Keywords
biomechanics - working length - plate strain
Introduction
Fractures with a very short proximal or distal juxta-articular fragment create particular
biomechanical challenges, complicating implant selection and placement. In veterinary
orthopaedics, distal radial fractures in toy and miniature breeds are a common example
of this, with the majority of these fractures reported to involve the distal third
of the radius, resulting in a very short distal fragment.[1]
[2] Plate fixation has been reported to have good overall success in these cases.[3]
A recent clinical report described use of either a 2.0 mm notched head locking compression
plate T-plate (NHTP) (LCP, Synthes GmbH, Oberdorf, Switzerland) or a 2.0 mm straight
locking compression plate (LCP) (Synthes GmbH, Oberdorf, Switzerland) for distal antebrachial
fractures in toy and miniature breed dogs weighing less than 6 kg.[1] Differences in design include differing plate geometry and dimensions, the number,
position and type of screws that can be placed in a very short juxta-articular fragment.
Given the design differences between the NHTP and equivalent-length straight LCP,
it is expected they will have different biomechanical properties. There is no information
on the comparable biomechanical performance of these two plates.
The 2 mm NHTP allows placement of locking screws in three stacked holes in a triangular
configuration within the plate head. These three screws can be placed in a 9 mm long
bone fragment. The 2 mm LCP permits placement of one locking screw and one cortical
screw in the equivalent length fragment. Of the 20 cases reported by Gibert and colleagues,[1] 10 cases were stabilized with only two screws placed in the distal fragment. Two
of those cases were stabilized with a straight LCP with only one locking screw and
one cortical screw placed. Hybrid fixation was performed in all dogs in one or both
fragments. Hybrid fixation is a term to describe the combination of both locking and
cortical screws within a single plate construct.[4]
[5]
Plate working length is one of the factors that affects construct stability.[6]
[7]
[8]
[9]
[10]
[11] Working length is affected by size of the fracture gap, plate stand-off distance,
distance between the innermost screws and load direction. The effect of working length
on construct stability and plate strain/stress is still the subject of some controversy.[11]
[12]
[13]
[14] A finite element analysis study demonstrated that a longer working length increased
plate stress in a 6 mm fracture gap model, though paradoxically reduced plate stress
in a 1 mm fracture gap model, during axial compression producing tension bending.[6] Much of the information on the effect of working length on plate strain/stress is
based on fracture gap models,[6]
[7]
[8]
[15]
[16] with no information from compressed fracture models.
The first objective of this study was to compare the biomechanical properties of a
2.0 mm NHTP construct to a 2.0 mm straight LCP construct, applied to a synthetic compressed
transverse fracture model, with a very short 9 mm fragment, using screw configurations
that simulated clinical application.[1] We hypothesized that the NHTP construct would be less stiff and have greater plate
strain in bending; however, it would be stiffer and have less strain in torsion testing
than the LCP construct.
The second objective was to compare the biomechanical properties of a second screw
configuration that created a longer working length than the original screw configuration.
We hypothesized that for each plate type, the screw configuration with the longer
working length would create a construct that was less stiff and have greater plate
strain in bending and torsion testing, compared with the original screw configuration
with the shorter working length.
