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ATTUNE ORS Abstracts

HOME

Introduction

Instability of the knee in total knee replacement patients has been reported during high

demand activities both through clinical observations and fluoroscopic evaluation.

Ploegmakers et al1, in an in vivo kinematic evaluation, cited implant design factors as a

determinant of knee instability. The objective of the current study was to compare anterior-

posterior (A-P) and internal-external (I-E) motions of the knee for four current TKR designs in

order to assess the influence of implant geometry on the inherent stability, motion, and

contact mechanics of the joint. Each design was assessed in two finite element (FE) models:

a laxity/stability test, and a full lower limb model during two high demand activities –

stepdown (high A-P force) and stance-phase gait (high I-E torque).

Methods

Implant design was quantified in terms of the

tibiofemoral (TF) conformity ratio, calculated

by dividing the femoral articular radius at 0°,

15°, 30° and 60° flexion by the radius of the

insert in the dwell point (Figure 1). To assess

tibiofemoral constraint, a finite element

model of the femoral component was

positioned in the dwell of the insert at 0°,

15°, 30° and 60° flexion under a compressive

load of 667N. A 5mm anterior translation or

10° internal rotation was applied to the

femur while the TF reaction force or torque

was measured.

Subsequently, dynamic simulations of

stepdown and stance-phase gait activities

were carried out in a FE model of the lower

limb (Figure 1). TF joint loads were taken from

in vivo telemetric data2 and a control system

was implemented to apply external loads at

the hip and ankle to create the

experimentally-measured loading condition

(compressive load, A-P force, I-E torque) at the

TF joint for each activity using the telemetric implant geometry.

The external loading condition was subsequently applied

directly in the model, and the simulation was carried out for

the four components, including current cruciate-retaining (CR)

and posterior-stabilizing (PS) designs from several

manufacturers. 6-DOF TF kinematics, medial and lateral

condyle lowest point, and contact mechanics were evaluated

for each design.

Results

Conformity ratios correlated well with laxity/constraint of the

components (r = 0.73 for translation tests; r = 0.78 for

rotation tests).

Trends during the dynamic activities were in

agreement with those predicted during the

laxity simulations; higher conformity

increased constraint and hence the loads

carried by the insert instead of the

surrounding soft tissue. These designs with

higher conformity had, in general terms,

higher contact area, and lower contact

pressure than the less conforming

components (Figure 2).

Both laxity tests and dynamic simulations

highlighted substantial variation in the

constraint provided by current implant

designs. The range of A-P and I-E motion for

the least constrained design was twice that

of the most constrained design during

dynamic activity (Figure 3).

Discussion

In the current analysis, each component was

analyzed under the same external loading

conditions with the same soft-tissue

representation, allowing for direct

comparison between components. Varying

ligament mechanics would alter the

magnitude of motions, but relative

performance of each implant would be

consistent.

Other factors, aside from geometry,

contribute to instability of the knee joint,

notably, ligamentous balance/tension. Some

knees, through natural mechanics or injury,

have a tendency towards instability.

Component designs with inherent geometric stability may aid in maintaining knee stability

during dynamic activity for these patients.

Significance

Understanding the variation in constraint provided by differing current implant designs may

aid clinicians in determining which type of implant is most appropriate, given the soft-tissue

quality of their patient, to provide adequate stability during activities of daily living while

maintaining range of motion.

THE INFLUENCE OF DESIGN ON

TKR MECHANICS DURING

ACTIVITIES OF DAILY LIVING

+1Fitzpatrick CK; 1,2Clary CW; 1Rullkoetter PJ

+1University of Denver, Denver, CO; 2DePuy Orthopedics, Inc., Warsaw, IN – [email protected]

Con

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Triathlon® NexGen® ATTUNE™ SIGMA®

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Figure 1: TF conformity ratios (left, top) and laxity

test reaction forces (left, middle) and moments (left,

bottom) to anterior and internal motions, respective-

ly – zero indicates post-cam impingement; FE model

of the lower limb (below); external loading profiles

implemented in the model to apply joint compressive

load, I-E torque and A-P force during gait and stepdown

activities (right)

hip load

ankleI-E

ankleflexion

Acknowledgements: This work was supported in

part by DePuy, a Johnson & Johnson company.

Poster #2034, from the ORS 2012 Annual Meeting,

February 2012, San Francisco, CA.

The third party trademarks used herein are the

trademarks of their respective owners.

Peak

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(M

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Triathlon® Nexgen® ATTUNE™ SIGMA®50

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Triathlon® NexGen® ATTUNE™SIGMA®

Figure 2: Contact area during stepdown (top left); peak

contact pressure for PS components during both ac-

tivities (top right); contact patch for PS components

at peak external torque during gait (center) and peak

posterior force during stepdown (bottom)

Figure 3: Medial and lateral A-P kinematics for each PS

(solid) and CR (dashed) component shown for gait (top)

and stepdown (bottom)

Stance-phase Gait Cycle (%)Late

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References

1. Ploegmakers et al., 2009, Knee 17:204-209.

2. Kutzner et al., 2010, J Biomech 43:2164-2173.

© DePuy Synthes Joint Reconstruction, a division of DOI 2013

0612-68-512 (Rev. 1) EO 3/13

www.depuysynthes.com

DePuy Orthopaedics, Inc.

700 Orthopaedic Drive

Warsaw, IN 46582

Telephone: (800) 366-8143

Fax: (800) 669-2530

ATTUNE ORS Abstracts HOME

Introduction

Fluoroscopic kinematic evaluation of total knee arthroplasty (TKA) has shown a sudden anterior shift of the tibiofemoral contact point, frequently of the medial femoral condyle1. It has been suggested this motion is tied to abrupt changes in the femoral sagittal radius of curvature (J-Curve) typical of traditional TKA. To evaluate the link between detailed implant geometry and joint mechanics, an experimental or computational model that effectively demonstrates the in vivo behavior is a necessity.

The purpose of the current study was to utilize a previously validated computational model of the Kansas knee simulator (KKS)2 to understand the influence of TKA geometry on the resulting joint mechanics and then as an iterative design-phase tool to develop implant geometry which improves dynamic mid-stance stability. To verify the predictions of the computational model, the new geometry was compared to an existing TKA in a cadaveric study utilizing the experimental simulator. This comparison enabled assessment of the accuracy of the computational model and illustrated whether the simulations were sensitive enough to appropriately differentiate subtle changes in implant design and the resulting kinematic patterns.

Methods

A previously validated specimen-specific finite element model of a cadaveric knee in the KKS2, including specimen-specific bony geometry and soft-tissue representations, was implanted with multiple prototype implant geometries. Design iterations implicated an abrupt change from the sagittal femoral distal radius to the posterior radius as responsible for the anterior slide seen in vivo. Based on this understanding, a gradually reducing sagittal femoral radius was developed and incorporated into the updated femoral design (Fig. 1).

Six cadaveric knees were implanted with a traditional multi-radius TKA design and mounted into the KKS3. A simulated deep knee bend (DKB) was performed on the knees between 10° and 100° flexion driven by a force applied to the quadriceps tendon to balance a body-weight force applied at the hip. The medial-lateral (M-L) translation and all rotations at the ankle were unconstrained. Subsequently, the traditional TKA was replaced with the updated TKA geometry, incorporating the gradually reducing sagittal femoral radius of curvature, and the cycle repeated. Knee motion was measured using an Optotrak 3020 (Northern Digital Inc., Waterloo, Ontario, Canada) and six-degree-of-freedom tibiofemoral kinematics described using a three-cylindrical open-chain model4. Additionally, the contact points between the insert and femoral component were approximated by identifying the lowest point on the femoral geometry along the superior-inferior (S-I) axis of the tibia.

Results

Both the computational and experimental simulators were able to identify key relationships between the implant shape and the contact mechanics, including the abrupt anterior slide of the femoral condyles of the traditional TKA at the transition from the distal to posterior sagittal radius of curvature (Fig. 3, left). In comparison, the gradually reducing femoral sagittal radius of curvature attenuated the anterior slide of the medial femoral condyle and led to a gradual posterior translation of the lateral condyle with knee flexion (Fig. 3, right). Although not statistically significant, the cadaveric knees on average experienced increased femoral rollback with the updated design.

Discussion

In vitro experimental and computational simulations are critical pre-clinical tools in the evaluation of new implant designs. The combined experimental and computational approach described here was able to relate subtle design changes in the sagittal femoral radius of curvature to A-P stability during a DKB and femoral rollback in flexion. While the models were able to identify and enable a solution to a clinically observed phenomenon, current and future work is focused on improving the fidelity and validation of the computational simulations to represent more sophisticated activities of daily living like gait, navigating stairs, and rising from a chair.

Significance

This study utilized computational and experimental knee simulations to identify the relationship between TKA implant shape and a clinically observed kinematic phenomenon and then enabled design changes to address the paradoxical motion.

IMPROVING DYNAMIC MID-STANCE STABILITY: AN EXPERIMENTAL AND FINITE ELEMENT STUDY1, 2Clary CW; 1Fitzpatrick CK; 3Maletsky LP; +1Rullkoetter PJ+1University of Denver, Denver, CO; 2DePuy Orthopedics, Inc., Warsaw, IN; 3University of Kansas, Lawrence, KS – [email protected]

Kansas Knee Simulator Computational Simulator

Figure 1: The Kansas knee simulator (left) and the computational representation of the simulator (right).

Traditional TKA (Multi-radius)

Abrubt radius change in mid-�exion

Updated TKA (Continuously Reducing Radius)

Gradual radius reduction through 90˚ �exion

Figure 2: Comparison of femoral sagittal curvatures for the traditional multi-radius TKA (left) and the updated design with a gradually reducing radius (right).

Traditional TKA(Multi-radius)

Updated TKA(Continuously Reducing Radius)

Knee Flexion (º) Knee Flexion (º)

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nsla

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Figure 3: A-P condylar translations measured in the KKS and predicted by the computational simulator for the PFC® SIGMA® System (left) and ATTUNE™ (right) Implants.

References

1. Dennis, CORR, 2003.

2. Balwin, J Biomech, in review.

3. Maletsky et al., J Biomech Eng, 2005.

4. Grood and Suntay, J Biomech Eng, 1983.

Acknowledgements: This work was supported in part by DePuy, a Johnson & Johnson company.

Poster #1044, from the ORS 2012 Annual Meeting, February 2012, San Francisco, CA.

© DePuy Synthes Joint Reconstruction, a division of DOI 2013 0612-62-512 (Rev. 1) EO 3/13


DePuy Orthopaedics, Inc.700 Orthopaedic DriveWarsaw, IN 46582Telephone: (800) 366-8143 Fax: (800) 669-2530

ATTUNE ORS Abstracts

HOME

IntroductionWhile recent focus on attaining deep flexion after total knee replacement (TKR) has focused on

the effects of femoral “rollback”1, the interplay between implant design, ligamentous constraint,

and maximum knee flexion remains unclear. In this study, the influence of implant shape, in

particular the shape of the posterior condyles and amount of femoral rollback, on MCL

elongation was assessed with a computational model of passive flexion and the results

substantiated through implantation into cadaveric specimens.MethodsThree prototype posterior stabilized rotating platform knee implants were formulated with different levels of mid-flexion femoral thickness, condylar height, and femoral rollback (Fig. 1). Implant 1 had an increased condyle height, decreased mid-flexion thickness, and increased rollback. Implant 2a had a neutral condyle height, neutral mid-flexion thickness, and neutral rollback. Implant 2b utilized the same femur as 2a, but with 1-mm less rollback. Implant 3 had a decreased condyle height, neutral mid-flexion thickness, and decreased rollback.

