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).

Metal-on-metal hip resurfacing has been introduced recently, due to its potential advantages of biomechanics and biotribology. However, a number of problems have been identified from clinical retrievals, including significant elevation of wear when the implant is mal-positioned. Our hypothesis is that implant mal-position and micro-lateralisation can result in edge contact, leading to increases in wear. The aim of this study was to investigate the combined effect of cup position and micro-lateralisation on the contact mechanics of metal-on-metal hip resurfacing prosthesis, in particularly to identify conditions which resulted in edge contact. Finite element (FE) method was used. A generic metal-on-metal hip resurfacing prosthesis was modelled. The bearing diameters of the femoral head and acetabular cup components were 54.49mm and 54.6mm respectively, with a diametral clearance between the head and the cup of 0.11mm. The resurfacing components were implanted into a hemi-pelvic hip joint bone model and all the materials in the FE model were assumed to be homogenous, isotropic and linear elastic (Udofia et al 2007). The FE models consisted of approximately 80,000 elements, which were meshed in I-DEAS (Version 11, EDS, USA) and solved using ABAQUS (Version 6.7-1, Dassault Systèmes). For this study, the femoral component was fixed with an inclination angle of 45° and an anteversion angle of 10°. The orientation of the acetabular cup was varied, using inclination angles of 35° and 65°, and anteversion angles between −10° to 30°. Contact at the bearing surface between the cup and femoral head was modelled using frictionless surface-based elements, simulating a fully lubricated situation, as coefficients of friction less than 0.1 would not have appreciable effects on the predicted contact mechanics. The femoral component was fixed into the femur (except the guide pin) using PMMA cement with an average thickness of approximately 1mm. The other contact interfaces in the FE model (cup/acetabulum, cement/bone and cement/femoral component) were all assumed to be rigidly bonded. The hip joint model was loaded through a fixed resultant hip joint contact force of 3200N, and was applied through medial, anterior muscle forces and subtrochanteric forces to simulate the mid-to-terminal stance phase (approximately 30% – 50%) of the gait cycle (Bergmann et al., 1993). Micro-lateralisation was modelled through displacing the femoral head laterally, up to 0.5mm, relative the centre of the cup. Edge contact was detected once the inclination angle became greater than 65°. The effect of ante-version was to further shift the contact area towards the edge of the cup, nevertheless no edge contact was found for ante-version angles up to 25° and inclination angles below 55°. However, when the micro-lateralisation was introduced, edge contact was detected at a much smaller inclination angle. For example, even with a micro-lateralisation of 0.5 mm, edge contact occurred at an inclination angle of 45°. This study highlights the importance of surgical techniques on the contact mechanics and tribology of metal-on-metal hip resurfacing and potential outcome of these devices

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

Introduction: Laboratory simulator and clinical retrieval studies of metal-on-metal (MOM) total hip replacements have shown that the metallic alloy, the femoral head radius, the clearance between the acetabular cup and femoral head and the cup thickness can influence the contact mechanics, the lubrication and the wear of the articulation. MOM hip resurfacing procedures have received significant attention recently. The purpose of the present study was to compare the contact mechanics between a MOM hip resurfacing implant and a MOM total hip replacement under identical conditions. Materials and Methods: A 50mm diameter DUROM. TM. MOM hip resurfacing prosthesis and a 28mm diameter Metasul. TM. MOM bearing system (Centerpulse Orthopedics, a Zimmer Company, Winterthur, Switzerland) were investigated. All implants were manufactured from wrought-forged high carbon cobalt chromium alloy (Pro-tasul 21WF. TM. ). The diameters of the DUROM. TM. femoral head and acetabular cup were 50mm and 50.145mm respectively, and the corresponding wall thickness of the acetabular component was around 4mm. The diameters of the Metasul. TM. femoral head and acetabular cup were 28mm and 28.12mm. Three-dimensional finite element models were created to simulate the contact between the bearing surfaces of both the femoral head and the acetabular cup fixed to a three dimensional anatomically positioned pelvic and femoral bone consisting of both cortical (with 1mm thickness) and cancellous regions. The load applied to both models was 3200N. Results: The maximum contact pressure at the bearing surfaces was found to be around 22MPa for the DUROM. TM. and the contact area between the femoral and acetabular components was predicted to be 237mm. 2. For the Metasul. TM. bearing under identical conditions, the maximum contact pressure and the contact area predicted were approximately 47MPa and 74mm. 2. respectively. Discussion: A large reduction in the contact pressure, which should improve overall tribological performances, was noted for the DUROM. TM. hip resurfacing prosthesis, as compared with the Metasul. TM. bearing. The main reasons for this reduction were the large diameter of the articulation and the small acetabular cup thickness of the DUROM. TM. system. In contrast, the Metasul. TM. bearing has a smaller head diameter, and relies on a polyethylene backing underneath the metallic cup inlay to reduce the contact pressure at the articulating surfaces