Materials and Methods
A synthetic, compressed, transverse fracture model was created using two Delrin (Delrin
Acetal Polymer: Polytech Plastics Australasia, Jandakot, WA, Australia) tubes 100 mm
in length, with a 12.5 mm outer diameter and a 7.4 mm inner diameter. The distal fragment
modelled was 9 mm in length, which is within the reported range in a previous clinical
series of distal radial fractures in dogs < 6 kg.[1]
The dimensions of the shaft of the NHTP are less than the LCP. The shaft of the NHTP
has a width of 5 mm and thickness of 1.3 mm, whereas the LCP has a width of 5.5 mm
and thickness of 1.5 mm.[17] The length of the 2 mm NHTP is comparable to the 8-hole 2 mm LCP, 54 mm and 55 mm
respectively. The NHTP allows compression of a fracture from any of the shaft combination
holes, all directing compression towards the head of the plate. The LCP allows compression
from any of the combination holes; however, each plate half compresses towards the
middle of the plate.[17]
Configuration 1
The NHTP and LCP were each applied as compression plates with no stand-off distance
using bicortical screws positioned as end of fragment screws.[18] The insertional torque applied to each screw was standardized using a 0.4 Nm torque
limiter (Torque Limiter, 0.4 Nm, with AO/ASIF Quick Coupling: Synthes GmbH), as recommended
for 2.0 mm screws.[17] For all constructs, the method of implant placement was performed following Arbeitsgemeinschaft
für Osteosynthesefragen (AO Foundation, Davos, Switzerland) recommendations.[19]
For the NHTP, three locking screws (Self-tapping Locking Screw Star Drive, Synthes
GmbH) were placed in the plate head in the short fragment. The short fragment was
arbitrarily defined as the distal fragment. The fracture was compressed with a single
cortical screw (Self-tapping Cortex Screw Star Drive: Synthes GmbH), placed as a compression
screw, in the shaft hole immediately proximal to the fracture, plate hole six. Plate
holes were numbered sequentially from proximal to distal in both plates. Two additional
locking screws (Self-tapping Locking Screw Star Drive, Synthes GmbH) were subsequently
placed in the proximal segment in plate holes one and five ([Fig. 1A]).
Fig. 1 (A) Configuration 1 constructs (short working length) for the notched head T-plate (NHTP)
and straight locking compression plate (LCP). The hollow black circle indicates a
locking screw. The solid black circle indicates a cortical screw in compression. (B) Configuration 2 (long working length) constructs.
For the LCP, a locking screw (Self-tapping Locking Screw Star Drive, Synthes GmbH,
Oberdorf, Switzerland) was placed immediately proximal to the fracture in hole six
followed by a single cortical screw (Self-tapping Cortex Screw Star Drive: Synthes
GmbH) placed as a compression screw in plate hole seven in the distal fragment. Further
locked screws were placed in holes one, five and eight achieving three locking screws
in the proximal fragment and one cortical screw and one locking screw in the distal
fragment ([Fig. 1A]).
Configuration 2: Longer Working Length
The longer working length constructs were created by leaving screws out immediately
adjacent to the fracture in the proximal fragment. For both plates, two locking screws
(Self-tapping Locking Screw Star Drive, Synthes GmbH) were placed in the first and
second most proximal shaft holes. A compression cortical screw was placed in the third
most proximal shaft hole in the NHTP. A third locking screw was placed in the third
most proximal shaft hole in the LCP ([Fig. 1B]). Screw placement in the short distal fragment was identical to screw configuration
1. As with screw configuration 1, both plates were applied in compression, using bicortical
screws with no plate stand-off distance. The sequence and method of screw placement
followed AO (AO Foundation) recommendations.[19]
A sample size of seven replicates of each plate was used, sufficient to detect an
effect size as small as 1.7 which is smaller than expected based on previous work
using Delrin models (power = 0.8, α = 0.05, variance = 10%).[7]
Non-destructive four-point quasi-static bending was conducted on a material testing
machine (Instron 5848 MicroTester, Norwood, Massachusetts, United States) with a 100 N
load cell, applying a constant bending moment of 0.6 Nm. The support rollers had a
200 mm gap and the load rollers had a 140 mm gap ([Fig. 2]). Each construct was pre-loaded (0.4 N), then ramp loaded for three cycles under
displacement control at 10 mm/min to a force of 40 N, to produce a peak bending moment
of 0.6 Nm in all tested planes. This bending moment was within the elastic limit of
the constructs.
Fig. 2 Four-point bending was conducted on a material testing machine, applying a constant
bending moment along the constructs about three different planes. (A) Compression bending. (B) Perpendicular bending. (C) Tension bending.
Each construct was sequentially tested in four-point bending about three different
planes. Testing was first performed simulating compression bending, with load applied
parallel to the screw axis along the compression surface of the construct ([Fig. 2A]). The construct was then rotated 90 degrees and testing repeated, simulating perpendicular
bending, with load applied perpendicular to the screw axis ([Fig. 2B]). The construct was again sequentially rotated 90 degrees and testing repeated,
to simulate tension bending, with load applied parallel to the screw axis ([Fig. 2C]).
