Abstract

Introduction

In hip arthroplasty, it has been shown that assembly of the femoral head onto the stem remains a non-standardized practice and differs between surgeons [1]. Pennock et al. determined by altering mechanical conditions during seating there was a direct effect on the taper strength [2]. Furthermore, Mali et al. demonstrated that components assembled with a lower assembly load had increased fretting currents and micromotion at the taper junction during cyclic testing [3]. This suggests overall performance may be affected by head assembly method. The purpose of this test was to perform controlled bench top studies to determine the influence of impaction force and compliance of support structure (or damping) on the initial stability of the taper junction.

Statement of Purpose

Mechanically assisted crevice corrosion of modular tapers continues to be a concern in total joint replacements as studies have reported increases in local tissue reactions¹. Two surgical factors that may effect taper seating mechanics are seating load magnitude and orientation.

In this study 12/14 modular taper junctions were seated over a range of loads and loading orientations. The goals of this study were to assess the effects of load magnitude and orientation on seating load-displacement mechanics and to correlate these to the pull-off load.

Burroughs et al showed that frictional torque increases with increasing head size in a simple in vitro model and showed differences in frictional torque with different polyethylene materials [1]. Therefore, the purpose of this study was to evaluate the influence of bearing material and bearing size on the frictional torque of hip bearings utilizing a more physiologically relevant hip simulator model.

A total of four hip bearing combinations (Crosslinked PE/CoCr, Conventional PE/CoCr, Crosslinked PE/Delta and Alumina /Alumina) with various bearing sizes were evaluated. The sizes tested in this study range from 22 mm to 44 mm; it is important to note that the study only evaluated bearing combinations (size and material combination) currently commercially available. A total of three samples per bearing combination were tested, with the exception of conventional PE, which included a total of 4 samples. A MTS hip joint simulator was used. All components were oriented anatomically with the femoral head mounted below on a rotating angled block which imparts a 23° biaxial rocking motion onto the head. Loading was held constant at each load level (500N, 1000N, 1500N, 2000N, 2450N) for at least two rotational cycles while all 3 axes of load and all 3 axes of moments were measured at 10 khz. Fresh Alpha Calf Fraction serum was utilized as a lubricant.

Results show that frictional torque increases with the increase of head size regardless of head material for all polyethylene combinations (p > 0.05), as shown in Figure 1 and 2. However, results showed no change in frictional behavior for the Alumina/Alumina combination regardless of the bearing size. The results of this test did not show any significant difference between crosslinked PE and conventional PE materials for sizes 28 mm and 32 mm when paired against a CoCr head (p > 0.05) (Figure 3). The Alumina/Alumina bearing combination had the lowest frictional torque among all the bearing material combinations evaluated in this study.

This data suggests that there is a strong correlation between increased head size and increased frictional torque (R² = 0.6906, 0.8847) for the polyethylenes evaluated here regardless of head material. No correlation can be concluded for the Alumina /Alumina bearing combination (R² = 0.0217). The combination of Alumina /Alumina seems to have the most favorable frictional properties. This data also suggests no effect on frictional properties regardless of the polyethylene material (crosslinked and conventional) for sizes 28 mm and 32 mm. The frictional torque values recorded in this study are different than those published by Burroughs et al [1]. This difference may be attributed to the testing methodology. The current study utilizes a hip simulator, which closely mimics the natural joint providing a more physiologically relevant model whereas the Burroughs et al study utilizes a single axis machine. It is important to understand that frictional behavior in hip bearings may be highly sensitive to bearing clearance, cup thickness, and stiffness, which may outweight the effect of head diameter. Further evaluation is necessary to isolate and investigate those parameters.

The dual mobility hip incorporates a femoral head mated within a spherical polyethylene liner which also has an unconstrained outer articulation with a polished metal shell. An additional wear surface is introduced at the outer articulation, however, the mobility of the polyethylene insert does allow for femoral-neck/acetabular-insert impingement by allowing the insert to displace upon contact. We evaluated the wear performance of a dual mobility hip during abrasive and impingement conditions independently. Three abrasive conditions were evaluated; abraded acetabular cup, abraded femoral head, and both abraded cup and head. Two impingement conditions were evaluated; impingement of the unconstrained acetabular insert against the femoral neck, and acetabular-insert/femoral-neck impingement when the insert becomes immobilized at the outer articulation.

