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Research Papers: Contact Mechanics

Wear Simulation of Metal-on-Metal Hip Replacements With Frictional Contact

[+] Author and Article Information
Lorenza Mattei

Researcher
e-mail: l.mattei@ing.unipi.it

Francesca Di Puccio

Assistant Professor
e-mail: dipuccio@ing.unipi.it
Department of Civil and Industrial Engineering
University of Pisa
Pisa 56122, Italy

1Corresponding author.

Contributed by the Tribology Division of ASME for publication in the JOURNAL OF TRIBOLOGY. Manuscript received August 13, 2012; final manuscript received December 16, 2012; published online March 18, 2013. Assoc. Editor: James R. Barber.

J. Tribol 135(2), 021402 (Mar 18, 2013) (11 pages) Paper No: TRIB-12-1126; doi: 10.1115/1.4023207 History: Received August 13, 2012; Revised December 16, 2012

Preclinical wear evaluation is extremely important in hip replacements, wear being one of the main causes of failure. Experimental tests are attractive but highly cost demanding; thus predictive models have been proposed in the literature, mainly based on finite element simulations. In such simulations, the effect of friction is usually disregarded, as it is considered not to affect the contact pressure distribution. However, a frictional contact could also result in a shift of the location of the nominal contact area, which can thus modify the wear maps. The aim of this study is to investigate this effect in wear prediction for metal-on-metal implants. Wear assessment was based on a purpose-developed mathematical model, extension of a previous one proposed by the same authors for metal-on-plastic implants. The innovative aspect of the present study consists in the implementation of a modified location of the nominal contact point due to friction, which takes advantage of the analytical formulation of the wear model. Simulations were carried out aimed at comparing total and resurfacing hip replacements under several gait conditions. The results highlighted that the adoption of a frictional contact yields lower linear wear rates and wider worn areas, while for the adopted friction coefficient (f=0.2), the total wear volume remains almost unchanged. The comparison between total and resurfacing replacements showed higher scaled wear volumes (wear volume divided by wear factor) for the latter, in agreement with the literature. The effect of the boundary conditions (in vivo versus in vitro) was also investigated remarking their influence on implant wear and the need to apply more physiological-like conditions in hip simulators. In conclusion although friction is usually neglected in numerical wear predictions, as it does not affect markedly the contact pressure distribution, its effect in the location of the theoretical contact point was observed to influence wear maps. This achievement could be useful for increasing the correlation between numerical and experimental simulations, usually based on the total wear volume. In order to improve the model reliability, future studies will be devoted to implement the geometry update by combining the present model to finite element analyses. On the other hand, further experimental investigations are required to get out from the wide dispersion of wear factors reported in the literature.

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Figures

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Fig. 1

(a) Model geometry. (b) Coordinate reference frames used in model implementation: fixed global frame Cg = {Og,x,y,z} and cartesian frame moving with the cup Cc = {Oc,xc,yc,zc} in the reference configuration, i.e., with null motion angles. The anteversion α and inclination β angles are constant and orient the cup with respect to the pelvis. The cartesian frame moving with the head Ch is not indicated since in this configuration it corresponds to Cc. (c) Spherical coordinates for the cup.

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Fig. 2

Simulated loading (left) and kinematic (right) conditions for the head. Top: hip simulator curves [30]. Bottom: in vivo curves [31].

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Fig. 3

(a) Frictionless contact: nominal contact point K aligned to the load direction l. (b) Frictional contact: nominal contact point K shifted in the opposite direction to the sliding velocity (c), calculated according to Eq. (26). (d) In red the contact area on the head surface, delimited by the angle ψmax (d). Note: in (a,b) the external load applied to the cup is represented in green, while the contact forces acting on the head in red. In the frictional contact case, an additional moment Mf must be applied to the head to overcome frictional resistance.

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Fig. 4

THR results for in vitro boundary conditions (A-BCs) assuming frictionless (top) and frictional contact (bottom): (a) K trajectories on the cup and the head; (b) linear wear depth on the cup; and (c) linear wear depth on the head

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Fig. 5

THR results for in vitro boundary conditions (B-BCs) assuming frictionless (top) and frictional contact (bottom): (a) K trajectories on the cup and the head; (b) linear wear depth on the cup; and (c) linear wear depth on the head

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Fig. 6

THR results for in vivo boundary conditions (C-BCs) assuming frictionless (top) and frictional contact (bottom): (a) K trajectories on the cup and the head; (b) linear wear depth on the cup; and (c) linear wear depth on the head

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Fig. 7

RHR results for in vitro boundary conditions (A-BCs) assuming frictionless (top) and frictional contact (bottom): (a) K trajectories on the cup and the head; (b) linear wear depth on the cup; and (c) linear wear depth on the head

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Fig. 8

RHR results for in vitro boundary conditions (B-BCs) assuming frictionless (top) and frictional contact (bottom): (a) K trajectories on the cup and the head; (b) linear wear depth on the cup; and (c) linear wear depth on the head

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Fig. 9

RHR results for in vivo boundary conditions (C-BCs) assuming frictionless (top) and frictional contact (bottom): (a) K trajectories on the cup and the head; (b) linear wear depth on the cup; and (c) linear wear depth on the head

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