For testing in non-destructive torsion, the proximal end of the construct was restrained
in a custom-made jig allowing free rotation. Axial load was applied to a jig screw
using a material testing machine (Instron 5567: Instron, Canton, Massachusetts, United
States) to create torque, resulting in a rotational displacement of ∼5.6 degrees/s,
to produce a peak vertical displacement of 5 mm (11.3 degrees torsion reached).
Measurement of Stiffness
All testing load and actuator displacement measurements were recorded. Bending and
torsional stiffness was determined from the slope of the linear elastic portion of
the load displacement curve.
Measurement of Strain
Three-dimensional digital image correlation was used to measure plate surface strain
during compression bending and torsion testing only.[20]
[21]
[22]
[23] Each construct was sprayed with a speckle pattern (uniform base white followed by
a black speckle) prior to testing.[20] The high definition recordings were captured with VicSnap software (VicSnap, Correlated
Solutions, North Carolina, United States).[20] Strain was measured for 12 regions of interest (ROI) along the plate surface using
Vic-3D software (VIC-3D, Correlated Solutions, North Carolina, United States) ([Fig. 3]).[20] The ROI were numbered sequentially, representative of analogous regions for both
plates. The regions along the axial solid plate section were marked by odd numbers
and the abaxial partial plate sections, by even numbers. The mean von Mises strain
was calculated from the linear line of best fit of the strain load graph from the
third cycle, for each sample tested.
Fig. 3 Region of interest from where strain was measured for both plate types using three-dimensional
digital image correlation. LCP, locking compression plate; NHTP, notched head T-plate.
Statistical Analysis
The stiffness for each four-point bending and torsion test was the response of interest
and confirmed to be normally distributed using the Shapiro–Wilk test and visual inspection
of Q-Q plots. The stiffness was summarized as mean (95% confidence interval, CI).
The stiffness in each plane of bending for screw configuration 1 was compared between
plate types using a t-test. The stiffness in each plane of bending was compared between screw configurations
within each plate type using t-tests. Equality of variances was tested and either a pooled or Satterthwaite test
of significance was used based on equality/non-equality variances, respectively, to
avoid type 1 error. Significance was determined at p ≤ 0.05.
The strain was verified as normally distributed using the Shapiro–Wilk test and visual
inspection of Q-Q plots, and summarized as mean (95% CI). The strain for screw configuration
1 was compared between plate types using a two-way analysis of variance including
the main effects of plate type and ROI, and their interaction. When there was significant
interaction (p ≤ 0.05), selected post-hoc, pairwise comparisons were made between plate types
at each region against a Bonferroni-adjusted p ≤ 0.005. The strain for configuration 1 and 2 was compared within each plate type
using a two-way analysis of variance including the main effects of screw configuration
and ROI, and their interaction. When there was significant interaction (p ≤0.05), selected post-hoc pairwise comparisons were made between screw configuration
for each ROI, within each plate type against a Bonferroni-adjusted p ≤ 0.005. SAS v9.4 (SAS Institute, Cary, North Carolina, United States) was used
for analysis.
Results
Comparison of Plate Type for Screw Configuration 1
The LCP was significantly stiffer than the NHTP in all three planes of bending (p ≤ 0.05, [Table 1]). The NHTP was significantly stiffer than the LCP in torsion (p ≤ 0.05, [Table 1]).
Table 1
Mean (95% confidence interval) stiffness in four-point compression, perpendicular
and tension bending (N/mm) and torsion (Nm/degree) with comparison between the NHTP
and LCP screw configuration 1
|
Plate
|
Compression bending
|
Perpendicular bending
|
Tension bending
|
Torsion
|
|
NHTP
|
36.0
(34.0–38.0)
|
63.7
(57.5–69.8)
|
50.17
(47.7–52.6)
|
0.8
(0.75–0.85)
|
|
LCP
|
48.33
(47.5–49.2)
|
66.8
(65.0–68.5)
|
83.7
(79.7–87.6)
|
0.68
(0.65–0.71)
|
|
p-Value
|
≤0.0001
|
0.0128
|
<0.0001
|
0.0026
|
Abbreviations: LCP, locking compression plate; NHTP, notched head T-plate.
