Creep dominates tensile fatigue damage of the cement–bone interface

ELSEVIER

Journal of Orthopaedic Research

Journal of Orthopaedic Research 22 (2004) 633-640

www.elsevier.com/locate/orthres

Creep dominates tensile fatigue damage of the cement-bone interface Do-Gyoon Kim *, Mark A. Miller, Kenneth A. Mann Depurtment of Orthopuedic S u r p r j . , Instirute ,for Hutiiun Perforniunce, S U N Y Upstcitc~Mcdiccil Uniwrsitj., 750 Eust A[/uiiis Slrerf. Sjwiciise. N Y 13210, USA Accepted 23 September 2003

Abstract Fatigue damage from activities of daily living has been considered to be a major cause of aseptic loosening in cemented total hip arthroplasty. The cement-bone interface is one region where loosening could occur, but to date the fatigue response of the interface has not been examined. Cement-bone specimens were prepared from fresh frozen human cadaver tissue using simulated in vivo conditions. Tensile fatigue tests to failure were performed in an environmental chamber. Loss of specimen stiffness (stiffness damage) and permanent displacement after unloading (creep damage) were found in all specimens. At failure, creep damage accounted for the majority (79.9 ? 10.6%) of the total strain damage accumulation at failure (apparent strain, c = 0.01 14 It 0.00488). A power law relationship between strain-damage rate and time-to-failure showed that the strain-damage rate was an excellent predictor of the fatigue life of the cement-bone interface. The S-N response of the interface was obtained as a function of the applied stress ratio and the initial apparent strain. The total motion between cement and bone (72.2 ? 29.8 pm) prior to incipient failure due to both stiffness and creep fatigue damage may be sufficient to result in fibrous tissue formation and contribute to eventual clinical loosening. 0 2004 Orthopaedic Research Society. Published by Elsevier Ltd. All rights reserved. Kqw~orc/.s:THA; Cement-bone interface: Tension: Fatigue; Creep

Introduction Clinical evidence has shown that aseptic loosening of total hip replacements becomes more frequent with time after implantation and that reduced levels of activity improves survivorship rates [2]. These observations suggest that curnulative fatigue damage from activities of daily living niay contribute to eventual clinical loosening. In the cemented hip implant system, aseptic loosening may be initiated by the mechanical failure of the stem-cement and cement-bone interfaces, as well as by fracture of the cement mantle [7]. To date, the fatigue damage accumulation of the cement surrounding implants has been the focus of both experimental and computational investigations [15,18]. Damage to the cement-bone interface has only been investigated using static loading [13]. But even with static loading, the postyield softening behavior of the interface [ 131 suggests that the damage process to the cement-bone interface is very different from bulk cement. *Corresponding author. Tel.: +313-916-8066; fax: +313-916-8064. E-muil uddress: [email protected] (D.-G. Kim).

The goal of this study was to determine the fatigue damage response of the cement-bone interface using carefully fabricated specimens from the proximal femur of cemented total hip replacements. Tensile fatigue loads were applied to the cement-bone interface, and the results were examined for two different failure mode characteristics: stiffness and creep damage. Stiffness dunzuge was defined here as the loss of structural stiffness of the stress-strain curve (tangent stiffness). Creep dumuge was defined as the permanent deformation after the specimen was unloaded. Using this approach we addressed three research questions: (1) during the fatigue damage process, does damage occur as stiffness damage, creep damage, or both?; (2) can the loss of mechanical integrity be quantified in terms of loading cycles and stress level?; and (3) can we predict the fatigue life of the cement-bone interface for a given applied stress level?

