Self-tapping ability of carbon fibre reinforced polyetheretherketone suture anchors.

XML Template (2014) [8.5.2014–3:15pm] //blrnas3/cenpro/ApplicationFiles/Journals/SAGE/3B2/JBAJ/Vol00000/140025/APPFile/SG-JBAJ140025.3d

(JBA)

[1–12] [PREPRINTER stage]

Article

Self-tapping ability of carbon fibre reinforced polyetheretherketone suture anchors

Journal of Biomaterials Applications 0(0) 1–12 ! The Author(s) 2014 Reprints and permissions: sagepub.co.uk/journalsPermissions.nav DOI: 10.1177/0885328214535274 jba.sagepub.com

´ ’Bra´daigh1 and Emer M Feerick1, Joanne Wilson2, Marcus Jarman-Smith2, Conchur M O J Patrick McGarry1

Abstract An experimental and computational investigation of the self-tapping ability of carbon fibre reinforced polyetheretherketone (CFR-PEEK) has been conducted. Six CFR-PEEK suture anchor designs were investigated using PEEK-OPTIMAÕ Reinforced, a medical grade of CFR-PEEK. Experimental tests were conducted to investigate the maximum axial force and torque required for self-taping insertion of each anchor design. Additional experimental tests were conducted for some anchor designs using pilot holes. Computational simulations were conducted to determine the maximum stress in each anchor design at various stages of insertion. Simulations also were performed to investigate the effect of wall thickness in the anchor head. The maximum axial force required to insert a self-tapping CFR-PEEK suture anchor did not exceed 150 N for any anchor design. The maximum torque required to insert a self-tapping CFR-PEEK suture anchor did not exceed 0.8 Nm. Computational simulations reveal significant stress concentrations in the region of the anchor tip, demonstrating that a re-design of the tip geometry should be performed to avoid fracture during self-tapping, as observed in the experimental component of this study. This study demonstrates the ability of PEEK-OPTIMA Reinforced suture anchors to self-tap polyurethane foam bone analogue. This provides motivation to further investigate the self-tapping ability of CFR-PEEK suture anchors in animal/cadaveric bone. An optimised design for CFR-PEEK suture anchors offers the advantages of radiolucency, and mechanical properties similar to bone with the ability to self-tap. This may have positive implications for reducing surgery times and the associated costs with the procedure. Keywords Suture anchors, polyetheretherketone, carbon fibre reinforced polyetheretherketone, rotator cuff, computational simulation, experimental insertion testing, self-tapping

Introduction The number of rotator cuff repairs is growing annually due to an aging population.1 In the USA, per patient costs typically associated with a rotator cuff repair is $17,427.2 Currently, suture anchors used to address rotator cuff repair are made from a variety of materials including titanium, polyetheretherketone (PEEK) or resorbable polymer/biocomposites, with each material having its own benefits such as strength – that allows self-tapping (titanium metal), radiolucency (polymer) and biointegration (resorbable polymer). The development of next generation suture anchor designs that offer a combination of the aforementioned benefits is of great interest. PEEK-OPTIMA Natural is one of the medical grades of PEEK and is commonly used in

sports medicine. It is a radiolucent, non-degradable, biomaterial with a Young’s modulus close to that of bone and offers excellent wear and fatigue properties.3 Medical grade variants of this biomaterial, containing carbon fibre reinforcement (CFR) are also available. Improved mechanical performance can be achieved via fibre reinforcement of PEEK, with the mechanical

1

Department of Mechanical and Biomedical Engineering, National University of Ireland Galway, Ireland 2 Invibio Biomaterials Solutions, Technology Centre-Hillhouse International, Thornton Cleveleys, Lancashire, United Kingdom Corresponding author: Emer M Feerick, Department of Mechanical and Biomedical Engineering, National University of Ireland Galway, Ireland. Email: [email protected]

Downloaded from jba.sagepub.com at Bobst Library, New York University on May 26, 2015


(JBA)


2

Journal of Biomaterials Applications 0(0)

properties of the composite material depending on the volume, length and alignment of these fibres. Short fibre reinforced PEEK consists of randomly aligned short fibres, producing an isotropic homogenous material. To date short CFR-PEEK has been considered for spinal applications such as interbody fusion4,5 and spinal arthroplasty.6 Short CFR-PEEK has also been considered for arthroscopic applications for the upper and lower extremity.7,8 A long carbon fibre reinforced (CFR) grade of PEEK is also available (PEEKOPTIMA Ultra Reinforced) and can be used to significantly increase stiffness and strength in directions of fibre alignment. Controlled alignment of the fibres can provide a broad range of anisotropic properties that can be tailored for a specific application. Examples include intramedullary rods and fracture plates.9,10 Volume fraction of fibres also determines mechanical properties of CFR-PEEK. Two standard % volume fractions are commercially available, i.e. 30% and 60%. A summary of the stiffness of natural, short CFR-PEEK and long CFR-PEEK are summarised in Table 1. The inclusion of carbon fibre within the polymer matrix significantly increases the strength of an unfilled PEEK polymer as illustrated in Table 1. The current study investigates the application of short CFR-PEEK for self-tapping suture anchors. For the remainder of the current study the term CFR-PEEK refers to short fibre reinforcement, unless otherwise stated. Hole preparation for insertion of suture anchors consists of a number of steps. Pilot holes involve drilling a cylindrical core, typically with a lower diameter than that of the anchor to be inserted. Tapping of holes creates an impression of the thread shape of the anchor to be inserted. Self-tapping refers to the ability of an anchor to be inserted without the need for tapping of the hole. Self-drilling refers to the ability of an anchor to be inserted without the need for drilling of a pilot hole. Both self-tapping and self-drilling allows for faster insertion with improved anchorage.11,12

