METHODS

Ten healthy subjects were instrumented and asked to perform six motions commonly associated with hip dislocation, including picking up an object, squatting, and low-chair rising. Femur-to-pelvis relative motions were recorded throughout for flexion/extension, abduction/adduction, and internal/external rotation.

A previously reported custom, validated hip ROM three-dimensional simulator was utilized. The user imports implant models, and sets parameters for pelvic tilt, stem version, and specific motions as defined by the subjects. Acetabular cup orientations for abduction and anteversion combinations were chosen. The software was then used to compute minimum clearances or impingement between the components for any hip position.

Graphs for acetabular cup abduction vs. anteversion were generated using a tapered wedge stem with a 132º neck angle, a stem version of 15°, and a pelvic tilt of 0°. The only variable changed was femoral head size. Head sizes reviewed were 32mm, 36mm, and a Dual-Mobility liner with an effective head size of 42mm. All femoral head sizes can be used with a 50mm acetabular cup.

Introduction

Correct acetabular cup placement is a critical step to prevent dislocation in the total hip arthroplasty. There are many mechanical alignment devices available but they are usually only referencing to the body long axis and the table therefore are lack of accuracy. Recently more accurate guide was achieved by image or imageless hip navigation system. But they add more cost, steps and time. The purpose of this study was to find a simple acetabular cup placement plane by selcting bonny land marks. The plane was adjusted with a fixed value by comparing it to native acetabular orientation in CT constructed 3D pelvi.

Materials and Methods

Figure 1 lists the study groups (minimum of n = 5). All groups were tested with a 50 mm Trident PSL shell (Stryker Orthopaedics, NJ) and a Ti6Al4V trunnion. Metal-on-Metal specimens were custom designed and manufactured, and are not approved for clinical use. The remaining groups consisted of commercially available products (Stryker Orthopaedics, NJ).

All groups were tested with the shells in deformed and undeformed states.

Deformed Setup: A two-point relief configuration was created in a polyurethane foam block (Figure 2) with a density of 30 lb/ft³ to replicate shell deformation due to press-fit [2]. The blocks were machined to replicate the press-fit prescribed in the shell's surgical protocol. Each shell was assembled into the foam block by applying an axial force at 5 mm/min until it was completely seated.

Undeformed Setup: Each shell was assembled in a stainless steel block with a hemispherical cavity that resulted in a line-to-line fit with the shell OD.

Frictional torque was measured using a physiologically relevant test model [3]. In this model, the specimen block was placed in a fixture to simulate 50° abduction and 130° neck angle (Figure 2). A 2450N side load was applied and the femoral head underwent angular displacement of ± 20° for 100 cycles at 0.75 Hz. The articulating surfaces were lubricated with 25% Alpha Calf Fraction Serum.

Peak torque was observed towards the end or the beginning of each cycle where the velocity of the femoral head approaches 0 and the head changes direction. This torque is referred as maximum static frictional torque. Specimen groups were statistically compared with a single-factor ANOVA test and a Tukey post-hoc test at 95% confidence level. Paired t-tests were performed to compare individual groups in deformed and undeformed states.

Methods:

Experimental wear rates for low and highly crosslinked polyethylene hip liners were obtained from a previously reported conventional hip wear simulator study [5]. A finite element model of the wear simulation for this design was constructed to replicate experimental conditions and to compute the wear coefficients that matched the experimental wear rates. We have previous described out this method of validation for knee wear simulation studies [6,7]. This wear coefficient was used to predict wear in a Dual-Mobility hip component (Fig 1).

Dual mobility total hip arthroplasty components, Restoration ADM (Fig 1), with highly crosslinked acetabular liners were experimentally tested: the control group was subjected to wear testing using the ISO 14242-1 waveform on a hip wear simulator. The microseparation group was subjected to a nominal 0.8 mm lateral microseparation during the swing phase by engaging lateral force springs and reducing the swing phase vertical force.

Methods:

Test coupon samples made from alloys' wrought bar stocks were polished then tested. Each material had 4 groups (n³3) that were tested in 4 aqueous solutions: 1) pH 7.4, 2) pH 5, 3) pH 2, and 4) pH 2 with 1 M of H₂O₂. Potential dynamic polarization was conducted using a Princeton Applied Research VMC Potentiostat, with an Ag/AgCl electrode as the reference electrode, and it resembled method in ASTM F 2129-08. The samples were evaluated with Scanning Electron Microscope (SEM).

Methods

Specimen Description Table 1 lists tested study groups (n≥3). Metal-on-Metal specimens were custom designed and manufactured, and are not approved for clinical use. The remaining groups consisted of commercially available products (Stryker Orthopaedics, NJ).