Each design was virtually implanted into a finite element (FE) model of the knee. The model included four non-linear spring elements

representing the anterior, central, posterior, and distal portions of the medial collateral

ligament (MCL), with the attachment site based on measured data from a cadaveric knee

implanted with the prototype designs. Passive flexion boundary conditions were applied to

the model, with the tibial component fixed in space while a compressive load of 500-N was

applied to the femur and flexed from 0° to 140° knee flexion with the remaining femoral

degrees of freedom unconstrained. Elongations of the MCL bundles were measured during

the simulation.

The prototype knee implants were manufactured and implanted into 8 cadaveric knees by eight

different orthopaedic surgeons. Four surgeons employed a “measured resection” technique while

the other four employed a “balanced gap” technique.2 After implantation, surgeons were blinded

to the implant in the knee and asked to quantify the MCL tension by palpation as either “loose”,

“normal”, or “tight” at 3 flexion angles (45°, 90°, and 120°) and to measure the maximum passive

flexion angle using a goniometer.ResultsAccording to the FE predictions, increased condylar height and increased femoral rollback led to increased elongation of the distal bundle of the MCL in deep flexion (Fig. 2). These predictions were corroborated by observations of half the surgeons (100% of the balanced gap surgeons) that implant 1 had an unacceptably tight MCL at 120° flexion (Fig. 3) and that fewer surgeons felt that implant 2a had a tighter MCL than implant 2b. On the contrary, while the model predicted that implant 1 had a more lax MCL at 45°, the surgeons were not able to distinguish this difference. Despite the observation that the MCL was unacceptably tight in implant 1 in flexion, there was not a significant difference in the measured terminal passive flexion.

DiscussionIn this study, a FE model was used to quantify changes in MCL elongation in deep flexion associated with implant shape. The predicted changes, on the order of 1%–2% of MCL elongation, were perceptible to orthopaedic surgeons performing a “balanced gap” technique, but not to surgeons performing a “measured resection” technique. While the differences in implant shape were subtle, the effect of these changes on the soft tissue function and the surgeon’s perception was clear. This is the first time such a correlation between implant shape and surgical technique has been documented.

Although it’s been shown that femoral rollback enables deep flexion by preventing bony

impingement of the insert on the posterior femoral bone1, this data suggests that too much

rollback coupled with increased condylar tip height can lead to excessive strain in the MCL in

deep flexion. While the cadaveric simulation did not illustrate a measureable change in

terminal flexion across designs, the surgeons almost unanimously preferred Implant 3,

leading to a new knee design (ATTUNE™ System, DePuy Orthopaedics, Inc., Warsaw, IN).SignificanceThis study clarifies the effect of implant shape on soft tissue function in deep flexion and the

interaction with surgical technique. Results of this work will enable implant designs and surgical

technique to improve high-flexion outcomes for patients receiving TKA.

MANAGEMENT OF MCL TENSION IN DEEP FLEXION: INFLUENCE OF IMPLANT SHAPE+1Clary, CW; 1Wyss, JG; 1Wright, AP; 1Bennett, TD; 1Auger, DD; 1Heldreth, MA

+1DePuy Orthopedics, Inc., Warsaw, IN – [email protected]

Figure 1: Three prototype implants with different fem-oral geometry. Table indicates level of condyle height, mid-flexion thickness, and rollback (+=increased, N=-neutral, -=decreased).

Mid-FlexionThickness

CondyleHeight

Height Thickness RollbackImp 1 + - +Imp 2a N N NImp 2b N N -Imp 3 - N -

References1. Banks et al., CORR, 2003.2. Dennis et sl., CORR, 2010.


MCL Elongation with Knee Flexion1.03

1.025

1.02

1.015

1.01

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Implant 1Implant 2a (Normal Rollback)Implant 2b (Reduced Rollback)Implant 3

0 20 40 60 80 100 120 140Flexion Angle (Deg)

Anterior MCLTight Distal MCL

Tight

MCL

Len

gth

(% In

itial

Len

gth)

Figure 2: Composite MCL elongation of the longest bundle of the MCL predicted by the FE model for the three prototype designs.

1

100

75

50

25

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f Su

rgeo

ns

(All)

% o

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(Bal

ance

d G

ap)

Implant:

45° Knee Flexion 90° Knee Flexion 120° Knee Flexion

2a 2b 3 1 2a 2b 3 1 2a 2b 3Figure 3: Surgeon assessments of MCL tension for the three prototypes (red=tight, green=normal, yellow=-loose) by all surgeons (top, n=8) and only balanced gap surgeons (below, n=4).




THE ATTUNE® KNEE SYSTEM: CONFERENCE ABSTRACTS

TABLE OF CONTENTS

ATTUNE GRADIUS™ Curve

Improving Dynamic Mid-Stance Stability: An Experimental and Finite Element Study

The Influence of Design on TKR Mechanics During Activities of Daily Living

Management of MCL Tension in Deep Flexion: Influence of Implant Shape

LOGICLOCK™ Tibial Base

Advanced Fixed Bearing TKA Locking Mechanism Minimizes Backside Micromotion

GLIDERIGHT™ Articulation

Comparison of Natural and Unresurfaced Patellofemoral Mechanics

The Effect of Surgical Variability and Patella Geometry on Extensor Efficiency in Total Knee Replacement

Materials and Sizing

Probabilistic Modeling of Femoral Component Overhang

Tibial Tray Design Factors Affecting Tibial Coverage after Total Knee Arthroplasty

The Effect of Tibial Tray Rotational Alignment on Asymmetry of the Resected Tibial Plateau

Extraction Study of Gamma-Irradiated UHMWPE Stabilized with a Hindered Phenol Antioxidant

Determination of Antioxidant Distribution in Powder and Molded UHMWPE Materials

Biocompatibility Study of Gamma-irradiated UHMWPE Stabilized with a Hindered-Phenol Antioxidant

Characterization of Gamma-Irradiated UHMWPE Stabilized with a Hindered Phenol Antioxidant

Wear Resistance

Multi-ADL Profiles in TKR Wear Testing Help Discriminate Between Wear Theories

Wear of a Total Knee Replacement with Antioxidant UHMWPE and Gradually Varying Sagittal Curvature

ATTUNE Knee Conference AbstractsHOME

Introduction

Fluoroscopic kinematic evaluation of total knee arthroplasty (TKA) has shown a sudden anterior shift of the tibiofemoral contact point, frequently of the medial femoral condyle1. It has been suggested this motion is tied to abrupt changes in the femoral sagittal radius of curvature (J-Curve) typical of traditional TKA. To evaluate the link between detailed implant geometry and joint mechanics, an experimental or computational model that effectively demonstrates the in vivo behavior is a necessity.

The purpose of the current study was to utilize a previously validated computational model of the Kansas knee simulator (KKS)2 to understand the influence of TKA geometry on the resulting joint mechanics and then as an iterative design-phase tool to develop implant geometry which improves dynamic mid-stance stability. To verify the predictions of the computational model, the new geometry was compared to an existing TKA in a cadaveric study utilizing the experimental simulator. This comparison enabled assessment of the accuracy of the computational model and illustrated whether the simulations were sensitive enough to appropriately differentiate subtle changes in implant design and the resulting kinematic patterns.

Methods

A previously validated specimen-specific finite element model of a cadaveric knee in the KKS2, including specimen-specific bony geometry and soft-tissue representations, was implanted with multiple prototype implant geometries. Design iterations implicated an abrupt change from the sagittal femoral distal radius to the posterior radius as responsible for the anterior slide seen in vivo. Based on this understanding, a gradually reducing sagittal femoral radius was developed and incorporated into the updated femoral design (Fig. 1).

Six cadaveric knees were implanted with a traditional multi-radius TKA design and mounted into the KKS3. A simulated deep knee bend (DKB) was performed on the knees between 10° and 100° flexion driven by a force applied to the quadriceps tendon to balance a body-weight force applied at the hip. The medial-lateral (M-L) translation and all rotations at the ankle were unconstrained. Subsequently, the traditional TKA was replaced with the updated TKA geometry, incorporating the gradually reducing sagittal femoral radius of curvature, and the cycle repeated. Knee motion was measured using an Optotrak 3020 (Northern Digital Inc., Waterloo, Ontario, Canada) and six-degree-of-freedom tibiofemoral kinematics described using a three-cylindrical open-chain model4. Additionally, the contact points between the insert and femoral component were approximated by identifying the lowest point on the femoral geometry along the superior-inferior (S-I) axis of the tibia.

Results

Both the computational and experimental simulators were able to identify key relationships between the implant shape and the contact mechanics, including the abrupt anterior slide of the femoral condyles of the traditional TKA at the transition from the distal to posterior sagittal radius of curvature (Fig. 3, left). In comparison, the gradually reducing femoral sagittal radius of curvature attenuated the anterior slide of the medial femoral condyle and led to a gradual posterior translation of the lateral condyle with knee flexion (Fig. 3, right). Although not statistically significant, the cadaveric knees on average experienced increased femoral rollback with the updated design.

Discussion

In vitro experimental and computational simulations are critical pre-clinical tools in the evaluation of new implant designs. The combined experimental and computational approach described here was able to relate subtle design changes in the sagittal femoral radius of curvature to A-P stability during a DKB and femoral rollback in flexion. While the models were able to identify and enable a solution to a clinically observed phenomenon, current and future work is focused on improving the fidelity and validation of the computational simulations to represent more sophisticated activities of daily living like gait, navigating stairs, and rising from a chair.

Significance


IMPROVING DYNAMIC MID-STANCE STABILITY: AN EXPERIMENTAL AND FINITE ELEMENT STUDY1, 2Clary CW; 1Fitzpatrick CK; 3Maletsky LP; +1Rullkoetter PJ

+1University of Denver, Denver, CO; 2DePuy Orthopedics, Inc., Warsaw, IN; 3University of Kansas, Lawrence, KS

Kansas Knee Simulator Computational Simulator

Figure 1: The Kansas knee simulator (left) and the computational representation of the simulator (right).

Traditional TKA (Multi-radius)

Abrubt radius change in mid-�exion

Updated TKA (Continuously Reducing Radius)

Gradual radius reduction through 90˚ �exion

Figure 2: Comparison of femoral sagittal curvatures for the traditional multi-radius TKA (left) and the updated design with a gradually reducing radius (right).

References

1. Dennis, CORR, 2003.

2. Balwin, J Biomech, in review.

3. Maletsky et al., J Biomech Eng, 2005.

4. Grood and Suntay, J Biomech Eng, 1983.






HOME

Traditional TKA(Multi-radius)

Updated TKA(Continuously Reducing Radius)

Knee Flexion (º) Knee Flexion (º)

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Figure 3: A-P condylar translations measured in the KKS and predicted by the computational simulator for the PFC® SIGMA® System (left) and ATTUNE™ (right) Implants.


Introduction

Instability of the knee in total knee replacement patients has been reported during high demand activities both through clinical observations and fluoroscopic evaluation. Ploegmakers et al1, in an in vivo kinematic evaluation, cited implant design factors as a determinant of knee instability. The objective of the current study was to compare anterior-posterior (A-P) and internal-external (I-E) motions of the knee for four current TKR designs in order to assess the influence of implant geometry on the inherent stability, motion, and contact mechanics of the joint. Each design was assessed in two finite element (FE) models: a laxity/stability test, and a full lower limb model during two high demand activities – stepdown (high A-P force) and stance-phase gait (high I-E torque).

Methods

Implant design was quantified in terms of the tibiofemoral (TF) conformity ratio, calculated by dividing the femoral articular radius at 0°, 15°, 30° and 60° flexion by the radius of the insert in the dwell point (Figure 1). To assess tibiofemoral constraint, a finite element model of the femoral component was positioned in the dwell of the insert at 0°, 15°, 30° and 60° flexion under a compressive load of 667N. A 5mm anterior translation or 10° internal rotation was applied to the femur while the TF reaction force or torque was measured.