Table 1. Mechanical properties of the natural and carbon fibre reinforced grades of PEEK-OPTIMA.3

Material PEEK OPTIMA natural PEEK-OPTIMA reinforced (short fibre) PEEK-OPTIMA ultra reinforced (long fibre)

Flexural modulus (GPa)

Tensile strength (MPa)

4 20

93 170

135

>2000

The objective of the current study is to investigate both the self-tapping and self-drilling ability of suture anchor designs manufactured using CFR-PEEK. Use of the term self-tapping for the remainder of the current article also refers to self-drilling. The development of a self-tapping CFR-PEEK suture anchor has positive implications in terms of reducing surgical procedure times, in addition to the aforementioned benefits associated with CFR-PEEK. Experimental insertion tests were conducted using SawboneTM polyurethane foam (bone analogue) (Sawbones, Malmo¨, Sweden13). Tests were conducted to investigate three thread designs with three driver shapes: square, hexagonal and hexalobe. Computational simulations were conducted to investigate the stresses exerted on the anchor and the surrounding foam when subject to a torque at a range of insertion depths. Additionally, simulations were conducted to investigate the stresses exerted on the tip of the anchor during insertion.

Methods and materials Experimental All anchors were fabricated from CFR-PEEK (PEEKOPTIMA Reinforced, Invibio Biomaterials Solutions Ltd., Lancashire, UK). The self-tapping anchors that were investigated experimentally are summarised in Figure 1. An initial thread design (A1H1) was based on representative generic geometries of existing metallic equivalents. Three driver geometry shapes were investigated for anchor A1H1: a square, hexagonal and a hexalobe driver. Two alternate pitch distances were also investigated. A square headed anchor with a pitch distance of 2.1 mm (referred to as A2H1); a square headed anchor with a pitch distance of 1.45 mm (referred to as A3H1). The original thread spacing was also used with an altered taper angle that resulted in a 4 mm non-tapered region at the top of the anchor (referred to as A1H2). Two tip thickness magnitudes were investigated (Tw ¼ 0.60 mm for A1H2 anchor designs and Tw ¼ 0.35 mm for all other anchor designs). A synthetic analogue of the cortical shell and cancellous bone under-layer was fabricated using bilayers of polyurethane of alternate densities (Sawbones, Malmo¨, Sweden13). The 2 mm top layer (40 pcf) represented the cortical bone shell, while the remainder of the block (15 pcf) represented the cancellous bone. This bi-layer composite structure represents a synthetic test platform for the assessment of the ability of a suture anchor design to undergo self-tapping insertion.14 Experimental testing was conducted in accordance with the ASTM standard F543 Annex 4.15 The axial displacement, torsion and axial load were all recorded




(JBA)

Feerick et al.

3

simultaneously. A Zwick (Zwicki-Line 5 kN, 20 Nm load cell) biaxial tester was used to conduct all testing. The experimental test setup is shown in Figure 2. The driver was attached to the upper moving cross head. The composite bone analogue structure was secured in an aluminium casing and the tip was placed in contact with the cortical bone surface. A continuous rotation and axial displacement was applied to each anchor. For each thread type (A1H1, A2H1 and A3H1) the axial loading rate was applied as per the pitch distance/min and a continuous torque loading of 5 revolutions/min.

for the cortical and cancellous bone. Model parameters are outlined in Table 2. Investigation of driver head geometry at three insertion depths. A schematic of the computational model developed for anchor torque simulations is presented in Figure 3. The driver was modelled as a rigid body as it is four orders of magnitude stiffer than the CFRPEEK anchor material. The bilayered Sawbone block was modelled with a 2 mm skin of 40 pcf foam. The remaining block volume was assigned the properties of 15 pcf foam. Simulating the self-tapping of suture anchors is not computationally feasible due to the

Computational Based on the experimental test results, several computational models were developed to gain further insight into the design features of the anchors. Simulations were conducted for three anchor head heights (r ¼ 0.5, 2.0 and 4.0 mm as shown in Figure 3) to investigate the stress distribution in the driver head of the anchors. The wall thickness was also investigated for square, hexagonal and hexalobe driver head geometries. Finally, a study of the stress levels in the anchor tip, responsible for initiating self-tapping of the anchor, during insertion was conducted. A parameter study of anchor tip diameter was conducted to evaluate tip strain levels for a given applied force. In all simulations the CFR-PEEK material was modelled as isotropic elastic and elastic material behaviour was also assumed

(a)

Figure 2. Experimental test set up for insertion testing.

