The Journal of Bone & Joint Surgery

The Journal of Bone & Joint Surgery, Volume 85, Issue 6

Scientific Article | June 01, 2003

A Biomechanical Analysis of Polyethylene Liner Cementation into a Fixed Metal Acetabular Shell

Geoffrey F. Haft, MD; Anneliese D. Heiner, PhD; Lawrence D. Dorr, MD; Thomas D. Brown, PhD; John J. Callaghan, MD

Investigation performed at the Departments of Orthopaedic Surgery and Biomedical Engineering, University of Iowa, Iowa City, Iowa

Geoffrey F. Haft, MD
Thomas D. Brown, PhD
Department of Biomedical Engineering, University of Iowa, 2181 Westlawn Building, Iowa City, IA 52242. E-mail address for G.F. Haft: geoff-haft@uiowa.edu. E-mail address for T.D. Brown: tom-brown@uiowa.edu

Anneliese D. Heiner, PhD
John J. Callaghan, MD
Department of Orthopaedic Surgery, University of Iowa, 200 Hawkins Drive, Iowa City, IA 52242. E-mail address for J.J. Callaghan: john-callaghan@uiowa.edu. E-mail address for A.D. Heiner: anneliese-heiner@uiowa.edu

Lawrence D. Dorr, MD
The Dorr Arthritis Institute, Centinela Hospital Medical Center, 501 East Hardy Street, Suite 300, Inglewood, CA 90301. E-mail address: centinela.appts@tenethealth.com.

In support of their research or preparation of this manuscript, one or more of the authors received grants or outside funding from DePuy, Warsaw, Indiana. In addition, one or more of the authors received payments or other benefits or a commitment or agreement to provide such benefits from commercial entities (DePuy and Howmedica, Rutherford, New Jersey). Also, a commercial entity (DePuy) paid or directed, or agreed to pay or direct, benefits to a research fund, foundation, educational institution, or other charitable or nonprofit organization with which the authors are affiliated or associated.

J Bone Joint Surg Am, 2003 Jun 01;85(6):1100-1110

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Abstract

Background: A common clinical scenario encountered by an orthopaedic surgeon is a patient with a secure cementless acetabular shell and a failed polyethylene liner. One treatment option is to cement a new liner into the fixed shell. The purpose of this study was to evaluate technical variables to improve the mechanical strength of such cemented liner constructs.

Methods: The contributions of shell texturing, liner texturing, and cement mantle thickness (between the liner and the shell) were evaluated by comparing torsional strength (among nine groups of constructs) and lever-out strength (among eight groups of constructs).

Results: Failure almost always occurred at the cement-liner interface. The two exceptions (failure at the shell-cement interface) occurred with a polished, untextured shell with no screw-holes. This finding indicates that if a shell has existing texturing (such as holes), further intraoperative scoring of the shell is unnecessary, but some sort of texturing is necessary to avoid construct failure at the shell-cement interface. Textured liners had significantly (a = 0.05) greater torsional and lever-out strength than untextured liners. The greatest construct strength occurred when liner grooves were oriented so as to oppose the applied loading. A 4-mm-thick cement mantle resulted in slightly greater torsional strength than a 2-mm-thick cement mantle, and a 2-mm-thick cement mantle resulted in considerably greater lever-out strength than a 4-mm-thick cement mantle, but these differences were not significant.

Conclusions: When cementing a liner into a well-fixed shell, a surgeon should ensure that both the shell and the liner are textured, as interdigitation of the cement with the shell and the liner is crucial to the mechanical strength of this construct.

Figures in this Article

Introduction

Surgeons who perform total joint arthroplasty are frequently faced with a patient who has a cementless acetabular shell that is securely fixed to the pelvis but the polyethylene liner has failed. Polyethylene wear and osteolysis associated with a secure cementless acetabular shell are becoming frequent clinical problems, as better designs of the component last into the second decade ¹ and less optimal designs fail earlier ^2-10 . In addition to wear, another complication is dislodgment of the modular liner. Numerous case reports over the last decade have established the ubiquity of this complication in association with a variety of acetabular components ^11-22 .