Methods Mechanical testing of the cement-bone interface was performed using carefully prepared cement-bone structures from proximal femur constructs. One major challenge in fatigue testing of these specimens

0736-02666 - see front matter 0 2004 Orthopaedic Research Society. Published by Elsevier Ltd. All rights reserved. doi: I0.1016/j.orthres.2003.09.007

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D.-G. Kim rt d.I Journal of Ortlioyuctlic Resrarch 22 ( 2 0 0 4 ) 6 3 3 4 4 0

was defining the appropriate applied cyclic load given that the strength of the specimens could vary widely. In this work, an initial nondestructive stiffness-strength relationship was developed to guide the magnitude of the applied fatigue loading. Thus, the testing procedure included specimen preparation, quasi-static tensile tests, tensile fatigue tests, substantial post-processing, and statistical analysis to address the three main research questions. Three pairs of fresh-frozen human femurs (63, 71, 85 years old) were broached for a Charnley stem followed by water lavage, brushing, and insertion of a distal plug. Femurs were obtained from the Anatomical Donor Program at Upstate Medical University. Venous pressure (25 mm Hg) [ I ] with a blood analog solution at 37 "C was applied to the periosteal bone surface through use of an external bladder surrounding the femur. This was done to simulate the thermal and hemodynamic conditions present during surgery [I]. The blood analog solution contained 0.08 wt'% polyacrylamide to simulate the non-Newtonian viscosity of human blood [I91 and 0.16 wt'%l CaCI? with 0.9 wt% NaCl [8] to maintain calcium levels in the bone tissue. Vacuum mixed cement (Simplex P. Stryker-Howmedica-Osteonics. Mawah, NJ) was introduced to the 37 "C femur in a retrograde fashion, followed by proximal pressurization and insertion of a mock polymethylmethacrylate (PMMA) stem. The stems were cast from PMMA to insure that sufficient cement would be available for attachment in the mechanical tests. Following cementing, the femur

constructs were placed in an incubator for 2 h at 37 "C, then stored at -21 "C until sectioning. A total of 41 specimens containing cement. the cement-bone interface, and bone were prepared using a previously described sectioning method [13]. Transverse slices at 10 mm intervals (Fig. la) were made from the level of the calcar to the distal tip of the stem using an abrasive cut-off wheel (Buehler, Lake Bluff, IL) with water irrigation. Specimens were sectioned further into 5 by 10 mm rectangular parallelepiped specimens (Fig. 1b) and were randomly assigned either to static (n = 16) or fatigue ( n = 25) test groups. Specimens were mechanically loaded in tension using a servo-hydraulic materials testing machine (Mini Bionix 11, MTS, Eden Prairie, MN). A circulating environmental chamber was used at 37 "C with calcium buffered solution (Fig. Ic). Grip to grip displacement was measured using a submersible linear variable differential transformer (LVDT, full scale range: & I mm, Sensotec, Columbus, OH), and the applied load was measured using an in-line load cell (400 N). The apparent applied stress (0)was computed as the applied load divided by the cross-sectional area of the specimen. Apparent strain ( i : ) was calculated by dividing the overall displacement by the initial grip-togrip distance (6.28 t 2.24 mm). Static tests to failure were performed under displacement control loading in an effort to develop a predictive relationship between specimen elastic stiffness and ultimate strength. Static test specimens

- lld

k Drain

Linear Sliders

Fig. 1. Transverse sections were made at 10 mm intervals of the cemented femoral constructs (a) and then sectioned further into parallelepiped sections (b). Cement-bone specimens were tested in an environmental chamber (c) with circulating buffered saline solution.

D.-G. Kim et ul. I Jounwl .f Ortkopcrc& Reseurdi 2-7 (2004) 633-640

635

T

(4

t'

Eyield

c- I

(b)

Apparent Strain (E)

Apparent Strain (e)

m

Time (sec)

Fig. 2. For static tests (a), plots of applied stress versus apparent strain were generated to determine the ultimate tensile strength of the interface. Fatigue tests (b) resulted in a cyclic response to applied stress and a change in apparent strain with time of cyclic loading (c).