The NHTP had significantly greater strain than the LCP during compression bending
at five of twelve ROI (p ≤ 0.005, [Table 2]), with no difference at the remaining regions.
Table 2
Mean (95% confidence interval) plate strain (mm/mm) at ROI during compression bending
and torsion with comparison between the NHTP and LCP screw configuration 1
|
Bending
|
ROI
|
|
1
|
2
|
3
|
4
|
5
|
6
|
7
|
8
|
9
|
10
|
11
|
12
|
|
NHTP
|
1.34 (1.27–1.47)
|
4.44 (4.1–4.77)
|
2.38 (2.03–2.73)
|
4.80 (4.21–5.40)
|
1.74 (1.46–2.03)
|
2.82 (2.67–2.98)
|
1.79 (1.43–2.15)
|
3.96 (3.27–4.66)
|
2.41 (2.03–2.80)
|
3.82 (3.19–4.45)
|
1.47 (1.26–1.68)
|
2.53 (2.01–3.06)
|
|
LCP
|
1.17 (1.11–1.24)
|
3.13 (2.52–3.74)
|
1.59 (1.48–1.70)
|
3.57 (3.44–3.70)
|
1.45 (1.31 -1.59)
|
3.22 (2.63–3.82)
|
1.25 (1.18–1.31)
|
3.50 (3.20–3.81)
|
1.90 (1.83–1.98)
|
3.24 (2.77–3.72)
|
1.93 (1.78–2.07)
|
2.53 (2.29–2.78)
|
|
Significance[a]
|
NS
|
[a]
|
[a]
|
[a]
|
NS
|
NS
|
NS
|
[a]
|
NS
|
[a]
|
NS
|
NS
|
|
Torsion
|
|
NHTP
|
2.74 (2.43–3.04)
|
12.52 (2.43–3.04)
|
3.29 (3.05–3.52)
|
13.67 (12.31–15.02)
|
3.16 (2.96–3.36)
|
17.61 (15.86–19.35)
|
3.59 (3.06–4.13)
|
16.95 (14.80–19.11)
|
7.96 (7.06–8.86)
|
21.32 (19.33–23.31)
|
5.92 (5.57–6.26)
|
12.40 (9.48–15.32)
|
|
LCP
|
3.52 (3.13–3.91)
|
11.90 (10.48–13.32)
|
3.77 (3.32–4.23)
|
15.36 (13.40–17.33)
|
5.16 (4.91–5.41)
|
14.93 (13.00–16.86)
|
5.93 (5.58–6.28)
|
21.45 (18.10–24.79)
|
7.57 (7.09–8.05)
|
18.25 (14.25–22.26)
|
6.59 (6.03–7.16)
|
17.37 (15.76–18.97)
|
|
Significance[a]
|
NS
|
NS
|
NS
|
NS
|
NS
|
[a]
|
NS
|
NS
|
NS
|
[a]
|
NS
|
NS
|
Abbreviations: LCP, locking compression plate; NHTP, notched head T-plate; NS, not
specified; ROI, regions of interest.
a Bonferroni adjusted p ≤ 0.005.
The NHTP had significantly greater strain than the LCP during torsion testing at two
of twelve ROI (p ≤ 0.005, [Table 2]), with no difference at the remaining regions.
Comparison of Screw Configuration 1 and 2: Notched Head T-plate
Configuration 1 was significantly stiffer in all three planes of bending and in torsion
(p ≤ 0.05, [Table 3]).