Several options are available to the surgeon in this situation. One option is to perform a complete acetabular revision, which may be the preferred treatment in patients who have poor fixation of the acetabular shell, malpositioning of the shell, or a very small shell in which a cemented liner would unduly compromise the polyethylene thickness. However, complete revision of the acetabular component is accompanied by serious potential complications, including pelvic discontinuity and severe bone loss. Typically, a 6 to 8-mm increase in the diameter of the replacement acetabular shell is required with this option, and cutting the original shell into pie-slice pieces (with the potential for particulate debris generation) is sometimes necessary. Also, although revision acetabular components demonstrate stable interfaces at five to ten years, they are more commonly associated with radiolucent lines than are primary components ^23,24 . Another option is simply to replace the damaged or worn liner with a new version of the original. However, there are several common situations in which this is not an option: the locking mechanism of the liner may be damaged, a replacement polyethylene liner may be unavailable or of questionable quality (as a result of gamma irradiation in air or a long or unknown shelf life), or a constrained liner may be needed to combat instability after liner removal and capsular débridement. In these latter situations, the surgeon's best option may be to leave the fixed shell in place and cement a replacement liner into the shell.

Heck and Murray ²⁵ , in 1986, were the first to describe, as far as we know, the cemented liner technique in a case report on the revision of a metal-on-metal prosthesis. While there have been only a few additional case reports on this technique in the orthopaedic literature ^{12,14,20,26-30} , discussions at national meetings have suggested that many surgeons throughout the United States have been cementing liners into fixed shells in recent years. In a retrospective review, seventeen patients with a cemented acetabular liner were followed for an average of 2.5 years (range, one to 4.7 years) ²⁷ . There were no radiographic changes in the bone-shell interface during the follow-up interval, but the first patient in whom this method was used had a revision because of failure of the cement-liner interface.

Given that surgeons are actively employing this technique, laboratory work is clearly necessary to help to determine the best construct that can be created. Previous studies of the strength of cemented liner constructs have investigated a few of the variables involved ^29-32 . The purpose of the present study was to expand upon these earlier studies and determine which surgeon-controlled variables would lead to the strongest mechanical construct when a polyethylene liner is cemented into a fixed acetabular shell. The specific objectives of the study were to determine the contributions of shell texturing, liner texturing, and cement mantle thickness to the overall mechanical strength of the construct.

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Fig. 1:: Acetabular shells (top) and liners (bottom) used for torsional and lever-out testing of cemented liner constructs. The scoring simulated intraoperative scratching that is done ostensibly to improve cement interdigitation; the channels are 2 mm wide and 1 mm deep. All implants were manufactured by DePuy Orthopaedics.

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Fig. 2-A:: Figs. 2-A and 2-B Torsional testing of the cemented liner specimens. Fig. 2-A Cutaway schematic for a smooth 4-mm liner with a cluster hole-unscored shell. Z indicates the symmetry axis of the liner.

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Fig. 2-B:: Representative recording, indicating yield and maximum torque values.

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Fig. 3-A:: Figs. 3-A and 3-B Lever-out testing of cemented liner specimens. Fig. 3-A Cutaway schematic for a circumferentially grooved and nubbed liner with a cluster hole-unscored shell. P = force.

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Fig. 3-B:: Representative recording, indicating yield and maximum moment values.

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Figs. 4-A and 4-B:: Effect of liner texturing in torsion testing (Fig. 4-A) and lever-out testing (Fig. 4-B). The cluster hole-unscored shell was common to all specimens in this evaluation. Asterisks indicate groups with significantly different average values (a = 0.05).

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Figs. 5-A and 5-B:: Effect of cement mantle thickness in torsion testing (Fig. 5-A) and lever-out testing (Fig. 5-B). The cluster hole-unscored shell was common to all specimens in this evaluation. No groups had significantly different average values (a = 0.05).

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TABLE I: Combinations of Shells and Liners Used in Torsion and Lever-out Testing*

*Groups were chosen on the basis of component availability. Five replicate shells and liners were tested in each combination. The components are designated by name and size. Cement mantle thicknesses for shell-liner combinations are also shown (in parentheses). All implants were manufactured by DePuy Orthopaedics (Warsaw, Indiana). T = torsion testing, and L = lever-out testing.

