Experimental and analytical validation of a modular acetabular prosthesis in total hip arthroplasty

BioMed Central

Journal of Orthopaedic Surgery and Research

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Open AcceResearch articleExperimental and analytical validation of a modular acetabular prosthesis in total hip arthroplastyFrancisco Romero1, Farid Amirouche1,2, Luke Aram1 and Mark H Gonzalez*1,2
Address: 1Department of Mechanical Engineering, Biomechanics Laboratory, University of Illinois at Chicago, 842 W. Taylor Street, Room 1039, Chicago, Illinois 60607, USA and 2Department of Orthopaedic Surgery, University of Illinois at Chicago, 1801 W. Taylor Street, Chicago, Illinois, USA

Email: Francisco Romero - [email protected]; Farid Amirouche - [email protected]; Luke Aram - [email protected]; Mark H Gonzalez* - [email protected]

* Corresponding author

AbstractA finite element model has been developed to predict in vivo micro motion between a modularacetabular cup and liner after cement less total hip arthroplasty. The purpose of this study is toexperimentally validate the model. Six LVDT sensors were used to monitor the micromotion ofthe liner when subjected to loading conditions ranging from 250 N to 5000 N. Deformations atpoints of interest for both the experiment and FEM were compared. Results of the FEM withdifferent coefficient of friction between the liner and the cup were investigated to correlate withthe experimental results.

BackgroundPolyethylene wear particles generated from the acetabularliner of an acetabular cup are a cause of the osteolysisassociated with acetabular cup failure. Wear generationcomes from the articulation of the femoral head and theliner and from the articulation of the liner and the metalshell. Wear generated by micromotion at the liner metalshell interface, termed backside wear, is potentially moredamaging because of the egress of the particles to the cupbone interface through the screw holes and the centralfenestration in the metal shell. These particles then createan inflammatory reaction that can produce osteolysis ofthe bone, leading to degradation of the interface andeventual failure. The etiology of micromotion betweenthe metal shell and polyethylene is multifactorial. Impor-tant factors include the contact stress [10], relative slidingdistances [11], conformity between contacting surfaces

and the number of screw holes in the acetabular shell[12]. A rigid locking mechanism is paramount in limitingmicromotion at the liner shell interface. Conformity ofthe liner shell interface and polishing of the metallic shellare new advances to limit fractional wear.

The loading of the hip joint is cyclical and occurs duringgait. Impact loading of the hip joint can also occur as aresult of subluxation and relocation of the metallic ball inthe polyethylene cup. Little work has been done on theeffect of physiologic loads on micromotion at the cupliner interface. The load patterns are particularly complexduring subluxation relocation because the initial contactat relocation may be eccentric creating a moment. The cyl-ical pattern of loading associated with level gait, stairclimbing and the impact from eccentric subluxation andrelocation are virtually impossible to reproduce physically

Published: 16 May 2007

Journal of Orthopaedic Surgery and Research 2007, 2:7 doi:10.1186/1749-799X-2-7

Received: 12 May 2006Accepted: 16 May 2007

This article is available from: http://www.josr-online.com/content/2/1/7

© 2007 Romero et al; licensee BioMed Central Ltd. This is an Open Access article distributed under the terms of the Creative Commons Attribution License (http://creativecommons.org/licenses/by/2.0), which permits unrestricted use, distribution, and reproduction in any medium, provided the original work is properly cited.

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in the laboratory. To this end a finite element model hasbeen created to investigate micromotion of the metal shellliner interface. The purpose of this experiment is to vali-date this model. The validation of the model quantifiesmicromotion of a loaded acetabular cup liner assemblywith linear variable differential transducer (LVDT) sen-sors.

Methods2.1. Experiment set-upSeven commercially available UHMWPE liners, 28 mminner diameter and 54 mm outer diameter, were used. Theexperiment set-up is represented in Figure 1.