Table 3
Mean (95% confidence interval) stiffness in four-point compression, perpendicular
and tension bending (N/mm) and torsion (Nm/degree) of NHTP and LCP with comparison
between screw configuration 1 and 2 for each plate type
|
NHTP
|
Compression bending
|
Perpendicular bending
|
Tension bending
|
Torsion
|
|
Configuration 1 (short working length)
|
36.0
(34.0–38.0)
|
63.7
(57.5–69.8)
|
50.17
(47.7–52.6)
|
0.8
(0.75–0.85)
|
|
Configuration 2 (long working length)
|
16.47
(14.9–17.9)
|
53.24
(49.8–56.6)
|
39.4
(38.2–40.6)
|
0.63
(0.60–0.67)
|
|
p-Value
|
≤0.0001
|
≤0.0001
|
≤0.0001
|
0.0003
|
|
LCP
|
Compression bending
|
Perpendicular bending
|
Tension bending
|
Torsion
|
|
Configuration 1 (short working length)
|
48.33
(47.5–49.2)
|
66.8
(65.0–68.5)
|
83.7
(79.7–87.6)
|
0.68
(0.65–0.71)
|
|
Configuration 2 (long working length)
|
26.99
(25.9–28.0)
|
70.16
(67.7–72.6)
|
77.69
(73.0–82.3)
|
0.53
(0.51–0.54)
|
|
p-Value
|
≤0.0001
|
0.0640
|
0.1
|
≤0.0001
|
Abbreviations: LCP, locking compression plate; NHTP, notched head T-plate.
Configuration 2 had significantly greater strain than configuration 1 during compression
bending at all ROI (p < 0.005), except one region where there was no difference ([Table 4]).
Table 4
Mean (95% confidence interval) plate strain (mm/mm) at ROI during compression bending
of the NHTP and LCP with comparison between screw configurations 1 and 2 (Configuration
1 and 2)
|
NHTP
|
ROI
|
|
1
|
2
|
3
|
4
|
5
|
6
|
7
|
8
|
9
|
10
|
11
|
12
|
|
Configuration 1 (short working length)
|
1.34 (1.27–1.47)
|
4.44 (4.1–4.77)
|
2.38 (2.03–2.73)
|
4.80 (4.21–5.40)
|
1.74 (1.46–2.03)
|
2.82 (2.67–2.98)
|
1.79 (1.43–2.15)
|
3.96 (3.27–4.66)
|
2.41 (2.03–2.80)
|
3.82 (3.19–4.45)
|
1.47 (1.26–1.68)
|
2.53 (2.01–3.06)
|
|
Configuration 2 (long working length)
|
3.53 (2.53–4.53)
|
9.89 (8.89–10.89)
|
4.50 (3.79–5.21)
|
9.46 (8.592–10.34)
|
3.59 (2.79–4.41)
|
8.61 (7.33–9.89)
|
3.08 (2.55–3.60)
|
6.40 (5.68–7.12)
|
3.94 (3.34–4.53)
|
7.83 (6.83–8.83)
|
3.44 (2.82–4.05)
|
5.55 (4.88–6.22)
|
|
Significance[a]
|
[a]
|
[a]
|
[a]
|
[a]
|
[a]
|
[a]
|
NS
|
[a]
|
[a]
|
[a]
|
[a]
|
[a]
|
|
LCP
|
|
Configuration 1 (short working length)
|
1.17 (1.11–1.24)
|
3.13 (2.52–3.74)
|
1.59 (1.48–1.70)
|
3.57 (3.44–3.70)
|
1.45 (1.31 -1.59)
|
3.22 (2.63–3.82)
|
1.25 (1.18–1.31)
|
3.50 (3.20–3.81)
|
1.90 (1.83–1.98)
|
3.24 (2.77–3.72)
|
1.93 (1.78–2.07)
|
2.53 (2.29–2.78)
|
|
Configuration 2 (long working length)
|
1.61 (1.53–1.68)
|
4.26 (3.52–5.01)
|
2.18 (1.99–2.38)
|
4.90 (4.58–5.23)
|
1.99 (1.78–2.20)
|
4.41 (3.59–5.22)
|
1.70 (1.63–1.78)
|
4.78 (4.43–5.14)
|
2.60 (2.55–2.66)
|
4.46 (3.74–5.18)
|
2.64 (2.41–2.88)
|
3.48 (3.07–3.89)
|
|
Significance[a]
|
NS
|
[a]
|
NS
|
[a]
|
NS
|
[a]
|
NS
|
[a]
|
NS
|
[a]
|
NS
|
[a]
|
Abbreviations: LCP, locking compression plate; NHTP, notched head T-plate; NS, not
specified; ROI, regions of interest.
a Bonferroni adjusted p ≤ 0.005.