Shells	Liners
Smooth 2-mm (Summit 50)	Smooth 3-mm (Duraloc 48)	Smooth 4-mm (Summit 46)	Vertically Scored (Summit 46)	Circumferentially Grooved and Nubbed (Ultima 44)
Cluster hole-unscored (Summit 54)	T, L (2.05 mm)		T, L (4.05 mm)	T, L (4.05 mm)	T, L (3.24 mm)
Cluster hole-scored (Summit 54)			T (4.05 mm)		T (3.24 mm)
Polished-no hole-unscored (Duraloc 54)		T, L (3.15 mm)			T, L (3.20 mm)
Polished-no hole-scored (Duraloc 54)		T, L (3.15 mm)			L (3.20 mm)

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TABLE I: Combinations of Shells and Liners Used in Torsion and Lever-out Testing*

Shells	Liners
Smooth 2-mm (Summit 50)	Smooth 3-mm (Duraloc 48)	Smooth 4-mm (Summit 46)	Vertically Scored (Summit 46)	Circumferentially Grooved and Nubbed (Ultima 44)
Cluster hole-unscored (Summit 54)	T, L (2.05 mm)		T, L (4.05 mm)	T, L (4.05 mm)	T, L (3.24 mm)
Cluster hole-scored (Summit 54)			T (4.05 mm)		T (3.24 mm)
Polished-no hole-unscored (Duraloc 54)		T, L (3.15 mm)			T, L (3.20 mm)
Polished-no hole-scored (Duraloc 54)		T, L (3.15 mm)			L (3.20 mm)

Materials and Methods

The contributions of the shell-cement and cement-liner interfaces to the cemented liner construct were evaluated with use of torsional and lever-out tests. Torsional and lever-out tests were chosen on the basis of previously published studies of mechanical testing of modular acetabular components ^33,34 and cemented all-polyethylene liners ³² as well as proposed mechanisms for liner dislodgment in failed modular acetabular components ^{11,14,16,18-20} . Push-out testing was rejected as a clinically unrealistic failure mode, despite its previous use in the testing of liner-locking mechanisms of newly introduced modular components ^33-35 . For both the torsional and lever-out tests, yield strength and maximum strength were determined.

Experimental groups were designated ( Table I ) to determine the effects of (1) unscored compared with scored acetabular shells, with and without holes; (2) smooth compared with textured liners; and (3) cement mantle thickness. Four different shells and five different liners were used to evaluate these variables ( Fig. 1 ), with five specimens for each combination studied. The groups were chosen on the basis of component availability. All implants were provided by DePuy Orthopaedics (Warsaw, Indiana).

The shells included Summit cluster hole size-54 (unmodified) and highly polished Duraloc Enduron size-54 (central-hole-eliminated) components. The shells were scored (by electron discharge machining) to simulate the intraoperative scratching that is often done ostensibly to improve cement interdigitation ^25,29,31 . The scoring consisted of channels centered about the cup apex. The channels had a width of 2 mm and a depth of 1 mm. The shells were designated as "cluster hole-unscored," "cluster hole-scored," "polished-no hole-unscored," and "polished-no hole-scored."

The liners used in the present study included Summit size-50, Summit size-46, Duraloc Enduron size-48, and Ultima size-44 implants. Equatorial nubs and ridges were removed from all liners except the Ultima liners (which were unmodified). The Summit liners and the Duraloc liners were axisymmetric. These liners were designated as "smooth" and were further designated by their cement mantle thicknesses (approximately 2, 3, and 4 mm for the Summit size-50, Duraloc Enduron size-48, and Summit size-46 liners, respectively). The Ultima liners, which had three circumferential grooves and a circle of six spacer nubs with a radius of 2 mm, were designated as "circumferentially grooved and nubbed." Summit size-46 liners were vertically scored with a specially ground cutter on a vertical computer numeric controlled (CNC) milling machine. The scoring consisted of a cross pattern centered about the cup apex, with 2-mm-wide and 1-mm-deep scores. The scored Summit liners were designated as "vertically scored."

The acetabular shells were potted with dental acrylic into cylindrical containers (torsional specimens) or rectangular containers (lever-out specimens) to simulate well-fixed shell backing. Care was taken to keep the potting medium out of the screw-holes and central impaction hole in the shells with cluster holes. For the scored shells that were to be evaluated with lever-out testing, the scores were aligned with the sides of the acrylic pot. For the shells with cluster holes that were to be evaluated with lever-out testing, the screw-holes were aligned so that they would be on the tensile side of the specimen (and aligned with the top of the acrylic pot).

