Mc for reasons not understood and with no guidelines available for selection of test parameters (
Table 1). In particular, the severity of edge-loading will depend to what degree the normal wear-pattern becomes truncated by the cup rim. This can only be determined if size and location of wear-patterns are known. In an earlier clinical study, we used an algorithm to predict wear-patterns and severity of cup edge-loading in MOM patients [
57]. This was extended to our 60 mm MOM simulator study to validate size and location of wear-patterns. The trial study with Inverted-cups used two tests with cyclic cup-inclinations −6° to 40° and 4° to 50° to verify “non edge-loaded” conditions and one with 24° to 70° inclination to verify “edge-loading” status [
60]. The present 10 Mc wear study satisfactorily demonstrated edge-loading, the resulting MOM wear-rates averaging 1.7 mm
3/Mc. Applying these new estimates back to the 5 Mc-datum provided overall wear-rate 2 mm
3/Mc. This was approximately 3-fold greater wear than our Anatomic study with non edge-loaded cups [
48]. These results served to validate the experimental design of the Inverted-cup test, and also confirmed our 3rd hypothesis.
It is well documented that protein contamination from serum lubricants can be a confounding problem in gravimetric wear assessments [
61,
62,
63,
64]. Our prior 5 Mc study revealed major contamination confounded the wear assessments [
60]. This was partly anticipated because we reported a similar cyclic phenomenon in 28 mm MOM studies [
63,
64]. We extended the 60 mm MOM study from 5 Mc to 10 Mc to provide more credibility to wear predictions in the Inverted-cup method. Possibly because 60 mm bearings were capable of greater frictional torque, there may have been a more conspicuous deposition of proteins that were well bonded onto CoCr surfaces (
Figure 15). However, even with this limitation, linear regression trends taken over 9-million cycles proved sufficiently robust for steady-state wear assessments. By extension backwards to the Y-axes, new estimates were determinable for head and cup run-in trends. This showed that best-fit for run-in wear transitioning into steady-state was at 0.75 Mc and 1.5 Mc for heads and cups, respectively. Therefore, hypothesis #2 stating that heads and cups would transition into steady-state phase before 1 Mc was not fully satisfied. Comparing 10 Mc to 5 Mc data (
Table 4), it was apparent that the short-duration test had underestimated run-in wear-rates for heads and overestimated that for cups (
Table 7). However, the largest discrepancy in the prior 5 Mc study was underestimating steady-state wear-rates in heads. The new data revealed that heads had higher run-in wear-rates but for a shorter duration than cups. As a result, both transitioned into their steady-state phases after approximately the same amount of wear was produced during run-in (
Figure 12B and
Figure 13B).
A major confounding effect in MOM simulator studies has been the appearance of unexplained, usually high wear-rates that appeared transiently for 1 to 3 Mc duration. Such wear phenomena, varyingly described as “runaway” or “break-away” (“BAW”) wear, appeared in non edge-loaded tests, and were estimated to occur in 15% to 40% of MOM bearings [
48,
50,
51]. A 50 mm MOM study [
50] demonstrated one of five bearings produced 3-times higher wear by 3 Mc with subsequent recovery over 2 Mc duration (
Figure 16A). A similar 40 mm MOM study noted adverse wear in three of eight bearings (38%), cup wear-rates being noted in excess of 30 mm
3/Mc [
51]. It was noted that one of three cups demonstrated a recovery by 2.5 Mc duration, producing minimal steady-state wear thereafter (
Figure 16B). This BAW phenomenon presented as a major confounding event, particularly when the typical simulator study has only small numbers of replicates, i.e.,
N = 2 to 8 bearings.
In the present study, MOM-1 bearing produced an overall wear-rate of 2.7 mm
3/Mc, approximately 60% greater than the other bearings (
Figure 14). Our laboratory noted that the station-1 cam-bearings were replaced early in the 5 Mc study, apparently indicating a problem that could have triggered adverse wear. However, with hindsight, we would now ascribe MOM-1’s trend to an extended “BAW run-in” phase to 5 Mc duration (
Figure 10), similar to those previously described (
Figure 16). It was also notable that the break-away wear-rates (BAW) described in the prior 60 mm Anatomic study [
48] were of similar magnitude, averaging 2.1 mm
3/Mc overall. However, these wear-rates were quite modest compared to data in some MOM studies (
Table 8). A surprising result was in the degree of wear experienced by 60 mm cups. At 5 Mc in the prior Anatomic study, wear in BAW-cups represented 85% of total, compared to 68% in standard cups (STD). At 5 Mc in the Inverted study, the corresponding cup wear ratios were BAW = 95% and STD = 70%, these satisfying hypothesis 4. Prior steep-cup studies did not provide individual head and cup wear-rates [
53,
55], so there were no other comparisons at 5 Mc duration. Given these examples of MOM wear variations, we are of the opinion that such “MOM outliers” should be included as part of the overall assessment.