(b)

Anchor Thread Geometry

q

W h Wi p

L

A1H1

A1H2

A2H1

A3H1

Wi (mm)

4.10

4.10

4.10

4.10

Wo (mm)

5.50

5.50

5.50

5.50

h (mm)

2.81

4.00

2.81

2.81

p (mm)

2.81

2.81

2.10

1.45

Th (mm)

2.00

2.00

2.00

2.00

Tw (mm)

0.35

0.60

0.35

0.35

L (mm)

15.2

15.2

15.2

15.2

8.6°

8.2°

8.6°

8.6°

Q

Li = 1.68 mm Lii= 2.69 mm

Hi= 1.50 mm Hii= 2.00 mm

Si= 1.8 mm Sii= 2.5 mm

Th Tw A1H1

A1H2

A2H1

A3H1

Figure 1. (a) Anchor geometry parameters investigated. Parameters held constant; internal head width (Wi), outer width (Wo), tip length (Th) and length (L). Varying parameters; head height (h), pitch distance (p), tip thickness (Tw) and taper angle (). (b) Driver shape geometries investigated for anchor A1H1; hexalobe, hexagonal and square driver geometries selected based upon ASTM F116 14.



(JBA)


4


extensive deformation and material damage during the insertion process. Therefore, for the purposes of this study, torsion of anchors were simulated at three different insertion depths (r ¼ 4.0, 2.0 and 0.5 mm, see Figure 3). The block was held fixed in position in the vertical direction at the base and in the horizontal direction at the sides. Hard contact was applied between the rigid body driver and the CFR-PEEK anchor. Full hard contact was implemented with relative sliding between the anchor and foam with a co-efficient of friction of 0.3.16 Investigation of wall thickness within anchor head. The ASTM standard F11617 for biomedical screws was used to select driver geometries for this study. Simulations were conducted to investigate the effect of altering the driver sizes and hence the wall thickness of anchors within this region. The sizes were altered

Compression and torque of anchors. Simulations were conducted to investigate whether axial loads were high enough to induce stress levels in the region of plastic deformation of the CFR-PEEK anchor. For all simulations the anchor was held fixed in the vertical position as shown in Figure 4(a). An axial load was applied to the driver and ramped up linearly to 300 N. Simulations were conducted for the three driver head geometries (square, hexagonal and hexalobe). A parameter study (Figure 4b) was conducted for multiple tip diameters under 150 N and 300 N loads in order to identify the required diameter such that strains in the tip do not exceed the elastic limit.

Results Anchor tip results

Rigid Body Driver

r

within the size ranges reported for each driver type included in the ASTM standard F116.17 Two driver sizes were investigated for each driver shape as shown in Figure 1(b).

CFPEEK Anchor 40 PCF 2mm Skin

15 PCF

Figure 3. Computational set up; simulations were conducted at different insertion depths (r ¼ 0.5, 2.0 and 4.0 mm).

The tip design with Tw ¼ 0.3 mm (n ¼ 6) failed during experimental insertion (in the region of 100 N), the tip underwent plastic deformation and failed as shown in Figure 5(b,c). However, tips with Tw ¼ 0.6 mm were capable of withstanding self-tapping insertion forces. Pilot holes were also investigated to evaluate the effect on axial force and torque during insertion. VonMises stress distribution during compressive axial loading of anchors is shown in Figure 6 for a square, hexagonal and hexalobe driver geometry. Two applied loads are considered, 150 N (Figure 6a–c) and 300 N (Figure 6d–f); a 150 N axial load was required to insert an anchor experimentally. For a 150 N axial load a stress concentration was predicted in the region of the anchor tip. When the applied load was increased to 300 N the maximum computed stress exceeded the ultimate compressive strength (309.7 MPa) approximately by a factor of 3. A parameter study was conducted to determine the tip diameter required to ensure that strains in the tip remain below the elastic limit.

Table 2. Material properties.

PEEK OPTIMAÕ reinforced 40 pcf foam 15 pcf foam

Elastic modulus

Compressive modulus

Ultimate compressive strength

Ultimate tensile strength

Ultimate shear strength

18 GPa 1 GPa 173 MPa

8.25 GPa 759 MPa 123 MPa

309.7 MPa 31 MPa 4.9 MPa

193.02 MPa 19 MPa 3.7 MPa

105 MPa 11 MPa 2.8 MPa



(JBA)


Feerick et al. (a) 300 N

5 (b)

Rigid Body Driver

The results of the parameter study are summarised in Figure 5(d). For a load of 150 N, a Tw ¼ 0.6 mm tip is required to prevent plastic strain within the tip. For a load of 300 N, a Tw ¼ 1.2 mm tip is required to prevent plastic strain within the tip. Following insertion testing, anchors were removed from the test blocks to examine the tip and anchor threads. Figure 7(a, b) shows some signs of compressive failure at the anchor tip. The re-design of the anchor tip to eliminate such failure is supported by the computational simulations of varying Tw. However, the key feature to highlight in the experimental images of Figure 7 is that the threads of the anchor remain undamaged post insertion. This supports the proposed use of this material for self-tapping anchors, as the threads are capable of withstanding insertion loads.

Load (150 N / 300 N)

Square

Tw

Hexagonal

Tw (mm) 0.6 Compression Modulus 8.25 GPa

0.8 1.0 1.2

Hexalobe

Anchor insertion Figure 4. (a) Loading and boundary conditions applied to OS, OH and OL anchors, (b) Tw diameter parameter study boundary conditions.

