Novel Design of Low-Cost Composite Shell and Backfill Tool for Stamping of HSS 590 Sheet Metal

Novel Design of Low-Cost Composite Shell and Backfill Tool for Stamping of HSS 590 Sheet Metal | Research Square window.SnipcartSettings = { analytics: { enabled: false } }; (function() { var accessVector = localStorage.getItem('access_vector') || ''; window.dataLayer = window.dataLayer || []; if (accessVector) { window.dataLayer.push({ user: { profile: { profileInfo: { snid: accessVector } } } }); } })(); (function(w,d,s,l,i){w[l]=w[l]||[];w[l].push({'gtm.start':new Date().getTime(),event:'gtm.js'});var f=d.getElementsByTagName(s)[0],j=d.createElement(s),dl=l!='dataLayer'?'&l='+l:'';j.async=true;j.src='https://www.googletagmanager.com/gtm.js?id='+i+dl;f.parentNode.insertBefore(j,f);})(window,document,'script','dataLayer','GTM-K279D39R'); Browse Preprints In Review Journals COVID-19 Preprints AJE Video Bytes Research Tools Research Promotion AJE Professional Editing AJE Rubriq About Preprint Platform In Review Editorial Policies Our Team Advisory Board Help Center Sign In Submit a Preprint Cite Share Download PDF Research Article Novel Design of Low-Cost Composite Shell and Backfill Tool for Stamping of HSS 590 Sheet Metal Madhura Athale, Taejoon Park, Ryan Hahnlen, Farhang Pourboghrat This is a preprint; it has not been peer reviewed by a journal. https://doi.org/ 10.21203/rs.3.rs-4652135/v1 This work is licensed under a CC BY 4.0 License Status: Published Journal Publication published 30 Oct, 2024 Read the published version in The International Journal of Advanced Manufacturing Technology → Version 1 posted 5 You are reading this latest preprint version Abstract Sheet metal stamping uses hardened steel tools, due to their advantages in hardness, resistance to deformation, and resistance to abrasion. However, these tools have limitations when it comes to prototype production volumes, due to the high cost and time required for tool fabrication. Forming tools fabricated with polymers using additive manufacturing (AM) offer an inexpensive alternative suitable for low volume production and prototyping. For successful implementation of polymer AM tooling in sheet metal forming, tool cost, tool life, and part accuracy are important metrics. A novel composite tool design concept consisting of two distinct components – an outer polymer AM shell, and inner backfill – to make up the composite tool is proposed. Experimental and numerical investigation of stamping of high strength steel sheets with the new tool design is presented. It is shown that the new composite tool design concept improves performance and is more economical compared with fully dense or solid AM polymer tools. 3D printing Anisotropy Additive Manufacturing (AM) Finite Element Analysis (FEA) Sheet metal forming AM tooling Figures Figure 1 Figure 2 Figure 3 Figure 4 Figure 5 Figure 6 Figure 7 Figure 8 Figure 9 Figure 10 Figure 11 Figure 12 Figure 13 Figure 14 Figure 15 Figure 16 Figure 17 Figure 18 Figure 19 1. Introduction Sheet metal forming is a popular manufacturing technique used in various industries as it is an economical way of mass production. Typical mass production forming toolsets are machined from tool steel and designed to last long enough to produce large volumes of parts, sometimes up to hundreds of thousands. Thus the cost of the machined steel dies can be easily justified in this scenario [ 1 ]. However, there are certain situations where these conventional forming tools fall short in the economics aspect – such as for low or medium volume production of custom parts or prototype parts [ 2 ]. The need to manufacture inexpensive yet effective dies quickly for prototyping has paved way for the use of polymer AM technologies for forming tool production [ 3 ]. Fused Deposition Modeling (FDM) or alternatively, Fused Filament Fabrication (FFF) is the most commonly used polymer AM technology for tool fabrication. Over the last two decades, several research works have been carried out that look into various aspects of this topic. It has been experimentally demonstrated that polymer AM tooling can successfully form Aluminum and low-strength steels for small production volumes [ 4 – 6 ]. Good friction behavior of polymers makes them attractive as tooling materials [ 7 ]. Tribological properties of polymer composite tools for contact with metals were studied by Park and Colton [ 8 ] as well as Liewald and de Souza [ 9 ]. Park and Colton also investigated the failure modes of polymer composite deep drawing dies, including fracture, wear and plastic deformation [ 10 ]. Kuo and Li reported that addition of Zirconia particles improved the wear performance of epoxy resin dies [ 11 ]. Many studies have shown that FDM parts have significant anisotropy in mechanical properties [ 12 – 14 ] with the raster direction being significantly stronger than the transverse direction [ 15 ]. Several factors such as layer-wise fabrication, fiber reinforcements, internal porosity etc. may cause this anisotropy [ 16 ]. However, most numerical investigations into the use of polymer AM tooling for sheet metal forming have used simple isotropic material models for the polymer tooling [ 9 , 10 , 17 ]. The authors previously demonstrated the importance of using anisotropic material models for accurately predicting the elastic-plastic deformation of FDM CF-Nylon dies during stamping of HSS sheets using finite element analysis [ 18 ]. An important issue with polymer AM tooling that is reported in published research is their elastic and plastic deformation during forming and resultant part deviations. Nakamura et al. performed V bending of steel sheets using FDM PLA tools and reported a higher geometrical deviation in final parts as compared to steel tools [ 19 ]. Durgun performed stamping of steel sheets with FDM Polycarbonate dies and reported that MC355 parts were outside the dimensional tolerance range due to tool deformations [ 20 ]. Similarly, Tondini et al. reported higher springback in V bending with polymer tools due to their low stiffness and elastic deformation [ 21 ]. In a previously published work [ 22 ], the authors showed that stamping 100 HSS590 sheets with glass fiber reinforced Polycarbonate dies resulted in a consistent reduced draw depth of 2mm across all parts as compared to steel tooling. Finite element analysis revealed the elastic deformation of the dies to be a potential cause of this [ 22 ]. Similar observations of lower resultant draw depth of up to 1.7 mm for steel blanks in case of CF-Nylon tooling compared to steel tooling were reported by Giorleo and Ceretti [ 23 ]. In contrast to elastic deformation of dies, most of the plastic deformation takes place in the first few stamping passes and reaches a steady state value [ 17 , 18 , 21 ]. Deviation in the final part dimension when using AM polymer tooling is an important aspect that needs to be further investigated. In this study, the authors have proposed a novel “shell and backfill” die design concept to mitigate some of the issues with using fully dense or solid AM tooling for sheet metal stamping reported in the previous work [ 22 ]. A “shell and backfill” type tool consists of an outer thin shell of the stamping die or punch produced by 3D printing of polymer composite material and backfilled with a less expensive material. This is expected to achieve higher cost savings than solid AM tooling. Geueke et al. used geometry optimization of solid AM polymer tooling for cup drawing by removing material strategically from low load areas in the tool for cost reduction [ 24 ]. However, such a design approach may only be applicable to relatively simple geometries and extension of it to more complex geometries would incur high computational costs during design. On the other hand, the shell and backfill design is also highly adaptable to any tool geometry. Additionally, using a backfill material that is stiffer than the polymer shell material, limits the elastic deformation of the tools during stamping. This reduces the possibility of insufficient draw depth or other geometric deviations in final parts as seen with solid AM polymer tooling. This paper is organized as follows: In Section 2, the shell and backfill tool design concept is explained in further detail, as well as material testing and characterization information is provided. Details of the finite element models and the experimental stamping setup are also given. Section 3 is dedicated to a case study of the shell and backfill tooling including tool design, finite element analysis, experimental stamping trials and results. Finally, Section 4 provides conclusions and future work. 2. Materials and Methods 2.1. Shell and backfill tooling In the previous research, glass fiber polycarbonate (GFPC) solid tools showed only 0.25 mm deformation in the highest stress area after stamping 100 parts from 1.539 mm HSS 590 sheets [ 22 ]. This indicates that the GFPC tools could possibly be used for stamping more parts, but it also implies that their potential may not be fully leveraged in a typical prototype or low volume production scenario. Moreover, the solid GFPC die set was able to achieve 11% cost savings compared to a conventional steel tool. In the innovative “shell and backfill” type tooling approach, an outer thin shell of the stamping die or punch is produced by 3D printing of polymer composite material, which is then backfilled with a less expensive material. Figure 1 shows schematic representation of the shell and backfill tool design. As most of the die volume is made up of the inexpensive backfill material, there is potential for additional cost savings of up to 50% as compared to the conventional steel tool cost, depending on the specific backfill material used. This approach takes advantage of the most important feature of 3D printing, that is to fabricate complex shapes rapidly, by fabricating the forming surfaces of the tools, while also reducing the printing time and material as well as machine cost by removing a large volume from the 3D printed part. Backfilling operation is very fast and requires minimal operator skill. 