Configuration 2 had significantly greater strain than configuration 1 in torsion at
five ROI (p < 0.005), with no difference at the remaining seven regions ([Table 5]).
Table 5
Mean (95% confidence interval) plate strain (mm/mm) at ROI during torsion of the NHTP
and LCP with comparison between screw configurations 1 and 2
|
NHTP
|
ROI
|
|
1
|
2
|
3
|
4
|
5
|
6
|
7
|
8
|
9
|
10
|
11
|
12
|
|
Configuration 1 (short working length)
|
2.74
(2.43–3.04)
|
12.52
(12.43–13.04)
|
3.29
(3.05–3.52)
|
13.67
(12.31–15.02)
|
3.16
(2.96–3.36)
|
17.61
(15.86–19.35)
|
3.59
(3.06–4.13)
|
16.95
(14.80–19.11)
|
7.96
(7.06–8.86)
|
21.32
(19.33–23.31)
|
5.92
(5.57–6.26)
|
12.40
(9.48–15.32)
|
|
Configuration 2 (long working length)
|
2.84
(2.36–3.33)
|
18.20
(14.31–22.08)
|
7.42
(6.69–8.15)
|
30.14
(28.03–32.24)
|
9.37
(8.64–10.11)
|
30.59
(29.83–31.35)
|
9.23
(8.37–10.08)
|
26.25
(20.75–31.74)
|
11.29 (10.84–11.75)
|
22.71
(14.23–31.18)
|
9.60
(9.10–10.09)
|
16.96
(12.80–21.12)
|
|
Significance[a]
|
NS
|
NS
|
NS
|
[a]
|
[a]
|
[a]
|
NS
|
[a]
|
NS
|
NS
|
NS
|
[a]
|
|
LCP
|
|
Configuration 1 (short working length)
|
3.52
(3.13–3.91)
|
11.90
(10.48–13.32)
|
3.77
(3.32–4.23)
|
15.36
(13.40–17.33)
|
5.16
(4.91–5.41)
|
14.93
(13.00–16.86)
|
5.93
(5.58–6.28)
|
21.45
(18.10–24.79)
|
7.57
(7.09–8.05)
|
18.25
(14.25–22.26)
|
6.59
(6.03–7.16)
|
17.37
(15.76–18.97)
|
|
Configuration 2 (long working length)
|
5.46
(4.87–6.06)
|
21.20
(17.97–24.42)
|
7.08
(6.69–7.46)
|
23.81
(18.56–29.05)
|
12.14
(10.82–13.46)
|
20.03
(14.26–25.80)
|
11.63
(11.15–12.11)
|
26.37
(20.41–32.33)
|
12.81
(12.04–13.58)
|
19.09
(15.98–22.21)
|
12.13
(11.59–12.68)
|
19.13
(14.93–24.44)
|
|
Significance[a]
|
NS
|
[a]
|
NS
|
[a]
|
[a]
|
NS
|
[a]
|
NS
|
NS
|
NS
|
NS
|
NS
|
Abbreviations: LCP, locking compression plate; NHTP, notched head T-plate; NS, not
specified; ROI, regions of interest.
a Bonferroni adjusted p ≤ 0.0005.
Comparison of Screw Configuration 1 and 2: Straight LCP
Configuration 1 was significantly stiffer in compression bending and in torsion (p ≤ 0.05, [Table 3]), with no difference in perpendicular, and tension bending.
Configuration 2 had significantly greater strain than configuration 1 during compression
bending at six ROI (p < 0.005), with no difference at the remaining six regions ([Table 4]).
Configuration 2 had significantly greater strain than configuration 1 during torsion
at four ROI (p < 0.005), with no difference at the remaining eight regions ([Table 5]).
Discussion
The results of this study confirmed our hypothesis that the NHTP construct would be
less stiff than the LCP construct in bending. Our hypothesis that the NHTP would be
stiffer in torsion was also confirmed. The greater stiffness in bending of the LCP
plate was reflected in the plate strain with the LCP having lower strain than the
NHTP. As hypothesized, the NHTP had significantly greater strain in bending than the
LCP. Our hypothesis that the NHTP would have lower plate strain than the LCP under
torsional loading was not supported. Despite the NHTP having greater stiffness under
torsional load than the LCP, the NHTP had higher plate strain than the LCP.