Wear testing was conducted using a hip stimulator. The simulator applied physiologic loading with a maximum load of 2450 N and serum as the lubricant. Components were abraded at the pole according to a published method. Abraded samples were tested at 0° of inclination. The unconstrained impingement condition was created by adjusting the femoral neck angle to achieve impingement with 45° of acetabular inclination. Neck to liner impingement can occur at either the superior or inferior surface of the femoral neck, with subsequent impingement occurring randomly as the insert is allowed to re-align itself throughout testing. The fixed impingement conditions was created by locking the outer bearing through fixturing and inducing impingement as previously described. Dual mobility control components were tested at 0° and 50° of inclination. Inserts were sequentially crosslinked GUR 1020 polyethylene.

Results are shown in Figure 1. Abrasion testing results correlated to a combination of friction at the abraded articulation and bearing size. Abrasion at only the inner bearing had a larger effect on wear when compared to abrasion of only the outer bearing. When both sides were damaged, femoral head abrasion led to an increase in friction and resistance to movement at the inner articulation, thereby forcing an increase in overall movement of the outer articulation. This increased the contact area subject to motion across a scratched metal surface, which increased the wear rate of the system. Unconstrained impingement samples impinged during the first cycle and then randomly throughout testing, while the fixed impingement samples had predictable impingement at the same location every cycle of testing. The unconstrained impingement model was designed to replicate an instance where the dual mobility hip would run in a near/intermittent impingement condition where the polyethylene insert displaces upon contact with the femoral neck. Unconstrained impingement wear rates were not statistically different than the ideally aligned control. The fixed impingement samples wore at a higher rate than the unconstrained impingement and control groups. The insert encountered resistance to movement upon impingement resulting in wear and deformation at the point of contact. Additional intended bearing wear was also generated by head sliding and translation of the load path upon impingement of the rim. Note that this condition is difficult to envision clinically and all wear rates, even under adverse conditions, were acceptably low.

Multi-directional motion at the ball-socket interface of a hip replacement joint has been discovered as a fundamental feature that determines the magnitude of wear for ultra-high molecular weight polyethylene (UHMWPE). The present study considers the wear of UHMWPE moving along a circular path with a uniform angular change rate of the velocity vector defined by the curvature of the sliding circle. It is apparent the as the sliding circle radius increases the motion is approaching more towards linear tracking. Therefore, wear rate per unit sliding distance would decrease with increasing the slidng circle radius. However, the sliding distance per cycle increases linearly with the radius of the circle, which would cause a proportional increase in the wear rate per cycle. We hypothesize that these two opposing effects on wear with respect to the changing radius of the sliding circle would cancel out each other leading to wear rate per cycle being independent of sliding distance.

Experiments were conducted on a hip simulator with a biaxial rocking motion that results in a circular sliding path at the polar region of the acetabular cup that experiences the highest contact stresses and wear. The radius of the sliding circle, r, depends solely on the radius of the femoral ball, R, and the biaxial rocking angle, a, such that r=Rsina. Two tests were conducted. The first test was run under standard conditions with a constant biaxial rocking angle of +/−23 and head diameters ranging between 28 mm and 44 mm. Acetabular components were machined from virgin non-crosslinked UHMWPE with inner diameters matching those of the femoral heads. For the 28 mm bearing, the cups were of standard hemispherical geometry. The larger cups were truncated by various degrees so that the nominal contact area remained exactly the same as that of the standard 28 mm hemispherical components. The second test was run with the standard 28 mm components and various biaxial rocking angles: +/−10, +/−15, +/−20 and +/−23. Both tests were run for a total duration of 2 million cycles with diluted alpha-serum as a lubricant and physiologic loading (peak load: 2450N) as described by Paul.

Volumetric wear at 2 million cycles for both tests are summarized in Figure 1. Fig. 2 shows a graphic representation of the total volumetric wear (DV) as a function of the sliding circle radius (r). Total volumetric wear is independent of the head diameter (2R), the biaxial-rocking angle (a) and the sliding circle radius (r). The total volumetric wear is proportional to the number of cycles and independent of the sliding distance per cycle. The clinically observed wear rate-ball diameter relationship, therefore, is not attributed to variations in sliding distance per walking step with differing ball head sizes.