( n = 16) were preconditioned with an applied tensile load of 0 -10N at 5 Hz using a sinusoidal waveform for 10 cycles. The last two cycles were used to assess the initial low-load stiffness ( k l O N ) by determining the slope of a least-squares fit through apparent stress versus apparent strain curves. The 10 N load level was chosen to provide an initial estimate of the specimen stiffness while applying a stress (0.2 MPa) that was non-destructive and always less than was applied for the cyclic tests. A single-cycle tensile test to failure was then conducted using a displacement-controlled rate of I m d m i n (Fig. 2a). The ultimate

tensile strength (uUll)was calculated as the maximum applied load divided by the cross-sectional area of the specimen. The yield strain ( C ? , ~ I was ~) calculated using the 0.2% offset method [ I I]. Finally, the ultimate strength versus low-load stiffness relationship was developed to serve as a predictor of specimen strength based on the specimen stiffness. Consistent with the static tests, low-load (10 N ) cyclic preconditioning was first performed to determine initial specimen stiffness (kloN) of the tensile fatigue specimens ( ) I = 25). The single-cycle

Table 1 Regression equations used to relate test parameters with input variables Test parameter

Regression equation

r-

I7

SEE

Figure

Denenden t

Indenendent

Ultimate strength, m1(MPa)

Low-load stiffness, kioN (MPa)

0.0032 k l O N+ 0.0202

0.87

16

0.285

Fig. 3a

Stiffness damage, k/ki

Fatigue life fraction, N / N f applied stress, (MPd)

(-0.538 + 0.1 4 8 ~ ~ ) ( N / N r )+ 0.943

0.53

17'

0.I 34

Fig. 4a

Creep damage,

Fatigue life fraction, N / N f applied stress fraction, u/cul,

0.0086

0.65

17"

0.0024

Fig. 4b

Total damage rate, dc,/dt ( S K I ) Creep damage rate, ds,/dt (s-I) Stiffness damage rate. dc,/dt (s -')

0.00595(d~,/dt)-I"IL4

0.92

17

I

0.00732(ds,/dt)-"

0.92

17

4

0.00202(d~,/dt)-""""~

0.86

17

il'

12.58 1(V/c7"Il 1- I 3

0.55

33

Ni! j 2 ?

Fig. 6a

1 . 7 4 10 ~ I*(C~)-~'~

0.91

33

N:' "I2

Fig. 6b

Time-to-failure,

cc

t r (s)

Cycles-to-failure, Nf

Applied stress frdction, a/aUl, Applied initial strain,

( ~ / ~ , . ) ( l 01171t060iln. l m,,~,11

9'52

216

Fig. 5

231'

7'x

CI

The correlation coefficient (r?)and standard estimate of the error (SEE) were used as measures of goodness of fit for the regression relationships. All correlations were significant at the p < 0.0001 level. *Regression relationship based on multiple data points from each test sample.

strength of each specimen was then predicted based on the relationship ~ )ultimate tensile strength ( m o l t ) between the low-load stiffness ( k l ~and determined above. The specimens were randomly assigned an applied stress level for fatigue testing that was between 50% and 80% (mean= 1.21 MPa; range: 0.2-2.95 MPa) of the estimated tensile strength. Each specimen was tested to interfacial fracture or 700,000 cycles under a sinusoidal loading pattern at 5 Hz with a stress ratio (lower stress/upper stress) of 0.1. Data were collected using a logperiod scheme to 10.000 cycles and every 10,000 cycles thereafter. Load and displacement data were collected for two Consecutive cycles at each sample interval. Several measures were used to quantify the structural response of the fatigue tests (Fig. 2b). A least-squares linear fit was made at each sampling interval using the data from the two consecutive loading cycles. From this, the slope was used to define the tangent stiffness ( X , ) of the construct for a given number of loading cycles, j . The stiffness damage was defined as ( k , / k l )and describes the degradation of stiffness from the initial loading cycle (j= I ) . The initial apparent strain (sI j was the strain that corresponded to the maximum applied stress measured at the first loading cycle. The creep damage ( c C ) was defined as the deformation at zero stress for each loading cycle. The deformation associated with the stiffness damage (8,) could be calculated by subtracting the deformation in the first cycle (t:,,j from the current cycle (r:\! j. The total damage ( r : , ) was defined as the sum of the creep damage and the deformation associated with the stiffness damage. During the three-phase damage process (Fig. 2c), the strain-damage rate (dcldf) was determined using a least-squares fit through the second linear phase. Because of the abrupt fracture of these load controlled tests. the apparent strain in the last few loading cycles was not considered in the analysis. In light of this observation, we defined the strain-to-failure (q)and corresponding time-to-failure (tl ) as failure criteria. These were calculated by taking the intersection of a line that passed through the second phase with a line that passed through the third phase [3]. All statistical analyses were performed using Statview (SAS Institute Inc., Cary, NC). A linear regression model was used to relate the static stiffness and strength of the cement-bone interface. Paired f-tests were used to compare magnitudes of stiffness and creep damage. We used both linear regression and power regression models to describe predictive relationships as shown in Table 1. The choice of regression model was determined using the best fit between the variables.