Methods:

The study was conducted using two methods:

Computer simulation study

265 pelvic CT scans consisting of 158 males and 107 females were converted to virtual 3-dimensional bones. The average native acetabular diameter was 52.0 mm, SD = 4.0 mm (males in 52.4 mm, SD = 2.8 mm and 46.4 mm, SD = 2.6 mm in females). Images were analyzed by custom CT analytical software (SOMA™ V.3.2)¹ and over-sized reaming was simulated. Four distinct points, located in and around the acetabular margins, were used to determine the reamer sphere. Points 1, 2, 3 were located at the inferior and inferior-medial acetabular margins, and Point 4 was located superiorly and posteriorly in the acetabulum to simulate a bony defect in this location, Point 4 was placed at 10%, 20%, 30%, 40%, 50% and 60% of the distance from the superior – posterior margin of the acetabular rim to the sciatic notch to simulate bony defects of increasing size. (Figure 1)

Radiographical study

Retrospective chart review of patient records for all cementless acetabular revisions utilizing jumbo cups between January 1, 1998 and March 30, 2012 at UCFS (98 patients with 57 men, 41 women). Jumbo cups: ≥66 mm in males; <62 mm in females. Reaming was directed inferiorly to the level of the obturator foramen to place the inferior edge of the jumbo cup at the inferior acetabulum. To determine the vertical position of the hip center, a circle was first made around both the jumbo and the contralateral acetabular surfaces using Phillips iSite PACS software. The center of this circle was assumed to correspond to the “hip center”. The height of the hip center was estimated by measuring the height of a perpendicular line arising from the interteardrop line (TL) and ending at the hip center

Methods and Materials

CoCr femoral heads were tested with trunnions that were machined with both a large and small taper geometry, replicating commercially available stem taper designs, V40 (small) and C (large) (Table-1, Stryker Orthopaedics, NJ).

The femoral heads were assembled onto the trunnions with a 2 kN axial force. A multi-axis test frame (MTS Corp, MN) was used to test the head-trunnion combination by dynamically loading with a torque of ± 5Nm and a constant axial load of 2450N for 1000 cycles at 1.5 Hz (Figure 1). Samples were submerged in 25% diluted Alpha Calf Fraction Serum (Hyclone, UT). Upon completion of the dynamic test, a static torque to failure test was performed where the axial force of 2450N was maintained and the trunnion was rotated to 40° at a rate of 3°/sec.

The torque required to rotate the trunnion by 1° was determined for each specimen. Also, the torsional resistance, defined as change in torque/change in angle in the linear region of the torque-angular displacement data curve, was calculated for all the specimens. A limitation associated with the static test was that at 1° rotation it was difficult to differentiate between rotation of the trunnion inside the femoral head and physical twisting of the trunnion. Specimen groups were compared with a single-factor ANOVA test and a Tukey post hoc test at 95% confidence level.

Methods

The novel shortened tapered wedge stem geometry was designed based on a morphological study of 556 CT scans. We then selected 10 discrete femoral geometries of interest from the CT database, including champagne fluted and stove pipe femurs. The novel and the predicate stems were virtually implanted in the bones in ABAQUS CAE. A total of thirty FEA models were meshed with 4 nodes linear tetrahedral elements. Bone/implant interface properties was simulated with contact surface and a friction coefficient of 0.35. Initial stability of each stem/bone assembly was calculated using stair-climbing loading conditions. The overall initial stability of the HA coated surface was evaluated by comparing the mean rotational, vertical, gap-opening and total micromotion at the proximal bone/implant interface of the novel and predicate stem designs.

To characterize the fit of the stem designs we analyzed the ratio of a distal (60 mm below lesser trochanter) and a proximal (10 mm above lesser trochanter) cross section. A constant implantation height of 20 mm above the lesser trochanter was used. The fit of the stems was classified as Type 1 (proximal and distal engagement), Type 2 (proximal engagement only) and Type 3 (distal engagement only).

Results

The mean % micromotion of the HA coated surface greater than 50 mm was lowest at 40.2% (SD 11.5%) for the novel tapered wedge stem compared to the clinically successful predicate stem design (Accolade TMAZ, Stryker, Mahwah, USA) at 44.9% (SD 13.2%) and its shortened version at 48.5% (SD 9.0%) as shown in Figure 1. Improved initial stability of the new stem was also confirmed for rotational, vertical and gap-opening micromotion. However, there was no statistically significant difference.

The novel tapered stem design showed a well balanced proximal to distal ratio throughout the complete size range. The novel tapered stem design showed a reduced percentage of distal engagements (2.8%) compared to the predicate standard stem (17.2%). In the 40 to 60 year old male group the distal engagement for the standard stem increases (28.2%), whereas the distal engagements for the novel stem remains unchanged (1.3%).

METHODS

Test methods were based on a previous study [3]. Five stems of the standard tapered wedge design (Accolade, Stryker Orthopaedics, Mahwah, NJ) and the novel stem (Accolade II, Stryker Orthopaedics, Mahwah, NJ) were implanted into a homogenous set of 10 synthetic femora (Figure 1) utilizing large left fourth generation

composite femurs (Sawbones, Pacific Labs, Seattle, WA). The six degrees-of-freedom (6 DoF) motions of the implanted stems were recorded under short-cycle stair-climbing loads. Minimum head load was 0.15 kN and the maximum load varied between 3x Body Weights (BW) and 6 BW. Loading began with 100-cycles of “normal” 3 BW and was stepped up to 4 BW, 5 BW & 6 BW for 50-cycles each. Prior to each load increase, 50 cycles of 3 BW loading was applied. This strategy allowed a repeatable measure of cyclic stability after each higher load was applied.

The 6 DoF micromotion data, acquired during the repeated 3 BW loading segments, were reduced to four outcome measures: two stem migrations (retroversion and subsidence at minimum load) and two cyclic motions (cyclic retroversion and cyclic subsidence). Data were analyzed using repeated measures ANOVA with a single between-subjects factor (stem type) and repeated measures defined by load-step (3 BW, 4 BW, 5 BW 6 BW).