The liners were then cemented into the shells. All specimens were cleaned with soap and water before cementing. The cement (Surgical Simplex P provided by Howmedica, Rutherford, New Jersey) was mixed by hand, and the liner was cemented five minutes after the cement-mixing was started. For the purpose of testing reproducibility, the thickness of the cement mantle and the centering of the liner within each shell were precisely controlled, with use of specially designed hemispherical spacers and an axial-torsional materials testing machine (model 858; MTS Systems, Eden Prairie, Minnesota). The hemispherical spacer was first placed in the shell, and the liner was lowered into the spacer with use of the materials testing machine until a compressive load was detected. The axial position of the materials testing machine with the liner resting in the spacer was noted, the liner was raised, and the spacer was removed. The cement was then poured into the shell, and the liner was lowered into the cement to the previously noted level. Excess cement was removed with use of a straight, narrow osteotome, taking care to avoid pulling cement out of the interface. The liner was kept at this level for fifteen minutes, and then the specimen (the potting medium, shell, cement, and liner) was carefully removed from the fixturing of the materials testing machine. The specimens were placed in an incubator at 37°C, and the cement was cured for four to five days before testing.

Torsional Testing

Nine experimental groups of five specimens each, for a total of forty-five specimens, were tested in torsion ( Table I ). The liners were rigidly fixed to the actuator of the materials testing machine with use of twelve small, circumferentially spaced screws. The potted acetabular shells were fixed to the load-torque cell of the materials testing machine ( Fig. 2-A ). An x-y stage allowed free horizontal motion. An axial load of 70 kg (690 N) was applied to the liner. The liner was then rotated about its symmetry axis ( Fig. 2-A ) at a rate of 1°/sec until failure occurred. Torque and rotation angle data were collected at 0.002-second intervals by the materials testing machine software (MTS) and analyzed with use of Excel software (Microsoft, Seattle, Washington). The failure interface was recorded and photographed.

Yield torque and maximum (ultimate) torque were determined from each torque-angle recording ( Fig. 2-B ). Yield torque was defined as either the first abrupt change of slope for the torque-angle recording or (to account for nonabrupt slope changes) the intersection of the torque-angle curve with a 0.01° offset initial slope, whichever was the smaller value. Each torque measurement was averaged for each shell-liner combination. We used two methods to examine our data. First, after testing the assumption of homogeneity of variance among the nine experimental groups with use of the Brown and Forsythe method ³⁶ , one-way analysis of variance was used to determine whether there was a significant mean difference (a = 0.05) between any of the shell-liner combinations. The Tukey-Welsch multiple comparison procedure was then used to determine which shell-liner combinations were significantly different from one another. Second, when the difference between the mean values did not quite reach significance with the Tukey-Welsch test, we used the common Student t test. The first method was chosen because of variance heterogeneity and because it provides a conservative estimate of significance. The second method ignores variance heterogeneity and thus provides a lower limit to the estimate of p. Statistical power was calculated for a = 0.05 for nine experimental groups of five specimens each and for the realized effect sizes and a benchmark effect size of one standard deviation ³⁷ .

Lever-out Testing

Eight experimental groups, with five specimens in each for a total of forty specimens, were evaluated with lever-out testing. The liners were attached to a 6.35-mm-thick (0.25-in) gripping ring, with use of twelve small circumferentially spaced screws ( Fig. 3-A ). A lever arm (a 3/8"-24 bolt) was screwed into the center of the bearing surface, and a retaining nut was screwed to the level of the ring. A low-melting-point (70°C) bismuth alloy (Cerrobend; Cerro Metal Products, Bellefonte, Pennsylvania) was then poured into the liner. For the smooth 3-mm liners, the vertically scored liners, and the circumferentially grooved and nubbed liners, eight holes were drilled into the bearing surface for interdigitation of the bismuth alloy, to improve purchase. The specimen was then attached to the load cell of the materials testing machine. Lever-out torque was applied by means of a cylindrical platen eccentrically contacting the lever arm. The platen was lowered at a rate of 1.33 mm/sec (corresponding to a liner rotation of approximately 1°/sec) until failure occurred. Force ( Fig. 3-A ) and displacement data were collected in 0.01-sec intervals by the materials testing machine software (MTS), were converted to moment and angular data, and were analyzed with Excel software (Microsoft). The failure interface was recorded and photographed.

