Joint hemiarthroplasty replaces one side of a synovial joint and is a viable alternative to total joint arthroplasty when one side of the joint remains healthy. Most hemiarthroplasty implants used in current clinical practice are made from stiff materials such as cobalt chrome or ceramic. The substitution of one side of a soft cartilage-on-cartilage articulation with a rigid implant often leads to damage of the opposing articular cartilage due to the resulting reductions in contact area and increases in cartilage stress. The improvement of post-operative hemiarthroplasty articular contact mechanics is of importance in advancing the performance and longevity of hemiarthroplasty. The purpose of the present study was to investigate the effect of hemiarthroplasty surface compliance on early in-vitro cartilage wear and joint contact mechanics. Cartilage wear tests were conducted using a six-station pin-on-plate apparatus. Pins were manufactured to have a hemispherical radius of curvature of 4.7 mm using either Bionate (DSM Biomedical) having varying compliances (80A [E=20MPa], 55D [E=35MPa], 75D [E=222MPa], n=6 for each), or ceramic (E=310GPa, n=5). Cartilage plugs were cored from fresh unfrozen bovine knee joints using a 20 mm hole saw and mounted in lubricant-containing chambers, with alpha calf serum diluted with phosphate buffer solution to a protein concentration of 17 g/L. The pins were loaded to 30N and given a stroke length of 10 mm for a total of 50,000 cycles at 1.2 Hz. Volumetric cartilage wear was assessed by comparing three-dimensional cartilage scans before and during wear testing. A two-way ANOVA was used for statistical analysis. To assess hemiarthroplasty joint contact mechanics, 3D finite element modelling (ABAQUS v6.12) was used to replicate the wear testing conditions. Cartilage was modeled using neo-Hookean hyper-elastic material properties. Contact area and peak contact stress were estimated. The more compliant Bionate 80A and 55D pins produced significantly less volumetric cartilage wear compared with the less compliant Bionate 75D and ceramic pins (p 0.05). In terms of joint contact mechanics, the more compliant materials (Bionate 80A and 55D) had significantly lower maximum contact stress levels compared to the less compliant Bionate 75D and ceramic pins (p < 0 .05). The results of this study show a relationship between hemiarthroplasty implant surface compliance and early in vitro cartilage wear, where the more compliant surfaces produced significantly lower amounts of cartilage wear. The results of the joint contact mechanics analysis showed that the more compliant hemiarthroplasty materials produced lower maximum cartilage contact stresses than the less compliant materials, likely related to the differences in wear observed. More compliant hemiarthroplasty surfaces may have the potential to improve post-operative cartilage contact mechanics by increasing the implant-cartilage contact area while reducing peak contact stress at the implant-cartilage interface, however, such materials must be resistant to surface fatigue and longer-term cartilage wear/damage must be assessed

Altered distal radioulnar joint contact (DRUJ) mechanics are thought to cause degenerative changes in the joint following injury. Much of the current research examining DRUJ arthrokinematics focuses on the effect of joint malalignment and resultant degenerative changes. Little is known regarding native cartilage contact mechanics in the distal radioulnar joint. Moreover, current techniques used to measure joint contact rely on invasive procedures and are limited to statically loaded positions. The purpose of this study was to examine native distal radioulnar joint contact mechanics during simulated active and passive forearm rotation using a non-invasive imaging approach. Testing was performed using 8 fresh frozen cadaveric specimens (6 men: 2 women, mean age 62 years) with no CT evidence of osteoarthritis. The specimens were thawed and surgically prepared for biomechanical testing by isolating the tendons of relevant muscles involved in forearm rotation. The humerus was then rigidly secured to a wrist simulator allowing for simulated active and passive forearm rotation. Three-dimensional (3D) cartilage surface reconstructions of the distal radius and ulna were created using volumetric data acquired from computed tomography after joint disarticulation. Using optically tracked motion data and 3D surface reconstructions, the relative position of the cartilage models was rendered and used to measure DRUJ cartilage contact mechanics. The results of this study indicate that contact area was maximal in the DRUJ at 10 degrees of supination (p=0.002). There was more contact area in supination than pronation for both active (p=0.005) and passive (p=0.027) forearm rotation. There was no statistically significant difference in the size of the DRUJ contact patch when comparing analogous rotation angles for simulated active and passive forearm motion (p=0.55). The contact centroid moved 10.5±2.6 mm volar along the volar-dorsal axis during simulated active supination. Along the proximal-distal axis, the contact centroid moved 5.7±2.4 mm proximal during simulated active supination. Using the technique employed in this study, it was possible to non-invasively examine joint cartilage contact mechanics of the distal radioulnar joint while undergoing simulated, continuous active and passive forearm rotation. Overall, there were higher contact area values in supination compared with pronation, with a peak at 10 degrees of supination. The contact centroid moved volarly and proximally with supination. There was no difference in the measured cartilage contact area when comparing active and passive forearm rotation. This study gives new insight into the changes in contact patterns at the native distal radioulnar joint during simulated forearm rotation, and has implications for increasing our understanding of altered joint contact mechanics in the setting of deformity