Subsequently, dynamic simulations of stepdown and stance-phase gait activities were carried out in a FE model of the lower limb (Figure 1). TF joint loads were taken from in vivo telemetric data2 and a control system was implemented to apply external loads at the hip and ankle to create the experimentally-measured loading condition (compressive load, A-P force, I-E torque) at the TF joint for each activity using the telemetric implant geometry. The external loading condition was subsequently applied directly in the model, and the simulation was carried out for the four components, including current cruciate-retaining (CR) and posterior-stabilizing (PS) designs from several manufacturers. 6-DOF TF kinematics, medial and lateral condyle lowest point, and contact mechanics were evaluated for each design.

Results

Conformity ratios correlated well with laxity/constraint of the components (r = 0.73 for translation tests; r = 0.78 for rotation tests).

Trends during the dynamic activities were in agreement with those predicted during the laxity simulations; higher conformity increased constraint and hence the loads carried by the insert instead of the surrounding soft tissue. These designs with higher conformity had, in general terms, higher contact area, and lower contact pressure than the less conforming components (Figure 2).

Both laxity tests and dynamic simulations highlighted substantial variation in the constraint provided by current implant designs. The range of A-P and I-E motion for the least constrained design was twice that of the most constrained design during dynamic activity (Figure 3).

Discussion

In the current analysis, each component was analyzed under the same external loading conditions with the same soft-tissue representation, allowing for direct comparison between components. Varying ligament mechanics would alter the magnitude of motions, but relative performance of each implant would be consistent.

Other factors, aside from geometry, contribute to instability of the knee joint, notably, ligamentous balance/tension. Some knees, through natural mechanics or injury, have a tendency towards instability. Component designs with inherent geometric stability may aid in maintaining knee stability during dynamic activity for these patients.

Significance

Understanding the variation in constraint provided by differing current implant designs may aid clinicians in determining which type of implant is most appropriate, given the soft-tissue quality of their patient, to provide adequate stability during activities of daily living while maintaining range of motion.

THE INFLUENCE OF DESIGN ON TKR MECHANICS DURING ACTIVITIES OF DAILY LIVING+1Fitzpatrick CK; 1,2Clary CW; 1Rullkoetter PJ

+1University of Denver, Denver, CO; 2DePuy Orthopedics, Inc., Warsaw, IN



The third party trademarks used herein are the trademarks of their respective owners.

References

1. Ploegmakers et al., 2009, Knee 17:204-209.

2. Kutzner et al., 2010, J Biomech 43:2164-2173.




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Figure 3: Medial and lateral A-P kinematics for each PS (solid) and CR (dashed) component shown for gait (top) and stepdown (bottom)


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Introduction

While recent focus on attaining deep flexion after total knee replacement (TKR) has focused on the effects of femoral “rollback”1, the interplay between implant design, ligamentous constraint, and maximum knee flexion remains unclear. In this study, the influence of implant shape, in particular the shape of the posterior condyles and amount of femoral rollback, on MCL elongation was assessed with a computational model of passive flexion and the results substantiated through implantation into cadaveric specimens.

Methods

Three prototype posterior stabilized rotating platform knee implants were formulated with different levels of mid-flexion femoral thickness, condylar height, and femoral rollback (Fig. 1). Implant 1 had an increased condyle height, decreased mid-flexion thickness, and increased rollback. Implant 2a had a neutral condyle height, neutral mid-flexion thickness, and neutral rollback. Implant 2b utilized the same femur as 2a, but with 1-mm less rollback. Implant 3 had a decreased condyle height, neutral mid-flexion thickness, and decreased rollback.

Each design was virtually implanted into a finite element (FE) model of the knee. The model included four non-linear spring elements representing the anterior, central, posterior, and distal portions of the medial collateral ligament (MCL), with the attachment site based on measured data from a cadaveric knee implanted with the prototype designs. Passive flexion boundary conditions were applied to the model, with the tibial component fixed in space while a compressive load of 500-N was applied to the femur and flexed from 0° to 140° knee flexion with the remaining femoral degrees of freedom unconstrained. Elongations of the MCL bundles were measured during the simulation.

The prototype knee implants were manufactured and implanted into 8 cadaveric knees by eight different orthopaedic surgeons. Four surgeons employed a “measured resection” technique while the other four employed a “balanced gap” technique.2 After implantation, surgeons were blinded to the implant in the knee and asked to quantify the MCL tension by palpation as either “loose”, “normal”, or “tight” at 3 flexion angles (45°, 90°, and 120°) and to measure the maximum passive flexion angle using a goniometer.

Results

According to the FE predictions, increased condylar height and increased femoral rollback led to increased elongation of the distal bundle of the MCL in deep flexion (Fig. 2). These predictions were corroborated by observations of half the surgeons (100% of the balanced gap surgeons) that implant 1 had an unacceptably tight MCL at 120° flexion (Fig. 3) and that fewer surgeons felt that implant 2a had a tighter MCL than implant 2b. On the contrary, while the model predicted that implant 1 had a more lax MCL at 45°, the surgeons were not able to distinguish this difference. Despite the observation that the MCL was unacceptably tight in implant 1 in flexion, there was not a significant difference in the measured terminal passive flexion.

Discussion

In this study, a FE model was used to quantify changes in MCL elongation in deep flexion associated with implant shape. The predicted changes, on the order of 1%–2% of MCL elongation, were perceptible to orthopaedic surgeons performing a “balanced gap” technique, but not to surgeons performing a “measured resection” technique. While the differences in implant shape were subtle, the effect of these changes on the soft tissue function and the surgeon’s perception was clear. This is the first time such a correlation between implant shape and surgical technique has been documented.

Although it’s been shown that femoral rollback enables deep flexion by preventing bony impingement of the insert on the posterior femoral bone1, this data suggests that too much rollback coupled with increased condylar tip height can lead to excessive strain in the MCL in deep flexion. While the cadaveric simulation did not illustrate a measureable change in terminal flexion across designs, the surgeons almost unanimously preferred Implant 3, leading to a new knee design (ATTUNE™ System, DePuy Orthopaedics, Inc., Warsaw, IN).

Significance

This study clarifies the effect of implant shape on soft tissue function in deep flexion and the interaction with surgical technique. Results of this work will enable implant designs and surgical technique to improve high-flexion outcomes for patients receiving TKA.

MANAGEMENT OF MCL TENSION IN DEEP FLEXION: INFLUENCE OF IMPLANT SHAPE+1Clary, CW; 1Wyss, JG; 1Wright, AP; 1Bennett, TD; 1Auger, DD; 1Heldreth, MA

+1DePuy Orthopedics, Inc., Warsaw, IN

Figure 1: Three prototype implants with different femoral geometry. Table indicates level of condyle height, mid-flexion thickness, and rollback (+=increased, N=neutral, -=decreased).

Mid-FlexionThickness

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References

1. Banks et al., CORR, 2003.

2. Dennis et sl., CORR, 2010.


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Figure 3: Surgeon assessments of MCL tension for the three prototypes (red=tight, green=normal, yellow=loose) by all surgeons (top, n=8) and only balanced gap surgeons (below, n=4).




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Introduction

A major concern in the use of modular knee implants has been particle generation from the backside of the UHMWPE tibial insert. Motion, commonly referred to as micromotion, of the tibial insert against the proximal tibial tray can generate microscopic particles that with time can propagate throughout the joint and lead to osteolysis, a condition that can promote bone resorption and finally, implant loosening1,2. This study was undertaken to develop a simple and sound approach to characterizing modular knees that could be easily reproduced in any orthopaedic laboratory having a 2-axis test machine.

Methods

Six competitive products; Optetrak® (n=1) (Exactech, Inc., Gainesville, FL 32653), NexGen® (n=3) (Zimmer, Inc., Warsaw, IN 46581), Journey® (n=1) (Smith & Nephew, Memphis, TN 38116), Advance® (n=1) (Wright Medical, Arlington, TN 38002), Scorpio (n=1) and Duracon® (n=2) (Stryker, Kalamazoo, MI 49002) were tested along with 2 DePuy products, SIGMA® XLK (n=5) with i2 locking mechanism, and ATTUNE™ AOX™ (n=3) with ATTUNE locking mechanism (DePuy Orthopaedics, Warsaw, IN 46581). A variety of insert locking techniques were employed across the samples. All specimens were taken from sealed packages.

Trays were cemented in appropriate potting fixtures using Ultracryl II epoxy (Masel, Bristol, PA 19007) and allowed to cure before milling 2 holes in the insert for a multi-directional load applicator. Holes were positioned such that the force will be applied at 0.050 inches from the tray/ insert interface about the insert center of rotation. Each construct was soaked in a 37C RO water bath for a minimum of 12 hours before testing.

Testing was conducted on an MTS 858 Bionix servo hydraulic test frame with TestStar IIm controller utilizing MPT software. Micromotion was measured using a Heindenhain model ST 1278 encoder fixed to a custom encoder fixture mounted to the tray fixture. A laptop computer equipped with a custom MatLab (MathWorks, Inc., Natick, MA 01760) v7.2.0.232 data acquisition program was used to acquire the encoder data. Microsoft® Excel 2000 was used to analyze and plot the micromotion data. The measurement system has resolution of ± 0.5 micrometers. A 0-N (Newton) compressive load was maintained for all test directions. The A/P micromotion test applied a 100-N anterior load through the holes in the tibia insert and reversed to a 100-N posterior load, and returned to the 0-N load position. Micromotion was defined as the measured displacement between the minimum and maximum load positions. The M/L micromotion test is identical to the A/P micromotion loading except the tibial tray/insert assembly and load applicator was rotated 90 degrees. For rotational measurements, a mark was placed on the insert to identify the measurement location of the encoder. The x-y coordinates of this point with respect to the insert rotational center needed to be determined to provide inputs for the angular micromotion calculation. This work was conducted on the Bridgeport mill in the machine shop equipped with a digital indicator, catalog 387538400 (Accu-Rite Companies, Inc., Jamestown, NY 14701). Equations were derived to solve for the angle of rotation, theta, given the measured amount of linear micro-motion, m., and the x-y location of the encoder. A MatLAB Version 7.2.0.232 (r2006a) program was written to solve for angle where the two equations were equal. For the RT micromotion test, the rotational orientation of the load applicator was used to apply a 1-N-m ramp torque to the insert in the counter-clockwise direction, followed by a 6-N-m ramp torque in the clockwise direction.

Results

AP, ML, R, and RT micromotion was characterized for each specimen. “R” is a compilation of the AP and ML micromotion and was calculated by taking the square root of the sum of the squares of AP and ML. Standard deviation was calculated for each sample, where applicable, and ranged between 0% and 47% of the mean.

Discussion

Measurement capability appears to range from constructs demonstrating extremely small micromotion to those exhibiting relatively large motions. A sample of 3 specimens was completely characterized within about an hour. This approach has the advantage over contemporaries3,4 by actuating the insert while avoiding insert clamping that can over-constrain and distort the insert. It features fixtures of limited complexity with greater stability through rigid fixture clamping. The Heindenhain encoders present a great improvement in accuracy over other measurement devices.

The means and standard deviations for designs having samples of n=3, or greater, are as follows: ATTUNE AOX (15 ± 1), SIGMA XLK i2 (16 ± 4), and NexGen (82 ± 14), with NexGen being statistically different (p <0.0001). The methodology shows the ability to distinguish between different manufacturer’s designs even with small sample sizes. Clearly, careful attention to design details can result in very significant differences in micromotion behavior.

ADVANCED FIXED BEARING TKA LOCKING MECHANISM MINIMIZES BACKSIDE MICROMOTIONLeisinger, S; Hazebrouck, S ; Deffenbaugh, D ; Heldreth, M

DePuy Orthopaedics, Inc. Warsaw, IN 46581

Figure 1

References1. Parks NL et al, Orthop Trans, 18, 1994;

2. Wasielewski RC et al, Clin Orthop, 345, 1997;

3. Taki, N et al, Soc For Biomat, 28th Annual Meeting Transactions, 2002;

4. Parks NL et al, Clin Orthop, 356,1998.

Presented at the ISTA Congress, 24th Annual Meeting, 2011, Bruges, Belgium.