The acetabular cup was placed into the acetabular shell asrecommended by the manufacturer. The acetabular cupwas then placed into a custom made jig that held the cuprigidly. (Figure 1 supporting jig). Six circumferential setscrews were tightened rigidly to the acetabular cup withinthe jig. The supporting jig was designed to hold theacetabular shell at a constant angle of 45° with respect tothe horizontal simulating a 45° inclination angle. For thisexperiment the anteversion angle was fixed at 15–20°.The acetabular cup is statically placed at 25 degrees ofanteversion

Two acetabular shells (66 mm outer diameter) were usedduring all the experiments. The first had only a polar fen-estration and the second had polar fenestration and threescrew holes. A sensor positioning plate containing six cal-ibrated LVDT sensors (AC-DC LVDTS 332, TRANS-TEKInc, Ellington, CT) was rigidly attached to the jig. TheLVDT sensors detect linear micromotion in the x, y and zdirection. The metal tip of each LVDT sensor was set tozero position when it came in contact with the liner beforeany loading conditions were applied. Care was taken tolocate the sensors' tips away from the liners' anti-rotationtabs. Three of the sensors contacted the liner perpendicu-larly to the outer flat surface, while the other three con-tacted the liner perpendicularly from the side (Figure 2).The displacements measured using the six LVDT sensorscorrespond to the local deformation of the polyethyleneliner at the points of contact.

The jig sensor assembly was positioned in an INSTRONmachine (Model 5579, Instron Corp., Canton, MA) forloading. Load was applied to the acetabular componentsthrough a 28 mm, commercially available femoral headattached to the movable crosshead of the INSTRONmachine. The femoral head was positioned in the acetab-ular liner and adjusted until a conforming contact wasachieved without applying significant load. The readingson the sensors were still showing zero displacements. Thisserved as the starting point for experimental measure-ments of load and micromotion.

The loading characteristics of each trial were set with theINSTRON computer controller (Merlin 5500 Series). Theloading characteristics applied to the liner (includingnumber of cycles, speed, and maximum load) were speci-fied. Each specimen (liner) was subjected to the loadingprofiles described in Table 1.

The LVDT sensors were connected through a Data Acqui-sition Card (DAQCard-A1-16E-4, National Instruments,Austin, TX) to a laptop computer. Using data analysis soft-ware (Lab-View, National Instruments, Austin, TX), a pro-gram was developed to transform the signals from theLVDT sensors into micromotion displacement values. Themicromotion values from the LVDT sensors were synchro-nized with the loads applied with the INSTRON machine.

The FEM was then used to generate values for micromo-tion at the points of attachment of the LVDT on the poly-ethylene liner. The comparison of results from theexperimental and computational models allowed us tovalidate the finite element model of the modular acetabu-lar component.

2.2. Modular acetabular component: CAD modelThree-dimensional computer models of the modularprosthetic components (femoral head, acetabular linerand acetabular shell) were developed using Pro/Engineer2000i2 (Parametric Technologies Inc., Needham, MA).Component dimensions used in the models were basedon precise measurements taken from commercially avail-able products (DePuy, Inc., Warsaw, IN) using handinstruments.

The model was created with a polar fenestration. The sameconsideration was made to model the liner. Geometricaldetails without structural significance were omitted inorder to import CAD models to the standard ANSYS for-mat (IGES). All 3-D models were converted to IGES for-mat (No-defeaturing option).

2.3. Finite Element modelA three dimensional 3595 elements FE model (Figure 3)of the femoral head, liner, and acetabular shell was devel-oped based on the CAD geometry previously createdusing ANSYS/LS-DYNA. The diameter of the articulatingsurface femoral head/liner measured 28 mm while theliner backside diameter measured 54 mm. A 12 mm diam-eter polar fenestration was placed in the metal acetabularshell which corresponds to the central insertion hole. Theacetabular shell was positioned with 45° inclination and25° anteversion [13].

The Finite Element mesh of the acetabular shell, liner andfemoral head consisted of eight-node hexahedron solidelements (Table 2). A polar mesh design was selected to



avoid the nodal penetration problem (Figure 4). Themesh pattern was kept the same in all the contacting areas(outer surface of the femoral head-liner front side andliner back-side and acetabular shell) so that, the nodesplaced in the contacting surfaces do not interfere with thetarget area. A cross section of the liner, acetabular shelland femoral head were first meshed due to axial symmet-ric conditions of the model. Planar elements were thenused having perfect control of the shape of the elements.The planar elements were revolved with a specific anglecreating a polar mesh design. The angle of revolution waschosen to avoid singularities that can occur when hexahe-

dron elements are converted to tetrahedron elements[14].