2.2. Materials In this study, the AM shell material was carbon fiber reinforced nylon (CF-Nylon) printed with the FDM technique. FDM CF-Nylon was chosen as the die shell material based on its excellent mechanical properties and demonstrated success as a tooling material for sheet metal stamping [ 18 ]. CF-Nylon mechanical characterization was performed using coupons machined from FDM printed blocks along four unique directions to capture the anisotropy in their mechanical properties as detailed in [ 18 ]. High strength steel sheets (HSS 590) with a grade of 590 MPa and a thickness of 1.539 mm were used throughout this study, as it is a common material grade and gauge used in the automotive industry [ 1 ]. HSS 590 material properties, including Young’s modulus, Possion’s ratio, 0.2%-offset initial yield stress, and hardening curve, were obtained from uniaxial tension tests performed along the rolling direction (RD) [ 18 ]. Concrete was considered as the backfill material in this study for its high stiffness compared to polymers, and its low cost, and ease of use. Mechanical properties of concrete, including average values for Young’s modulus and Poisson’s ratio listed in Table 1 , were obtained from various online sources [ 25 ]. Compressive strength for concrete was obtained from the manufacturer’s specification. Commercially available high-strength concrete mix with a cured strength of 27.5 MPa was used for this study [ 26 ]. Mechanical properties of materials used in this study are given in Table 1 and Table 2 . Table 1 Mechanical properties of materials. Property Material Concrete HSS 590 RD * Young’s Modulus ( E ) (GPa) 24 206.5 Poisson’s ratio ( ν ) 0.21 0.31 Yield stress ( σ y ) (MPa) -- 435.2 Strength (MPa) 27.5 709.8 * Rolling direction Table 2 Averaged uniaxial compression results with deviations for CF-Nylon [ 18 ]. Property Orientation X Z XY XZ Young’s Modulus ( E ) (GPa) 2.20 ± 0.20 1.85 ± 0.08 3.95 ± 0.52 1.6 ± 0.12 Poisson’s ratio ( ν ) ν xy = 0.55 ± 0.05 ν xy = 0.14 ± 0.03 ν planar = 0.09 ± 0.01 ν planar = 0.31 ± 0.02 ν xz = 0.25 ± 0.04 ν xy = 0.12 ± 0.04 Yield stress ( σ y ) (MPa) 26.4 ± 2.51 29.2 ± 0.76 45.5 ± 1.92 25.8 ± 2.81 2.3. Numerical analysis Finite element (FE) analysis was performed using the commercial solver ABAQUS/Explicit throughout this study for stamping simulations. Tool components including the AM polymer shell and backfill were modeled as deformable bodies and meshed with linear tetrahedral elements (C3D4). HSS 590 blanks were meshed with linear quadrilateral shell elements. In each case, a steel blank holder was modeled as a rigid body. HSS 590 was modeled with von Mises yield criterion since the difference in the measured stress-strain curves was negligible in uniaxial tensile tests along 0, 45, and 90 degrees to the rolling direction. Isotropic hardening rule was assumed based on the stress-strain curve measured in the uniaxial tension test. The hardening curve was extrapolated beyond the uniform elongation limit using a power-law-type hardening equation. To account for the severe anisotropy of the CF-Nylon FDM shell, orthotropic elasticity and Hill 1948 yield criterion were assumed with the tetragonal symmetry constraints. Additionally, an isotropic hardening rule was assumed to represent the anisotropic expansion of the yield stress preserving the initial anisotropy. Material model parameters for the FDM CF-Nylon were calibrated based on experimental data as detailed in [ 18 ]. Although concrete is typically brittle under the uniaxial compression, it demonstrates brittle-to-ductile transition, increased strength, strain hardening, and increased failure strain under compressive hydrostatic pressure [ 27 , 28 ]. In other words, under confinement, when the hydrostatic pressure in the concrete is high, it exhibits higher load-carrying capacity than its nominal compressive strength. Studies have shown that applying hydrostatic pressure of 100% or 200% of the nominal compressive strength, increased the compressive strength of concrete significantly – by 200–400% [ 29 ]. In sheet metal stamping, the tools experience mostly compressive stresses. As the concrete backfill is surrounded by the polymer composite shell, the compressive stresses arising during the stamping process are unlikely to lead to catastrophic failure of the concrete. For simplification of the material modeling, the concrete backfill was assumed as isotropic elastic – perfectly plastic material. As the perfect plasticity model does not consider the increased yield stress and hardening due to the hydrostatic pressure, this concrete backfill modeling can be seen as a type of worst-case scenario assumption. In other words, if the simulation results show high hydrostatic pressure comparable to nominal stress in the critical regions, it can be expected that in reality, the tool performance will be better due to higher yield stress and hardening behavior of concrete. Additionally, depending on the confining pressure, high strains in concrete even up to 5% may not result in failure [ 27 , 29 ]. 