(a)

Experimental results. Axial load–displacement and torque–displacement curves are shown for each thread

Tip Failure for Tw = 0.3 mm 120

Load (N)

100 80 60 Tip Bending and Failure

40 20 0 0

2

1

Displacement (mm) Th Fail 01 Th Fail 04 Average

Th Fail 02 Th Fail 05

Th Fail 03 Th Fail 06

(d) Plastic Strain Load (N)

Tw (mm)

Tw

150 N

300 N

0.6

0.043

0.394

0.8

0.000

0.064

1.0

0.000

0.017

1.2

0.000

0.000

Figure 5. (a) Load displacement curves for Tw ¼ 0.3 mm, (b,c) images of failure observed for Tw ¼ 0.3 mm, (d) finite element results for plastic strain levels for Tw values of 0.6–1.2 mm.



(JBA)


6

Journal of Biomaterials Applications 0(0) (a)

(b)

OS: 150 N

(c)

OH: 150 N

OL: 150 N

S, Mises 300 275 250 225 200 175 150 125 100 75 50 25 0

(e)

(d) OS: 300 N

(f) OH: 300 N

OL: 300 N

Figure 6. VonMises stress for axial loading of anchors; (a–c) anchors subjected to 150 N compressive force; (d–f) anchors subjected to 300 N compressive force.

design in Figure 8. The A1H1 anchor required a mean axial load of 108.4 N and a torque of 0.57 Nm for complete insertion. The A1H2 anchor required a mean axial load of 105.3 N and a torque load of 0.75 Nm for complete insertion. The A2H1 anchor required a mean axial load of 79.5 N and a torque load of 0.61 Nm for complete insertion. The A3H1 anchor required a mean axial load of 138.7 N and a torque load of 0.6 Nm for complete insertion. Results of a one way ANOVA Tukey (SNK) analysis of the four anchor types are summarised in Table 3. Insertion torques for the four thread types revealed no significant difference between all groups. Insertion tests using a pilot hole required significantly lower axial loads, while the insertion torque was unaffected by the presence of a pilot hole. Computational results Investigation of driver head geometry at different insertion depths. VonMises stress is presented for all simulations that were conducted under applied torques to

analyse the potential failure in the region of the driver head. The stress distribution within the anchor at three heights are summarised in the graph of Figure 9(a). A 10 rotation was applied to all anchors using a rigid body driver. The stress distribution within the anchor head for each of the driver shapes (for r ¼ 0.5 mm) investigated is shown in Figure 9(b–d). Simulations predict that the hexalobe anchor is subject to lower stress than the square and hexagonal drivers for all heights. At an anchor height of 0.5 mm (r ¼ 0.5 mm) the predicted max stress in the hexalobe anchor is 20% lower than square and hexagonal anchors. The VonMises stress within the anchor under a 10 applied rotation is presented in Figure 10 for all anchor designs at an anchor head height of r ¼ 0.5 mm. The computed stress in the material is lowest for the A3H1 anchor (Figure 10(vi)). This is due to the fact that the A3H1 anchor contains a greater number of threads, increasing the contact surface area and lowering the stress concentrations within the anchor. The A1H1 anchor computed the highest stresses.




(JBA)

Feerick et al.

7 The elevated stress concentrations occur in the corners of the square driver head as illustrated in Figure 10(i). All simulations compute a max shear stress of roughly 28 MPa in the peri-prosthetic foam as shown in Figure 10(vii–xi). No difference is observed between computed shear stress for the A1H1 (square, hexagonal and hexalobe driver) anchors, as the thread geometry is the same for each (Figure 10(vii–ix)). A lower stress is predicted within the 2 mm foam layer for the A1H2 anchor

Table 3. Results of one way ANOVA Tukey (SNK) analysis of the four anchor types for axial insertion forces during self-tapping. Groups A1H1 A1H1 A1H1 A1H2 A1H2 A2H1

O; Axial:108.4N, Torque:0.57Nm n=11

0.6

100

0.5

80

0.4

60

0.3

40

0.2 Axial Torque

0 0

2

4

6

8

10

12

14

Axial Load (N)

120

0.7

120

0.6

100

0.5

80

0.4

60

0.3

40

0.1

20

0.0

0

16

0.2 Axial Torque

0

2

4

Displacement (mm)

(c)

0.6

100

0.5

80

0.4

60

0.3

40

0.2 Axial Torque

20 0 4

6

8

10

Displacement (mm)

12

14

16

Axial Load (N)

0.7

2

8

10

12

14

0 16

SP3; Axial:138.7 N, Torque:0.6Nm n=4

140 Torque (Nm)

Axial Load (N)

(d)

0.8

120

0

6

0.1

Displacement (mm)

SP2; Axial:79.5N, Torque:0.61Nm n=4

140

0.8

140

0.7 Torque (Nm)

Axial Load (N)

140

20

ST; Axial:105.3N, Torque:0.75Nm n=4

(b) 0.8

0.8 0.7

120

0.6

100

0.5

80

0.4

60

0.3

40

0.1

20

0.0

0

Torque (Nm)

(a)

p < 0.0001 No statistical significance No statistical significance No statistical significance p < 0.0001 p < 0.01

A1H2 A2H1 A3H1 A2H1 A3H1 A3H1

Torque (Nm)

Figure 7. Two examples of anchors with blunted tips (Tw ¼ 0.6 mm) removed post insertion tests.

vs. vs. vs. vs. vs. vs.

