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This study developed a new titanium cervical spine expandable corpectomy cage (ECC) that consisted of two superior/inferior patient-specific endplate-conformed contact surfaces (PECSs) fabricated through 3D printing and a standardized centrally expandable mechanism (CEM) manufactured using traditional machining. Fatigue testing was conducted to evaluate whether the components of ECC could be assembled using different manufacturing methods to meet the functional testing requirements in compliance with FDA regulations. The titanium 3D-printed superior/inferior PECSs were designed based on the CT images/CAD system to assembly with a CNC milling/precision wire cutting cuboid CEM (14 mm square cross-section) expansion driven by a set of sliders interlocked via key and keyway mechanisms for the ECC of two-segmental vertebral bodies. The ECC underwent FDA-compliant static/dynamic compression, shear, torsion, subsidence tests and finite element (FE) analysis for understanding the stress differences on endplates between PECS and flat-shaped endplate contact surface (FECS). Stiffness/yield strength were found to be 2127 ± 146 N/mm/8547 ± 1213 N for the compression test, 1158 ± 139 N/mm/ 2561 ± 114 N for the shear test, and 1.22 ± 0.32 Nm/deg/16.5 ± 0.58 Nm for the torque test. Some failure samples showed that fixation screws were fractured/loosened under the shear test. The endure limits were 1600 N, 350 N and 1.0 N*m under compression, shear and torsion cyclic load tests, respectively. The results of the static axial compression and static/dynamic compression-shear testing did not meet the acceptance criteria of ISO 23,089. The PECS with higher stiffness value showed that better subsidence resistance was achieved than FECS and was consistent with that the maximum stress value and distribution for the FECS were more harmful than those for the PECS models. This study demonstrated that mechanical testing is essential when assembling 3D-printed components with CNC-manufactured parts in multi-component medical implants, as mismatches in interface behavior may result in mechanical failure. The results of FE analysis and subsidence tests indicated that PECS can decrease stress concentration between the ECC and endplates to present better performance for subsidence resistance.
Introduction
Anterior cervical corpectomy and fusion (ACCF) is one of the most effective approaches for treating various cervical disorders, such as degenerative spinal diseases, trauma, primary or metastatic tumors, spondylitis, posterior longitudinal ligament ossification, and anterior spinal compression [1, 2, 3–4]. However, anterior column reconstruction after corpectomy presents challenges, often requiring the use of autografts, allografts, or vertebral implant devices. The titanium mesh cage (TMC) and expandable corpectomy cage (ECC) are currently the most common clinical used titanium alloy implant devices for ACCF [1, 2, 5, 6–7].
The TMC requires trimming the intervertebral cage height to match the patient’s intervertebral height needs. The hollow device interior allows for backfilling with local bone grafting, enhancing bone fusion and structural stability [6, 8]. Also, TMCs lack height adjustability and limited contact area with bone may result in a stress concentration that causes instability and increases risks for tilting and subsidence [2]. The ECC usually included superior and inferior curved/plat surfaces in contact with the patient’s endplate and an intermediate height-adjustable mechanism offered adjustable height/correction for deformities, providing solid anterior support, and accommodating differences in vertebral heights, thus offering excellent initial stability [7]. Compared to the TMC, ECCs also reduce the risk for damage to the vertebral endplates. However, endplate flattening during surgery is needed to match the planar design of the typical ECC and flat superior/inferior endplate contact surfaces may increase stress concentration points to elevate the risk for subsidence. Therefore, a cage with patient-specific endplate-conformed contact surface (PECS) was proposed that can decrease endplate interface stress, increase cervical stability in flexion/extension movements and protect adjacent segments by reducing intra-disc stress and facet loading [9, 10].
Although metal additive manufacturing (AM)(also noted 3D-printing) techniques are well established for building complicated 3D constructions from computer-aided design (CAD) models [11]. Integrating of computed tomography (CT) image reconstruction, CAD and metal 3D-printing enabled the effective fabrication of PECS with intricate, tailored geometries for clinical applications. However, a suitable ECC design not only should have patient-specific PECS but needs to focus on maintaining height adjustability while ensuring simple surgical procedures with better bone grafting to enhance bone fusion.