A linear isotropic material was chosen to model theUHMWPE liner for the proposed experiment. For a studyof wear and fatigue a non linear material will be moreappropriate. The material properties of the liner werebased on the information obtained in the available litera-ture [15], where the Young's Modulus = 975 MPa, andPoisson's Ratio = 0.46. A rigid material was chosen for theacetabular shell and the femoral head components

A linear contact between implant components was mod-eled using an automatic surface-to-surface contact (ASTS).The acetabular shell was fully constrained, assuming arigid union between the acetabular shell and the acetabu-

Modular Acetabular components set-up and LVDT sensors' positioningFigure 2Modular Acetabular components set-up and LVDT sensors' positioning.

Experiment set-upFigure 1Experiment set-up. Orientation of the acetabular compo-nents (acetabular shell, liner and femoral head) with respect to applied load.

Finite Element Model of the acetabular shell, liner and femo-ral headFigure 3Finite Element Model of the acetabular shell, liner and femo-ral head. The liner locking mechanism was simulated con-straining all degrees of freedom of the nodes located at the same positions as locking tabs of the real-life liner.

Contacting areas (Acetabular Shell/liner and Liner/femoral head) involved in the Finite Element ModelFigure 4Contacting areas (Acetabular Shell/liner and Liner/femoral head) involved in the Finite Element Model.




Table 2: Material Properties

Component N° element FEM mat. Actual mat.

Femoral Head 501 Rigid Cobalt-ChromeLiner 1799 E = 975 MPa

ν = 0.46ρ = 960 Kg/m3

UHMWPE

Acetabular Shell 720 Rigid Titanium

Comparison between the material properties Finite Element model and real-life components material properties.

Table 1: Loading profile applied to each liner specimen.

Loading case Speed (mm/min) Cycles Maximum Load (N)

1 5 20 2502 5 20 3503 5 20 4504 5 20 5505 5 20 6506 5 20 7507 5 20 8508 5 20 9509 5 20 105010 5 20 115011 5 20 125012 5 20 135013 5 20 145014 5 20 155015 5 20 165016 5 20 175017 5 20 185018 5 20 195019 5 20 205020 5 20 215021 5 20 225022 5 20 235023 5 20 245024 5 20 255025 10 25 265026 10 25 275027 10 25 285028 10 25 295029 10 25 305030 10 25 315031 10 25 325032 10 25 335033 10 25 350034 10 25 375035 10 25 400036 10 25 425037 10 25 450038 10 25 475039 10 25 500040 10 25 600041 0.5 2 10000

TOTAL 882

In each loading case the load range applied goes from 0 N to the maximum load specified for each loading case.


lum. The femoral head was constrained with respect to therotational degrees of freedom. Perfect conformity betweenfemoral head and liner as well as between liner andacetabular shell was assumed. Anti-rotation features suchas the equatorial locking tabs provide the greatest restraintto keep the liner in the proper position within the shell[16]. When load is applied to a modular acetabular com-ponent with locking constraints, the liner locking tabsshare the greater part of the applied load. In order to sim-ulate the six locking equatorial tabs of the liner, a total oftwelve nodes of the liner FE model were fully constrained(Figure 3).

2.4. FE validationA similar loading profile to the one used in the experi-ment was created in ANSYS/LS-DYNA in order to validatethe experimental results in terms of stability/micromotionof the liner. The load was applied at the center of the fem-oral head and varied from no load up to 1775 N appliedlinearly. The maximum load was based on previous stud-ies in gait [17-21] where the hip joint forces were evalu-ated. In order to reproduce the experimental conditions asclosely as possible both the orientation of the appliedload (vertical) and the position of the acetabular compo-nent (45° inclination) were the same in the computermodel and the experiment set-up. (Figure 5)

Published reports of the friction coefficient between themetal and polyethylene ranged from μ = (0.083 to 0.2).[[15],15a] The finite element model was run with frictioncoefficient values that varied between the given range andresults of the FE with the minimum and maximum coeffi-cient of friction were used to compare the analytical solu-tion with the experimental data. Further discussion isprovided in the section below.