Significance (p value)

0.2 Axial Torque

0

2

4

6

8

10

12

14

0.1 0.0 16

Displacement (mm)

Figure 8. Load–displacement and torque–displacement curves for each thread type; (a) A1H1 thread type, (b) A1H2 thread type, (c) A2H1 thread type and (d) A3H1 thread type.




(JBA)

8

Journal of Biomaterials Applications 0(0) Hexagonal

Hexalobe

Square

(a) 120

Max Stress (Mpa)

110 100 90 80 70 60 0

1

2

4

3

r (mm) S, Mises 25.00 22.92 20.83 18.75 16.67 14.58 12.50 10.42 8.33 6.25 4.17 2.08 0.00

(b) Square

(d) Hexalobe

(c) Hexagonal

Figure 9. Shear stress in the direction of applied torque at anchor driver interface; (a) max stress for each driver type at different insertion depths, r ¼ 0.5, 2 and 4 mm; shear stress distribution at the anchor driver interface for (b) square, (c) hexagonal and (d) hexalobe.

S, Mises

(i)

(ii)

A1H1

A1H1

A1H1

(iii)

Square Driver

Hexagonal Driver

Hexalobe Driver

(vii) A1H1

(viii) A1H1

(ix)

(iv)

(v)

(vi)

A1H2

A2H1

A3H1

25.00 22.92 20.83 18.75 16.67 14.58 12.50 10.42 8.33 6.25 4.17 2.08 0.00

S, Mises

A1H1

(x)

(xi)

(xii)

A1H2

A2H1

A3H1

5.00 4.58 4.17 3.75 3.33 2.92 2.50 2.08 1.67 1.25 0.83 0.42 0.00

Figure 10. (i–vi) Stress (vonMises, MPa) distribution in each anchor design under 10 driver rotation; (vii–xii) Stress (vonMises, MPa) distribution in each block of Sawbone foam under 10 driver rotation (at r ¼ 0.5 mm).

(Figure 10(x)) due to the extended non-tapered region with h ¼ 4 mm, as shown in Figure 1. The computed shear stress, within the 2 mm 40 pcf layer, is higher for the A2H1 and A3H1 anchor designs (Figure 10 (v,vi)). This is due to the fact that the lower pitch

magnitude for each design leads to a larger number of threads in contact with this region. Investigation of wall thickness within anchor head. The computed vonMises stress for alternate driver geometry



(JBA)


Feerick et al.

9 (a)

(b) S, Mises

S, Mises 25.00 22.92 20.83 18.75 16.67 14.58 12.50 10.42 8.33 6.25 4.17 2.08 0.00

25.00 22.92 20.83 18.75 16.67 14.58 12.50 10.42 8.33 6.25 4.17 2.08 0.00

A1H1 Square Driver: 1.80 mm WT: 1.15 mm

(c)

(d)

S, Mises

S, Mises

25.00 22.92 20.83 18.75 16.67 14.58 12.50 10.42 8.33 6.25 4.17 2.08 0.00

25.00 22.92 20.83 18.75 16.67 14.58 12.50 10.42 8.33 6.25 4.17 2.08 0.00

A1H1 Hexagonal Driver: 1.5 mm WT: 1.30 mm

(e)

(f)

S, Mises

S, Mises

25.00 22.92 20.83 18.75 16.67 14.58 12.50 10.42 8.33 6.25 4.17 2.08 0.00

A1H1 Hexalobe Driver: 1.68 mm WT: 1.21 mm

25.00 22.92 20.83 18.75 16.67 14.58 12.50 10.42 8.33 6.25 4.17 2.08 0.00

A1H1 Square Driver: 2.50 mm WT: 0.80 mm

A1H1 Hexagonal Driver: 2.0 mm WT: 1.05 mm

A1H1 Hexalobe Driver: 2.69 mm WT: 0.71 mm

Figure 11. VonMises Stress (MPa) for varying wall thickness at anchor head height of 0.5 mm. (a) A1H1 anchor with square driver thickness 1.8 mm and wall thickness of 1.15 mm, (b) A1H1 anchor with square driver thickness 2.5 mm and wall thickness of 0.8 mm, (c) A1H1 anchor with hexagonal driver thickness of 1.5 mm and wall thickness of 1.3 mm, (d) A1H1 anchor with hexagonal driver thickness of 2.0 mm and wall thickness of 1.05 mm, (e) A1H1 anchor with hexalobe driver diameter of 1.68 mm and wall thickness of 1.21 mm and (f) A1H1 anchor with hexalobe driver diameter of 2.69 mm and wall thickness of 0.71 mm.

sizes is shown in Figure 11. The maximum stress computed within the anchor head reduced from 69.3 MPa to 46.9 MPa as the square driver wall thickness increased from 1.8 mm to 2.5 mm. The computed stress concentrations at the thread–core interface increased with decreasing wall thickness (Figure 11b). As the hexagonal driver thickness was increased from 1.5 mm to 2.0 mm (Figure 11c,d) the max stress computed within the head was reduced from 58.1 MPa to 51.5 MPa. A further reduction in wall thickness may reduce the maximum stress to similar level computed for the square diver geometries. The results for hexalobe driver geometry sizes are summarised in Figure 11(e,f). As the hexalobe driver wall thickness was increased from 1.68 mm (Figure 11e) to 2.69 mm (Figure 11f) the stress concentrations in the anchor surrounding the driver edges were reduced. The computed max stress within the anchor head was reduced from

172.5 MPa to 42.1 MPa as the wall thickness reduced from 1.21 mm to 0.71 mm. As the wall thickness is increased the risk of failure due to splitting of the anchor head is reduced.