the contact interface between the cement and bone (Fig. lb). A strong positive correlation (r2 = 0 . 8 7 , ~< 0.0001) was found between the construct stiffness (kloN ) from the initial low-load tests and the ultimate tensile strength of the static tests (Fig. 3 ) . Given the strong correlation, this approach was used to estimate the single-cycle strength for the remainder of the fatigue specimens. The estimated single-cycle strength of the fatigue specimens ranged from 0.25 to 4.12 MPa (mean 1.83 MPa) and these were randomly assigned stress levels corresponding to 50-80'%1 of the estimated strength. The strain damage due to fatigue loading (Fig. 2c) exhibited a standard three-phase response: a rapid early rise, a constant rate region, and a rapid final rise to complete fracture. The stiffness damage (as measured by decreasing k / k l ) varied widely for these specimens (Fig. 4a) before incipient failure occurred. Note that the final phase of rapid decrease in stiffness was not included as this occurred in very few loading cycles. The creep

A o (MPa) 1.037

J

;

;

;

0.0

0.2

0.4

0.6

"' 0.0

(4

Results All of the static specimens ( n = 16) and 17 of the 25 fatigue specimens fractured. All fractures occurred at

:;A\

I

0.8

1.0

Fatigue Life Fraction, NIN,

0.016 W"

.G- 0.012

E tj al

?C 5 a a

5

0.008

0.004 0.0 y 0.0

(b)

0

I' 0

I

200 40 0 600 Stiffness, k, (MPa)

800

Fig. 3. The static tensile tests resulted in a strong positive correlation between the initial cyclic stiffness ( X I O N ) and ultimate tensile strength (mu,,). The linear regression equation was used to predict the strength of the fatigue specimens.

0.7 0.7 0.5

I

0.2

0.4

0.6

0.8

1.0

Fatigue Life Fraction, NIN,

Fig. 4. The stiffness damage ( k l k l ) with cyclic loading (a) indicated reduced stiffness as fatigue damage accumulated. A modest but significant correlation was found between the applied cyclic stress magnitude (Am) and stiffness damage parameters. Creep damage strain (c') increased with fatigue loading cycles (b). Specimens with higher applied stress fractions (m/ot,l,) tended to have greater creep damage magnitude. Note that the third (final) phase of the damage response is not shown.

D.-G. Kim el ul. I Joitrnul of Ortliopuedic~Rrsrcircli 22 (2004) 6 3 3 4 4 0

O

N

0

creep

Fractured Not fractured (estimated) I

10.’

I I

i

I I

I

I

lo-*

IO-~ 10.~ Strain damage rate, dddt (sec’)

10.~

Fig. 5. A strong negative correlation was found between time-to-failure and strain-damage rate. Raw data for total damage rate of the frdCtured specimens are shown together with a linear regression. The specimens that were not fractured are also included with estimates of time-to-fracture based on the strain-damage rate of the specimens. The contributions of stiffness damage and creep damage to the total damage are shown as straight regression lines. The number of cyclesto-failure was five times the time-to-failure given a cyclic test frequency of 5 Hz.

+ Static-fractured

2 .