2.4. Experimental setup The same stamping tool geometry used in the previous study with solid AM tooling was also used in this study to maintain consistency. Figure 2 shows the components of the Universal Formability tool (UFT) – punch, primary die with die inserts, and blank holder – and their dimensions. Square sheets of HSS 590 with 470 mm side length and 1.539 mm thickness were used for stamping. A “baseline” stamping trial was first performed with HSS 590 blanks using conventional steel tooling for comparing the performance of shell and backfill tools against a reference. A steel blank holder was used during all the stamping trials. All forming trials utilized a 300-ton servo press programmed with a crank slide motion. The blank holder force was maintained at 200 kN for the duration of the draw phase. Total draw depth was set at 65 mm with a speed of 14 strokes per minute. These parameters were chosen based on prior experience with the sheet metal material and relevance to mass-production manufacturing. For all trials, the HSS 590 blanks were lubricated with the same mass-production relevant water-based lubricant and selected blanks were etched for post-forming strain analysis. All process parameters were kept consistent between the steel tools and composite tools. ATOS 3D scans of the shell and backfill tools were performed before and after the stamping trial to get a quantitative measure of the dimensional change of tools due to plastic deformation, damage, and wear. ATOS and ARGUS scans of the AM-stamped parts were also performed and compared against those from steel-stamped parts to evaluate the accuracy and consistency of the tools. 3. Case study: CF-Nylon + Concrete tool 3.1. Tool design: shell thickness optimization Simulation conditions To minimize computational cost, the optimum thickness of the outer shell die was determined by performing FE simulations for cup drawing with a simple geometry shown in Fig. 3 . The conditions for cup drawing were designed to closely mirror those of UFT forming. The punch and die radii for cup drawing were selected to be similar to the tightest tool radius in the UFT die. Additionally, the 3.4 mm gap between punch and die was maintained. For the initial circular blank, the same 1.539 mm thick HSS 590 sheet was employed. As for the shell and backfill materials, FDM-produced CF-Nylon and concrete were used, respectively. To further save computational cost, a quarter model was utilized for the FE simulations. Three different shell thicknesses were considered for the optimization simulations: 2.5 mm, 5 mm, and 10 mm. Figure 4 shows the simulation model for different shell thicknesses. The contact condition at the interface between the shell and backfill has uncertainties in terms of friction and bonding. The surface of the CF-Nylon shell, produced by the FDM process, may exhibit irregularities, and the interfacial bonding can become extremely weak due to dry shrinkage and capillary water evaporation during the concrete curing process. Therefore, the simulations need to account for any possible interface condition and select the most suitable shell thickness. Two extreme scenarios were considered for the FE simulations to include all possible friction conditions: 1) smooth contact to simulate no friction or bonding/adhesion between the shell and backfill, and 2) bonded interface to simulate infinite friction or perfect bonding between the shell and backfill. The bonded interface condition was simulated using tie constraints between nodes on the surfaces of shell and backfill that are in contact with each other. With 3 different shell thicknesses and 2 different interface conditions, a total of 6 unique simulation conditions were examined. Simulation results FE simulations of cup drawing using the shell and backfill tools showed that the bonding between the shell and backfill at the interface has a significant effect on the tool performance in terms of resultant stress, strain, and plastic deformation. Critical areas in the tools such as the punch and die corner radii as well as the corner radii at the shell and backfill interface were examined for different shell thicknesses and two extreme interface conditions as described previously. Figure 5 shows a comparison of the effective plastic strain (PEEQ) on the CF-Nylon punch shell for all 6 simulation cases. The