A suitable ECC design must not only maintain a high degree of adjustability, but also simplify the cooperation with surgical instruments as much as possible. Additionally, it is preferable to retain a substantial internal space to enhance the space available for artificial bone filling and promote bone fusion. However, utilizing metal 3D printing manufacturing methods may present challenges related to assembly and component fitting due to ECCs typically feature complex structural designs comprising multiple components. Considering the fixed height adjustment specifications, it may be an economically feasible method to assemble the centrally expandable mechanism (CEM) in the middle of the ECC using traditional processing and manufacturing, but this approach must first address the mechanical issues arising after the assembly of the upper and lower 3D-printed PECS.
A new ECC consisting of a traditional manufactured CEM and metal patient-specific 3D-printed superior/inferior PECS was developed. This study attempts to evaluate the mechanical feasibility of the ECC assembled with traditionally processed CEM and titanium 3D-printed PECS under in vitro static/dynamic/subsidence mechanical testing ensuring compliance with FDA regulations. Additionally, finite element (FE) analysis was also employed to verify the ECC contact stress distribution using PECS and flat-shaped endplate contact surface (FECS) in the superior/inferior regions.
Materials and methods
PECS design and 3D-printing manufacture
One patient was selected to accomplish cervical spine computed tomography scanning (CT; Gold Seal Optimal CT660 scanner, GE Healthcare, Chicago, IL, USA) with 0.625 mm interval to maintain clearer image quality for easy cervical spine endplate morphology measurements. The patient’s image was approved by the Institutional Review Board (IRB) of Tri-Service General Hospital Songshan Branch (Protocol Title: Morphological analysis of Cervical spine based on computer tomography scan; IRB Number: C202405107). Contour extraction of different anatomical features and image reconstruction were performed in reverse engineering software (Geomagic Design, Geomagic Inc., Morrisville, North Carolina, USA). Considering that the common locations for two-segmental cervical vertebral bodies of C4 and C5 were replacement. A superior/inferior PECS solid model of the ECC was obtained to perform Boolean operations between C3/C6 images using a cube 14 mm in size. The inferior C3 endplate surfaces and the superior C6 endplate were extracted as the PECS superior and inferior contour surfaces, respectively (Fig. 1(a)). Four screw holes and countersink screw holes were designed at the superior/inferior PECS that can be matched with the fixation screw positions on the CEM surfaces to fix the PECS to the CEM (Fig. 1(b)). An additional cross slot was cut into the middle area of the superior/inferior PECS to provide space for bone graft filling and facilitate bone graft material passing through into the expandable mechanism. Vertical anterior side wings were designed at the anterior position of the superior/inferior PECS to facilitate compatibility with experimental clamps in the following static/dynamic functional mechanical tests (Fig. 1(b)).
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Fig. 1
(a) PECS illustration of the generation process including CT image reconstruction and CAD Boolean operation with a 14 mm cube block. The ECC is shown assembled by CEM and PECSs. (b) Detailed design including countersink and screw holes on the PECS surface and potential side wall design for further clinical application
The PECSs were fabricated using a metal 3D printer (AM400, Renishaw, Gloucestershire, UK) using titanium alloy powder (Ti6Al4V powder with average grain size of 30 μm) (Fig. 2(a)) [11]. The 3D printing machine was operated with a laser power of 400 W, a scanning rate of 0.6 m/s, and an exposure time of 125 s. Post-processing of 3D-printed PECS components included acid etching to eliminate residual powder and ultrasonic cleaning. No additional thermal or mechanical finishing processes were applied in this study (Fig. 2(a)) [11].