Discussion and conclusions In summary, the self-tapping capability of PEEK OPTIMA reinforced suture anchors has been confirmed for insertion into synthetic polyurethane foam. However the optimum design has yet to be identified with several key design considerations being identified in this study: . Anchor tips should be redesigned with a larger contact area to prevent plastic deformation and failure of the tip. This is also supported by the experimental results for anchors with Tw ¼ 0.3 mm, where tip



(JBA)


10


bending/failure was observed during insertion. Currently unreinforced PEEK suture anchors require pre-tapped holes or the assistance of a metallic punch/anchor tip as the material is not capable of self-tapping.18–20 However, the use of CFR-PEEK could potentially overcome this limitation, making self-tapping possible due to the improved mechanical properties.3 . Wall thickness within the head of the anchor determines the failure strength of the anchor under the applied torque during insertion of the anchor. An optimum wall thickness and driver geometry must be designed to maximise contact area between the driver and the anchor, ensuring that stresses within the wall and anchor threads do not exceed the strength of the material. . Anchors with a greater number of threads, such as the A2H1 and A3H1 anchors, are subject to lower stress levels during insertion due to larger contact areas. However, anchors with a higher number of threads require slower insertion rates and higher number of turns during insertion. Both of these limitations could prove problematic for surgery times and lead to potential heat generation in the surrounding bone tissue during insertion.21 . Pilot holes reduce the axial force required before the anchor begins to self-tap, consequently reducing the stress in the anchor tip. The current study suggests that in the absence of pilot holes, thicker anchor tips should be used to avoid localised plastic deformation and tip fracture. Some studies have shown that drilling of pilot holes can lead to elevated heat profiles in the surrounding bone versus the compression insertion of a self-tapping insertion.21 Such elevated temperatures can lead to osteonecrosis of the tissue surrounding the implant.21,22 Pilot holes also lead to removal of a greater volume of material during hole preparation which could potentially reduce the pullout strength of the screw/anchor.12,23 A number of commercially available self-tapping titanium suture anchors require the penetration of the cortical shell with a punch prior to self-tapping of the anchor.24 This is analogous to the generation of a pilot hole and the current study demonstrates the potential benefits of such an approach for the insertion of CFR-PEEK anchors. The experimental testing was conducted using a Zwick biaxial tester; this ensured precise control and measurement of axial and torque loading. In a clinical environment a large variability in the applied loads could be encountered. A recent study has reported difficulties in control of applied forces during surgery.25 The ability to maintain axial and torque loads within a certain approved range for a given design may be required to

ensure that the anchor does not become overloaded during insertion. In the case of the anchors examined in the present study, the axial load should not exceed 150 N and the torque should not exceed 0.6 Nm. The synthetic bone used in the current study consists of a bi-layered polyurethane test block (Sawbones, Malmo¨, Sweden13). A dense (40 pcf) 2 mm layer represents the outer cortical bone shell, while the underlying porous material (15 pcf) represents the cancellous bone. A similar bi-layer composite structure has previously been used to assess of the ability of absorbable suture anchor designs to undergo self-tapping insertion.14 Previous investigations report that the stiffness of Sawbone synthetic bone falls within the range reported for cadaveric bone samples.26–28 The aim of the present study is to provide a robust and reliable preliminary assessment of the mechanical performance of CFR-PEEK anchors during insertion. Controlled test conditions are essential for such a mechanical investigation, requiring the use of rigorously characterised and repeatable Sawbone substrate structure. Such an approach is commonly used for pilot studies of orthopaedic devices.28–34 The use of cadaveric bone is hindered by significant inter-specimen variability, prohibiting a reliable repeatable assessment of the device performance. The current study identifies a number of key design parameters for CFR-PEEK suture anchors. Building upon the current study, a redesigned suture anchor should next be developed and tested, again using the controlled Sawbone substrates. Having reliably established the performance mechanical characteristics of the device, a follow-on investigation using cadaveric bone should be completed. Finally, in vivo clinical trials could be undertaken. The current study demonstrates the importance of computational analysis for optimisation of the next generation of PEEK anchors, identifying locations of stress concentrations and key geometric features. Future studies should be performed to develop more advanced computational models to further evaluate design and insertion of anchors. In particular, cracking and fracture of cortical bone should be incorporated into finite element models to analyse the mechanics of self-tapping tip insertion.35,36 Additionally, plastic deformation and crushing of trabecular bone should be considered in detailed simulations of anchor insertion37,38 and computational evaluation of suture anchor performance should include the application of physiological loading and prediction of fatigue failure.39 In conclusion, the present study demonstrates the self-tapping and self-drilling ability of CFR-PEEK suture anchors for insertion into polyurethane foam bone analogue. Critical design parameters have been identified, with the anchor tip and driver head



(JBA)


Feerick et al.