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Introduction

The decision to resurface the patella during total knee arthroplasty (TKA) remains controversial, as recent clinical results for unresurfaced patellae have been on par with resurfaced.1 Differences in congruency between the natural patella and the femoral component affect kinematics and contact mechanics, and may alter contact stresses on the articular surface of the patellar cartilage2, potentially leading to early degeneration or eventual resurfacing.

A femoral component with perfectly anatomic geometry may aid in restoring natural kinematics to the joint, but contact mechanics are affected not just by geometry, but also by the material properties of the contacting surfaces. The purpose of this study was to compare mechanics of the natural patella articulating against five femoral representations; natural femoral cartilage, three current Co-Cr femoral component designs (LCS®, SIGMA®, and ATTUNE™ Posterior-Stabilized Components, DePuy Orthopaedics, Inc., Warsaw, IN), and an idealized natural Co-Cr component, in order to assess the influence of geometry and material properties on patellofemoral (PF) mechanics.

Methods

A population of 10 subject-specific explicit finite-element models of the PF joint were developed from MR scans of normal knees. Hexahedral meshes of the femoral and patellar cartilage were created using an automated custom-scripted algorithm and morphing approach and represented as deformable (E = 12 MPa, v = 0.45). The analysis was performed using Abaqus/Explicit (Simulia, Providence, RI) five times for each subject with the patella articulating against: deformable femoral cartilage, LCS, SIGMA, and ATTUNE Co-Cr Components, and a ‘natural’ Co-Cr component. The natural component was modeled as femoral cartilage geometry with femoral component material properties (Figure 1). Due to the greater stiffness of CoCr relative to cartilage, each femoral component was modeled as a rigid body. The extensor mechanism and retinacula of the knee were represented by 2D fiber-reinforced membranes with a 1000 N ramped load distributed among the quadriceps (vasti and rectus femoris) as the knee was flexed from full extension to 120°.3 Six-degree-of-freedom kinematics and contact mechanics were evaluated.

Results

Predicted PF kinematics and contact mechanics illustrated differences between the femoral representations. Kinematics for the natural cartilage and the natural component were identical, while differences were present between the natural knee and the femoral components, particularly in internal-external (I-E) rotation (Figure 2). I-E tracking of the patella when articulating against the dome and anatomic components was opposite to the natural knee.

The SIGMA Component resulted in the highest peak contact pressure throughout flexion (Figure 3). The LCS and ATTUNE Components maintained better contact in early flexion than the natural component, as the trochlear grooves of these two components extended further superior, providing a larger contact surface. Once the patella left the conforming trochlear groove, differences between implant components decreased. Contact area as a function of flexion was similar for each of the implants. Contact pressure and area patterns were similar between the natural knee and the natural Co-Cr component; however, contact pressure was consistently 1 MPa higher and contact area 20% lower for the natural Co-Cr component than with the natural knee for each subject throughout flexion (Figure 4).

Discussion

Using a relatively small population of subject-specific models, unresurfaced joint mechanics were impacted by congruency of the femoral representation. In early and mid flexion, conformity between the LCS and ATTUNE Components and the unresurfaced patella resulted in lower peak contact pressure compared to the less conforming SIGMA Design. In deeper flexion, the differences were less. The natural Co-Cr component maintained a larger contact area and lower pressure with the patella in deeper flexion.

Despite kinematic agreement, there were substantial and consistent differences in contact mechanics between the natural femur and natural Co-Cr component, attributable to material property differences between metal and cartilage surfaces. Component designs with a more anatomic distal and posterior geometry may improve contact mechanics in later flexion; however, material property differences and the constraints of a PS design limit the available improvement.

Significance

Implant design does influence unresurfaced patellofemoral mechanics in early and mid flexion, but the soft-hard bearing surface increases patellar contact pressures compared with the natural, even with perfectly anatomic geometry.

COMPARISON OF NATURAL AND UNRESURFACED PATELLOFEMORAL MECHANICS+1Deacy, J D; 1Fitzpatrick, C K; 1Laz, P J; 1Rullkoetter, P J

+1University of Denver, Denver, CO

Figure 1: Unresurfaced patella articulated against natural femoral cartilage, natural femoral geometry, LCS, SIGMA, and ATTUNE components (left to right).

References1. Whiteside et al., 2003, CORR, 410:189-98.2. Thompson el al., 2001, J Arthroplasty, 16:607-12. 3. Farahmand et al., 1998, J Orthop Res, 16(1):136-43.



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Figure 3: PF contact mechanics (mean and ± 1 standard deviation) of the unresurfaced patella and five femoral representations.




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Introduction

Extensor mechanism function is critical to the outcome for total knee replacement patients, although the role of surgical variability and implant design on extensor efficiency is not well understood. Some authors have documented improvement in TKR patients’ abilities to perform a sit-to-stand activity and attributed the improvement to the femoral design1, while other authors have shown that extensor efficiency is more closely tied to patella-femoral mechanics2. Although it is generally assumed that maximizing extensor efficiency is desirable, it has also been shown that over-stuffing of the patella-femoral joint can restrict range of motion3. Therefore, the current research had two objectives: 1) to determine the relationship between various extensor moment arm metrics and the quadriceps force required to perform a deep knee bend, and 2) understand how surgical variability and implant design influence extensor efficiency.

Methods

A validated computational model of the Kansas Knee Simulator was used to study extensor mechanism function during a deep knee bend (Figure 1)4. A virtual model of the knee implanted with total knee replacement components (ATTUNE® Knee System, DePuy Synthes Joint Reconstruction, Warsaw, IN) was flexed from full extension to 130° flexion by applying a variable load to the quadriceps tendon balancing a compressive load applied at the hip4. During the deep knee bend, various moment arms were calculated, including the perpendicular distances from the medial and lateral contact points (MCP, LCP) and the helical axis (HA) to the line-of-action of the quadriceps muscle (QM) and the patella ligament (PL) (Figure 1). In addition, an “effective” moment arm was calculated by dividing the incremental quadriceps elongation by the corresponding increase in knee flexion through the flexion range. To assess the role of surgical variability on extensor efficiency, multiple surgical scenarios were assessed, including increasing and decreasing the overall thickness of the patella by 2-mm and assessing ± 3-mm of patella alta/baja. In addition, the influence of the patella articular geometry was assessed by comparing a domed to an anatomic style patella which has increased sagittal conformity with the femur. Linear regressions were performed to compare the % change in extensor moment arms from the neutral condition with the change in quadriceps force through the flexion cycle.

Results

In general, the “effective” moment arm measurement was the only metric that consistently quantified the observed changes in quadriceps force through the full flexion range for all of the knee conditions (Table 1). While other metrics, including the distance from the contact points and helical axis to the line-of-action of the quadriceps muscle, explained variations through portions of the flexion cycle, interaction with changes in the other metrics led to low R2 values across the full flexion cycle. Patella thickness had the largest influence on extensor efficiency in knee extension, where a 2-mm increase in thickness caused a 3.6% reduction in quadriceps force at 30° knee flexion (Figure 2). The increased efficiency was attenuated with increased flexion. Conversely, patella alta/baja had the biggest influence on extensor efficiency in mid-flexion, where 3-mm of patella baja caused a 4.3% increase in the required quadriceps force at 60°. The anatomic-style patella increased the extensor efficiency in deeper flexion with a 4.7% reduction in quadriceps force at 90° flexion. None of the knee conditions studied had a significant influence on the quadriceps force at maximum knee flexion, but did change the overall quadriceps excursion required to reach 120° flexion by ±3.4% (thick/thin patella), ±1.5% (patella alta/baja), and 1% (anatomic patella).

Discussion

The relationship between surgical technique, patella geometry, and extensor mechanism efficiency is complex. Surgical steps taken to maximize extensor efficiency may require increased quadriceps excursion and a reduction in maximum active knee flexion. If extensor weakness is a concern, these results suggest avoiding elevation of the joint line (pseudo patella baja), using a thicker patella, or using an anatomic-shaped patella would enhance the extensor efficiency. When quantifying the extensor moment arms, it is necessary to recognize the inherent interactions between the various moment arm metrics and that a composite metric like the “effective” moment arm was necessary to quantify the observed changes in quadriceps force. Future work will identify surgical technique and implant design factors which maximize extensor efficiency during the appropriate range of motion to improve function during common activities of daily living while minimizing the overall extensor excursion required to achieve maximum flexion.

Significance

This study quantifies the effects of surgical technique and patella design on extensor mechanism function after total knee replacement to aide surgeons in optimizing their intra-operative decisions to improve patient outcomes.

THE EFFECT OF SURGICAL VARIABILITY AND PATELLA GEOMETRY ON EXTENSOR EFFICIENCY IN TOTAL KNEE REPLACEMENTChadd Clary, Ph.D. – DePuy Synthes Joint Reconstruction*, Warsaw, INAZ Ali – DePuy Synthes Joint Reconstruction, Warsaw, INAbe Wright – DePuy Synthes Joint Reconstruction, Warsaw, INClare FitzPatrick, Ph.D. – University of Denver, Denver, COPaul Rullkoetter, Ph.D. – University of Denver, Denver, CO

References: 1. Mahoney et al, Journal of Arthroplasty, 2002, 2. Ward et al, Knee, 2011,3. Bengs and Scott, Journal of Arthroplasty, 2006, 4. FitzPatrick et al, CMBBE, 2012.

Poster #1959, from the 2014 ORS 60th Annual Meeting, March 15-18th, 2014, New Orleans, Lousiana.

Table 1: The slope and R2 values for the regression between the % change in moment arm and corresponding % change in quadriceps force. Moment arm metrics which accurately predicted changes in quadriceps force (slope < -0.5 and R2 > 0.5) were highlighted.

Figure 2: The % change in the quadriceps and the “effective” extensor moment arm for the various knee conditions through the flexion range.

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MLP-QM -1.10 0.96 -0.77 0.95 -0.48 0.49 -0.28 0.22 -0.39 0.72

Figure 1: Computational Kansas Knee Simulator model with a graphical representation of the moment arm metrics calculated during a flexion cycle.

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DePuy (Ireland)Loughbeg RingaskiddyCo. CorkIrelandTel: +35 (321) 491-4278Fax: +35 (321) 491-4199

DePuy Orthopaedics, Inc.700 Orthopaedic DriveWarsaw, IN 46582USATel: +1 (800) 366-8143Fax: +1 (800) 669-2530


© DePuy Synthes 2014-2015. All rights reserved. DSUS/JRC/0614/0264(1) 09/15

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References

1. Ormonde, el al, “Overhang of the Femoral Component in Total Knee Arthroplasty: Risk Factors and Clinical Consequences”, JBJS(A) 2010.


Figure 3: Probability (p) excess implant overhang for SIGMA Component size increments 2–5. Yellow indicates p > 5%.

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Introduction

The adequacy of femoral implant fit in total knee arthroplasty (TKA), often characterized by the extent of overhang, is a matter of continuous debate. In this work we present a computational method for determining the probability of excess overhang or underhang (beyond an acceptable threshold value). Unlike most previous work which relied on simple landmark data or resection contours, we consider the entire distal femur, real implant models, and clinically accurate resection surfaces. The technique is demonstrated using a commercially available implant and a modified design intended to reduce overhang.