The placement of three locking screw in the short fragment in the NHTP did not overcome
the presumed lower stiffness of the smaller shaft of that plate, compared with two
axially positioned screws in the LCP with larger shaft dimensions. Factors that can
influence the bending resistance of a plate construct under identical load conditions
are the presence of a fracture gap, the modulus of elasticity of the plate material,
the cross-sectional dimensions of the plate, the plate stand-off distance, the number
of screws per fracture fragment, the working leverage of the screws in each fragment
created by the distance between the end of fragment screws and the working length.[6]
[13]
[18] Cross-sectional plate dimension is one of the key determinants of the resistance
of a plate to bending loads, which is calculated as Area Moment of Inertia (AMI),
and torsion loads, calculated as Polar Moment of Inertia (PMI). In our model, fracture
compression was observed to be maintained in tension and perpendicular bending though
not in compression bending or in torsion. Both the plates are 316L stainless steel
and were applied with no stand-off from the bone and with identical working lengths.
So, the relevant factors of influence in this model are the plate design/AMI and PMI,
the number of screws per fracture fragment and the working leverage of the screws.
The NHTP is a different shape with different dimension to the LCP. The LCP has a larger
AMI and PMI than the NHTP and so would be expected to have greater resistance to bending,
eccentric axial loading and torsion. We were particularly interested in the effect
on construct stability of being able to place three locking screws in the head section
of the NHTP in a very short juxta-articular segment compared with one locking screw
and one cortical screw in the LCP and whether this would overcome the effect of different
plate dimensions.
Stoffel and colleagues[6] showed that for the same working length, three screws per fracture fragment provided
significantly increased resistance to eccentric axial compression when compared with
two screws per fragment in a fracture gap model stabilized with 4.5 mm titanium LCP.
Conversely, Pearson and colleagues[7] using 3.5 mm straight LCP in a synthetic fracture gap model showed that for the
same plate working length, three screws per fracture fragment did not significantly
increase construct stiffness in tension bending compared with two screws; however,
they did significantly increase axial stiffness and perpendicular bending stiffness.
In our study, the three screws in the head of NTHP plate in the short fragment have
the same axial working leverage as the two screws in the LCP distal fragment ([Fig. 1A] and [B]). Working leverage is the distance between the inner and outermost screws of a fracture
fragment.[18] This may have contributed to the apparent lack of benefit of three screws compared
with two screws in resisting bending. This was not the case for torsional load resistance.
The increased torsional resistance of the NHTP over the LCP can likely be attributed
to the increased number of screws in the distal fragment and the abaxial screw position,
creating a longer perpendicular leverage arm than the axially positioned LCP screws.
The longer perpendicular working leverage likely mitigated the reduced PMI of the
NHTP plate due to the smaller shaft cross-sectional area than the LCP plate. This
finding was supported by Stoffel and colleagues,[6] who found that stiffness under torsional load in a synthetic fracture gap bone model,
stabilized with a standard LCP, increased significantly with more screws (up to four)
per fragment. Similarly, in a study by Freeman and colleagues,[24] torsional stiffness of hybrid constructs was most affected by the number of screws,
where mean stiffness increased at least 33% with four screws in each fracture fragment
versus three per fragment.
Our study cannot discern the effect of screw type on increased torsional resistance
of the NHTP, since our screw selection for each plate type was fixed. Previous studies
testing a variety of different locking devices have suggested locking constructs have
advantages over unlocked screws and plates.[25]
[26]
[27] Freeman and colleagues[24] showed that replacement of three unlocked screws with locked screws significantly
increased the torsional stiffness of the construct by 24%. This finding was supported
by Gordon and colleagues,[26] who investigated the effect of combinations of locking and cortical screws on the
torsional properties of locking-plate constructs. They found that the LCP construct
was significantly stiffer in torsion than the non-LCP construct, with a mean 17% increase
in torsional resistance after addition of a locking screw to a non-locking construct.[26]
Under identical load, plate strain was greater along the NHTP than the LCP at five
ROI in compression bending. This is not surprising, biomechanical studies have shown
a stiffer construct will have reduced plate strain.[15]
[20]
[28]
[29] The plate strain was greater along the NHTP than the LCP at two ROI when loaded
in torsion. This was surprising given the NHTP was stiffer during torsional loading
than the LCP. The reason for this is not clear; however, it may reflect an interaction
between the reduced PMI of the smaller cross-sectional area of the NHTP and the increased
perpendicular screw leverage arm of the abaxial positioned head screws.