11

geometries being of particular importance. The experimental-computational approach presented in this study provides a framework for the optimisation of PEEK self-tapping suture anchor design. Acknowledgements The authors would like to acknowledge the Medical Engineering Design and Innovation Centre (MEDIC) at the Cork Institute of Technology (CIT) for providing access to the Zwicki-Line biaxial tester at their facility, for the experimental insertion testing of the current study.

Conflict of interest The authors declare that there is no conflict of interest.

Funding This work was commissioned by Invibio Biomaterials Solutions, Thornton Cleveleys, Lancashire, United Kingdom.

References 1. Tempelhof S, Rupp S and Seil R. Age-related prevalence of rotator cuff tears in asymptomatic shoulders. J Shoulder Elbow Surg 1999; 8: 296–299. 2. AAOS. American Academy of Orthopaedic Surgeons, Research Statistics on Rotator Cuff Repairs, National Inpatient Sample, 1998–2004. 2004. 3. Kurtz SM and Devine JN. PEEK biomaterials in trauma, orthopedic, and spinal implants. Biomaterials 2007; 28: 4845–4869. 4. Ferguson S, Visser JA and Polikeit A. The long-term mechanical integrity of non-reinforced PEEK-OPTIMA polymer for demanding spinal applications: experimental and finite-element analysis. Eur Spine J 2006; 15: 149–156. 5. Brantigan JW. A carbon fiber implant to aid interbody lumbar fusion. Two-year clinical results in the first 26 patients. Spine 1993; 18: 2106–2117. 6. Kurtz SM. Applications of polyaryletheretherketone in spinal implants: fusion and motion preservation. PEEK Biomater Handb 2011: 201–220. 7. Lovald S and Kurtz SM. Applications of polyetheretherketone in trauma, arthroscopy, and cranial defect repair. PEEK Biomater Handb 2011: 243–260. 8. Scholes SC and Unsworth A. The wear properties of CFR-PEEK-OPTIMA articulating against ceramic assessed on a multidirectional pin-on-plate machine. Proc IMechE, Part H: J Eng Med 2007; 221: 281–289. 9. Ponnappan RK, Serhan H, Zarda B, et al. Biomechanical evaluation and comparison of polyetheretherketone rod system to traditional titanium rod fixation. Spine J 2009; 9: 263–267. 10. Steinberg EL, Rath E, Shlaifer A, et al. Carbon fiber reinforced PEEK Optima—A composite material biomechanical properties and wear/debris characteristics of CF-PEEK composites for orthopedic trauma implants. J Mech Behav Biomed Mater 2013; 17: 221–228.

11. Porto MA, Silva P, Rosa R, et al. Experimental in vivo acute and chronic biomechanical and histomorphometrical comparison of self-drilling and self-tapping anterior cervical screws. Eur Spine J 2012; 21: 956–963. 12. O¨ktenog˘lu BT, Ferrara LA, Andalkar N, et al. Effects of hole preparation on screw pullout resistance and insertional torque: a biomechanical study. J Neurosurg 2001; 94: 91–96. 13. PacificResearchLaboratories. Sawbones biomechanical test materials. Pacific Research Laboratories, 2012. 14. Meyer DC, Fucentese SF, Ruffieux K, et al. Mechanical testing of absorbable suture anchors. Arthroscopy 2003; 19: 188–193. 15. ASTM. ASTM F543 - 13e1 Standard specification and test methods for metallic medical bone screws. ASTM International, 2013. 16. Grant JA, Bishop NE, Go¨tzen N, et al. Artificial composite bone as a model of human trabecular bone: the implant–bone interface. J Biomech 2007; 40: 1158–1164. 17. ASTM. ASTM F116 - 12 Standard Specification for Medical Screwdriver Bits. ASTM International, 2012. 18. Efird C, Traub S, Baldini T, et al. Knotless single-row rotator cuff repair: a comparative biomechanical study of 2 knotless suture anchors. Orthopedics 2013; 36: e1033–e1037. 19. Barber F, Herbert MA, Hapa O, et al. Biomechanical analysis of pullout strengths of rotator cuff and glenoid anchors: 2011 update. Arthroscopy 2011; 27: 895–905. 20. Barber FA, Herbert MA, Beavis RC, et al. Suture anchor materials, eyelets, and designs: update 2008. Arthroscopy 2008; 24: 859–867. 21. Markovic´ A, Misˇ ic´ T, Milicˇic´ B, et al. Heat generation during implant placement in low-density bone: effect of surgical technique, insertion torque and implant macro design. Clin Oral Implants Res 2013; 24(7): 798–805. 22. Pandey RK and Panda SS. Drilling of bone: a comprehensive review. J Clin Orthop Trauma 2013; 4: 15–30. 23. Chen L-H, Tai C-L, Lee D-M, et al. Pullout strength of pedicle screws with cement augmentation in severe osteoporosis: a comparative study between cannulated screws with cement injection and solid screws with cement prefilling. BMC Musculoskelet Disord 2011; 12: 33. 24. Burkhart SS. Arthroscopic rotator cuff repair:Indications and technique. Oper Tech Sports Med 1997; 5: 204–214. 25. Stoesz MJ, Gustafson PA, Patel B, et al. Surgeon perception of cancellous screw fixation. J Orthop Trauma 2014; 28(1): e1–e7. 26. Poukalova M, Yakacki CM, Guldberg RE, et al. Pullout strength of suture anchors: effect of mechanical properties of trabecular bone. J Biomech 2010; 43: 1138–1145. 27. Patel PS, Shepherd DE and Hukins DW. Compressive properties of commercially available polyurethane foams as mechanical models for osteoporotic human cancellous bone. BMC Musculoskelet Disord 2008; 9: 137. 28. Hausmann JT. Sawbones in biomechanical settings - a review. Osteo Trauma Care 2006; 14: 259–264. 29. Feerick EF, Kennedy J, FitzPatrick DP, et al. Computational Investigation of metallic and carbon fibre PEEK fracture fixation devices for three-part



(JBA)


12

30.