two interface conditions, namely perfect bonding and no bonding, generated reverse trends in the amount of plastic strain as a function of the shell thickness. Increasing shell thickness led to higher accumulated plastic strain near the punch corner of the CF-Nylon shell in the perfect bonding interface condition. Conversely, in the no bonding case, as the shell thickness increased, the amount of accumulated plastic strain in the CF-Nylon shell decreased. As the interface condition is uncertain, the optimal shell thickness should be chosen based on the lowest accumulated plastic strain in both extreme interface conditions. In this case, 5 mm thickness was found to be optimal. For the 5 mm shell thickness case, the no bonding interface condition was found to have maximum plastic strains of 2–3% in the shell. As forming tools mostly experience compressive loads, the higher value of 2–3% plastic strain is still below failure strain [ 18 ]. However, the effect of shell thickness and interface condition on the backfill material should be confirmed before finalizing the optimal shell thickness. Figure 6 shows the effective plastic strain in the concrete punch backfill for the 6 simulation cases. The backfill concrete demonstrated an increase in the accumulated plastic strain as the shell thickness increased in both interface conditions. Unlike the CF-Nylon shell shown in Fig. 5 , the concrete backfill exhibited higher plastic strain in the perfect bonding interface case. This is the result of a more effective transfer of load from the shell to the backfill when there is bonding between the two. In the no bonding interface condition, shear load and plastic deformation remains isolated within the shell. In general, some bonding between the two materials is desired so that the shell does not experience very high deformations which may lead to part deviations or tool failure. Considering the resultant plastic strain on both shell and backfill materials under the two interface conditions, 5 mm shell thickness was found to be optimal. However, if there is good bonding between AM material and concrete, the concrete material is likely to develop internal cracks. As the backfill is completely surrounded by the polymer shell, risk of catastrophic failure due to concrete was considered to be low. A similar analysis was performed for the cup drawing die, considering FDM Nylon shell and concrete backfill to obtain the optimized die shell thickness. Shell thickness of 5 mm was again found to be the most suitable. 3.2. Experimental stamping trial Based on the shell thickness optimization presented in the previous section, tool shells with 5 mm wall thickness were fabricated with CF-Nylon using FDM technique for the UFT forming. A track width of 0.5 mm in the build plane (X-Y) and a layer height of 0.254 mm along the build direction (Z) were used to make all tools. In this case, the punch travel is in Z direction as well. Each layer was printed using a single perimeter pass and filled with an alternating raster orientation of ± 45 degrees from the X axis. Printed tools were then backfilled with concrete and allowed to cure for the specified amount of time. Figure 7 shows the tools before and after the backfilling operation. The original steel toolset for the UFT geometry cost $ 20,500. The FDM CF-Nylon shell for the punch and die combined cost $ 7,100 and the concrete backfill cost $ 10. The CF-Nylon shell and concrete backfill toolset achieved 65% cost reduction compared to the conventional toolset. As the same steel blank holder was used for both toolsets, it was left out of the cost comparison. Stamping of HSS 590 blanks was performed with a steel blank holder using process parameters described in Section 2.4. A total of 39 parts were stamped with the shell and backfill toolset. 3.3. Numerical analysis of stamping FE simulations of the stamping of HSS 590 blanks with shell and backfill tooling were performed using the Abaqus/Explicit commercial solver. Relevant simulation conditions and material model details are given in Section 2.3. Figure 8 shows a sectional view of the Abaqus simulation model used for the FE analysis. Boundary conditions for controlling the displacement of punch and die were determined considering the physical attachment points where tools were attached to the press plates during the stamping trials indicated by red circles in Fig. 9 (a). Specifically, displacement boundary conditions, including punch displacement in Z direction and fixed displacements in X and Y directions ( \({U}_{X}={U}_{Y}=0\) ), were applied to the corresponding elements highlighted in red in Fig. 9 (b). Additionally, similar to the shell thickness optimization simulations in Section 3.1, two interface conditions corresponding to perfect bonding and no bonding were considered to evaluate the effect of the interface between shell and backfill. 