Introduction. Hemiarthroplasty is a treatment option for comminuted fractures and non-unions of the distal humerus. Unfortunately, the poor anatomical fit of off-the-shelf distal humeral hemiarthroplasty (DHH) implants can cause altered cartilage contact mechanics. The result is reduced contact area and higher cartilage stresses, thus subsequent cartilage erosion a concern. Previous studies have investigated reverse-engineered DHH implants which reproduce the shape of the distal humerus bone or cartilage at the articulation, but still failed to match native contact mechanics. In this study, design optimization was used to determine the optimal DHH implant shape. We hypothesized that patient-specific optimal implants will outperform population-optimized designs, and both will optimize simple reverse-engineered designs. Methods. The boney geometries of six elbow joints were created based on cadaver arm CT data using a semi-automatic threshold technique in 3D Slicer. CT scans were also obtained with the elbows denuded and disarticulated, such that the high contrast between hydrated cartilage and air could be exploited in order to reconstruct cartilage geometry. Using this 3D model data, finite element contact models were created for each elbow, where bones (distal humerus, proximal ulna and radius) were modelled as rigid surfaces covered by non-uniform thickness layers of cartilage. Cartilage was modelled as a Neo-Hookean hyperelastic material (K = 0.31 MPa, G = 0.37 MPa), and frictionless contact was assumed. In order to simulate hemiarthroplasty, the distal humerus cartilage surface was replaced by either a rigid surface in the shape of the subchondral bone (bone reverse engineered or BRE design), or a surface offset from the bone by some distance, which was defined parametrically and modified by an optimization algorithm. Simple flexion-extension with constant balanced muscle loads was simulated in ABAQUS (Fig 1), and resulting contact areas and contact stresses were calculated. For each specimen, the contact mechanics of the intact and DHH reconstructed joints were calculated. A design optimization algorithm in Matlab was used to determine the optimal offset distance which resulted in contact stress distributions on the ulna and radius which most closely resembled their intact conditions. This procedure was repeated in order to generate specimen-optimal offsets, as well as population-optimal offsets. Results. The population-optimal offset distance was 0.72 mm; whereas the specimen-optimal offsets ranged from 0.52 to 1.04 mm. Compared to the BRE design, which is effectively an offset distance of 0 mm, contact area generally increased at both the ulna (Fig 2) and radius (Fig 3) when either optimized design was used. On average, the specimen-optimal implant designs yielded only slightly larger contact areas than the population-optimal offsets, and only at mid-flexion (40–60 deg). Neither optimization strategy increased contact areas to those of the intact joint. Conclusions. Design optimization is a promising technique for improving patient-specific implants by offering customization in terms of contact mechanics, instead of simply reproducing osseous geometry. In this study, our models predict a large increase in contact area if optimal offsets are used when designing subject-specific DHH, and a population-optimal offset distance seems to be just as good as a subject-optimal offset. For any figures or tables, please contact authors directly (see Info & Metrics tab above).

Introduction. The process of wear and corrosion at the head-neck junction of a total hip replacement is initiated when the femoral head and stem are joined together during surgery. To date, the effects of the surface topography of the femoral head and metal stem on the contact mechanics during assembly and thus on tribology and fretting corrosion during service life of the implant are not well understood. Therefore, the objective of this study was to investigate the influence of the surface topography of the metal stem taper on contact mechanics and wear during assembly of the head-neck junction using Finite Element models. Materials and Methods. 2D axisymmetric Finite Element models were developed consisting of a simplified head-neck junction incorporating the surface topography of a threaded stem taper to investigate axial assembly with 1 kN. Subsequently, a base model and three modifications of the base model in terms of profile peak height and plateau width of the stem taper topography and femoral head taper angle were calculated. To account for the wear process during assembly a law based on the Archard equation was implemented. Femoral head was modeled as ceramic (linear-elastic), taper material was either modeled as titanium, stainless steel or cobalt-chromium (all elastic-plastic). Wear volume, contact area, taper subsidence, equivalent plastic strain, von Mises stress, engagement length and crevice width was analyzed. Results. Titanium tapers showed largest wear volume throughout all simulations, followed by stainless steel and cobalt-chromium. A larger head taper angle resulted in an increase of the wear volume for all taper materials while the increase of the plateau width resulted in a decrease of the wear volume. Taper subsidence, von Mises stress and equivalent plastic strain followed the same trends. Contact area was largest for the models with a large plateau width for all taper materials. Other taper parameters had little effect on contact area. A pure increase of the angular mismatch (AM) resulted in the strongest decrease of the engagement length, while a combined increase of the AM and plateau width showed only a moderate decrease. The smallest effect concerning the engagement length was found when a combined increase of the profile peak height and AM was simulated. Crevice width was largest for a pure increase of the AM and for a combined increase of the AM and profile peak height for all taper materials. Discussion. This study showed that depending on the surface topography and material of the stem taper, wear and taper mechanics during assembly could be affected. For the examined surface topographies wear is distinctively elevated by increasing the AM and the profile peak height due to the resulting higher mechanical loading. More parameter studies under in vivo loading and the study of other taper surface parameters like the peak-to-peak distance have to be conducted to get a deeper insight into taper mechanics and wear effects. However, this study demonstrates the importance of good manufacturing practice of components for hip replacement systems to guarantee reproducible taper mechanics. For any figures or tables, please contact authors directly

Purpose. Capitellum hemiarthroplasty is an emerging concept. The current metallic capitellar implants have spherical surface shapes, but the native capitellum is not spherical. This study evaluated the effect of capitellar implant shape on the contact mechanics of the radiocapitellar joint when articulating with the native radial head. Method. Eight paired radii and humeri were potted in a custom jig. Articular casts were made with medium-viscosity resin while 85 N of axial load was applied to the reduced radiocapitellar joint at 0, 45, and 90 of elbow flexion, and at neutral, 50 pronation and 50 supination at each flexion angle. The native radiocapitellar articulation was compared to capitellar hemiarthroplasties of two surface designs (anatomical and spherical). Contact area and shape (circularity) were determined. Circularity was defined as the ratio of the minor axis and major axis of the shape. Results. At 0 of flexion, the anatomical hemiarthroplasty had a contact area of 52–70% that of the native articulation (p=0.03), while the spherical hemiarthroplasty had a contact area 40–42% that of the native articulation (p=0.003). At 45 of flexion, both hemiarthroplasties displayed contact area <53% that of the native joint (p<0.007). At 90 of flexion, the hemiarthroplasties had contact areas ranging from 40–70% that of the native articulation (p=0.1). The two capitellar implants had similar contact areas at all flexion angles tested (p>0.05). The contact shape of the native radiocapitellar articulation was ellipsoid, with a range of circularity values from 0.530.19 to 0.720.16, depending on the flexion and rotation angle. At 0 and 90 flexion, there was no difference in contact shape between the native articulation, the anatomical, or spherical implant (p>0.05). At 45 flexion, the anatomical implant contact was less circular than either the native articulation (p=0.006) or the spherical hemiarthroplasty (p=0.002). Conclusion. Metallic capitellar hemiarthroplasty causes a significant reduction in contact area at 0 and 45 elbow flexion, which may have important long-term implications for wear of the radial head cartilage. This reduction is similar to previous reports, which have evaluated the effect of metallic radial head hemiarthroplasty articulating with the native capitellum. More compliant alternative materials are needed to improve the contact characteristics of metallic capitellar hemiarthroplasties. Although the anatomical hemiarthroplasty was created from a detailed morphological study of the capitellum, the anatomical implant failed to completely reproduce the contact native shape. The theoretical advantages of a more anatomical capitellar implant shape may not be realized clinically, suggesting a spherical implant, which is easier to manufacture and implant, may be adequate for patient application. Further studies are required to delineate the effect of this altered contact morphology on implant function and radial head wear in-vivo