Methodology

The methodology is based upon a model which describes variation in 3D femoral geometry constructed using sample bones segmented from computed tomography images. Using a principal component analysis (PCA) based surface registration technique optimized for shape model construction, the segmented data is consolidated into a point distribution model (PDM) with a size variable (PC1) and several additional parameters (i.e. PC2, PC3, etc.) which represent the distal femur’s major modes of shape variation. (Fig. 1)

The PDM is used as the basis for a Monte-Carlo style simulation where thousands of sample femurs are generated and virtually implanted to assess the overhang characteristics of implant designs. Surface landmarks, anchored to the PDM, are used to automatically fit and align each implant. Overhang and underhang characteristics are reported as scalar values along the implant edge adjacent to the resection surfaces of the bone, indicating the probability of excess overhang greater than a user specified tolerance value. The shape modeling and Monte-Carlo process were implemented with Arthron: a morphometric analysis interface based on the open-source Visualization Toolkit (VTK). (Fig. 2)

Results & Discussion

The commercial implant and modified design were evaluated using a femoral PDM constructed from the left knees of thirty-one Caucasian and forty-five Japanese subjects. The pre-existing commercial design (PFC® SIGMA® Knee System, DePuy Orthopaedics, Inc.) features seven non-uniform size increments (1p5, 2, 2p5, 3, 4, 5, 6) and a fixed inter-condylar notch width. The 4N component was not included. A newly developed commercial design (ATTUNE™ Knee System, DePuy) consisted of ten uniform size increments, an inter-condylar notch width proportional to size, and a narrower medial-anterior flange. To compare the two designs, one thousand femurs were automatically generated from Latin hypercube sampling of the PDMs control parameters. During the simulation, the probability of excess overhang was tabulated along the resection edge of each implant size. The overhang limit k was defined to be proportional to the medial-lateral bone width ML (the distance between the lateral and medial femoral epicondyles). We specified the limit to be k = (3/70) ML [mm], representing a 3mm overhang limit for a 70mm wide bone.

Conclusions

Results indicate a significant reduction in the incidence of excess overhang on the medial-side of the anterior flange and around the intercondylar notch with the modified implant design (Fig. 3–4). Future work will include analysis of patient variation factors and surgical factors (i.e. implant alignment) and their effects on implant conformity.

Significance

Recent evidence links excess overhang to poor patient prognoses including soft-tissue irritation and reduced joint mobility1; we present a computational method for predicting and comparing the excess overhang of femoral component designs beyond an acceptable threshold.

PROBABILISTIC MODELING OF FEMORAL COMPONENT OVERHANG+1Courtis, RP; 2Heldreth, M; 1FitzPatrick D,

+1University College Dublin, Ireland, 2DePuy Orthopaedics, Inc., Warsaw, IN

-3 PC2 3

-3 PC3 3

Figure 1: Varying PDM control parameters PC2 and PC3, affecting anterior-posterior length and inter-condylar bone width.

Figure 2: Femoral components are automatically sized and aligned using landmarks anchored to the bone surface; conformity is estimated using the shortest distance from the implant edge to the resected bone and mapped. Under-hanging edges are colored red, over-hanging edges blue.

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HOMEATTUNE Knee Conference Abstracts

TIBIAL TRAY DESIGN FACTORS AFFECTING TIBIAL COVERAGE AFTER TOTAL KNEE ARTHROPLASTYChadd Clary, Ph.D. - DePuy Synthes Joint Reconstruction – Warsaw, USADaren Deffenbaugh - DePuy Synthes Joint Reconstruction – Warsaw, USAFilip Leszko - DePuy Synthes Joint Reconstruction – Warsaw, USAPatrick Courtis - DePuy Synthes Joint Reconstruction – Warsaw, USA

Introduction

Adequate coverage of the resected tibial plateau with the tibial tray is necessary to reduce the theoretical risk of tibial subsidence after primary Total Knee Arthroplasty (TKA). Maximizing tibial coverage is balanced against avoiding excessive overhang of the tray causing soft tissue irritation, and establishing proper tray alignment improving implant longevity and patella function1. Implant design factors, including the number of tray sizes, tray shape, and tray asymmetry influence the ability to cover the tibial plateau2. Furthermore, rotating platform (RP) tray designs decouple restoring proper tibial rotation from maximizing tibial coverage, which may enhance the ability to maximize coverage. The purpose of the current study was to assess the ability of five modern tray designs (Figure 1), including symmetric, asymmetric, fixed-bearing, and RP designs, to maximize coverage of the tibial plateau across a large patient population.

Methods

Lower limb computed-tomography scans were collected from 14,791 TKA patients and the tibia was segmented. Virtual surgery was performed with an 8-mm tibial resection (referencing the high side) made perpendicular to the tibial mechanical axis in the frontal plane, with 3° posterior slope, and aligned transversely to the medial third of the tibial tubercle. An automated algorithm placed the largest possible tray on the plateau, optimizing the ML and AP placement (and I-E rotation for the RP tray), to minimize overhang. The largest sized tray that fit the plateau with less than 2-mm of tray overhang was identified for each of the five implant systems. The surface area of the tibial tray was divided by the area of the resected plateau and the percentage of patients with greater than 85% plateau coverage was calculated.

Results

The percentage of patients with greater than 85% plateau coverage across the tray designs ranged from 17.0% to 61.4% (Fig. 1). The tray with the greatest number of size options (Tray 4, 10 sizes) had the best coverage among the fixed-bearing trays. The RP variant of the same tray had the best overall coverage. Tibial asymmetry did not significantly improve the overall tibial coverage across the patient distribution for both asymmetric designs. Incorporating a broader medial condyle improved fit along the posterior medial corner for Tray 2, but increased the average under-hang along the posterior lateral plateau offsetting any improvement in total coverage.

Discussion

This analysis represents the most comprehensive assessment of tray coverage to date across a large TKA-patient population. Large variations exist in the size and shape of the proximal tibia among TKA patients3. Developing a tray design which provides robust coverage despite this variation remains challenging. This analysis suggests that tibial asymmetry may not robustly improve coverage. Conversely, incorporating an increased number of tray sizes and utilizing an RP implant to decouple coverage from alignment may provide the most reliable solution for maximizing coverage across the patient population.

Significance


References

1. Hofmann et al. Orthopade, 2003;32(6):469-76.

2. Wernecke et al., Journal of Orthopaedic Surgery, 2012;20(2):143-7.

3. Fitzpatrick et al., Proc Inst Mech Eng H, 2008;222(6):923-32.

Presented at the ISTA Congress, 26th Annual Meeting, 2013, West Palm Beach, FL.

© DePuy Synthes Joint Reconstruction, a division of DOI 2014 0612-76-514 EO 01/14



Figure 1: Tibial plateau coverage for five current TKA tray designs. Designs 1 and 2 have an asymmetric medial plateau while designs 3 through 5 are symmetric about the M-L axis. Trays 4 and 5 have identical outer profiles, but tray 5 has a RP insert which allows flexibility in the rotational alignment of the tray on the bone. The colored border around the tray periphery indicates the average under-hang of the tray (in mm) across the patient population.

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THE EFFECT OF TIBIAL TRAY ROTATIONAL ALIGNMENT ON ASYMMETRY OF THE RESECTED TIBIAL PLATEAUChadd Clary, Ph.D. - DePuy Synthes Joint Reconstruction – Warsaw, USAAlex Schenher - DePuy Synthes Joint Reconstruction – Warsaw, USALuke Aram - DePuy Synthes Joint Reconstruction – Warsaw, USAFilip Leszko - DePuy Synthes Joint Reconstruction – Warsaw, USAMark Heldreth - DePuy Synthes Joint Reconstruction – Warsaw, USA

Introduction

Appropriate transverse rotation of the tibial component is critical to achieving a balance of tibial coverage and proper tibio-femoral kinematics in Total Knee Replacement (TKR), yet no consensus exists on the best anatomic references to determine rotation. Historically, surgeons have aligned the tibial component to the medial third of the tibial tubercle1, but recent literature suggests this may externally rotate the tibial component relative to the femoral epicondylar axis (ECA) and that the medial border of the tubercle is more reliable2. Meanwhile, some TKR components are designed with asymmetry of the tibial tray assuming that maximizing component coverage of the resected tibia will result in proper alignment. The purpose of this study was to determine how different rotational landmarks and natural variation in osteoarthritic patient anatomy may affect asymmetry of the resected tibial plateau.

Methods

Pre-operative computed-tomography scans were collected from 14,791 TKR patients. The tibia and femur were segmented and anatomic landmarks identified: tibial mechanical axis, medial third and medial border of the tibial tubercle, PCL attachment site, and the surgical ECA of the femur. Virtual surgery was performed with an 8-mm resection (referencing the high side) made perpendicular to the tibial mechanical axis in the frontal plane, with 3° posterior slope, and transversely aligned with three different landmarks: the ECA, the medial border, and medial third of the tubercle. In each of these rotational alignments, the relative asymmetry of the medial and lateral plateaus was calculated (Medial AP / Lateral AP) (Fig. 1).

Results

Rotational alignment of the tibial component to the ECA, medial border, and medial third of the tubercle resulted in progressive external rotation of the tibial tray on the bone. Alignment to the medial border and medial third of the tubercle resulted in average 0.9°±5.7° and 7.8°±5.3° external rotations of the tray relative to the ECA, respectively (Fig. 2). Greater external rotation of the tibial implant relative to the bone increased the appearance of tibial asymmetry (Fig. 3). Referencing the medial border and medial third of the tubercle resulted in apparent tibial bone asymmetry of 1.10±0.10 and 1.12±0.10, respectively.

Discussion

Assuming the ECA is the appropriate rotational reference to re-establish appropriate kinematics2, alignment to the medial border of the tubercle resulted in the most favorable tray alignment. However, there was a great deal of variation between the relative position of the ECA and the tubercle across the patient population. Rotational alignment to either the medial border or medial third of the tubercle resulted in external tray alignment relative to the ECA of greater than 3 degrees for 36% and 84% of patients, respectively. In addition, increased tray asymmetry (broader medial plateau) necessitates relative external rotation of the tray on the bone reducing the flexibility of intra-operative rotational adjustment. Tray asymmetry greater than 1.10 (the asymmetry of the resected tibia when aligned to the ECA) may result in external mal-rotation for a significant portion of the patient population.

References1. Insall, Surgery of the Knee, 1993.

2. Akagi et al. Clin Orthop Relat Res, 2005;Jul(436):172-6.

Presented at the ISTA Congress, 26th Annual Meeting, 2013, West Palm Beach, FL.

Figure 2: The rotation between the femoral ECA and the medial border (red) and medial third (blue) of the tibial tubercle across a population of osteo-arthritic patients.

Figure 3: The apparent medial to lateral AP tibial asymmetry of the cut bone when aligned to the femoral ECA (black), medial border (red) and medial third (blue) of the tibial tubercle across a population of osteo-arthritic patients.

Figure 1: Rotation of the tibial component was set perpendicular to the femoral ECA (black), towards the medial border of the tubercle (red), and towards the medial third of the tubercle (blue). In each of these rotational alignments, the ratio of the A-P dimensions of the medial and lateral plateaus was calculated.

© DePuy Synthes Joint Reconstruction, a division of DOI 2014 0612-77-514 EO 01/14



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EXTRACTION STUDY OF GAMMA-IRRADIATED UHMWPE STABILIZED WITH A HINDERED PHENOL ANTIOXIDANTKing, R.; Sharp, M.; Narayan, V.S.DePuy Synthes Joint Reconstruction, a division of DePuy Orthopaedics, Inc. Warsaw, IN.

Introduction

Gamma-irradiated Ultra-High Molecular Weight Polyethylene (UHMWPE) stabilized with a hindered phenol antioxidant was developed as an alternative to gamma-irradiated, re-melted UHMWPE.1,2 Hindered phenol antioxidants have demonstrated good biocompatibility and elution-retention profiles, for example octadecyl 3,5-di-tert-butyl-4-hydroxyhydrocinnamate has been used in commercial orthopedic polyethylene implant devices for twenty years.3

Tetrakis[3-(3,5-di-tert-butyl-4-hydroxyphenyl) propionate], PBHP, is a solid hindered phenol antioxidant characterized by a high anti-oxidation efficiency. The sterically bulky structure of the molecule would contribute to restricted mobility of PBHP within the UHMWPE matrix.