ResultsFigure 6 shows the micromotion measured for each of theseven liner specimens studied. Each graph corresponds toone liner specimen. The six curves in each graph representthe maximum micromotion achieved for each of the sixLVDT sensors. The maximum physiological loads experi-enced in hip are within a range of 1700 to 2200 N. Asdescribed in this study, the maximum load applied toeach liner specimen during the validation experiment was10,000 N, which is beyond the attainable force betweenthe acetabular liner and the femoral head in the body. Theobjective of carrying the experiment beyond 2200 N is toinvestigate how the liner responds to higher loads andwhether there is drastic change in deformation (plasticdeformation) where certain peak forces are achieved.

As described previously, a loading profile, similar to theone in the experiment, was used in the FE model. Accord-ing to the experimental set-up, the micromotion resultscorresponded to six specific points, where the sensors' tipscame in contact with the liner. The corresponding pointswere identified in the FE model. Because the LVDT sensorsonly measure displacements along their longitudinal axes,those directions were identified with simple trigonomet-ric calculation in the FE model through superposition ofthe experimental axes. The micromotions (displace-ments) were computed in the FE model of the nodes posi-tioned at the reference points, along the identifieddirection.

Figures 7 and 8 indicate the maximum and minimumdeformation (micromotion) values, for all the seven lin-ers analyzed, where each sensor was compared to theresult obtained with the finite element model. The FEmodel results are shown for two values of coefficient offriction (μ = 0.2 and μ = 0.083). As expected, the sensorspositioned perpendicular to the outer flat surface of theliner recorded the most significant micromotion. This isattributed to the fact that this is the direction along whichthe liner is less constrained. The three sensors positionedperpendicularly to the side of the liner recorded very littlemicromotion. This micromotion was produced becauseof local deformations from the bulging effect of the poly-ethylene when subjected to compressive loads.

Coordinate system used as a reference for the loads in the FE modelFigure 5Coordinate system used as a reference for the loads in the FE model. The CS is shown with respect to the acetabular shell inclination angle.




Maximum micromotions achieved for each of the seven liner specimen analyzedFigure 6Maximum micromotions achieved for each of the seven liner specimen analyzed.

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In all the experimental data collected from the seven spec-imens the result obtained in the FE model, for both fric-tion coefficients, are between the maximum andminimum values obtained in the experiment. This is afirst check on the validity of the model and its analyticalprediction. Considering the average values for the experi-ment results (based on the maximum and minimum

micromotion values collected for each load), the FEmodel with the lowest value of showed a closer fit anddepicts a more realistic response. When modeling con-formity it is important to note that the actual liner and cupdo not possess such a characteristic, in fact there are gapsdue to the geometrical difference of the two components.These clearances will make the friction between the linerand the cup less pronounced; hence a relaxed value in theFE model is more appropriate. Our validation supportssuch a scenario.

DiscussionThis study highlights two main points. One the experi-ment provides an insight into the stability and micromo-tion of the liner/cup interface. Indeed it provides themeans by which we can quantify the micromotion/loadinterface in relation to the liner and locking mechanismdesigns. Second these experiments along with the FEmodel becomes an important tool in identification ofweak spots and areas where both stress and wear can beidentified. The objective of this research while focused onseveral specimens the design of the locking mechanism ofthe liner with respect to the cup is the same. In future stud-ies we intend to perform a much larger study where differ-ent locking mechanisms designs can be compared underdifferent loads and moments.