1.3

0 Fatigue-fractured

D

.-

0

Fatigue-not fractured

0.5 -

6

0.4

oOOo

!

1 (a)

1o2 1o4 Cycles-to-Failure, Nf

1o6

+ Static-fractured

0 Fatigue-fractured

0

Fatigue-not fractured

C

6

3 0.003._ c ._

-C

0.002 1

1 (b)

10’

1o4 Cycles-to-Failure, Nf

631

damage (Fig. 4b) followed a power-law response with early increases in creep damage followed by a near linear response. The strain associated with stiffness damage ( E ~ , 0.00245 f 0.00176) was significantly smaller (p < 0.0001) than the strain associated with creep damage (cC, 0.00898 k 0.00371) as tested using a paired t-test. Creep damage accounted for 79.9k 10.6‘%1 of the total damage accumulation (Et, 0.01 14 & 0.00488) at failure. Displacement at failure due to stiffness damage (16.34 & 12.7 pm), creep damage (56.4f20.2 pm), and total damage (72.2f 29.8 pm) were found. In terms of predictive relationships for stiffness and creep damage, a two-step approach was used to include the effects of applied stress level. For stiffness damage, a linear regression was fit to each of the stiffness damage responses, and the slope of this relationship was found to correlate with the applied stress level (Ao,r‘ = 0.3 19, p = 0.0181). From this, the overall stiffness damage was defined as a function of the applied stress level and fatigue life fraction (Table 1). Overall, this fit was moderate ( r 2 = 0.53). A similar approach was used for the creep damage. The creep damage behavior was fit to a power law, and the exponent was found to correlate with the applied stress ratio (o/oult, r2 = 0.247, p = 0.0424). Creep damage was then defined as a function of the applied stress ratio and fatigue life fraction (Table 1) with a moderate overall fit (r’ = 0.65). A strong inverse relationship was found between the strain-damage rate (dcldt) and the time-to-failure (Fig. 5). Using power-law relationships, between 86% and 92% of the variability in time-to-failure could be determined based on the stiffness, creep, or total damage rate (Table 1). This strong correlation was used to estimate the number of cycles-to-failure for the eight non-fractured specimens. The non-fractured specimens would be expected to fail at a total number of loading cycles greater than 700,000 cycles. Using this approach, the number of cycles-to-failure for the non-fractured specimens ranged up to 3,884,800 cycles. An S-N response curve was developed by plotting the number of cycles-to-failure ( N f ) as a function of the applied stress ratio (o/oult)(Fig. 6a). Static data were included by plotting the ratio of the actual applied stress to the single-cycle predicted strength. A moderate inverse relationship ( r 2= 0.55) was found consistent with typical S-N response. An improved S-N response was found (Fig. 6b) when the initial apparent strain ( ~ 1 ) was used in place of the stress fraction (r2 = 0.91). In this case, the yield strain was used as the initial apparent strain ( E , ) for the case of single-cycle loading.

106

Fig. 6 . The “S-N” response of the test specimens (a) indicated increased number of cycles-to-failure with decreased applied stress ratio. An improved regression relationship was found when the initial apparent strain ( i : , ) versus cycles-to-failure was used (b).

Discussion During the tensile fatigue damage process of the cement-bone interface, there was a loss of structural

D.-G.

638

Kiiii C ~ Iul.