3.4. Results During the experimental stamping trial with shell and backfill tools, 39 parts were formed before stopping due to failure of the tool shell that accumulated over the course of the trial as shown in Fig. 10 . Progressively increasing wrinkling was observed in the flange region of the stamped parts, which finally caused die failure at the location of the largest wrinkle due to the penetration of the wrinkle into the polymer die shell as shown in Fig. 11 . Both the punch and die were observed to be bulging outwardly on the outer vertical walls. This indicates insufficient adhesion or bonding between the shell and backfill. Due to the outward bulging, the punch sustained visible damage on the outer walls rubbing against the blank holder during each stamping pass. ATOS 3D scans were conducted to measure the deformation of the tools after the stamping trials. Figure 12 shows a maximum bulging of 1.18 mm on the die wall, while Fig. 13 shows bulging of 1.11 mm to 4.38 mm at different locations on the punch wall. ATOS 3D scans in both figures show the comparison of the deformed tool surfaces (after stamping 39 parts) with respect to the original tool surfaces (before the stamping trial). Here the outward direction from the tool surface is assigned positive sign. In other words, positive values indicate the tool surface has expanded outward compared to the original tool surface. One of the motivations behind using the shell and backfill die design was to mitigate the issue of the shallow draw as seen with the solid AM polymer tooling [ 22 ]. By using a backfill material with a higher elastic modulus, the elastic deformation of the tools is reduced, resulting in stamping deeper parts as intended. Figure 14 shows a cross-section of the formed part after #1 trial using this tooling and a comparison with that of a steel-stamped part as obtained from ATOS 3D scanning. No draw deficit was observed with the new CF-Nylon shell and concrete backfill tooling. However, a small amount of flattening of the stamped part at the die insert nose region is observed as compared to the steel-stamped part due to the deformation of the polymer die at that location. Final part strains compared between experimental values obtained through ARGUS strain analysis and those from simulation results show good agreement, providing a basic validation of simulation results as shown in Fig. 15 . The ARGUS grid pattern was washed out from some regions on the part during stamping where strain measurements could not be obtained. Differences between simulations and experiments are attributed to different averaging methods and averaging areas used between ARGUS and simulations. Figure 16 shows the deformation of the CF-Nylon die shell walls at the end of the stamping stroke with no bonding interface condition between the shell and backfill. Like the experimental observations, the die shell walls showed outward deformation or bulging. The excessive wrinkling of stamped parts seen in the experimental trials (refer to Fig. 11 ) is thought to result from the uneven bulging of the shell and the consequent uneven distribution of blank holding force. This claim is supported by simulation results showing non-uniform deformation of the top surface of the die shell in the stamping direction. Figure 17 shows a comparison of die top surface deformation in the Z direction (stamping direction) between the two simulation cases of perfect bonding and no bonding. The perfect bonding case shows negligible deformation of the die surface near the flange area. However, the no bonding case shows highly non-uniform deformation, which could create uneven gaps in between the die and blank holder, resulting in non-uniform blank holder force in different areas of the flange, promoting wrinkling. Actual location and height of wrinkles on the formed part could not be accurately predicted with the simple material model used for the blank, as wrinkling and formability of the formed part were not the focus of this study. Additionally, the simulation was able to qualitatively predict the deformation of the die insert area, which is the highest-stress area in the tool. Figure 18 shows a comparison of experimental die scan results after 39 stamping and simulation results of the same at the end of the first stamping pass. Accurate quantitative prediction of plastic deformation depends on the material model for AM polymers which should account for various effects such as anisotropy, tension-compression asymmetry, and strain rate sensitivity. Successful qualitative and quantitative prediction of FDM CF-Nylon tool deformation was demonstrated through the use of appropriate material models previously [ 18 ]. In the case of shell and backfill die design, the plastic deformation is also dependent on the material model used for the backfill and the unloading behavior of the backfill. To assess the damage to the concrete backfill, Fig. 19 shows the hydrostatic pressure and effective plastic strain (PEEQ) in the insert region of the die at the end of the stamping simulation. The maximum hydrostatic pressure is 196 MPa, which is more than 7 times the nominal compressive strength of concrete. As mentioned in Section 2.3, when hydrostatic pressure is twice the nominal compression strength, the compressive strength can increase up to four times compared to scenarios without hydrostatic pressure. Additionally, the maximum PEEQ in the concrete backfill of the die was 2.1%, which is well within the safe zone for such high confining pressure. 