For the same nominal contact area between a ball and a socket, the total volumetric wear of UHMWPE is independent of the ball diameter, the biaxial rocking angle and the sliding circle radius. In other words, the total volumetric wear is proportional to the number of cycles and independent of the sliding distance per cycle.

Pin-on-disk studies have demonstrated the role that cross-shear plays in polyethylene wear. It has been found that applying shear stresses on the polyethylene surface in multiple directions will increase wear rates significantly compared to linear sliding. Hip and knee joint replacements utilize polyethylene as a bearing surface and are subjected to cross-shear motions to various degrees. This is the mechanism that produces wear particles in hip and knee arthroplasty bearings and if excessive may lead to osteolysis, implant loosening, and failure. The amount of cross-shear is dependent on the bearing diameter and the angular motion exerted onto the bearing due to the gait of the patient. This study will determine the effect of sliding curvature (angular change per linear sliding distance) on the wear rate of polyethylene. Virgin polyethylene blocks were machined with a 28mm diameter bearing surface and against 28mm cobalt chromium femoral heads in a hip simulator. Dynamic loading was applied simulating walking gait but the motion differed between testing groups. Typical walking gait testing utilizes 23° biaxial rocking motion, in this study, 10°, 15°, 20°, and 23° biaxial rocking motions resulting in various sliding curvatures. Sliding motion path is described in Figure 1 and is a function of the bearing radius and the rocking angle. With increased rocking angle, the sliding distance reduces per cycle and the sliding path becomes more curved (more angular change per linear distance of sliding). Despite a significant increase in sliding distance at higher rocking angles, wear rates were relatively unchanged and ranged from 57mm3/mc to 62mm3/mc. Wear rates per millimeter increased exponentially with reduced sliding arc radius (smaller rocking angle) as shown in Figure 2. This study suggests that wear of polyethylene is highly dependent on sliding path curvature. The sliding path is largely a function of the bearing diameter and the patient activity. Large bearing diameter implants have been recently introduced to increase joint stability. Sliding distance increases proportional to the bearing radius which has led to some concerns regarding increased wear in larger bearings. However, in vitro wear studies have not shown this trend. Increased bearing diameter also increases the sliding path curvature which this study has shown to cause a reduction in wear roughly proportional to the radius of the bearing. Therefore, the increase in wear due to sliding distance is offset by the reduction in wear caused by the sliding curvature resulting in no significant change in wear with increased bearing diameter. Curved sliding path causes a change in surface shear direction which has been shown to increase wear of polyethylene. This study confirms that increased cross-shear in the form of more angular change per linear sliding distance can increase wear of polyethylene exponentially

Radial head fractures with fragment displacement should be reduced and fixed, when classified as Mason II type injuries. We describe a method of arthroscopic fixation which is performed as a day case trauma surgery, and compare the results with a more traditional fixation approach, in a case controlled manner.

We prospectively reviewed six Mason II radial head fractures which were treated using an arthroscopic reduction and fixation technique. The technique allows the fracture to be mobilised, reduced, and anatomically fixed using headless screws. All arthroscopic surgeries were conducted as day-cases. We retrospectively collected age and sex matched cases of open reduction and fixation of Mason II fractures using headless screws.

The arthroscopic cases required less analgesia, shorter hospital admissions, and had fewer complications. The averaged final range of follow-up, at 1 year post-operation was 15 to 140 degrees in the arthroscopic group and 35 to 120 degrees in the open group. The Mayo Elbow Performance Score was 95/100 and 90/100 respectively. No acute complications were noted in the arthroscopic group, and a radial nerve neuropraxia [n=1], superficial wound infection [n=1], and loose screw [n=1]. Two patients of the arthroscopic group required secondary motion gaining operations [n=1 arthroscopic anterior capsulectomy for a fixed flexion contracture of 35 degrees, and n=1 loss of supination requiring and arthroscopic radial scar excision]. Three patients in the open group required secondary surgery [n=2 arthroscopic anterior capsulectomy for fixed flexion deformities, and n=1 arthroscopic anterior capsulectomy for fixed flexion deformities, and n=1 arthroscopic radial head excision for prominent screws, loss of forearm rotation, and radiocapitellar arthrosis pain].

The technique of arthroscopic fixation of Mason II radial head fractures appears to be valid, with respect to anatomical restoration of the fracture, minimal hospital admission, reduction in analgesia requirement, fewer complications, and a decreased need for secondary surgery.