[See PDF for image]
Fig. 2
(a) Superior/inferior PECSs manufactured using metal 3D printing; (b) CEM manufactured using CNC milling and precision wire cutting; (c) The ECC assembled by CEM and PECSs (left: expanded and right: not expanded)
Design and manufacture of ECC expandable mechanism and instrument
In the adjustable CEM of the ECC, a two-segment expandable vertebral body fusion cage (consisting of two vertebrae bodies and three intervertebral discs) was designed in this study. The cage dimensions were set to be 14 mm in length, 14 mm in width, and 25 mm in height, with an expandable height of 13 mm after clinical consideration and comparison with commercially available products. The mechanism was comprised of an outer frame, a set of sliders, and a central driven screw. The set of sliders consist of upper, middle, and lower parts, with the middle slider engaged between the upper and lower sliders using a key and keyway. The driven screw passes through the outer frame and is secured at the bottom of the frame by a pin. The screw can be rotated to drive the middle slider to move back and forth. The engaged upper and lower sliders in turn outward from the frame to expand the entire cage when the middle slider is moved forward and back. The expansion mechanism can return to its original state when the screw is driven in the anti-rotation direction (Fig. 3).
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Fig. 3
(a) ECC structure illustration not yet expanded, left: cross-sectional structure and right: ECC dimension and instrument driven; (b) ECC structure illustration expanded, left: cross-sectional structure and right: 13 mm of expansion range for the ECC and instrument driven
The CEM device was fabricated using Ti6Al4V for all components and instruments. The manufacturing processes involved CNC machine tools and electrical discharge machining. Slider assembly and screw thread high precision manufacturing was necessary to ensure that the mechanism operates smoothly without mechanical binding (Fig. 2(b)).
Expandable cage assembly and mechanical tests
The superior/inferior 3D-printed PECS was aligned with the top/bottom traditional machining CEM surface and secured using 4 screws with a diameter of 2.5 mm and a length of 3 mm (Fig. 2(c)). This expandable ECC after assembly was then expanded to a maximum height of 38 mm using a screwdriver, serving as the worst-case scenario for functional testing. Static/dynamic tests were conducted according to the setup specifications outlined in ASTM F2077-22 [12]. According to the regulations, the ECC must undergo static/dynamic axial compression tests, static/dynamic compression shear tests, and static/dynamic torsion tests. The testing machine used in this study was the INSTRON 8874 (Illinois Tool Works Inc., America). In the static tests, the clamps were made using stainless steel, axial compression/compression shear (specimen is inclined at 27°) tests were performed at a rate of 5 mm/min until the ECC exhibited obviously rupture/failure or a maximum force decrease exceeding 20% (Fig. 4(a) and (b)). In the static torsion tests, a preload of 100 N was applied and the ECC was twisted at a rate of 60°/min until rupture/damage occurred or the maximum torque decreases by more than 20% (Fig. 4(c)). Force (torque)-displacement (angle) curves were recorded for all three static tests to measure ECC stiffness (N/mm) and yield strength (yield load, N) (yield torque, Nm).
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Fig. 4
Biomechanical test setup for (a) Axial compression; (b) Compression shear and (c) Torsion
In the dynamic tests, the metal clamps were replaced with ABS polymer material according to the regulations. In the axial compression/compression shear dynamic tests, a frequency of 5 Hz was applied during testing until the sample passage 5 million cycles with no damage and the load range value recorded. This was noted as the endure limit. In the dynamic torsion tests a preload of 100 N was applied and followed with a frequency of 5 Hz for torsional torque until the sample passed 5 million cycles with no damage cycles. This was recorded as the torque endure limit.
The subsidence test evaluated the ECC resistance to subsidence with the PECS and FECS by conformity with the ASTM F2267-22 standard testing method [13]. The superior/inferior FECS samples were modified with the PECS to have a flat-shaped contact surface and manufactured using CNC milling for ECC assembly (Fig. 5(a)). The test was also conducted using an INSTRON 8874, with the setup method identical to the previous static axial compression test. However, artificial bone clamping blocks (15 pcf, manufactured by Sawbones, 0.24 g/cc) and medium carbon steel clamping blocks were used for measuring Ks and Kd, according to the standards, respectively (Fig. 5(b)). The Ks and Kd were defined as the stiffness (unit: N/mm) in force-displacement curves during ECC testing using artificial bone and medium carbon steel clamping blocks. The final stiffness Kp presented the resistant capability to subsidence, calculated according to the following formula.