Introduction. Total knee arthroplasty (TKA) is a well proven surgical procedure. Squat and gait motions are common activities in daily life. However, squat motion is known as most dissatisfying motion in activities in daily life after total knee arthroplasty (TKA). Dissatisfaction after TKA might refer to muscle co-contraction between quadriceps and hamstrings. The purposed of this study was to develop squat and gait simulation model and analyses the contact mechanics and quadriceps and hamstring muscle stability. We hypothesized that squat model shows larger contact forces and lower hamstring to quadriceps force ratio than gait model. Materials and Methods. Squat motion and gait model were simulated in musculoskeletal simulation software (AnyBody Modeling System, AnyBody Technology, Denmark). Subject-specific bone models used in the simulation were reconstructed from CT images by Mimics (Materialize, Belgium). The lower extremity model was constructed with pelvis, femur, tibia, foot segments and total knee replacement components: femoral component, tibial insert, tibial tray, and patella component [Fig.1]. The muscle model was consisted of 160 muscle elements. The TKR components used in this study are PS-type LOSPA Primary Knee System (Corentec Co., Ltd, Republic of Korea). Force-dependent kinematics method was used in the simulation. The model was simulated to squat from 15° to 100° knee flexion, in 100 frames. Gait simulation model was based on motion capture and force-plate system. Motion capture and force-plate data were from grand challenge competition dataset. Results / Discussion. Patellofemoral contact forces ranged from 0.18 to 3.78 percent body weight (%BW) and from 0.00 to 1.36 %BW during squat motion and gait cycle, respectively. Patellofemoral contact forces calculated at 30°, 60°, and 90° flexion during squat motion were 0.53, 1.93, and 3.22 %BW, respectively. Wallace et al. also reported patellofemoral contact forces at 30°, 60°, and 90° flexion, which were 0.31, 1.33, 2.45 %BW during squat motion. Our results showed similar results from other studies, however the squat model overestimated the patellofemoral contact forces. Contact stiffness in the simulation model might affected the overestimated contact forces. Hamstring to quadriceps force ratio ranged from 0.32 to 1.88 for squat model, and from 0.00 to 2.54 for gait model. As our hypothesis, squat motion showed larger patellofemoral contact forces. Also, mean hamstring to quadriceps force ratio of squat model were about half than the mean hamstring to quadriceps force ratio of gait model. From the results, possibility exists that unbalanced force of quadriceps and hamstring can affect dissatisfaction after TKA while squat motion is the most dissatisfying motion after TKA. However, muscle stability is not the only factor that can affect dissatisfaction after TKA. In future study, more biomechanical parameters should be evaluated to find meaningful dissatisfying factor after TKA. Conclusion. In conclusion, TKA musculoskeletal models of squat and gait motion were constructed and patellofemoral contact force / hamstring to quadriceps force ratio were evaluated. Patellofemoral mechanics were validated by comparison of previous study. Additional studies are needed to find dissatisfying factor after TKA

Introduction. Malrotation of the femoral component is a cause of patellofemoral maltracking after TKA. Its precise effect on the patellofemoral (PF) mechanics has not been well quantified. The aim of this study was to investigate the effect of malrotation of the femoral component on PF initial contact area, initial contact pressure and wear after 4 million full gait cycles in TKA using a knee simulator. Moreover, the influence of the counterface material (CoCr or OxZr) on PF wear was also investigated. Materials & Methods. Femoral components (FCs) were cemented onto specially designed fixtures, allowing positioning of the FC in different angles of axial rotation. Patellar buttons and FCs were then mounted in a Prosim knee simulator. Patellofemoral contact mechanics. Seven axial rotation configurations were tested: neutral (FC parallel to the epicondylar axis), 2.5° endo- and exorotation, 5° endo- and exorotation and 7.5° endo- and exorotation. Patellar contact location, contact area and contact pressure were measured dynamically during 20 gait cycles with a Tekscan sensor covering the patella collecting data at a rate of 100 frames per second. Patellofemoral wear. For three alignments (neutral, 5° endo- and exorotation), a PF wear test of 4 million cycles in bovine serum (diluted to 40%) was done with three CoCr and three OxZr components on conventional ultra-high molecular weight polyethylene (UHMWPE, density: 0.93mg/mm. 3. ). Every 0.5 million cycles the test lubricant was replaced, the patellar samples were cleaned and dried and polyethylene wear was measured gravimetrically. A linear regression model was used to calculate the wear rate of each patellar sample. Aggregate wear rates were determined for each test condition by pooling the measurements of all three patellar samples. Results. For all six endorotation and exorotation configurations, the contact area was significantly lower and the contact pressure significantly higher than the neutral position (p < 0.001, Figs 1 and 2). In the patellofemoral wear test, the highest average wear rate was found in the group of endorotated CoCr femoral components (0.54 mm. 3. /Mcycle), but this is still only 11% of a typical tibiofemoral wear rate with the same CoCr component (5 mm. 3. /Mcycle). The following trends in the average wear rates could be observed: the average wear rate for CoCr (0.34 mm. 3. /Mcycle) was higher than for OxZr (0.19 mm. 3. /Mcycle) and the average wear rate for 5° endorotation (0.35 mm. 3. /Mcycle) was higher than for 5° exorotation (0.21 mm. 3. /Mcycle) and neutral alignment (0.23 mm. 3. /Mcycle) (Figs 3 and 4). None of these differences reached statistical significance (p=0.05), though. Discussion. Our results indicate that both internally and externally malrotated femoral components significantly decrease contact areas and significantly increase contact pressures in the patellofemoral joint. These significant changes in contact pressure didn't translate in significant changes in wear, however. Overall, patellofemoral wear is very small compared to tibiofemoral wear, in all the configurations that we investigated. Based on our results, we can conclude that clinical problems with patellar maltracking after femoral component malrotation seem not to be related to increased wear, but rather to pain and patellar instability