31.

32.

33.

34.

Journal of Biomaterials Applications 0(0) proximal humeral fractures. Med Eng Phys 2012; 35(6): 712–722. Cawley DT, Kelly N, Simpkin A, et al. Full and surface tibial cementation in total knee arthroplasty: a biomechanical investigation of stress distribution and remodelling in the tibia. Clin Biomech 2012; 27(4): 390–397. Kelly N, Cawley DT, Shannon FJ, et al. An investigation of the inelastic behaviour of trabecular bone during the press-fit implantation of a tibial component in total knee arthroplasty. Med Eng Phys 2013; 35(11): 1599–1606. Bert JM and McShane M. Is it necessary to cement the tibial stem in cemented total knee arthroplasty. Clin Orthop 1998; 356: 73–78. Completo A, Fonseca F and Simoes JA. Finite element and experimental cortex strains of the intact and implanted tibia. J Biomech Eng 2007; 129: 791–797. Luring C, Perlick L, Trepte C, et al. Micromotion in cemented rotating platform total knee arthroplasty: cemented tibial stem versus hybrid fixation. Arch Orthop Trauma Surg 2006; 126: 45–48.

35. Feerick EM, Liu X and McGarry P. Anisotropic modedependent damage of cortical bone using the extended finite element method (XFEM). J Mech Behav Biomed Mater 2013; 20: 77–89. 36. Feerick EM and McGarry JP. Cortical bone failure mechanisms during screw pullout. J Biomech 2012; 45: 1666–1672. 37. Kelly N, Harrison N, McDonnell P, et al. An experimental and computational investigation of the post-yield behaviour of trabecular bone during vertebral device subsidence. Biomech Model Mechanobiol 2013; 12: 685–703. 38. Kelly N and McGarry JP. Experimental and numerical characterisation of the elasto-plastic properties of bovine trabecular bone and a trabecular bone analogue. J Mech Behav Biomed Mater 2012; 9: 184–197. 39. Ryan G, McGarry P, Pandit A, et al. Analysis of the mechanical behavior of a titanium scaffold with a repeating unit-cell substructure. J Biomed Mater Res B 2009; 90: 894–906.


Comparison and Analysis on Mechanical Property and Machinability about Polyetheretherketone and Carbon-Fibers Reinforced Polyetheretherketone.

Multilevel surface engineering of nanostructured TiO2 on carbon-fiber-reinforced polyetheretherketone.

The influence of nominal stress on wear factors of carbon fibre-reinforced polyetheretherketone (PEEK-OPTIMA® Wear Performance) against zirconia toughened alumina (Biolox® delta ceramic).

Effect of surface roughness on osteogenesis in vitro and osseointegration in vivo of carbon fiber-reinforced polyetheretherketone-nanohydroxyapatite composite.

Mechanical properties of all-suture anchors for rotator cuff repair.

Adverse events associated with biodegradable lactide-containing suture anchors.

Recycling Suture Limbs from Knotless Suture Anchors for Arthroscopic Shoulder Stabilization.

Anatomic Hip Capsular Reconstruction With Separate Suture Anchors.

Dust from carbon fibre.

Carbon fibre microelectrodes.

Characterization of short-fibre reinforced thermoplastics for fracture fixation devices.

Ultrasonic assembly of anisotropic short fibre reinforced composites.

Cytotoxicity of coated and uncoated fibre-reinforced composites.

Outcome of repair of chronic tear of the pectoralis major using corkscrew suture anchors by box suture sliding technique.

Fibre Bragg Gratings in Embedded Microstructured Optical Fibres Allow Distinguishing between Symmetric and Anti-Symmetric Lamb Waves in Carbon Fibre Reinforced Composites.

Carbon fiber-reinforced carbon as a potential implant material.

Carbon nanotubes reinforced composites for biomedical applications.

Traumatic dislocation of posterior tibial tendon by avulsion of flexor retinacular release. Reconstruction with suture anchors.

Acute distal biceps ruptures: single incision repair by use of suture anchors.

Repair of acute injuries of the lateral ligament complex of the ankle by suture anchors.

Reprint of: The deadman theory of suture anchors: observations along a south Texas fence line.

Increasing pullout strength of suture anchors in osteoporotic bone using augmentation--a cadaver study.

Biomechanical Evaluation of Classic Solid and All-Soft Suture Anchors for Medial Patellofemoral Ligament Reconstruction.

Arthroscopic Fixation of Comminuted Glenoid Fractures Using Cannulated Screws and Suture Anchors.