Yield moment and maximum (ultimate) moment were determined from each moment-angle recording ( Fig. 3-B ). Yield moment was defined as the intersection of the moment-angle curve with a 0.05° offset initial slope. Each moment measure was averaged for each shell-liner combination. Again, after testing the assumption of homogeneity of variance among the eight experimental groups with use of the Brown and Forsythe method ³⁶ , one-way analysis of variance was used to determine whether there was a significant mean difference (a = 0.05) between any of the shell-liner combinations. The Tukey-Welsch multiple comparison procedure was used first, and then the Student t test was used to determine which shell-liner combinations were significantly different from one another. Statistical power was calculated for a = 0.05 for eight experimental groups of five specimens each and for the realized effect sizes and a benchmark effect size of one standard deviation ³⁷ .

Results

Shell texturing, liner texturing, and the cement mantle thickness each affected both the torsional and the lever-out strength of the cemented liner constructs (see Appendix). As an a priori measure, for an effect size of one standard deviation, power was 0.997 for the torsional measures and 0.995 for the lever-out measures. For the actual one-way analyses of variance, power was 1.000 for all four measures. Visible failure occurred only at the cement-liner interface, with the exception of only two specimens. The two exceptions were both polished-no hole-unscored shells; one involved failure of the shell-cement interface with a circumferentially grooved and nubbed liner, and the other was a combined failure of the cement-liner and shell-cement interfaces with a smooth 3-mm liner. Because of differences in the design features among the implants, not all cemented liner constructs could be directly compared; therefore, only certain subsets are discussed below.

Effect of Shell Texturing

The groups were not compared with respect to the effect of shell texturing because failure did not occur at the shell-cement interface except in the two specimens discussed above. As long as the shell had some type of texturing, whether existing features (screw-holes) or intraoperative scoring, no failure occurred at the shell-cement interface.

Effect of Liner Texturing

The effect of liner texturing was studied with the cluster hole-unscored shells as a constant factor; the smooth 4-mm liners, vertically scored liners, and circumferentially grooved and nubbed liners were compared. In torsion, yield torque was significantly higher for the vertically scored liners (a = 0.05) ( Fig. 4-A ). Maximum torque was significantly higher (a = 0.05) for the two textured liners (the vertically scored liners and the circumferentially grooved and nubbed liners), and these two liners were not significantly different from one another ( Fig. 4-A ). In the lever-out test, the yield and maximum moments were significantly higher for the circumferentially grooved and nubbed liners (a = 0.05) ( Fig. 4-B ).

Effect of Cement Mantle Thickness

The effect of cement mantle thickness was studied with the cluster-hole-unscored shell as a constant factor; the smooth 4-mm and smooth 2-mm liners were compared. In torsion, yield and maximum torque were slightly higher for smooth 4-mm liners, but the difference was not significant ( Fig. 5-A ). In the lever-out test, the yield moment for the 4-mm cement mantle was a mean (and standard deviation) of 6.63 ± 2.36 N-m, whereas the 2-mm cement mantle resisted with a mean of 22.85 ± 6.41 N-m ( Fig. 5-B ). The difference between the means (16.2 N-m) did not meet the requirement of a mean difference of 17.5 N-m for significance with use of the Tukey-Welsch test. However, when the conventional Student two-tailed t test was used, the p value was 0.0031. For the maximum moment in the lever-out test, the 4-mm cement mantle resisted with a mean (and standard deviation) of 23.14 ± 6.32 N-m, whereas the 2-mm cement mantle resisted with a mean of 42.37 ± 4.40 N-m ( Fig. 5-B B). The difference between these means (19.2 N-m) also did not meet the requirement for a mean difference of 21.6 N-m for significance with use of the Tukey-Welsch test; however, when examined with the Student two-tailed t test, the p value was 0.0008 ( Fig. 5-B ).

Discussion

The increasing use of the cemented liner technique has generated a compelling need to compare the strength of this construct with that of modular acetabular components. Meldrum and Hollis ³¹ established that the strength of a cemented liner construct was similar to that of modular components in lever-out and push-out testing. The present study evaluated the mechanical strength of the shell-cement-liner construct in torsional and lever-out tests, as a function of individual factors (shell texturing, liner texturing, and cement mantle thickness) that characterize the construct. The results of the current study offer guidance to a surgeon who chooses to use a cemented liner technique.