The clinical relevance of laboratory wear tests is quite difficult to assess. A hip simulator study represents a complex undertaking while still representing an over-simplification of patient conditions. A major body of work over the years identified fluid-film lubrication as the ideal condition for success in MOM bearings [
12,
13,
16]. However, MOM bearings failing in patients have demonstrated adverse wear mechanisms that can produce significantly high volumes of CoCr debris within a few years [
21,
22,
23]. The radiographic presentation of steeply-inclined cups has frequently been criticized, many studies citing “edge loading” as major evidence for adverse MOM wear in patients [
21,
22,
24,
27,
29,
30,
31,
49]. It is therefore interesting that three very different simulator studies of edge-loaded cups [
53,
55,
60] produced MOM wear-rates elevated compared to control data (
Table 1) but still within the range (0.5 to 2 mm
3/Mc) considered typical of many MOM simulator studies [
65]. In prior 3rd-body wear studies with CoCr and Ti6Al4V particulates, wear-rates > 3 mm
3/Mc turned the MOM lubricants black [
66]. This was not experienced in this 10 Mc study of edge-loaded bearings.
This 10 Mc study is the first Inverted-cup model with test inclinations varying each cycle, simulating a patient achieving edge-loading only intermittently during the activities of daily living (mode-II clinical risk). Our suspicion was that the Inverted-cup model might produce less wear than steep-cup studies running with full-time edge-loading (mode-II clinical risk). A 40 mm Anatomic-cup study run to 6 Mc duration appeared the most comparable data [
55]. The 1.9 mm
3/Mc wear-rate in 40 mm cups inclined 60° appeared almost identical to our 60 mm Inverted-cup study with 2 mm
3/Mc overall wear-rate (
Figure 14: 5 Mc duration) However, the question was how comparable were two tests using different MOM designs, very different cup diameters and differing test inclinations? The MOM algorithm predicts that the wear-pattern angle in a 40 mm cup would be A
c = 63.3° (Equation (5)). The cup-face angle (CF) for a 40 mm Cormet cup (Corin Group PLC, Cirencester, UK) was published elsewhere as 159° [
31]. Equation (2) provides the cup-profile angle as
P = 10.5° and from Equation (3), the critical cup inclination was estimated as
Ic = 48°. Therefore, the 35°, 50° and 60° fixed-cup inclinations would have test conditions given by, (i) MOS = 13°, (ii)
E = 2° and (iii)
E = 12°, respectively. The latter two angles represented edge-loading condition of EL = 3% and 19% (Equation (4)). Edge-loading conditions in our Inverted-cup study varied from MOS = 15° to EL = 9% and peaked at EL = 34% in each duty cycle (
Table 4). While such similarities in two studies are intriguing, they remain purely speculative until more simulator data is accumulated. However, these two simulator studies with full-time and intermittent edge-loading conditions did not produce the much-cited adverse wear blamed on steeply-inclined cups. In fact, both showed the typical run-in and steady-state wear typical of stable MOM wear trends. This suggested that the appearance of adverse wear mechanisms may be triggered by additional circumstances.
Hip impingement in particular represents an unavoidable and we believe a major clinical risk for production of large 3rd-body metal particulates. As illustrated (
Figure 17), when the femoral implant of either CoCr or Ti6Al4V alloy can impact against the acetabular cup, the CoCr liner rim (
Figure 17: #1) may notch the neck (
Figure 17: #2), releasing either CoCr or Ti6Al4V particles [
58,
67,
68]. The resulting destabilization of the hip-joint allows large hip muscles to force head rotation (or migration #3) over the cup rim (#4) during such episodes, creating large linear scratches on heads (#5) [
58] and releasing large quantities of 3rd-body CoCr particles [
58,
69]. Frequent head subluxations may also produce gross erosion of the liner rim (#4 [
48]). We termed this phenomenon “repetitive sub-clinical subluxation” (RSS) because our patients were generally unaware of it [
70]. It is both unpredictable and of unknown severity and consequences. Therefore we hypothesize that patient activities may result in hip micro-separation and impingement episodes that release large CoCr particles into the joint space [
58,
68]. We know that introduction of metallic particles greatly accelerates wear of MOM bearings, consistently turning simulator lubricants black in color at wear-rates greater than 3 mm
3/Mc [
66]. With ensuing activities of daily living, we further propose that the resulting 3rd-body wear mechanisms produce adverse wear until the CoCr particles are either worn small enough over time to escape or become ionized [
58]. However, continuing hip-impingement episodes will initiate further release of metal particulates.
In summary, the most disturbing result from this review of both standard and steep-cup wear studies is that MOM wear-rates in laboratory studies have varied widely for reasons not well understood. Even with purposefully matched bearings studied in expert laboratories, MOM wear performance has proved unpredictable and quite erratic [
48,
50,
51,
54,
56]. Thus, the large range of MOM wear measured in laboratory studies may in itself be a warning that the risk of adverse wear conditions in MOM patients is very high. These combined observations suggest a new hypothesis, that MOM bearings are extremely sensitive to external influences, be they simulator artifacts, or patient related.