I Joirrriul

of’ Orrliopudic

(cyclic) stiffness and increased permanent deformation upon unloading. Creep damage dominated the failure response, accounting for nearly 80Y0 of the total strain accumulated at failure. Overall, the test specimens exhibited a three-phase damage process with initial rapid changes in stiffness and creep damagc, followed by a steady damage rate regime, and ending in a rapid increase in damage to complete failure of the specimen. Due to the wide variation in cement-bone interface morphologies used, a wide range in structural responses was found. Predictive models that were developed to describe the first two phases of the damage response were modestly successful in capturing this damage response. Further, predictions of fatigue life were also modestly successful, and these exhibited standard S-N type responses. Interestingly, if the initial apparent strain was used as the S-N model input, the errors associated with the S-N prediction decreased substantially. The most important limitation of the present study was that the work was performed in situ with no possibility for biological processes to contribute to the damage process. Specimens retrieved at autopsy that were well functioning at the time of the patient’s death have shown intact cement-bone interfaces with little intervening fibrous tissue [lo]. Osteolysis and fibrous tissue formation at the cement-bone interface, secondary to debris production [ 121, would certainly change the mechanics of the interface. Therefore, long-term biological effects on the cement-bone interface are unclear, and the work reported here is most relevant to the early post-operative condition. The cement-bone composite structures tested here consisted of four unique zones (Fig. lb) including the

Rrscw’ch 22 (2004) 6 3 3 4 4 0

bulk PMMA cement, an interdigitated cement-bone region, a so-called contact interface at the extent of cement penetration into the bone, and cortical bone. The fatigue response of the structure depended on the sum of damage from each of these zones. Creep damage for cement [I71 and cortical bone [5] from cyclic stress application on the order of 3 MPa were estimated to be orders of magnitude smaller than the creep damage found here. This suggests that the bulk cement and cortical bone do not contribute appreciably to the damage response and remain linearly elastic during the test. Damage must then localize at the contact interface or interdigitated region (see Fig. Ib); the LVDT measurements would reflect the damage of these two components. We believe that the creep damage found in this study is not creep in the traditional sense but rather permanent deformation due to sliding between cement and bone at the contact interface and within the interdigitated region. Stiffness damage is possibly due to loss of stiffness from microcracks and loss of connectivity between the cement and bone. In a preliminary study conducted in our lab, tensile fatigue tests of cementbone constructs that contained only the interdigitated region indicated that substantial damage could occur to this layer. Thus damage to the interdigitated region, in addition to the contact interface, may contribute to the overall damage process of the cement-bone interface. Fluorescence microscopy of fatigued cement-bone specimens (Fig. 7) supports this observation. Fatigue cracks emanating from trabecular bone were often seen indicating that irreversible damage had occurred in the interdigitated region. Further work is needed to assess and delineate the role of the contact interface and interdigitated region in the damage process.

CEMENT

INTERDIGITATED ZONE

CONTACT IINTERFACE

CORTICAL BONE

Fig. 7. Fatigue fractured test specimen (a) indicating failure at the extent of cement penetration into the bone. The inset figure (b) illustrates microcracks (arrows) emanating from the trabecular bone (TB).

The fatigue response of the cement-bone interface with three distinct damage phases had a form that was similar to the fatigue response of trabecular bone in compression and PMMA cement loaded in tension [3,17]. However, the relative contribution of creep damage to the overall damage with the cement-bone interface (-80‘%1) is substantially higher than that of bulk trabecular bone (26%) [3] or PMMA cement (50‘%1) [17]. This suggests that the components of the cementbone interface may contribute differently to the fatigue damage in the composite structure. From a clinical perspective, loss of fixation at the cement-bone interface due to mechanical loading is poorly understood. To date, damage models of the cement-bone interface have usually been limited to static loading cases, and these have been applied to simplistic structural models [ 141. Efforts to understand the damage process due to fatigue loading has improved our understanding of bulk cement damage [18], but these models have by necessity not included a damage model for the cement-bone interface. The damage models developed here would be an important addition to understanding cement-bone damage in the context of the entire cemented femoral construct. As a first estimate of cement-bone failure for the cemented stem construct, one can compare the fatigue results found here with maximum tensile stresses across the cement-bone interface as determined using the finite element method [4]. For a Charnley type stem, a maximum tensile stress for normal gait of 1.1 MPa was found, and this was 61‘%1of the average single-cycle strength found for the cement-bone interface. From this, approximately 10,000 cycles would be sustained before failure in the peak stress region. This suggests that local failure of the cement-bone interface is possible. However, loss of local cement-bone interface stiffness due to fatigue loading was not considered here and may have an important effect on changing the local stress distribution. Further efforts are needed to understand the interaction between load transfer and cement damage in these complicated structures, including mixed-mode loading conditions. In this study, the total possible inducible displacement (72.2k 29.8 pm) due to both stiffness and creep damage was similar in magnitude to micromotions that can result in fibrous tissue formation between implants and bone (50-150 pm) [16]. Well-fixed components often do not exhibit fibrous tissue formation at the cementbone interface [9]. In contrast, a synovial-like membrane can form at the cement-bone interface in loose components [6]. However, the actual motion between the cement and bone due to fatigue damage in both of these cases is unknown. Based on the present results, one could anticipate that the permanent damage developed due to fatigue could result in interfacial gaps and sub-