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Fig. 5
(a) FECS design concept modified from PECS; (b) Illustrations of subsidence test for left: used carbon steel clamping block and right: used Sawbone clamping block
1
PECS and FECS FE analysis
Previous cervical reconstructed CT image model was selected to reconstruct a FE model model from C3 to C6. This reconstruction model included cortical bone, cancellous bone, and endplates but was simplified to remove C4 and C5 bodies and all relative ligaments because the propose was only to compare the endplate stress distribution for the ECC used with PECS and FECS.
The ECC used with PECS and FECS were implanted at the C4 and C5 positions according to the ACCF approach to perform the FE simulations (Fig. 6(a) and (b)). The FE model was generated with quadratic ten-node tetrahedral structural solid elements and a total of 203,620/192,586 elements and 351,144/330,258 nodes for PECS/FECS in the FE package (ANSYS Workbench v18.2, ANSYS Inc., PA, USA) for simulation (Fig. 6(c) and (d)). Cortical, cancellous bones, endplate, and Ti6Al4V ECC were defined with linear elastic and isotropic properties adopted from the relevant literature (Table 1) [11, 14].
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Fig. 6
FE analysis for (a) ECC solid models, left: with using PECS and right with using FECS; (b) FE mesh models and loading/boundary conditions for left: with using PECS and rught: with using FECS
Table 1. Materials properties used in FE analysis
Material | Young’s modulus (MPa) | Poisson’s ratio | Reference |
|---|---|---|---|
Cortical bone | 12,000 | 0.3 | [11, 14] |
Cancellous | 100 | 0.2 | |
Endplate | 24 | 0.25 | |
Ti6Al4V | 110,000 | 0.3 |
Nodes on the C6 inferior endplate bottom were constrained in all directions as the boundary conditions. Besides applying axial loads of 50 N on the C3 superior endplate, extension, lateral bending and axial rotation with 2 N-m, 1 N-m and 1 N-m were individually applied as the three different load conditions (Fig. 6(c) and (d)) [15, 16]. Contact elements with friction (friction coefficient as 0.2) were used to simulate the PECS/FECS at the cage/endplate interfaces. Maximum von Mises stresses and stress distribution on the C3 interior endplate and C6 superior endplate were recorded for comparison to understand the mechanical responses between different contact surfaces.
Results
According to ASTM F2077 standard, static axial compression, compression shear and torsion tests were conducted with a sample size of three per group. Force-displacement curves for axial compression/compression shear tests and torque-angle curves for torsion tests were obtained (Fig. 7). Stiffness and yield strength were calculated and it was found that the average stiffness and yield strength were 2127 ± 146 N/mm and 8547 ± 1213 N for axial compression test, 1158 ± 139 N/mm and 2561 ± 114 N for shear test, and 1.22 ± 0.32 Nm/deg and 16.5 ± 0.58 Nm for torque test (Table 2). In the sample failure observations, the central screw deformed due to the outer frame generating concave deformation with no large damage under axial compression testing. (Fig. 8(a)). No obvious damage was found to the expansion structure/endplate contact surface but the fixation screws were fractured or pulled out under shear compression testing (Fig. 8(b)). In static torque testing, outer frame showed light torsional deformation and some fixation screws loosening at the CEM/PECS contact surface (Fig. 8(c)).