Our second hypothesis, that for both plate types, the long working length constructs
would be less stiff and have higher plate strain than the short working length constructs
in bending and in torsion was confirmed with the exception of perpendicular and tension
bending for the LCP. The working length was altered in this study, by maintaining
the same number and type of screws in each fracture fragment and only changing the
screw location in the proximal fragment. This avoided confounding the effect of changing
the working length with a concurrent change in number of screws.
In a compressed, reduced fracture model, load-sharing is maintained under tension
and perpendicular bending; however, load-sharing is not maintained under compression
bending and under torsional loading. During both compression bending and torsional
loading, loss of contact or gapping of the Delrin bone model occurs and consequently
the working length becomes the distance between the innermost screws closest to the
fracture line and load is resisted by the implant alone.[30] Increasing the working length by leaving three screw holes vacant adjacent to the
fracture line in the proximal fragment reduced mean compression bending stiffness
for the NHTP (36–16.5 N/mm) and LCP (48.3–27 N/mm), by ∼54 and 44% respectively and
mean torsional stiffness for the NHTP (0.8–0.63 Nm/degree) and LCP (0.68 to 0.52 Nm/degree)
by ∼21 and 22% respectively. These findings are consistent with other studies that
demonstrated greater construct stiffness with shorter working lengths.[6]
[7]
[8]
[9]
[10]
[11]
[31]
Plate strain was determined at twelve ROI using digital image correlation for both
plate types under compression bending and torsion. At no ROI, mean plate strain was
reduced by increasing working length under either load direction. This is not surprising,
since both compression bending and torsion produce gapping of the fracture and loss
of load-sharing which immediately transfers load resistance to the implants alone,
resulting in reduced construct stiffness. Previous biomechanical studies have shown
that plate strain is inversely related to construct stiffness.[15]
[20]
[28]
[29]
[32]
[33]
Despite these previous biomechanical studies, the relationship between plate strain
and working length remains controversial in the literature.[11]
[12]
[14]
[34] This controversy is generated from work using a non-compressed 1 mm fracture gap
model where plate stress was calculated in an analytical fine element analysis model
loaded under axial compression producing tension bending.[6] In that modelling, increasing the working length paradoxically reduced plate strain
that was explained by early interfragmentary contact across the transcortical gap
creating load-sharing and thereby increased construct stiffness and decreased plate
strain. While the authors acknowledged this confounding factor, we believe that these
results have been misinterpreted in the literature.[11]
[13] We believe the fracture conditions modelled by Stoffel and colleagues[6] are unlikely to be realistic for a clinical situation. Eccentric axial loading producing
only tension bending in a small gap, transverse fracture without concurrent torsional
and compressive loads would be an unlikely clinical scenario. Simple transverse fractures
are typically managed with compression to achieve absolute stability. Furthermore,
the transcortical bone contact modelled in that study would produce high interfragmentary
strain, initiating bone resorption at the fracture gap.[35]
[36]
[37]
[38] The consequent gap widening, necessary to reduce interfragmentary strain to levels
where fibrous tissue could begin to form, could be expected to prolong fracture healing
and promote further plate deformation. In such a situation, successful healing could
only occur if fracture biology in that case was powerful enough to produce a healing
time that was within the fatigue life and capacity of the implant.[39]
[40]
Conclusion
In this experimental model of a compressed simple transverse fracture with a juxta-articular
9 mm distal fragment, a 2.0 mm LCP with two hybrid screws in the short fragment was
significantly stiffer than a 2.0 mm NHTP with three locking screws in the short fragment
in compression, tension and perpendicular bending but not torsion. Extending the working
length of each construct by omitting locking screws adjacent to the fracture significantly
reduced construct stiffness and increased plate strain.