The objective of this study is to verify low extractability of PBHP from gamma-irradiated, PBHP-stabilized UHMWPE. Aggressive extraction condition of refluxing hexane for extended duration was employed. Post-extracted samples were characterized by assessment of oxidation index and ester / ketone index.

Materials and Methods

Thin, fixed-bearing tibial components were machined from compression-molded UHMWPE, GUR 1020 (Ticona), as well as GUR 1020 wherein 0.075% (W/W) PBHP was incorporated (AOX™). Both sets of components were cleaned and dried prior to vacuum foil packaging and gamma irradiation to a nominal dose of 75 KGy (Steris).

Four tibial components from these sample sets were each cut into two halves. Four test groups from both materials, each consisting of two half tibial components, were categorized as follows: (1) controls, (2) samples after 72 hour refluxing hexane extraction, (3) samples after 72 hour refluxing hexane extraction, followed by 2 week 70 degrees Celsius / 5-atm oxygen accelerated aging per ASTM F 2003, (4) samples after 72 hour refluxing hexane extraction, followed by 6 week 80 degree Celsius / air accelerated aging.3

200-micron film specimens microtomed through thickness of the articulating surface section of tibial samples were used in Fourier Transform Infrared Spectroscopy (FTIR) analysis. The oxidation index was calculated based on the procedure specified by ASTM F 2102.

Conventional oxidation index measurements are based on the broad FTIR carbonyl peak between 1650 to 1850 cm-1 that include ester, ketone, aldehyde and carboxylic acid functional groups.4 In this study, ester index and ketone index were calculated based on absorption peak heights at 1738 cm-1 and 1718 cm-1 respectively, and normalized with the reference peak height at 1368 cm-1. PBHP is characterized by IR absorption peaks for both ester and ketone functional groups. In this study, ester index was used to assess PBHP content in gamma-irradiated UHMWPE while ketone index was used to monitor oxidative degradation.

Results

Unmodified GUR 1020 registered low levels of oxidation index post-gamma irradiation and post-72 hour refluxing hexane extraction. However, two week accelerated aging raised oxidation levels moderately while 6 week accelerated aging increased the oxidation index substantially (Table 1). On the other hand, AOX Polyethylene registered a measurable carbonyl peak initially, that can be attributed to the ester and ketone functionalities in PBHP. This shows up in a higher calculated value of oxidation index. There are no significant changes in oxidation index post-hexane extraction and post-hexane extraction plus acceleration aging, indicating oxidative stability after exposure to the adverse conditions. Both ester and ketone index profiles for AOX Polyethylene show relatively constant levels of the antioxidant regardless of exposure to the hexane extraction and subsequent accelerated aging (Figure 1). Ester index data indicate that there is minimal elution tendency for PBHP in radiation-crosslinked UHMWPE matrix. There are no significant increases in ketone index for AOX Polyethylene before and after accelerated aging point to oxidative stability. In contrast, both ester and ketone index profiles for 75 KGy GUR 1020 point to oxidative degradation post accelerated aging. The 75 KGy GUR 1020 control registered high surface ester index and ketone index after 6 week accelerated aging.




0.20

0.15

0.15

0.05

0.000 1000 2000 3000

Depth (microns)

Este

r In

dex

4000 5000 6000

As isHexane Ext.2 Week Acc. Aging6 Week Acc. Aging

Figure 1: Ester index profile of 75 KGy, 0.075% PBHP GUR 1020

3.0

2.5

2.0

1.5

1.0

0.5

0.0

Depth (microns)

Este

r In

dex

0 1000 2000 3000 4000 5000 6000


Figure 2: Ester index profile of 75 KGy GUR 1020

0.20

0.15

0.15

0.05

0.00

Depth (microns)

Ket

on

e In

dex

0 1000 2000 3000 4000 5000 6000


Figure 3: Ketone index profile of 75 KGy, 0.075% PBHP GUR 1020

3.0

2.5

2.0

1.5

1.0

0.5

0.0

Depth (microns)

Ket

on

e In

dex

0 1000 2000 3000 4000 5000 6000


Figure 4: Ketone index profile of 75 KGy GUR 1020

75 KGy GUR 1020

Ester index

Ketone index AOX

Ester index

Ketone index

Post gamma irradiation 0.007 0.029 Post gamma irradiation 0.023 0.041

After 72 hour refluxing hexane extraction

0.013 0.040After 72 hour refluxing hexane extraction

0.024 0.038

After the hexane extraction and 2 week acc. aging

0.006 0.137After the hexane extraction and 2 week acc. aging

0.020 0.039

After the hexane extraction and 6 week acc. aging

0.650 1.191After the hexane extraction and 6 week acc. aging

0.022 0.032

Table 2: Summary data for ester index and ketone index

References1. Narayan, V.; et al. (2009). Trans. 55th. Orthopaedic

Research Society, 462.

2. King, R.; et al. (2009). Trans. 55th. Orthopaedic Research Society, 19.

3. Wroblewski, B.M.; et al. (2005). Journal of Bone Joint Surgery, 87-B, 1220-1221.

4. Oral, E.; et al. (2006). Trans. 52nd. Orthopaeidc Research Society, 665.

Presented at the Orthopaedic Research Society, 56th Annual Meeting, Poster No. 2286, March 2010, New Orleans, LA

75 KGy GUR 1020 AOX

Post-gamma irradiation 0.036 +/- 0.008 0.097 +/- 0.028

After 72 hour refluxing hexane extraction

0.064 +/- 0.027 0.087 +/- 0.015

After the hexane extraction and 2 week accelerated aging

0.105 +/- 0.017 0.089 +/- 0.017

After the hexane extraction and 6 week accelerated aging

1.567 +/- 1.996 0.080 +/- 0.015

Table 1: Mean oxidative index data of gamma-irradiated UHMWPE

Examination of mean ester and ketone index data for gamma-irradiated UHMWPE shows that ketone index is more sensitive in detecting onset of oxidative degradation after 2 week accelerated aging (Table 2). Nevertheless, there is significant increase in ester index for high-level oxidation after 6 week accelerated aging. PBHP content in UHMWPE is preferably monitored by ester index for reduction in incidence of oxidation product interference.

Discussion

This study demonstrates the stability of an antioxidant under severe extraction condition of refluxing hexane extraction for 75 KGy, 0.075% PBHP-stabilized GUR 1020. The post-extraction accelerated aging study indicates that 75 KGy, 0.075% PBHP-stabilized GUR 1020 provides long-term oxidative stability even under a severe chemical environment. This study also demonstrates that review of ester index and ketone index provides valuable insights in assessment of the hindered phenol antioxidant content and onset of oxidative degradation in crosslinked UHMWPE.

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© DePuy Synthes Joint Reconstruction, a division of DOI 2014 0612-23-509 Rev. 2 EO 02/14



Introduction

Due to compromised oxidative stability of gamma irradiated UHMWPE; a post-irradiation treatment has been used to decrease the reactive free radical concentration. Addition of antioxidants (AO) in UHMWPE prior to irradiation is an alternative method for improving the oxidative stability of irradiated UHMWPE. The AO molecules form dormant free radical traps when they react with the free radicals that are created during gamma irradiation, thus stops further oxidative damage. However, these antioxidants also reduce the available free radical concentration to undergo crosslinking reactions. The characterization of AO concentration and distribution during the different processing stages of UHMWPE (powder and molded form) therefore becomes critical as this determines the outcome of competing reactions of free radicals with themselves and with AO molecules. This paper describes the characterization of AO concentration in medical grade UHMWPE powders and molded films.


An antioxidant, Pentaerythritol tetrakis [3-(3,5-di-tert-butyl-4-hydroxyphenyl) propionate], (PBHP) stabilized Ultra High Molecular Weight Polyethylene (UHMWPE GUR 1020 resin powder, Ticona, US) samples were prepared at MediTECH Fort Wayne. The PBHP blend is consolidated into compression molded sheets (CMS).

Ultraviolet absorption measurements of PBHP-containing GUR 1020 powder samples were carried out using a Perkin-Elmer Lambda 35. Three different GUR 1020 powder batches were prepared with 0.065, 0.075 and 0.085 % (W/W) of PBHP for testing using the UV-Vis spectrometer to establish the uniform distribution of PBHP. The details of the instrument and method used were published previously1.

A number of samples in the concentration range, 0.01, 0.03, 0.045, 0.06, 0.07, 0.09, and 0.12% (W/W), are employed for constructing the FTIR calibration curve based on the spectral peak at 1230 cm-1.

Results and Discussion

Thirty random samples were taken from different locations of three powder batches and the results are shown in Figure 1. The linear dependence (R2=0.99) of these concentrations along with the narrow standard deviations within 30 data points confirm the precise control of PBHP distribution in the blending process.

FTIR spectroscopy was used to characterize the PBHP concentration and distribution in molded plaques. The height of phenol peak located at 1230 cm-1 is plotted against the concentration of PBHP loading and regression analysis performed to generate the curve function for the master curve given in Figure 2. The equation, y=0.9448x+0.0034 is used to calculate the PHBP concentration. PBHP containing GUR 1020 CMS is sectioned from different locations for FTIR analysis (Figure 3).

The diffusion of other antioxidants, i.e. Vitamin E has been well established2. Twenty data points were collected from each location in order to establish the distribution of PBHP and the results are shown in (Figure 4).

DETERMINATION OF ANTIOXIDANT DISTRIBUTION IN POWDER AND MOLDED UHMWPE MATERIALSSenyurt, A. F.*; Sharp, M.; Warner, D.1; Narayan, V.S.DePuy Synthes Joint Reconstruction, a division of DePuy Orthopedics, Inc. Warsaw, IN

3.3

3.2

3.1

3.0

2.9

2.8

2.7

2.6

2.5

2.4

0.06 0.065 006 0.075

PBHP % in GUR 1020 (W/W)

y = 31.6x + 0.5463

R2 = 0.99

Ab

s (r

.u.)

0.08 0.085 0.09

Figure1: The linear dependence of PBHP concentration in GUR 1020 powders at different loadings.

References1. Senyurt, A. et. al. (2009). Poster Number 460,

Orthopedic Research Society.

2. Bellare, A. & Yau, S. (2009). Poster Number 457-458, Orthopedic Research Society.

3. Muratoglu, O. et. al. (2009). Biomaterials I. Paper Number 21, Orthopedic Research Society.


Figure 3: The schematic representation of PBHP containing UHMWPE CMS; FTIR sample locations are indicated as E1:End1; C: Center; and E2: End2.

Figure 4: The distribution of PBHP concentrations. 5 films (800 um) are microtomed from each location and 4 scans are recorded from each film. (Batch #8180856, batch size 1267 lbs)

1 m

5 cm E1 C E2

8 cm

0.08

0.07

0.06

0.05

0.04

0.03

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P in

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R 1

020

(W/W

)

0.14

0.12

0.08

0.06

0.04

0.02

0.00

0 0.02 0.04 0.06

Theoretical [% PBHP in GUR 1020 W/W]

y = 0.9448x + 0.0034

R2 = 0.9901

Mea

sure

d %

PB

HP

in G

UR

102

0 (W

/W)

0.08 0.1 0.12 0.14

Figure 2: The FTIR (1230 cm-1) calibration curve generated with PBHP containing GUR 1020 films (800 um).

Conclusion

In this study, the uniform distribution of PBHP in GUR 1020, both in powder and CMS form was established with UV-Vis and FTIR spectrometers respectively. This is in contrast to the distribution gradient reported when Vitamin E is diffused into consolidated polyethylene3.

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Introduction

Oxidative stability is vital for longevity of articulating-surface implant devices made from gamma-irradiated Ultra-High Molecular Weight Polyethylene (UHMWPE). The orthopaedic industry has employed both gamma-irradiated, re-melted UHMWPE and gamma-irradiated, antioxidant-stabilized UHMWPE to improve implant clinical performance1,2.