Shell texturing, in any form, prevented failure at the shell-cement interface. As long as the shell had holes or was scored, visible failure occurred only at the cement-liner interface. This result suggests that the practice of intraoperative scoring of the acetabular shell, to improve the strength of the cement-shell interface, is unnecessary provided that the shell has screw-holes or other existing texturing. The data reported by Dunlop et al. ³⁸ and the surgical technique described by LaPorte et al. ²⁹ also support the concept that, as long as a shell has screw-holes, additional scoring is not needed. Avoiding scoring of the shell prevents the creation of metal particulate debris (a likely source of third-body wear) during the scoring process. However, if a shell lacks texturing features, a surgeon should score the shell to avoid failure at the shell-cement interface.

Regarding the effect of cement thickness, the present results suggest that untextured (smooth) liners with a 2-mm-thick cement mantle had greater construct strength in the lever-out test than did the constructs with the 4-mm-thick cement mantle. We say this because there may be differing opinions as to the type of statistical test that is appropriate for these data. The most conservative test, the Tukey-Welsch test, just missed significance (the test provided no p values), whereas the less restrictive Student two-tailed t test provided p values (yield moment, 0.00031; maximum moment, 0.0008) that were highly significant. Under torsional loading, the differences between the liners with a 4-mm-thick cement mantle and those with a 2-mm-thick mantle did not appear to be significant. Bensen et al. ³⁰ observed that the mean lever-out strength for cemented liners with a 4-mm cement mantle was 37 N-m, whereas the liner with a 2-mm cement mantle would not lever out before the polyethylene yielded about the hole into which the lever-out rod was inserted (moments as high as 68 N-m were recorded). In torsion, the 2-mm-thick and the 4-mm-thick cement mantles demonstrated no significant differences, although a trend toward better performance was noted for the thicker (4-mm) cement mantle. A possible explanation may be that interface stresses change minimally as the cement thickness changes. In a finite element study of cemented liners under axial loading, Kurtz et al. ³⁹ found a change of only 9% in cement compressive principal stress as the cement mantle thickness was changed between 0.5 and 2.0 mm. With no compelling data pointing toward a particular cement thickness, a surgeon is faced with a choice. One option is to choose a thicker cement mantle. For example, to ensure a complete cement mantle of at least 3 mm throughout (and to prevent the liner from bottoming out), a 4-mm-thick differential (in the radius) between the liner and the shell is probably a good starting point. By erring on the side of caution and cementing a slightly undersized liner, the surgeon can ensure adequate liner positioning, containment of the component within the shell, and at least minimal cement mantle thickness throughout the construct. (In the present laboratory trials, cement mantle thickness and centering were more meticulously controlled than is practical intraoperatively.) The other option is to choose a thinner cement mantle, which means a thicker liner will be used, resulting in the availability of more polyethylene for wear and for any necessary liner scoring ³² . However, a thin cement mantle could lead to bottoming out of the liner or (in an effort to avoid bottoming out) placement of a proud liner. It should be noted, though, that if a liner with spacer nubs was used, bottoming out and malpositioning of the liner would not be problems.

The most important variable for the surgeon to control is liner texturing. In almost every cemented liner combination tested, failure occurred at the cement-liner interface. In torsion, the cemented liner constructs with a vertically scored liner had the highest average yield torque and maximum torque; this difference was significant as compared with the other constructs, with one exception for each measure (both exceptions having a circumferentially grooved and nubbed liner) (see Appendix). In lever-out testing, constructs with a circumferentially grooved and nubbed liner had a significantly greater yield moment and maximum moment than any other constructs (see Appendix). Thus, when the groove orientation was directed so as to oppose the applied loading, there was a much greater resistance to failure. These results suggest that if a smooth liner is to be used, the surgeon should texture it with a series of orthogonal grooves prior to cementing it in place. This conclusion is consistent with the data reported by Oh ³² , who tested the effect of grooves on the torsional strength of a cemented all-polyethylene liner and found that the addition of 1-mm-deep grooves to a liner increased torsional strength from approximately 28 to 154 N-m. This conclusion is also supported by Dunlop et al. ³⁸ , who measured the torsional strength of cemented liner constructs as 2.4 to 14.6 N-m with a smooth liner, 15.6 to 44.4 N-m with a scored liner, and 40.4 to 78.4 N-m with an all-polyethylene liner.