Scaphoid fractures are a common injury accounting for more than 58% of all carpal bone fractures(1,2). Biomechanical studies have suggested that scaphoid mal-union may lead to altered carpal contact mechanics causing decreased motion, pain and arthritis(1,2). The severity of mal-union required to cause deleterious effects has yet to be established. This limits the ability to define surgical indications or impacts on prevention of posttraumatic arthritis. Computed tomography has been shown to be a useful in determining the 3D implications of altered bony alignment on the joint contact mechanics of surrounding joints. The objective of this study was to report mid-term follow-up image-based outcomes of patients with scaphoid mal-unions to determine the extent to which arthritic changes and decreased joint space is present after a minimum of 4 years following fracture. Participants (n=14) who had previously presented with a mal-united scaphoid fracture (indicated by a Height:Length Ratio >0.6) between November 2005 and November 2013 were identified and contacted. A short-arm thumb spica case was used to treat X patients and X required surgical management. Baseline and follow-up CT images, were performed with the wrist in radial deviation and positioned such that the long axis of the scaphoid was perpendicular to the axis of the scanner. Three-dimensional inter-bone distance (joint space), a measure of joint congruency and 3D alignment, was quantified from reconstructed CT bone models of the distal radius, scaphoid, lunate, capitate, trapezium and trapezoid from both the baseline and follow-up scans(3). Repeated measures ANOVA was used to detect differences in contact area (mm2) between baseline and follow-up CT's for the radioscaphoid, scaphocapitate and scaphotrapezium-trapezoid joint. The average age of participants was 43.1 years (16–64 years old). There was significant loss of joint space, indicated by a greater joint contact area 3–4 years post fracture, between baseline and follow-up reconstruction models, at the scaphocapitate (mean difference: 21.5±146mm2, p=0.007) and scaphotrapezoid joints (mean difference: 18.4 ±28.6mm2, 0.042). Significant differences in the measured contact area was not found for the radioscaphoid (0.153) and scaphotrapezium joints (0.72). Additionally, the scaphoid, qualitatively, appears to track in the vorsal direction in the majority of patients following fracture. Increased joint contact area in the scaphocapitate and scaphotrapezoid joint 3–4 years following fracture results from decreased 3D joint space and overall narrowing. Joint space narrowing, while not significantly different for all joints examined, was reduced for all joints surrounding the scaphoid. Decreased joint space and increased contact area detectable within this short interval might be suggestive of a trajectory for developing arthritis in the longer term, and illustrates the potential value of these measures for early detection. Longer term follow-up and correlation to clinical outcomes are needed to determine the importance of early joint space narrowing, and to identify those most at risk

Pain and disability following wrist trauma are highly prevalent, however the mechanisms underlying painare highly unknown. Recent studies in the knee have demonstrated that altered joint contact may induce changes to the subchondral bone density and associated pain following trauma, due to the vascularity of the subchondral bone. In order to examine these changes, a depth-specific imaging technique using quantitative computed tomography (QCT) has been used. We've demonstrated the utility of QCT in measuring vBMD according to static jointcontact and found differences invBMD between healthy and previously injured wrists. However, analyzing a static joint in a neutral position is not necessarily indicative of higher or lower vBMD. Therefore, the purposeof this study is to explore the relationship between subchondral vBMDand kinematic joint contact using the same imaging technique. To demonstrate the relationship between kinematic joint contact and subchondral vBMDusing QCT, we analyzed the wrists of n = 10 participants (n = 5 healthy and n = 5 with previous wrist trauma). Participantsunderwent 4DCT scans while performing flexion to extension to estimate radiocarpal (specifically the radiolunate (RL) and radioscaphoid (RS)) joint contact area (JCa) between the articulating surfaces. The participantsalso underwent a static CT scan accompanied by a calibration phantom with known material densities that was used to estimate subchondral vBMDof the distal radius. Joint contact is measured by calculatinginter-bone distances (mm2) using a previously validated algorithm. Subchondral vBMD is presented using mean vBMD (mg/K2HPO4) at three normalized depths from the subchondral surface (0 to 2.5, 2.5 to 5 and 5 to 7.5 mm) of the distal radius. The participants in the healthy cohort demonstrated a larger JCa in the RS joint during both extension and flexion, while the trauma cohort demonstrated a larger JCa in the RL during extension and flexion. With regards to vBMD, the healthy cohort demonstrated a higher vBMD for all three normalized depths from the subchondral surface when compared to the trauma cohort. Results from our preliminary analysis demonstrate that in the RL joint specifically, a larger JCa throughout flexion and extension was associated with an overall lower vBMD across all three normalized layers. Potential reasoning behind this association could be that following wrist trauma, altered joint contact mechanics due to pathological changes (for example, musculoskeletal trauma), has led to overloading in the RL region. The overloading on this specific region may have led to a decrease in the underlying vBMD when compared to a healthy wrist. However, we are unable to conclude if this is a momentary decrease in vBMD that could be associated with the acute healing phase following trauma given that our analysis is cross-sectional. Therefore, future work should aim to analyze kinematic JCa and vBMD longitudinally to better understand how changes in kinematic JCa over time, and how the healing process following wrist trauma, impacts the underlying subchondral bone in the acute and longitudinal phases of recovery