Hindered phenol antioxidant (HPAO) has been demonstrated as an effective antioxidant for gamma-irradiated UHMWPE3. Gamma-irradiated, HPAO-stabilized UHMWPE provides favorable performance properties in comparison with gamma-irradiated, re-melted UHMWPE4. The objective of this study is to verify biocompatibility of gamma-irradiated, HPAO-stabilized UHMWPE.

This is a two-part study in which the first part evaluates effects of antioxidant content, radiation dose, accelerated aging and physical form on biocompatibility of UHMWPE implant material. The second part of this study covers comprehensive biocompatibility tests on a specific formulation.


Poly-phenol of tetrakis [3-(3,5-di-tert-butyl-4-hydroxyphenyl) propionate], PBHP, was used in this study as a model hindered phenol antioxidant. PBHP-containing UHMWPE was compression molded from GUR 1020 (Ticona) powder containing 0.075 % to 3.0 % (wt. / wt.) of PBHP. For powder form of HPAO-containing UHMWPE, PBHP was solution coated on GUR 1020 resin. Gamma-irradiation of UHMWPE samples was performed in vacuum foil pouch for nominal doses from 75 KGy to 100 KGy (Steris).

Table I shows material candidates prepared for four key screening tests in the first part of the biocompatibility study: (1) cytotoxicity study using the ISO elution method of 1 x Minimum Essential Medium extract, (2) ISO maximization sensitization study using 0.9% saline and sesame oil extracts, (3) ISO intracutaneous study using 0.9% saline and sesame oil extracts, (4) USP / ISO systemic toxicity study using 0.9% saline and sesame oil extracts.

The second part of the biocompatibility study covers comprehensive biocompatibility tests on 75 KGy, 0.075% PBHP-stabilized UHMWPE in consolidated form based on ISO 10993: (1) Cytotoxicity study using the ISO elution method with a Minimum Essential Medium (MEM) extract, (2) Murine local lymph node assays using DMSO and 0.9% saline extracts, (3) ISO intracutaneous study using 0.9% saline and sesame oil extracts, (4) ISO systemic toxicity study using 0.9% saline and sesame oil extracts, (5) Genotoxicity study using bacterial reverse mutation method with DMSO and 0.9% saline extracts, (6) Genotoxicity study using mouse lymphoma assays of DMSO and RPMI serum free extracts, (7) Mouse peripheral blood micronucleus study using 0.9% saline and sesame oil extracts, (8) ISO muscle implantation study – 2 weeks, (9) ISO muscle implantation study – 12 weeks, (10) ASTM hemolysis study, (11) 4-week systemic toxicity study in rats following subcutaneous implantation, (12) 26-week systemic toxicity study in rats following subcutaneous implantation

Standard HDPE implant samples were used as negative controls in all animal implantation studies. All comprehensive biocompatibility tests were performed in conformance with good laboratory practices.

Results

The screening biocompatibility tests showed that a combination of high PBHP content (0.3%) and high radiation dose (100 KGy) in UHMWPE does not induce adverse biological response. After accelerated aging, 100 KGy, 0.3% PBHP-containing UHMWPE likewise passed all screening biocompatibility tests. In powder form, gamma-irradiated, 0.075% PBHP-coated UHMWPE showed acceptable biocompatibility as its gamma-irradiated UHMWPE counterpart. The screening tests on UHMWPE powders shed light on the worst-case scenario when surface-coated PBHP is subjected to complete extraction by DMSO or sesame oil. Yet, there were no detectable adverse effects associated with 75 KGy-irradiated, 0.075% PBHP-containing UHMWPE powder.

The comprehensive biocompatibility study encompassed three categories of tests: (1) aqueous and lipophilic medium extractions followed by in-vitro cell culture studies, (2) aqueous and lipophilic medium extractions followed by in-vivo injection studies, and (3) long-term animal implantation studies. 75 KGy-irradiated, 0.075% PBHP-stabilized UHMWPE has passed all category 1 and 2 tests. It also shows comparable biocompatibility as HDPE control in the long-term animal implantation studies. Microphotographs of hematoxylin/eosin-stained muscle tissues surrounding the control implant (Figure 1L) and the test implant after 12-week rabbit implantation (Figure 1R) showed comparable characteristics. Moreover, microphotographs of hematoxylin/eosin-stained subcutaneous tissues surrounding the control implant (Figure 2L) and the test implant (Figure 2R) showed no signs of adverse effects after 26-week rat implantation.

Discussion

This biocompatibility study, based on ISO 10993 test protocols, demonstrates that there is no evidence of any adverse effects associated with 75 KGy, 0.075% PBHP-stabilized UHMWPE when used as an implant material in various pre-clinical settings. This finding is in line with clinical observations of biocompatibility associated with medical implant devices made from crosslinked polyethylene stabilized with another hindered phenol-type antioxidant5.

BIOCOMPATIBILITY STUDY OF GAMMA-IRRADIATED UHMWPE STABILIZED WITH A HINDERED-PHENOL ANTIOXIDANTKing, R.*; Arscott, E.; Narayan, V.DePuy Synthes Joint Reconstruction, a division of DePuy Orthopedics, Inc. Warsaw, IN

PBHP content

Radiation dose

Physical form

Accelerated aging

High AO content / Radiation dose

0.30 % 100 KGy Consolidated None

Accelerated aging 0.30 % 100 KGy Consolidated6-week 70° C/5

atm Oxygen

Physical form 0.075 % 75 KGy Powder None

Physical form None 75 KGy Powder None

Table I - Assessment of various parameters on biocompatibility

Figure 1L

Figure 2L

Figure 1R

Figure 2R

References1. McKellop, H., et al. (1999). Journal of Orthopedic

Research, 17(2), 157-167.

2. Mori, A., et. al. (2002). Trans. 48th Orthopedic Research Society, 1041.

3. King, R., et. al. (2009). Trans. 55th Orthopedic Research Society, 463.

4. King, R., et. al. (2009). Trans. 55th Orthopedic Research Society, 19.

5. Wroblewski, B.M., et. al. (2005). Journal of Bone and Joint Surgery, 87-B, 1220-1221.


Acknowledgement: All biocompatibility tests were performed at NAMSA.

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Introduction

Oxidative degradation in Ultra-High Molecular Weight Polyethylene (UHMWPE) has been the leading cause for delamination and fracture in air-packaged, gamma-sterilized articulating surface components. Post-radiation re-melting process has been used effectively to quench residual free radicals and enhance UHMWPE oxidative stability.1 Mori et al. have reported that addition of α-tocopherol to UHMWPE powder prior to consolidation improved its oxidative stability after gamma sterilization.2 Oral et al. have reported that incorporation of α-tocopherol to gamma-irradiated UHMWPE block further enhances its mechanical properties and fatigue strength.3

One of the concerns with α-tocopherol, a liquid antioxidant, has been its potential tendency to migrate out of the UHMWPE matrix during storage and under loading conditions. Therefore, the question of whether α-tocopherol is the most effective antioxidant for UHMWPE implants may be debated.

Hindered Phenol Antioxidants (HPAO’s) have been used to stabilize polyethylene for decades. Typical applications include food-contact package materials. The objectives of this study were to assess efficacy of a representative HPAO as an alternative antioxidant for UHMWPE and to investigate mechanical properties of such a gamma-irradiated, HPAO-stabilized UHMWPE.


An antioxidant-stabilized UHMWPE was made from GUR 1020 (Ticona) powder containing 0.075 wt % of a typical HPAO, compression molded and gamma-irradiated at a nominal dose of 75 KGy (Steris) in a vacuum foil pouch. Two control materials were used in this study: (1) conventional UHMWPE made from GUR 1020 powder, compression molded and gamma-irradiated at 40 KGy in a vacuum foil pouch, and (2) crosslinked UHMWPE made from GUR 1020 powder, ram extruded, gamma-irradiated at 50 KGy in a vacuum foil pouch, and then remelted at 155° C to quench residual free radicals. This material is commercially available as XLK.

Uniaxial tension tests were performed on Type IV specimens per ASTM D 638 and double-notched Izod impact tests were performed per ASTM F 648. The fatigue crack propagation tests were performed based on ASTM E647. Crystallinity and melting point data were generated per ASTM F 2102 while swell ratio data were generated per ASTM F 2214. The wear tests were performed using a bi-directional AMTI pin-on-disk tester with 90% bovine serum as lubricant.

The oxidation resistance of each material group was measured by accelerated aging respective samples for 2 weeks at 70° C in a 5-atm oxygen chamber, followed by analysis using a Fourier Transform Infrared Spectroscopy (FTIR) microscope. An oxidation index was calculated by normalizing the area under the carbonyl vibration to the area under the 1370 cm-1 absorbance. Each HPAO molecule contains four ester bonds and it contributes to FTIR carbonyl absorption and thus to the calculated oxidation index. A corrected oxidation index for post-accelerated-aged, gamma-irradiated, HPAO-stabilized UHMWPE was calculated by subtracting the baseline HPAO carbonyl absorption from the measured oxidation index average.

Results

Table I lists tensile mechanical data, impact strength data and stress intensity factor at crack inception measured during crack propagation experiments for all three material groups. Combination of high radiation dose and HPAO presence has no adverse effect on mechanical properties of gamma-irradiated, HPAO-stabilized UHMWPE; while DNI data for gamma-irradiated, HPAO-stabilized UHMWPE is typical of that of highly crosslinked UHMWPE. Nevertheless, 75KGy, HPAO-stabilized UHMWPE shows improved fatigue crack growth resistance in comparison with that of 50 KGy, re-melted UHMWPE.

Discussion

In this study, it has been demonstrated that HPAO’s provide excellent oxidation stability to gamma-irradiated UHMWPE without employing a remelting process to quench free radicals. This allowed the gamma-irradiated, HPAO-stabilized UHMWPE to provide better mechanical properties and fatigue resistance relative to the gamma-irradiated, remelted UHMWPE. Biocompatibility of HPAO-stabilized UHMWPE has been recently verified per the ISO 10993 protocol. Therefore, the study provides strong justification for the employment of a hindered phenol-stabilized UHMWPE for orthopaedic devices with enhanced performance characteristics relative to current UHMWPE materials.

CHARACTERIZATION OF GAMMA-IRRADIATED UHMWPE STABILIZED WITH A HINDERED PHENOL ANTIOXIDANTKing, R.; Narayan, V.S.; Ernsberger, C.; Hanes, M. DePuy Synthes Joint Reconstruction, a division of DePuy Orthopedics, Inc. Warsaw, IN

40 KGy UHMWPE

50 KGy Re-melted UHMWPE

75 KGy HPAO Stabilized UHMWPE

Crystallinity, % 60.0 ± 0.4 54.1 ± 1.0 58.7 ± 0.7

Melting point, °C 136.6 ± 0.2 134.5 ± 0.4 138.2 ± 0.3

Swell ratio 3.65 ± 0.06 3.50 ± 0.01 3.73 ± 0.04

POD wear, mg/MC 8.38 ± 0.13 4.23 ± 0.21 5.55 ± 0.34

Average oxidation index after acc. aging

0.506 ± 0.227 0.045 ± 0.065 0.000* (*after correction)

Table II - Thermal, chemical and wear characteristics of crosslinked UHMWPE

References1. McKellop, H.; et al., (1999) Journal of Orthopedic

Research, 17(2), 157-167.