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Fig. 7
(a) Load-displacement curve under axial compression test; (b) Load-displacement curve under compression shear test; (c) Torque-degree curve under torsion test
Table 2. The biomechanical in vitro test result of the yielding load and stiffness of our ECC and the ISO 23,089 standard acceptance criteria under compression, compression-shear, torsion and subsidence test
Test mode | Comparative parameter | ISO 23,089 | This study |
|---|---|---|---|
Static axial compression | Stiffness (N/mm) | 5097 | 2127 ± 146 |
Yield load (N) | 5450 | 8547 ± 1213 | |
Static compression-shear | Stiffness (N/mm) | 1492 | 1158 ± 139 |
Yield load (N) | 1464 | 2561 ± 114 | |
Static torsion | Stiffness (Nm/degree) | 0.3 | 1.22 ± 0.32 |
Yield load (Nm) | 3.1 | 16.5 ± 0.58 | |
Dynamic axial compression | Runout load (N) | 1500 | 1600 |
Dynamic compression-shear | Runout load (N) | 679 | 350 |
Dynamic torsion | Runout torque (Nm) | ± 1.0 | ± 1.0 |
Subsidence | Block stiffness, Kp (N/mm) | 257 | 546.71 ± 93.7 |
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Fig. 8
Sample failure modes under static loads for (a) The central screw deformed due to the outer frame generated concave deformation with no large damage under axial compression test; (b) No obvious damage was found to the expansion structure/endplate contact surface but the fixation screws were fractured or pulled out under shear compression test; (c) Outer frame showed light torsional deformation and some fixation screws loosening at the CEM/PECS contact surface
In dynamic axial compression and shear compression testing, it was found that dynamic cyclic loads ranging from 160 N to 1600 N and 35 N to 350 N respectively could withstand 5 million cycles of testing. Similarly, a fatigue torque of 1.0 N*m endured 5 million cycles of dynamic torque loading (Table 2). No significant damage was observed after completing 5 million cycles of dynamic axial compression and torsion testing. However, damage occurred at the PECS side wing plate, a fractured set of slider structures were found, caused by loosening of the fixing screws under shear load (Fig. 9).
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Fig. 9
Sample failure mode under dynamic shear loads with damage occurring at the PECS side wing plate, a set of fractured slider structures were found, caused by the loosening of fixing screws
The subsidence test result showed that the Ks and Kd values were 137.77 ± 5.54 N/mm and 2024.63 ± 476.51 N/mm for FECS, respectively, and corresponding values for PECS are 455.65 ± 62.41 N/mm and 2852.27 ± 287.37 N/mm. Therefore, the calculated Kp values using question 1 were 148.38 ± 4.02 N/mm for FECS and 546.71 ± 93.7 N/mm for PECS and indicated that the subsidence resistance of PECS was significantly better than that of FECS (Table 2).
The biomechanical FE analysis result showed that the maximum stress values at the C3 inferior and C6 superior endplates under torsion, extension and lateral bending for the FECS implantation model were all higher than those for the PECS models (Fig. 10). The stress values on the C6 superior endplate were also significantly higher than those on the C3 inferior endplate for all different loading conditions. Figure 10 showed the stress distribution and ECC/endplate contact positions at the C3 inferior and C6 superior interfaces under all load conditions for the FECS and PECS cage implantations.
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Fig. 10
Status of stress distribution and contact areas for ECC used with PECS and FECS under all load condition simulations. Red asterisk at the figures of stress distributions indicated the locations of maximum stress values. Black circle area at the figures of contact area showed the possibility of superior/inferior heavy contact areas of the EEC cage under all loads
Discussion
In ACCF treatment, ECC provides primary stability and deformity correction, anterior support against axial loads, and good contact surface with adjacent vertebral endplates to promote early fusion [2]. Therefore, it is more favored than autologous bone grafts with fewer complications [1, 2]. Actually, the ECC invention represents a step towards personalized medicine since vertebral heights is patient-specific. However, it is regrettable that commercially available ECCs do not have customized contact surfaces with endplates; most of them have flat or fixed-angle plates. This lack of good fitness between the ECC and endplates might lead to improper stress concentration and resulting in ECC subsidence or vertebral endplate fracture.