Introduction. Total ankle replacement (TAR) is surgically complex; malalignment can arise due to surgical technique or failure to correct natural varus/valgus malalignment. Across joint replacement, malalignment has been associated with pain, component edge loading, increased wear and higher failure rates. Good component alignment is considered instrumental for long term TAR success. The conforming surface geometry of mobile bearing TARs leaves no freedom for coronal plane malalignment. The aim of this study was to investigate the biomechanical effect of coronal alignment on a mobile bearing TAR. Methods. Three TARs (Zenith, Corin Group) were tested under five coronal malalignment angles from 0–10° in a single station electromechanical knee simulator applying a typical ankle gait profile. As swing phase load is critical to TAR contact mechanics but will vary depending on the joint laxity. Swing loads of 100N, 300N and 500N were investigated. A positive control test with a swing load of 1000N was also studied, and was expected to eliminate the majority of lift off effects. Under each condition, the version was allowed to move freely while tests were performed, and the version profile under each alignment angle was recorded. Each test was carried out for 600 cycles in 25% bovine serum. Under the same loading conditions, but without lubrication, a Tekscan sensor recorded data from two cycles to assess the change in contact pressure and area at the five coronal angles. Results. Across the three TARs the effect of the swing phase load varied the biomechanics with a similar pattern. The high swing load of 1000N eliminates the majority of version while with 100N swing loads the TAR abducts for the length of the swing phase only realigning when the force increases, the extent dependent on the malalignment angle. At both 300N and 500N swing loads there is an oscillation apparent which changes the contact mechanics. The Tekscan results (Figure 1b) show changes in the contact area at three points in the load cycle; swing, the lower peak and the peak load (Figure 1a). With any degree of malalignment, component lift off is highly prevalent under lower swing phase loads of 100–300N. As the swing load is increased, this effect is only noticeable at greater malalignment angles. Discussion. The observed component lift off results in edge loading and peak pressures occurring at the insert edges. When the insert is 10 degrees coronally malaligned and the insert is brought fully into contact, the peak pressure reaches 16–18MPa, a pressure similar to the yield stress of polyethylene. The high contact pressures will likely elevate the wear and may increase the potential for polyethylene failure. Conclusion. Biomechanical testing has shown the malalignment of a total ankle replacement combined with the joint tension may change the contact mechanics and potentially increase wear. For any figures or tables, please contact authors directly (see Info & Metrics tab above).

Introduction:. Most cases of hip osteoarthritis (OA) are believed to be caused by alterations in joint contact mechanics resulting from pathomorphologies such as acetabular dysplasia and acetabular retroversion. Over the past 13 years, our research group has focused on developing approaches for patient-specific modeling of cartilage and labrum in the human hip, and applying these approaches to study hip pathomorphology. The long term objective is to improve the understanding of the etiology of OA related to hip pathomorphology, and to improve diagnosis and treatment. The objectives of this presentation are to provide a summary of our subject-specific modeling approach, and to describe the results of our analysis of hips from three populations of subjects: normal, traditional dysplastic, and retroverted. Methods:. A combined experimental and computational protocol was used to investigate contact mechanics in ten normal subjects (normal center edge angles (CEA), no history of hip pain), ten subjects with hip pain secondary to acetabular dysplasia (CEA less than 25°), and ten patients with a radiographic crossover sign, pain and clinical exams consistent with acetabular retroversion. CT arthrography was used to image cartilage and bone. Volumetric image data were segmented and discretized, and subject-specific finite element models were produced using validated methods [Fig. 1]. Boundary and loading conditions were obtained from instrumented implant and gait data. Contact mechanics were evaluated on the acetabular cartilage and labrum. Labrum contact area and peak contact stress were evaluated. Cartilage contact area, peak and average contact stress were evaluated in six anatomical regions in the acetabulum. Results:. Hip contact patterns were subject-specific, but distinct patterns emerged in the groups. Dysplastic hips had a larger contact area in the lateral region of the acetabulum, while normal hips demonstrated a more distributed contact pattern. The labrum in dysplastic hips supported significantly more load than the labrum in normal hips in all activities [Fig. 2]. Contact in retroverted hips tended to be focused medially and superiorly [Fig. 3]. Retroverted subjects had smaller contact stress and area in most regions. Discussion:. The differences in labrum mechanics between the normal and dysplastic groups provide clear support for the mechanical importance of the acetabular labrum in dysplastic hips. There were only minor differences in cartilage contact stress and area between normal and dysplastic groups, because of a lateral shift in the location of contact and subsequent loading on the acetabular labra in the dysplastic hips. The larger labrum load support and contact area in dysplastic hips indicates that the labrum compensates for the shallow acetabula. Clinically, this may account for the pattern of OA onset in dysplastic hips. The results for the retroverted group do not support the commonly held belief that concentrated posterior loading in retroverted hips leads OA because there were lower contact stresses and areas in the posterior regions of retroverted hips. Further, these results suggest that rim trimming may be appropriate for retroverted hips. The preferred surgery likely depends on subtle patient specific aspects of hip pathoanotomy in both retroverted and dysplastic hips

Introduction. Current methodologies for designing and validating existing THA systems can be expensive and time-consuming. A validated mathematical model provides an alternative solution with immediate predictions of contact mechanics and an understanding of potential adverse effects. The objective of this study is to demonstrate the value of a validated forward solution mathematical model of the hip that can offer kinematic results similar to fluoroscopy and forces similar to telemetric implants. Methods. This model is a forward solution dynamic model of the hip that incorporates the muscles at the hip, the hip capsule, and the ability to modify implant position, orientation, and surgical technique. Muscle forces are simulated to drive the motion, and a unique contact detection algorithm allows for virtual implantation of components in any orientation. Patient-specific data was input into the model for a telemetric subject and for a fluoroscopic subject. Results. For both stance and swing phase, the model predicted similar patterns and magnitudes compared to telemetry (forces) and fluoroscopy (kinematics). During stance phase, the model predicts 2.5 xBW of maximum hip force while telemetry predicts 2.3 xBW, yielding 8.7% error (Figure 1a). During swing phase, the model predicts 1.1 xBW maximum hip force, similar to telemetry (Figure 1b). During stance phase, the model predicts 1.3mm of hip separation (sliding) compared to 1.6mm for fluoroscopy, yielding 18.8% error (Figure 1c). During swing phase, the model predicts 1.9mm of separation compared to 1.7mm for fluoroscopy, yielding 11.8% error (Figure 1d). The model was also used to assess component placement, version, and optimal positioning compared to live surgery, producing very promising results. Conclusion. The model has proven accurate in predicting kinematics and forces. Therefore, forward solution mathematical modeling can be used to efficiently evaluate new component designs, positioning and technique differences, patient-specific scenarios, and any specific contribution towards THA outcomes that cannot be controlled in vivo. For any figures or tables, please contact the authors directly