2. Mori, A.; et al. (2002). Trans. 48th. Orthopedic Research Society, 1014.

3. Oral, E.; et al. (2005). Trans. 51st. Orthopedic Research Society. 1661.

Presented at the Orthopaedic Research Society, 55th Annual Meeting, Poster No. 19 , 2009, Las Vegas, NV

40 KGy UHMWPE

50 KGy Re-melted UHMWPE

75 KGy HPAO Stabilized UHMWPE

Yield tensile strength (MPa) 23.0 ± 0.3 21.5 ± 0.3 23.2 ± 0.2

Ultimate tensile strength (MPa) 40.6 ± 0.8 40.1 ± 1.4 46.1 ± 1.4

Elongation at break, % 339 ± 5 305 ± 5 335 ± 11

Double-notched Izod impact strength, MPa/m2 93.7 ± 1.0 80.2 ± 1.2 73.6 ± 3.7

ΔKincep

, MPa(m)**1/2 1.76 ± 0.05 1.57 ± 0.00 1.88 ± 0.00

Table I - Mechanical properties of various crosslinked UHMWPE

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Introduction

Total knee replacement (TKR) wear testing is a key stage in the pre-clinical analysis of new implant designs. Experimental and computational methods are used, but understanding of wear remains limited, with only empirical predictive models available. Due to limited data and inherent uncertainties, multiple competing theories of polyethylene wear exist. It is important to resolve these different ideas of wear mechanisms in order to progress the science of TKR tribology. A number of researchers have performed testing using different activities of daily living (ADLs) e.g. stair ascent1 and descent.2 These more aggressive profiles may help to better differentiate between wear theories. Here, we consider how different wear theories would respond to different ADL profiles. Multiple designs are compared to avoid implant-specific artifacts and ensure the conclusions are more general.

Methods

The baseline model is a previously demonstrated3 model of a force-driven knee wear simulator (MSC.ADAMS, MSC Software, CA). Asymmetric spring constants were used based on a resected ACL / retained PCL.4 Four fixed-bearing posterior-stabilized implants were compared: ATTUNE™ (DePuy Orthopaedics, Inc.), NexGen® (Zimmer), P.F.C.® SIGMA® (DePuy Orthopaedics, Inc.) and Triathlon® (Stryker), to evaluate the effect of different model variables across various designs, and so ensure any observed trends were not simply design-specific. Three force-driven activity profiles were compared, based on literature-derived data: ISO gait5, stair ascent1 and stair descent.2 Custom code (MATLAB, Mathworks, CA) was used to implement three wear models: 1st generation (Archard)6, 2nd generation (with A/A+B7 cross-shear term to account for multi-directional sliding) and 3rd generation8 (taking account of time-history to capture variations in wear rate with sliding-path scaling, such that smaller paths will have a higher wear rate per unit sliding distance).9 Both the 2nd & 3rd generation models were also evaluated with and without10 a contact pressure (CP) term.10 Linear wear depth and overall wear volume were compared for the three different wear theories, for each implant and for each ADL profile.

Results

As anticipated by experimental data, the wear rates for the alternative stair-based ADL profiles were generally higher than for gait across different designs, although this was not a universal trend (fig. 1). The 1st generation Archard model only anticipated higher wear for stair ADLs if a CP term was included; without a CP term, the 1st generation model was not a sensitive predictor. Differences between the 2nd & 3rd generation models (fig. 2) were smaller than differences between models with & without CP (fig. 3). Agreement between 2nd and 3rd generation models gives confidence in the consistency of more recent theories. Models with linear CP dependency showed higher wear rates for stair ADLs relative to gait (on average 3× higher for stair descent). Stair descent consistently showed the greatest wear when CP was included; without the CP term this trend was less strong for both 2nd and 3rd generation models, although the increase in multi-directional sliding for stair profiles still resulted in higher wear on average compared to ISO gait.

Discussion

Polymer wear is known to be a complex multi-phenomenon process; refining models with more accurate theories is challenging. This study emphasizes that resolving the role of CP in wear is the more statistically important question than further refinement of cross-shear theories. The results we present here suggest that data from such ‘aggressive’ ADL profiles most clearly differentiates differences in the underlying wear models, and with sufficient supporting experimental data this may enable identification of the best wear theories. Presently, experimental studies disagree on the effect of aggressive ADLs on wear rate; some suggest considerably increased wear2, others do not.1 More experimental testing with a wider range of ADLs may help to clarify this state of knowledge, and so help further develop future wear theories. Such studies need to be based on multiple different implant designs, to avoid artifacts of design-dependency.

Significance

This work aims to identify limitations to the current understanding of polymer wear. If the mechanisms of polymer wear can be better understood, this will directly impact and improve the design of future orthopaedic implants, leading to lower-wearing devices available for clinical use.

References

1. Cotrell et al 2006 J.Biomed.Mater.Res.B 78(1),p15.

2. Benson et al 2002 Proc.Inst.Mech.Eng.H 216(6),p409.

3. Strickland et al 2009 J.Biomech 42(10),p1469.

4. Haider et al 2002 Trans. 48th ORS.

5. ISO Standard 14243-1.

6. Archard 1953 J.App.Phys 24(8),p981.

7. Turell et al 2003 Wear 255(7-12),p1034.

8. Strickland et al 2010 WCB 6th Meeting.

9. Dressler et al 2011 Wear 10.1016/j.wear.2011.06.006

10. Ernsberger et al 2007 Trans. 53rd ORS.






MULTI-ADL PROFILES IN TKR WEAR TESTING HELP DISCRIMINATE BETWEEN WEAR THEORIES+*Strickland, A.M, *Taylor, M

*University of Southampton, Southampton, UK

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Figure 3: Comparison of volumetric wear rate for cross-shear based wear model with (left) and without (right) CP term.

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Introduction

An estimated 46 million Americans suffer from osteoarthritis.1 As tissue engineered treatments and other options remain on the horizon, total knee arthroplasty (TKA) continues to be an efficacious solution for severe osteoarthritis and has been increasingly applied to younger more active patients.2 Consequently, new TKA devices are emerging that address previous challenges to long term success such as momentary instability, backside wear of modular devices, and oxidation of UHMWPE. The objective of this study was to evaluate the in vitro wear of a knee system with a novel femoral condyle geometry based upon a gradually changing sagittal curvature, refined tibio-femoral articulation, a new tibial insert locking mechanism, and an anti-oxidant containing poly material.


Wear testing was conducted on midsized SIGMA® and ATTUNE™ Fixed Bearing Cruciate Retaining TKA Components (DePuy Orthopaedics, Inc., Warsaw, IN) with a 10mm composite thickness. SIGMA Implants included components with CoCr femoral components, CoCr polished tibial trays, and curved tibial inserts with “XLK” moderately crosslinked (50 kGy) remelted compression molded GUR 1020 UHMWPE. ATTUNE Implants included cast CoCr femoral components, CoCr polished tibial trays, and “AOX™” Tibial Inserts machined from compression molded GUR 1020 UHMWPE with Covernox™ (DePuy), a solid antioxidant. The AOX Insert is irradiated to 80 kGy due to the attenuation of the crosslinking process by the antioxidant. No remelting is involved. Tibial inserts were presoaked 42 days in water at room temperature prior to testing. Wear simulation was conducted (n=3 each group) using a six station wear simulator (AMTI, Watertown, MA) at 1Hz using controlled inputs (Table 1). Flexion-extension (FE) and axial force profiles were adopted from ISO 14243-3.3 Anterior-posterior (AP) translation and internal-external (IE) rotation inputs were based on natural knee kinematics.4 Testing was performed in 25% bovine calf serum (Hyclone, Logan UT) at 37±2°C supplemented with sodium azide and EDTA. Tibial inserts were weighed on an analytical balance (XP205, Mettler-Toledo, Columbus, OH) prior to testing and then every 0.5 million cycles (Mcyc), which corresponded to the lubricant change intervals. Weights of two loaded soak controls were used to compensate for fluid uptake in the wear specimens. Wear rates were calculated by linear regression and then compared between groups using a student t-test (α=0.05). Surface finish of the metal components was evaluated pre- and post-simulation using a Talysurf Series II Profilometer (Taylor Hobson, West Chicago, IL). Serum samples were collected during the test to characterize the UHMWPE debris per ASTM 1877-05 using Low Angle Laser Light Scattering (LALLS) and quantitative Scanning Electron Microscope Elemental X-ray analysis (SEM/EDAX).

Results

The wear rate of the ATTUNE Knee Design with AOX Polyethylene was significantly less than the SIGMA Knee Design with XLK Polyethylene (Figure 1). The contact scar on the proximal surface of the tibial insert from both groups displayed similar burnishing and mild scratching. Contact scars on the distal bearing surfaces of both groups displayed a glossy-burnished appearance under the high load area, typically seen when articulated against a polished CoCr tibial tray. Femoral and tibial components experienced little change in surface roughness (Table 2). In addition, the tibial trays showed no evidence of the stippling that is commonly seen with tibial tray/insert combinations having appreciable micromotion.5 The shape of the UHMWPE wear particles were very similar (Table 3) and the mean diameters were not statistically significant although the ATTUNE AOX polyethylene was larger.

Discussion

The ATTUNE Knee System wore 50% less than the clinically successful SIGMA Knee System under these specific laboratory conditions. The difference in wear is primarily the effect of design as the AOX Material was developed to have similar wear properties to XLK.6 The antioxidant used in the AOX Material provides excellent long term oxidative stability7 and, interestingly, necessitates a higher irradiation dose to achieve the same level of crosslink because the antioxidant actively competes with the crosslinking process. The locking mechanism has been shown to allow very little motion8, which contributes to limiting backside wear.5 The motions of the simulation were displacement controlled; however, the different sagittal curvatures create distinctly different contact kinematics between the two devices. Current and ongoing studies are investigating the different cross-shear conditions to elucidate the potential wear benefit from the gradually changing sagittal curvature of the ATTUNE System compared to the discrete changes in radius of the SIGMA System.

Significance

This wear simulation evaluated a novel TKA device with features that are thought to improve patient tibio-femoral kinematics, wear, and long term oxidative stability. The in vitro wear results indicate an improvement compared to a current clinically successful TKA.

References

1. Helmick, et al, Arthritis Rheum, 2008.

2. Healy, et al, JBJS, 2008.

3. ISO 14243-3:2004 Wear of total knee-joint prostheses.

4. McEwen, et al, J Biomech, 2005.

5. McNulty, et al, ORS, 2005.

6. King, et al, ORS, 2009.

7. Narayan, et al, ORS, 2010.

8. Leisinger, et al, ISTA, 2011.

Acknowledgements: Many thanks to Grant Magner for his invaluable assistance.

Poster # 0959, from the ORS 2012 Annual Meeting, February 2012, San Francisco, CA.

Femoral Ra (um) Tibial Tray Ra (um)

Pre-test Post-test Pre-test Post-test

SIGMA 0.052±0.005 0.076±0.028 0.049±0.001 0.024±0.004

ATTUNE 0.053±0.001 0.050±0.008 0.054±0.004 0.049±0.006

Table 2: Surface roughness (Ra) of femoral components and tibial trays.

ATTUNE AOX SIGMA XLK

Mean Diameter (um) 118.8 64.4

Aspect Ratio 2.10 2.06

Roundness 0.56 0.56

Perimeter (um) 3.56 3.17

Table 3: Particle analysis of wear debris.

Figure 1: Wear for SIGMA XLK and ATTUNE AOX Knee Systems. Inset shows wear rate. Error bars represent ±1 standard deviation.

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Peak Axial Force 584 lbf 584 lbf

Flexion-Extension Angle 0° – 58° 0° – 58°

Internal-External Rotation -5° – 5° -2° – 6°

Anterior-Posterior Translation

10mm 5mm

Axial Force Offset 4.8mm medial offset

Table 1: Summary of customized inputs compared to ISO 14243-3.

WEAR OF A TOTAL KNEE REPLACEMENT WITH ANTIOXIDANT UHMWPE AND GRADUALLY VARYING SAGITTAL CURVATURE+MR Dressler1, S Swope1, J Tikka1, C Hardaker1, M Heldreth1, T Render1

+1DePuy Orthopaedics, Inc., Research, Warsaw, IN

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