In order to accommodate the individual anatomical variations among patients, there is a necessity to enhance the current ECC to enable height expansion and also achieve true customization in conforming to the endplate curvature. However, achieving precise fitting surfaces for the ECC and the superior/inferior endplates through traditional mechanical machining poses challenges in terms of cost-effectiveness and accuracy within a designated timeframe. Recent advancements in technology, particularly the integration of imaging processing, CAD, and metal 3D printing, offer promising solutions for creating patient-specific metal implants tailored to each patient’s anatomical structure and surgical requirements [11, 17, 18]. This approach ensures optimal support and fitting, thereby facilitating the bone healing process.
Consequently, this study proposed dividing ECC design and manufacturing into two parts. The CEM can be standardized and produced using traditional machining techniques, while the surfaces interfacing with the endplate can be fabricated using integrated imaging process/CAD/3D printing technologies to create superior/inferior PECS components. These two components can then be assembled with the CEM of the ECC.
However, it is essential to conduct a cervical computed tomography (CT) scan first to capture the endplate surface adjacent to the intended replacement vertebral bodies when applying our designed ECC for individual patient treatment. PECS components are then meticulously designed, manufactured, and can be assembled with the CEM. These preliminary technical efforts may be critical factors, as skilled techniques are essential to effectively implement the design and 3D printing manufacture of the proposed PECS in this study.
To simplify the surgical instrumentation and reduce surgical steps, the CEM design incorporates a sliding mechanism, that relies on the central screw rotation to move the central slider forward and backward. The vertical movement of the upper and lower sliders is facilitated by the central slider through a key and keyway mechanism that enables mechanism expansion. The surgical instrument was simplified with an external concave clamp head and internal screwdriver in a single instrument. The instrument external clamp head can be securely attached to the front groove of the cage to place the cage in its designated location. The internal screwdriver can be connected to the central driven screw of the cage to expand the cage (Fig. 1). Apart from CNC milling for fabricating the outer frame and central screw, precision wire cutting technology was employed for slider machining to mitigate potential mechanical binding resulting from manufacturing discrepancies in the key and keyway mechanism. The smooth interplay between the three sliders is paramount for effective CEM expansion.
In order to comply with the FDA’s regulatory standards for medical device marketing and clinical use, the CEM and PECSs included in the ECC were made using biocompatible Ti6Al4V as materials. The most important thing was that the functional performance met certain standards. We used the mechanical standards provided by ISO23089 to judge whether the ECC functionality met the standards because it was unavailable to perform substantial equivalence at the current stage [19]. The ISO23089 standard delineated the mechanical assessment requisites for intervertebral body fusion devices (IBFDs) employed in spinal fusion procedures, thereby establishing a baseline for performance testing throughout IBFD development. Table 2 delineated the 5th percentile values for IBFDs of static/dynamic axial compression, compression shear, torsion, and subsidence. These data were distilled from IBFD mechanical tests previously cleared by the FDA through the 510(k) process. Specifically, they represented outcomes from mechanical testing of single-piece metallic designs and did not encompass testing data for IBFDs featuring integrated fixation or expandable functionalities. The ECC we developed for ACCF treatment was much larger than the IBFD in height. The bending moment it endured also increased, and the mechanism was more complex. Therefore, it would have been a higher robust benchmark for our ECC if the standard of ISO23089 standards had been used.
Since the entire ECC is assembled from PECS made through a 3D-printing process and CEM components made using traditional CNC methods, the tests revealed that the stiffness in static axial compression and compression-shear tests, as well as the dynamic runout load (endure limit), did not meet regulatory standards based on the results of various tests. This phenomenon was also observed in the failure modes of the static torsion tests, where ineffective screw fixation led to relative micromotion (torsion) between the PECS and CEM. In the dynamic shear tests, screw loosening caused the applied force to shift to the PECS side walls, leading to their fracture. Additionally, in one sample, uneven internal stress distribution in the CEM after screw loosening resulted in the lower slider fracture. The causes of these failures were related to the manufacturing process and structural design. For example, the concave deformation of the CEM outer frame hindered the rotation of the device screws, reducing the stiffness in the static axial compression tests. Future designs may benefit from increasing the outer frame thickness or adopting alternative 3D printing techniques—such as electron beam melting (EBM), which utilizes different powder size distributions and energy input—to enhance structural strength, improve interfacial compatibility, and reduce performance variability. Otherwise, post-processing treatments such as hot isostatic pressing, surface polishing, or optimized build orientation may mitigate stress concentrations and enhance bonding strength at assembled interfaces. These strategies will be considered in future iterations of implant design.