Introduction. Current pre-clinical testing is performed using knee wear simulators with standardized walking profiles. Differences in generated damage patterns to those observed on retrieved liners have been explained with the absence of activities other than walking, less severe loading conditions, and a discrepancy in the simulator's tibiofemoral contact mechanics and in vivo knee excursion. While it has been recognized that rotational alignment of the knee may also drive the location and shape of wear scars, to the best of our knowledge this parameter has not been investigated in knee simulator studies. Methods. Here, we use patient specific gait as input to the simulation to approximate the patient specific contact mechanics. Kinematic and kinetic input data was obtained from gait analysis of a patient with a MGII (Zimmer Inc.) prosthesis at 11 years post-op using the point cluster technique for tibiofemoral kinematics, and a mathematical model for internal force calculations. Using the identical type of prosthesis on the simulator, wear tests were conducted in displacement mode on a closed-loop controlled station. Because x-rays of the patient suggested an internal rotation of the tibial tray, it was varied form 0–10° and the effect on location and wear scar dimension was assessed. Results were compared with the retrieved liner (obtained after 13 years in vivo). Results. The simulator inputs generated from the gait data were compared with ISO 14243–3 (Figure 1). The first contact force peak of the patient was significantly lower, while second contact force peak similar to ISO. There were minimal differences in the flexion/extension profiles. For stance phase, the anterior/posterior translation and internal/external rotation kinematics did not show similar patterns, but they did fall within similar ranges from zero. There was little similarity for the swing phase. The total wear scar area of the retrieval was measured to be 919.8 mm. 2. The average total wear scar of the tested components was 853.0 ± 59.8 mm. 2. (p= 26.28%) The outcome values of the tested components compared to the retrieval are shown in Figure 2. All offsets produced a smaller wear scar than the retrieval, but the 7° offset produced the closest area which was within 1 mm. 2. of the retrieval. The 7° offset also had the closed centroid offset angle, which was within 0.2° of the retrieval (Figure 3). On the retrieval, a small wear scar was observed on the anterior- medial aspect of the intracondylar eminence (not shown). Among the tested components, the 7° and 10° offsets recreated damage at this location. Discussion. Rotational alignment affected the wear scar size by as much as 15% in this study. Only, the 7° offset produced outcome values very similar to the retrieval, highlighting the importance of rotational mismatch for wear. It should be noted that ± 10° of rotational mismatch is clinically well tolerated [5] and therefore may occur frequently. All tested components had smaller wear scar areas than the retrieved liner. This suggests that other activities other than walking may have contributed to wear in vivo

Introduction. Dual mobility (DM) implants provide increased stability and range-of-motion through the use of a large diameter mobile liner articulating against an acetabular shell. However, recent studies have reported that such contemporary large head prostheses can directly impinge against the local soft tissues leading to anterior hip pain. To address this drawback, a novel Anatomically Contoured Dual Mobility (ACDM) liner was developed that maintains the outer spherical geometry over an approximately hemispherical portion and then contours inward the distal profile of the DM liner for soft-tissue relief. The extent of the inner profile encapsulating the small diameter head is increased to provide more coverage of the head and maintain the inner head pullout force. We hypothesized that the ACDM liner for soft-tissue relief would not affect retention of the small diameter inner head or liner-acetabular load-bearing contact area. Methods. A finite element model to evaluate head retention and contact mechanics was created with a rigid acetabular shell, a plastically deformable UHMWPE DM liner, a rigid femoral head and a rigid femoral stem. For the head retention analysis, the extent of head coverage (Fig. 1) was optimized to match the inner head pullout force of a conventional DM liner. Contact mechanics of a conventional DM and ACDM liner were analyzed at the maximum joint load of three activities: gait, deep-knee bend and chair sit. One set of simulations was completed with the mobile liner and head axes aligned and another with the axes mal-aligned so that the mobile liner rim was adjacent to the femoral stem neck and the potential area of contact was away from the mobile liner apex. This allowed a broader range of potential contact to be assessed including what was determined to be a worst-case alignment. Results. The head extraction force of the conventional mobile liner with 224° of coverage was 909 N (Fig. 2). The ACDM liner with 232° of head coverage and an 8 mm reduced radius had an inner head extraction force of 901 N. The contact simulation results were practically identical for the ACDM liner and the conventional DM liner. Contact stresses between the two designs differed by less than 3.6%. In most cases, contact area (Fig. 3) was virtually equal with a slightly higher contact area in the ACDM (∼6.3%). In two of three worst-case liner orientations, the contact area between the shell and liner was found to be slightly lower for the ACDM liner (∼15%). The contact area and its distribution in all cases were found to be sufficient. Conclusion. This study showed that the novel anatomically contoured dual mobility liner maintains adequate inner head retention and articular contact area. The ACDM liner matched the head retention capacity of the conventional DM by slightly increasing the coverage of the inner head. The ACDM articular contact area was comparable to that of a conventional DM and satisfactory in all cases. These results demonstrate that a soft-tissue friendly design can be achieved while providing adequate head retention and load-bearing contact area