The current investigation has prompted us to suggest areaswhere we can improve both the FE model as well as learn-ing further how the liners designs affect the results ofmicromotion and wear. These include: the conformity ofthe contact surfaces, geometry and tolerances of the CADmodel, representation of the liner locking mechanism andmaterial properties. Conformity allows distribution ofload more along the equatorial surface of the load andmight increase back wear when the material expand andslide along the surface at he interface of the cup/liner. Theconformity might also play a significant role in load trans-fer to the cup and hence changing the micromotion con-dition at the cup/bone interface. The cup bone interface isa subject of another investigation currently being con-ducted at our Laboratory.

The results of the validation show that the correlation ofthe micromotion data given by the FE model and theexperimental data depend upon the position of each sen-sor. The sensors positioned along the flat surface of theliner recorded values that closely matched the model.Among those three sensors, the one positioned further tothe axis of application of the load registered the bestresults. The other two sensors positioned in the flat sur-face of the liner (red and purple) matched closely theresults only in the load ranges between 1500 N and 1750N. Furthermore, the results obtained for the three sensorspositioned perpendicularly to the flat surface of the liner

Maximum and minimum micromotions values, among all the specimens analyzed, recorded for each sensorFigure 7Maximum and minimum micromotions values, among all the specimens analyzed, recorded for each sensor. Comparison with the results obtained with the FE model considering two different friction coefficients.

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are in the same range of the values obtained from theexperiment. Therefore, it is clear that the sensors subjectedto higher bulging/deformation showed a slight deviationfrom those in the FE model. These variations might be aconsequence of the assumptions made in selection of pol-yethylene material. A more precise material model wouldcertainly describe the bulging effects more accurately. Forexample a more accurate description of the material usedin the model would be a piecewise linear isotropic plastic-

ity material [16]. This planned in our future researchdevelopment.

Another important assumption made in the model is thetreatment of the liner's locking tabs. As describe earlier,the locking tabs were simulated constraining all thedegrees of freedom of the nodes placed in the same posi-tion as the actual locking tabs. This assumption states thatthe nodes simulating the locking tabs were completelyfixed (constrained) to the acetabular shell and may inducehigh stress concentration areas. In reality those tabs aredesigned to have significant gap around them when theliner is fully seated. Surely this factor might permit a moreeven stress distribution in the elements surrounding theconstrained nodes of the locking tabs. This might result ina smaller peak stress values at the nodes.

The last hypothesis considered in the present study is con-cerned with the conformity of the contact areas. Perfectconformity was assumed between liner's backside andacetabular shell inner-surface and between liner's front-side surface and femoral head. In the actual acetabularcomponents certain gaps are allowed between contactingsurfaces for both design and manufacturer tolerances. Thestress distributions and micromotions achieved in theliner might be different if those gaps were considered andthe results obtained might be closer than the one obtainedin this study.

ConclusionThe objective of the present study was to validate a FEmodel of modular acetabular prosthesis with data col-lected experimentally. The model is preliminary and givesa reasonable approximation of values of micromotionobtained experimentally. Further work is needed toenhance the model development and improve the accu-racy of the model.

AbbreviationsE (Young Modulus): MPa

Kg (Kilogram): Unit of mass (SI)

m (meter): Unit of Length (SI)

MPa (Mega Pascal): Unit of Pressure (SI)

N (Newton): Unit of Force (SI)

min (minute): Unit of Time equal to 60 seconds.

mm (millimeter): Unit of Length equal to one 10-3 meters.

sec (Second): Unit of Time (SI)

Maximum and minimum micromotions values, among all the specimens analyzed, recorded for each sensorFigure 8Maximum and minimum micromotions values, among all the specimens analyzed, recorded for each sensor. Comparison with the results obtained with the FE model considering two different friction coefficients.

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μ (mu): Friction Coefficient (Unit-less)

μm (micron):Unit of Length equal to 10-6 meters

ρ (rho): Symbol for density. Density units Kg/m3

ν (nu): Poisson's ratrio (Unit-less)

AcknowledgementsTodd Render and Paul Lewis, DePuy (Johnson and Johnson), Warsaw, Indi-ana.

This experiment was supported in part from unrestricted research funds from DePuy.

Jeff Goldstein. University of Illinois at Chicago. Lab Assistant.

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Experimental and analytical validation of a modular acetabular prosthesis in total hip arthroplasty

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