stantial movement during loading to induce fibrous tissue formation. Acknowledgements This work was supported by National Institutes of Health Grant AR42017. We would like to thank Stryker-Howmedica-Osteonics for donating the cement used in this study. References Benjamin JB, Gie GA, Lee AJ, Ling RSM, Volz RG. Cementing technique and the effects of bleeding. J Bone Joint Surg 1987; 69B:6204. Berry DJ, Harnisen WS. Cabanela ME. Morrey BF. Twentyfive-year survivorship of two thousand consecutive primary Charnley total hip replacements. J Bone Joint Surg 2002;84A: 171-7. Bowman SM, Guo XE. Cheng DW, Keaveny TM, Gibson LJ, Hayes WC, et al. Creep contributes to the fatigue behavior of bovine trabecular bone. J Biomech Eng 1998;120:647-54. Chang PB, Mann KA, Bartel DL. The effects of proximal bonding, stem geometry. and neck length on cemented femoral stem performance. Clin Orthop Re1 Res 1998:355:57-69. Cotton JR, Zioupos P. Winwood K, Taylor M. Analysis of creep strain during tensile fatigue of cortical bone. J Biomech 2003;36: 943-9. Goodman SB, Ma PH. Song Y , Lee K , Doshi A, Rushdieh B. et al. Loosening and osteolysis of cemented joint arthroplasties: a biologic spectrum. Clin Orthop Re1 Res 1997;337:149-63. [7] Gruen TA. McNeice G M , Amstutz HC. Modes of failure of cemented stem-type femoral components: a radiographic analysis of loosening. Clin Orthop 1979:141:17-27. [8] Gustafson MB, Martin RB, Gibson V. Storms DH, Stover SM, Gibeling 3, et al. Calcium buffering is required to maintain bone stiffness in saline solution. J Biomech 1996;29:11914. [9] Jasty M , Maloney WJ, Bragdon C R , Haire T, Harris WH. Histomorphological studies of the long-term skeletal responses to well-fixed cemented femoral components. J Bone Joint Surg 1990: 72A: 1220-9. [lo] Jasty M, Maloney WJ. Bragdon CR, O’Conner DO, Haire T, Harris WH, et al. The initiation of failure in cemented femoral components of hip arthroplasties. J Bone Joint Surg 3991;73B: 551-8. Keaveny TM. Wachtel EF. Ford CM, Hayes WC. Differences between the tensile and compressive strengths of bovine tibia1 trabecular bone depend on modulus. J Biomech 1994;27:113746. Lu JX. Huang ZW. Tropiano P, Clouet D’Orval B. Remusat M, Dejou J. et al. Human biological reactions at the interface between bone tissue and polymethylmethacrylate cement. J Mater Sci: Mater Med 2002;13:803-9. Mann KA, Ayers DC, Werner FW, Nicoletta RJ, Fortino MD. Tensile strength of the cement-bone interface depends on the amount of bone interdigitated with PMMA cement. J Biomech 1997;30:339 4 6 . Mann KA, Damron LA. Predicting the failure response of cement-bone constructs using a non-linear fracture mechanics approach. J Biomech Eng 2002; 124:462-70. McCormack BAO, Prendergast PJ. Microdainage accumulation in the cement layer of hip replacements under flexural loading. J Biomech 1999;32:467-75.

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