Complications of the patellofemoral (PF) joint remain a common cause for revision of total knee replacements. PF complications, such as patellar maltracking, subluxation, dislocation and implant failure, have been linked to femoral and patellar component alignment. Computational analyses represent an efficient method for investigating the effects of patellar and femoral component alignment and loading on output measures related to long term clinical success (i.e. kinematics, contact mechanics) and can be utilized to make direct comparisons between common patellar component design types. Prior PF alignment studies have generally involved perturbing a single alignment parameter independently, without accounting for interaction effects between multiple parameters. The objective of the current study was to determine critical alignment parameters, and combinations of parameters, in three patellar component designs, and assess whether the critical parameters were design specific. A dynamic finite element (FE) model of an implanted PF joint was applied in conjunction with a 100-trial Monte Carlo probabilistic simulation to establish relationships between alignment and loading parameters and PF kinematics, contact mechanics and internal stresses (Figure 1). Seven parameters, including femoral internal-external (I-E) alignment, patellar I-E, flexion-extension (F∗∗∗∗∗E) and adduction-abduction (A-A) rotational alignment, and patellar medial-lateral (M-L) and superior-inferior (S-I) translational alignment, as well as percentage of the quadriceps load on the vastus medialis obliquus (VMO) tendon, were perturbed in the probabilistic analysis. Ten output parameters, including 6-DOF PF kinematics, peak PF contact pressure, contact area, peak von Mises stress and M-L force due to contact, were evaluated at 80 intervals during a simulated deep knee bend. Three types of patellar component designs were assessed; a dome-compatible patellar component (dome), a medialized dome-compatible patellar component (modified dome), and an anatomic component (anatomic). Model-predicted bounds at 5 and 95% confidence levels were determined for each output parameter throughout the range of femoral flexion (Figure 2). Traditional sensitivity analysis, in addition to a previously described coupled probabilistic and principal component analysis (probabilistic-PCA) approach, were applied to determine the relative importance of alignment and loading parameters to knee mechanics in each of the three designs. The dome component demonstrated the least amount of variation in contact mechanics and internal stresses, particularly in the 30–100° flexion range, with respect to alignment and loading variability. The modified dome had substantially reduced M-L contact force when compared with the dome. The anatomic design, while wide bounds of variability were predicted, had consistently greater contact area and lowered contact pressure than the dome and modified dome designs. The anatomic design also reproduced more natural sagittal plane patellar tilt than the other components. All three designs were most sensitivity to femoral I-E alignment. Thereafter, sensitivity to component alignment was design specific; for the anatomic component, the main alignment parameter was F-E, while for the domed components it was a combination of F-E and translation (M-L and S-I) (Figure 3). Understanding the relationships and design-specific dependencies between alignment parameters can add value to surgical pre-operative planning, and may help focus instrumentation design on those alignment parameters of primary concern

Background. Aseptic loosening is the leading cause of total knee arthroplasty (TKA) failure in the long term, of which osteolysis from polyethylene wear debris remains a problem that can limit the lifetime of TKA past the second decade. To help speed up design innovations, our goal was to develop a computational framework that could efficiently predict the effect of many sources of variability on TKA wear—including design, surgical, and patient variability. Methods. We developed a computational framework for predicting TKA contact mechanics and wear. The framework accepts multiple forms of input data: patient-specific, population-specific, or standardized motions and forces. CAD models are used to create the FEA mesh. An analytical wear model, calibrated from materials testing (wheel-on-flat) experiments, is fully integrated into the FEA process. Isight execution engine runs a design of experiments (DOE) analysis with an outcome variable, such as volumetric wear, to guide statistical model output. We report two DOE applications to test the utility of the computational framework for performing large variable studies in an efficient manner: one to test the sensitivity of TKA wear to the femoral center of rotation, and the second to test the sensitivity of TKA wear to gait input perturbations. Results. Using this method, we demonstrated that choice of femoral center of rotation matters, and that although volumetric wear was most sensitive to variation in flexion/extension peaks, no one kinematic factor dominates TKA volumetric wear variability. Conclusion. The two DOE applications represent initial first attempts to study variability in component alignment and input waveforms across large solution spaces. The computational framework will be most useful if it can be used in a TKA design setting, where new innovations can be tested as soon as they are developed to see if they are worthy of further mechanical testing

Introduction. Lipped liners have the potential to decrease the rate of revision for instability after total hip replacement since they increase the jumping distance in the direction of the lip. However, the elevated lip also may reduce the Range of Motion and may lead to early impingement of the femoral stem on the liner. It is unclear whether the use of a lipped liner has an impact on the level of lever-out moments or the contact stresses. Therefore, the aim of the current study was to calculate these values for lipped liners and compare these results to a conventional liner geometry. Materials and Methods. 3D Finite Element studies were conducted comparing a ceramic lipped liner prototype and a ceramic conventional liner both made from BIOLOX. ®. delta. The bearing diameter was 36 mm. To apply loading, a test taper made of titanium alloy was bonded to a femoral head, also made from BIOLOX. ®. delta. Titanium was modeled with a bilinear isotropic hardening law. For the bearing contact a coefficient of friction of both 0.09 or 0.3 was assumed to model a well and poorly lubricated system. Frictionless contact was modeled between taper and liner. Pre-load was varied between 500 N and 1500 N and applied along the taper axis. While keeping pre-load constant, lever-out force was applied perpendicular to the taper axis until subluxation occurred. Liners were fixed at the taper region. Lever-out moment, equivalent plastic strain and von Mises stress of the taper, bearing contact area and contact area between taper and liner was evaluated. Results. With increasing pre-load, larger lever-out moment, equivalent plastic strain, contact area between taper and liner and bearing contact area was found for both liner designs. However, von Mises stresses were nearly constant but slightly exceeded yield strength of titanium. For all evaluated parameters almost no differences were found between the liner designs. Lever-out moments were comparable for both designs ranging from 4.5–10.5 Nm for the lipped liner and 4.4–10.2 Nm for the conventional liner. The increase of the coefficient of friction strongly affected lever-out moments, equivalent plastic strain and contact area between taper and liner. The other parameters were not affected by varying the coefficient of friction. Discussion. This study did not find significant differences in the lever-out behavior of the lipped acetabular liner compared to the conventional liner design. The inner geometry of the lipped liner is comparable to the conventional liner inner geometry. Therefore, contact area showed no significant differences and contact mechanics are identical in the current setup leading to similar results of both liner designs. For both designs small plastic deformations in the contact point of the taper were found at the contact region between liner and taper. However, the investigated mechanical parameters did not differ between the two investigated liner types. For any figures or tables, please contact authors directly