The U.S. Food and Drug Administration requires testing of the strength of liner capture mechanisms for all commercially available modular acetabular components. The results of the present study of cemented liner constructs with textured liners compare favorably with those in previously published studies of modular acetabular components. For the circumferentially grooved and nubbed liner cemented into the cluster hole-unscored shell in the present study, the average maximum lever-out moment was 146 N-m. Tradonsky et al. ³⁴ determined the lever-out strength of eight contemporary modular components; the results ranged from a high of 77.3 N-m (Duraloc; DePuy, Warsaw, Indiana) to a low of 4.9 N-m (Triloc; DePuy), with a median of 37.5 N-m (Omnifit; Osteonics, Allendale, New Jersey). Bailey et al. ³³ determined that the lever-out strength of a Durasul Inter-Op liner (Sulzer Orthopedics, Austin, Texas) was 65.5 N-m.

Davidson et al. ⁴⁰ determined the frictional torque of the interface of a cobalt-chromium-alloy head articulating against an ultra-high molecular weight polyethylene shell with use of a nonrocking biaxial hip-joint simulator. With a 32-mm-diameter head, 5000-N maximum load, and water as a lubricant, the frictional torque measured an average of 0.94 N-m. This value is lower than even the lowest average maximum torque measured for the cemented liners (6.1 N-m for a smooth 3-mm liner cemented into a polished-no hole-unscored shell) and considerably lower than the average maximum torques measured for the cemented liners that had texturing (51.9 to 65.7 N-m) (see Appendix).

The current study has several limitations. First, the process used to cement the liners into the shells was an idealized representation of the actual intraoperative procedure. Cement thicknesses, liner centering, and creation of reproducible scores in shells and liners were carefully controlled to avoid adding confounding variables to the model, but they clearly represented a best-case scenario. Second, a larger number of specimens in each shell-liner combination could have provided a greater statistical power; for instance, the differences in outcome measures between the 2-mm and 4-mm cement mantle thicknesses might have reached significance with a larger sample size. Third, fatigue failure, which would have been a potentially informative mode of testing, was not considered in the current study; liner fatigue damage and subsequent failure could result from numerous instances in which the yield strength of the liner is exceeded. Fourth, it was not possible to obtain completely axisymmetric shells, to directly test the hypothesis that at least some shell texturing is necessary; the designs of the shells in this study included an outer liner-locking mechanism, which provided some stability in torque and lever-out testing. However, most shells in patients who need an acetabular revision include these liner-locking mechanisms, making the results in the present study clinically applicable. Finally, the shells and liners tested came from only one company; the quantitative values might be somewhat different with shells and liners from a different company. However, if similar variables of shell and liner texturing were studied, the qualitative results and subsequent conclusions obtained should be similar, regardless of which company's implants were considered.

Given the importance of liner texturing, the intraoperative time limitations for adequately performing the texturing, and the potential for weakening a liner by "over-texturing" it with a high-speed burr, the authors recommend that manufacturers consider producing polyethylene liners specifically designed for cementing into a shell. The features of these liners would resemble the backside designs of cemented all-polyethylene acetabular components. The liners should have vertical grooves to increase torsional strength at the cement-liner interface and circumferential grooves to increase the lever-out strength. The liners should also incorporate spacers to prevent bottoming out or malpositioning of the liner in the shell; these spacers can also contribute to the strength of the cement-liner interface. Trial liners, which the surgeon can use to ensure a proper fit in a variety of shell types, are also crucial. Along these lines, manufacturers should also consider making a constrained liner with these characteristics, as the cemented liner technique is gaining popularity in patients with recurrent dislocations ^26,27 . Given the increasing prevalence of patients who will require cemented acetabular liners in the coming years, the availability of a line of prefabricated textured liners would be useful in revision hip surgery.

Appendix

Results of cemented liner constructs tested in torsion and lever-out are available with the electronic versions of this article, on our web site at www.jbjs.org (go to the article citation and click on "Supplementary Material") and on our quarterly CD-ROM (call our subscription department, at 781-449-9780, to order the CD-ROM).

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Topics

cementation ; metals ; torque ; musculoskeletal torsion, function ; polyethylene ; torsion

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Geoffrey F. Haft, Anneliese D. Heiner, Lawrence D. Dorr, Thomas D. Brown, John J. Callaghan; A Biomechanical Analysis of Polyethylene Liner Cementation into a Fixed Metal Acetabular Shell. The Journal of Bone & Joint Surgery. 2003 Jun;85(6):1100-1110.

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