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

Background: Increased contact stress with a femoral resurfacing prosthesis implanted in the medial femoral condyle and a non-functional meniscus is of concern for potential deleterious effects on tibiofemoral contact mechanics. Methods: Peak contact pressures were determined in seven fresh frozen human cadaveric specimens using a pressure sensitive sensor placed in the medial compartment above the menisci. A knee simulator was used to test each knee in static stance positions (5°/15°/30°/45°) and through 10 dynamic knee-flexion cycles (5°–45°) with single body weight ground reaction force (GRF) which was adjusted to the living body weight of the cadaver donor. All specimens were tested in three different conditions: Untreated knee (A); Flush implantation of a 20mm resurfacing prosthesis (HemiCAP. ®. ) in the weight bearing area of the medial femoral condyle (B); Complete radial tear at the posterior horn of the medial meniscus with the femoral resurfacing device in place (C). Results: On average, flush device implantation resulted in no statistically significant differences when compared to the untreated normal knee. The meniscal tear resulted in a significant increase of the mean maximum peak contact pressures by 63%, 57%, and 57% (all P ≤ 0.05) at 15°, 30° and 45° static stance positions and 78% (P ≤ 0.05) through the dynamic knee flexion cycle. No significant different maximum peak contact pressures were observed at 5° stance position. Conclusion: Possible effects of reduced meniscal tissue and biomechanical integrity of the meniscus must be considered in an in-vivo application of the resurfacing device

Syndesmotic ankle lesions involve disruption of the osseous tibiofibular mortise configuration as well as ligamentous structures stabilizing the ankle joint. Incomplete diagnosis and maltreatment of these injuries is frequent, resulting in chronic pain and progressive instability thus promoting development of ankle osteoarthritis in the long term. Although the pathogenesis is not fully understood, abnormal mechanics has been implicated as a principal determinant of ankle joint degeneration after syndesmotic ankle lesions. Therefore, the focus of this presentation will be on our recent development of a computationally efficient algorithm to calculate the contact pressure distribution in patients with a syndesmotic ankle lesion, enabling us to stratify the risk of OA development in the long term and thereby guiding patient treatment.

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

Metal-on-metal hip resurfacing arthroplasty is a conservative procedure that is becoming an increasingly popular option for young arthritic patients most likely to undergo a secondary procedure in their lifetime. The stability of the acetabular component is of particular concern in these patients who show an increased risk of failure of the cemented acetabular cups in conventional total hip replacements. The purpose of this study was to examine the initial stability of a cementless interference press-fit acetabular cup used in hip resurfacing arthroplasty and implanted into ‘normal’ versus poor quality bone. Also examined was the effect of the press-fit procedure on the contact mechanics at the cup-bone interface and between the cup and femoral head. A finite element (FE) model of the DUROM resurfacing (Zimmer GmbH) was created and implanted anatomically into the hip joint, which was loaded physiologically through muscle and subtrochanteric forces. The FE models included: a line-to-line, 1mm and 2mm interference press-fit cup. Also considered were two FE models based on the 1mm press-fit cups, in which the material properties of the cancellous and cortical bone tissues were reduced by 2 and 4 times, to represent a reduction in bone quality as seen with age or disease. Increasing the cup-bone interference resulted in a sig-nificant reduction in implant micromotion. All the pressfit models showed predicted cup-bone micromotion below 50 micrometers. This would ensure adequate initial stability and encourage secondary fixation through bone in-growth. The predicted acetabular stresses were found to increase with the amount of press-fit, however, there was no suggestion of a fracture. These stresses would further contribute to securing the cup. Reducing the bone quality showed an increase in the predicted micromotion and increased bone strain. Micromotion was below 50 micrometers, but the predicted compressive bone stresses, necessary for additional implant fixation, was reduced. This implied that poor quality bone would provide unsuitable support medium for the implant. The bearing surface contact mechanics were little affected by the amount of pressfitting

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

Aims. To fully quantify the effect of posterior tibial slope (PTS) angles on joint kinematics and contact mechanics of intact and anterior cruciate ligament-deficient (ACLD) knees during the gait cycle. Methods. In this controlled laboratory study, we developed an original multiscale subject-specific finite element musculoskeletal framework model and integrated it with the tibiofemoral and patellofemoral joints with high-fidelity joint motion representations, to investigate the effects of 2.5° increases in PTS angles on joint dynamics and contact mechanics during the gait cycle. Results. The ACL tensile force in the intact knee was significantly affected with increasing PTS angle. Considerable differences were observed in kinematics and initial posterior femoral translation between the intact and ACLD joints as the PTS angles increased by more than 2.5° (beyond 11.4°). Additionally, a higher contact stress was detected in the peripheral posterior horn areas of the menisci with increasing PTS angle during the gait cycle. The maximum tensile force on the horn of the medial meniscus increased from 73.9 N to 172.4 N in the ACLD joint with increasing PTS angles. Conclusion. Knee joint instability and larger loading on the medial meniscus were found on the ACLD knee even at a 2.5° increase in PTS angle (larger than 11.4°). Our biomechanical findings support recent clinical evidence of a high risk of failure of ACL reconstruction with steeper PTS and the necessity of ACL reconstruction, which would prevent meniscus tear and thus the development or progression of osteoarthritis. Cite this article: Bone Joint Res 2022;11(10):739–750