INTRODUCTION. Component positioning of an artificial hip joint plays a key role in durability of implant. Despite the fact that a number of numerical, experimental and clinical studies have been carried out to investigate the effects of cup inclination on polyethylene wear, steep inclination has been reported to be associated with both high and low volumetric wear. Moreover, how cup anteversion affects wear and its interaction with inclination are still unclear. To address these knowledge gaps, in this study wear and contact mechanics of a hip joint under various cup positioning has been investigated by using FEA (Finite Element Analysis). METHOD. A Pinnacle. ®. Marathon neutral liner 36×56mm was chosen to model the wear and creep over 3 million cycles (mc) based on the Archard's law and modified time hardening model in ANSYS, respectively. Central composite design of response surface method was used to generate 9 FEA runs, where the operative inclination angles varied from 35º, 45º to 55º and operative anteversion angles differed amongst 0º, 15º and 30º. The range of cup angles were chosen to be similar to the Lewinnek “golden” safe zone for dislocation. The gait cycle as specified in ISO 14242-1 was applied to the femoral head. RESULT. Edge contact takes place with steep inclination and anteversion, resulting in a decreased contact area (wear contour) which intersects with cup rim, consequently increasing the contact pressure and wear penetration depth. According to the FEA result, maximum wear rate, contact pressure, wear depth and Mises stress were found in the model with highest inclination and anteversion angles. However, within the range of positioning angles studied, the predicted volumetric wear rates only vary from 13.68 mm. 3. /mc to 13.92 mm. 3. /mc, which are comparable to the corresponding wear rates measured in vitro, i.e. 12.9±3.8 mm. 3. /mc (inclination = 35º, anteversion = 0°) and 15.4±5.0 mm. 3. /mc (inclination = 55º, anteversion = 0°). DISCUSSION. Increased wear rate, contact pressure, wear depth and stress were predicted with a combination of steep cup inclination and anteversion, resulting from the edge contact condition. While the increase in stress and wear was modest within the range of positions tested, combinations of higher inclination and version are not uncommon and would be expected to result in more significant increases. Hence in a total hip replacement surgery it is desirable to ensure that the contact area lies well within the bearing area by avoiding excessive inclination and anteversion angles

Introduction. The number of young and more active patients requiring total knee replacement (TKR) is increasing. Preclinical evaluation and understanding the long-term failure of TKR is therefore important. Preclinical wear simulation of TKR is usually performed according to the International Standards Organization (ISO) recommendations. Two international standards for preclinical wear simulation of TKRs have been developed so that the anterior-posterior (AP) translation and internal-external (IE) rotation can be driven in either force or displacement control. However, the effects of using different control regimes on the kinematics and wear of the same TKR have not been investigated. The current study investigated the kinematics, contact mechanics and wear performance of a TKR when running under ISO force and displacement control standards using an experimentally validated computational model. Materials/Methods. Three different ISO control standards were investigated using a size C Sigma curved TKR (DePuy, UK), with moderately cross-linked UHMWPE curved inserts; ISO-14243-3-2004, ISO-14243-3-2014 and ISO- 14243-1-2009. Axial force and flexion-extension angle are common for the three standards. AP and IE motions are displacement controlled in ISO-14243-3-2004 and ISO-14243-3-2014, with the only difference being a reversal of AP polarity between the two standards, and are force controlled in ISO-14243-1-2009. The test setup and soft tissue constraints were defined in accordance with ISO recommendations. The wear model was based on the modification of Archard's law where the wear volume is defined as a function of contact area, sliding distance, cross-shear and contact stress. The simulation framework has been independently validated against experimental wear rates under three different standard and highly demanding daily activities (Abdelgaied et al. 2018). Results. Reversing AP in the displacement control ISO-2014, compared to ISO-2004, resulted in high contact stresses of more than 70 MPa in the posterior direction. The predicted AP and IE from the force control ISO-2009 were in different directions and magnitudes to ISO-2014 AP and IE. The predicted wear rates were 1.8, 2.0, and 5.5 [mm. 3. /mc] for ISO-14243-3-2004, ISO-14243-3-2014 and ISO-14243-1-2009 respectively. Discussion. Reversing AP in the displacement control ISO-2014, without revising the femoral centre of rotation, resulted in high stress edge loading in the posterior direction, due to femoral rollback, and more than 10% increase in wear rate compared to ISO-2004. The predicted AP and IE from the force control ISO-2009 had different polarities and magnitudes to the corresponding displacement control ISO-2014 AP and IE. In addition, the predicted wear rate under the force control ISO-2009 was more than double that measured under displacement control standards due to the increased AP and IE motions predicted under the force control standard. In addition to the previous validation of the model, the predicted wear rate under the force control ISO-2009 of 5.5 mm. 3. /mc was within the 95% confidence limits of the reported experimental wear rate for the same TKR of 4.71±1.29 mm. 3. /mc (Johnston et al. 2018) which gives more confidence in the model. Conclusion. The study showed significant differences between ISO force and displacement control standards and between ISO displacement standards with different AP polarities. These differences should therefore be considered when choosing a control regime for preclinical simulation of TKR

Introduction. The purpose of this study was to experimentally evaluate impingement and dislocation of total hip replacements while performing dynamic movements under physiological-like conditions. Therefore, a hardware-in-the-loop setup has been developed, in which a physical hip prosthesis actuated by an industrial robot interacts with an in situ-like environment mimicked by a musculoskeletal multibody simulation-model of the lower extremity. Methods. The multibody model of the musculoskeletal system comprised rigid bone segments of the lower right extremity, which were mutually linked by ideal joints, and a trunk. All bone geometries were reconstructed from a computed tomography set preserving anatomical landmarks. Inertia properties were identified based on anthropometric data and by correlating bone density to Hounsfield units. Relevant muscles were modeled as Hill-type elements, passive forces due to capsular tissue have been neglected. Motion data were captured from a healthy subject performing dislocation-associated movements and were fed to the musculoskeletal multibody model. Subsequently, the robot moved and loaded a commercially available total hip prosthesis and closed the loop by feeding the physical contact information back to the simulation model. In this manner, a comprehensive parameter study analyzing the impact of implant position and design, joint loading, soft tissue damage and bone resection was implemented. Results. The parameter study revealed a generally high dislocation risk for the seating-to-rising with adduction scenarios. Improper implant positioning or design could be compensated by adjusting prosthesis components correspondingly. Gluteal insufficiency or lower joint loading did not result in higher impingement or dislocation risk. However, severe malfunction of the artificial joint was found for proximal bone resection. Discussion. Previous testing setups ignored the impact of active muscles or relied on simplified contact mechanics. Herein, total hip replacement stability has been investigated experimentally by using a hardware-in-the-loop simulation. Thereby, several influencing factors such as implant position and design as well as soft tissue insufficiency and imbalance could be systematically evaluated with the goal to enhance joint stability