A critical issue identified was the incompatibility between 3D-printed and CNC-machined components when assembled, as there were problems with the transmission of compressive/shear force, and torque at the interfaces. This was mainly because the components were not a single material, making it ineffective to transmit stress and causing relative motion between the PECS and CEM prevented forces from being effectively transferred to the sliding mechanism. Therefore, it was recommended mechanical testing is essential when assembling 3D-printed components with CNC-manufactured parts in multi-component medical implants. While using 3D printing for the entire ECC could be a viable solution, the complexity of combining multiple mechanical components poses challenges. The precision and surface smoothness of the components produced by 3D printing may be crucial factors for smooth operation and mechanical transmission after assembly, which will require further investigation.
The subsidence test results revealed that the Kp value of PECS ECC was approximately 2.13 times (547Nmm/257Nmm) that of FECS, indicating that the PECS has significantly better resistance to subsidence compared to the FECS. A similar finding was observed in the FE analysis results, where the maximum von Mises stress on the superior and inferior endplates using FECS ECC was relatively high, regardless of the type of load conditions, and stress concentration was also relatively severe (the asterisk symbol represents stress concentration at contact points). Figure 10 illustrates the contact status between the two ECCs and the endplates, explaining the tendency for point contact and consequent stress concentration between FECS ECC and the endplates. On the other hand, there was a larger surface contact area between PECS ECC and the endplates, resulting in a more uniform stress distribution. These findings can be explained accordingly in Fig. 10 and were consistent with the literature [20, 21]. However, it was important to note that these values obtained from FE analysis can only be used to assess numerical trends. The current model was not entirely accurate and does not account for the relative remnants of bone and ligaments, thus the obtained data may be relatively high.
Conclusions
This study concluded that an adjustable standardized CEM with key/keyway mechanism manufactured by conventional machining was proposed and assembled to superior/inferior titanium 3D-printed PECSs to achieve an ECC device to fulfill unmet clinical needs. However, the assembled ECC could not fully pass the mechanical testing results required by FDA’s ISO 23,089 standard. This study demonstrated that mechanical testing is essential when assembling 3D-printed components with CNC-manufactured parts in multi-component medical implants, as mismatches in interface behavior may result in mechanical failure. The results of FE analysis and subsidence tests indicated that PECS can effectively disperse and avoid stress concentration between the ECC and endplates and presents better performance for subsidence resistance.
Acknowledgements
N/A.
Author contributions
Conceptualization: Shih-Chieh Shen, Chun-Li Lin; Fund raising: Chun-Li Lin; Investigation: Shih-Chieh Shen, Shao-Fu Huang, Wei-Hsiang Sun; Methodology: Shih-Chieh Shen, Shao-Fu Huang, Wei-Hsiang Sun; Formal analysis: Shao-Fu Huang, Wei-Hsiang Sun; Writing – original draft: Shih-Chieh Shen, Chun-Li Lin; Writing – review & editing: Chun-Li Lin.
Funding
This study is supported in part by NSTC project 108-2622-E-010 -001 -CC2, 110-2222-E-032 -003 -MY2 and 113-2327-B-A49 -001, Taiwan.
Data availability
No datasets were generated or analysed during the current study.
Declarations
Ethics approval and consent to participate
N/A.
Consent for publication
Not applicable.
Competing interests
The authors declare no competing interests.
Publisher’s note
Springer Nature remains neutral with regard to jurisdictional claims in published maps and institutional affiliations.
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