Summary. This study describes the use of a quasi-static, 6DOF knee loading simulator using cadaveric specimens. Muscle force profiles yield repeatable results. Intra-articular pressure and contact area are dependent on loading condition and ACL integrity. Introduction. Abnormal contact mechanics of the tibiofemoral joint is believed to influence the development and progression of joint derangements. As such, understanding the factors that regulate joint stability may provide insight into the underlying injury mechanisms. Muscle action is believed to be the most important factor since it is the only dynamic regulator of joint stability. Furthermore, abnormal muscle control has been experimentally linked to the development of OA [Herzog, 2007] and in vivo ACL strain [Fleming, 2001]. However, the individual contributions to knee joint contact mechanics remain unclear. Thus, the purpose of this study was to examine the effects of individual muscle contributions on the tibiofemoral contact mechanics using an in-vitro experimental protocol. Methodology. Contact mechanics of 6 fresh frozen cadaver knee specimens were evaluated using the UofO Oxford knee loading device. Various combinations of quadriceps-hamstring co-contraction ratios were applied to the knee while it was “suspended” between the hip and foot components of the device. Loads of six muscle groups were computed using a hill-type musculoskeletal model [Buchanan, 2004]. Simulated ground reaction forces were also applied to the knee to represent force profiles of weight acceptance during gait as it has been shown to produce peak knee joint force in the gait cycle [Shelburne et al., 2006]. For respective medial and lateral joint compartments, the mean contact area (MC-CA and LC-CA), mean contact pressure (MC-CP and LC-CP), peak pressure (MC-PP and LC-PP), and centre of force displacement (MC-COFD and LC-COFD) were determined using a 4011 piezoelectric sensor form Tekscan (Tekscan Inc. Boston, MA). Additionally, the ACL was resected and measurements were repeated. Pearson correlations (r) examined the reliability of measurements as well as the effect an ACL transection on articular loads. Results. Positive correlations were computed for the following: COFD with intact ACL (r=0.99), COFD with resected ACL (r=0.82), MC-COFD pre vs. post ACL- resection (0.91). Furthermore, preliminary results indicated a positive correlation between MC-CA and ACL integrity (r=0.97). Discussion. The repeatability of the measured dependant variables validates the use of the knee-loading device. Interestingly, contact mechanics are more variable post ACL resection for a given muscle loading condition, indicating a decrease in knee joint stability. Also, the COFD is dependent on the different ratios of muscle loads applied to the knee, which demonstrates the importance of muscle action to the modulation of contact forces

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).

Objective. Full-thickness cartilage defects are commonly found in symptomatic knee patients, and are associated with progressive cartilage degeneration. Although the risk of defect progression to degenerative osteoarthritis is multifactorial, articular cartilage defects change contact mechanics and the mechanical response of tissue adjacent to the defect. The objective of this study was to quantify changes in intra-tissue strain patterns occurring at the defect rim and opposing tissue in an experimental model mimicking in vivo cartilage-on-cartilage contact conditions. Methods. Macroscopically intact osteochondral explants with smooth surfaces were harvested form the femoral condyles of 9 months old bovine knees. Two groups were tested; reference group with intact cartilage (n=8) and defect group with a full thickness cylindrical defect (diameter 8 mm) in one cartilage surface from each pair (n=8). The explants with defect articular surface and the opposing intact cartilage were compressed at ∼0.33 times body weight (350N) during cycles of 2s loading followed by 1.4s unloading. In plane tissue deformations were measured using displacement encoded imaging with stimulated echoes (DENSE) on a 9.4T MRI scanner. A two-sample t-test was used to assess statistical significance (p<0.05) of differences in maximal Green-Lagrange strains between the defect, opposing surface and intact reference cartilage. Results. Strain levels were elevated in the cartilage neighbouring the defect rim and in the opposing articulating surface. Similar to intact cartilage, compressive and tensile strains presented a depth dependent variation. The maximal strains profiles were highest in the superficial zone and decreased with depth for all explants, except for the shear strains in the cartilage opposing the defect which were constant. The maximal tensile strain in the middle and superficial zone were significantly higher for the defect cartilage (3.97±1.99% and 4.52±2.04%) compared to the intact reference (1.91±1.13% and 2.53±1.27%), indicating that the defect edges are bulging towards the defect. The shear strains were significantly higher (∼1.5x) throughout cartilage depth of the defect rim compared to the intact reference cartilage. However, in the cartilage opposing the defect, shear strains were significantly lower (∼0.5x) compared to the intact cartilage representing less matrix distortion. No significant difference in maximal compressive strains were observed between the opposing intact and defect at all cartilage depths. Conclusions. Presence of isolated full thickness cartilage defects will affect the cartilage deformations. Even under pure compressive loading alone, the altered contact mechanics resulted in excessive strains at tissue adjacent to the defect potentially damaging the cartilage and inducing tissue degeneration

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