4. Conclusions and future work The shell and backfill die design is a novel approach for further lowering the cost of polymer AM tooling for sheet metal forming applications. Solid AM GF-PC toolset achieved 11% cost reduction, while the new CF-Nylon + Concrete shell and backfill toolset achieved 65% lower cost compared to steel tools. Inexpensive backfill materials such as cement can achieve high cost-savings but may not be the most ideal backfill material to meet all performance criteria and recycling feasibility. The contact and bonding condition at the interface between the shell and backfill has a significant impact on the tool performance in terms of tool longevity as well as the final part geometry. The presence of bonding between the shell and backfill is essential for preventing buckling or bulging of the shell away from the backfill, which may lead to failure of the shell. Additionally, shell bulging may result in non-uniform distribution of the blank holding force on the flange region, promoting wrinkling in the part. Polymer materials for the shell often do not have sufficient bond strength with the various backfill materials under consideration. To address this issue, various mechanical interlocking mechanisms such as bolts, dovetails, etc. to attach the shell and backfill together need to be explored. This is also expected to improve part accuracy and tool life by eliminating she shell and backfill separation. Additionally, recyclable and energy-friendly materials such as metal alloys with low melting temperatures will be considered as potential backfill materials for future work. Declarations Competing Interests The authors have no relevant financial or non-financial interests to disclose. Funding This work was supported by Honda Development and Manufacturing of America, LLC (AWD-115703). Author contributions Madhura Athale : Methodology, Software, Validation, Formal analysis, Investigation, Writing – original draft, Visualization; Taejoon Park : Conceptualization, Methodology, Writing – Review & editing; Ryan Hahnlen : Supervision; Farhang Pourboghrat : Writing – review & editing, Supervision, Project administration, Funding acquisition Acknowledgements The authors wish to thank Honda Development and Manufacturing of America, LLC. for their support of this project through the award AWD-115703: An inverse design methodology to fabricate low-cost agile tools for manufacturing lightweight automotive components. References Hahnlen R, Pourboghrat F, Park T, Hoffman B, Athale M (2021) accessed October 27,. Additive Manufacturing for Sheet Metal Forming Tools. 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Int J Adv Manuf Technol 121:6973–6989. https://doi.org/10.1007/s00170-022-09801-0 Nakamura N, Mori K, Abe Y (2020) Applicability of plastic tools additively manufactured by fused deposition modelling for sheet metal forming. Int J Adv Manuf Technol 108:975–985. https://doi.org/10.1007/s00170-019-04590-5 Durgun I (2015) Sheet metal forming using FDM rapid prototype tool. Rapid Prototyp J 21:412–422. https://doi.org/10.1108/RPJ-01-2014-0003 Tondini F, Basso A, Arinbjarnar U, Nielsen CV (2021) The Performance of 3D Printed Polymer Tools in Sheet Metal Forming. Metals 11:1256. https://doi.org/10.3390/met11081256 Athale M, Park T, Hahnlen R, Pourboghrat F (2023) Design, performance, and cost savings of using GF-PC additively manufactured tooling for stamping of HSS 590 sheet metal. J Manuf Process 101:1–14. https://doi.org/10.1016/j.jmapro.2023.05.072 Giorleo L, Ceretti E (2022) Deep drawing punches produced using fused filament fabrication technology: Performance evaluation. J Manuf Process 84:1–9. https://doi.org/10.1016/j.jmapro.2022.09.054 Geueke M, Frohn-Sörensen P, Reuter J, Padavu N, Reinicke T, Engel B (2021) Structural optimization of additively manufactured polymer tools for flexible sheet metal forming. Procedia CIRP 104:1345–1350. https://doi.org/10.1016/j.procir.2021.11.226 Concrete Properties n.d. https://www.engineeringtoolbox.com/concrete-properties-d_1223.html (accessed March 21, 2023) Concrete Mix | QUIKRETE (2024) Cement and Concrete Products n.d. https://www.quikrete.com/productlines/concretemix.asp Vu XH, Malecot Y, Daudeville L, Buzaud E (2009) Experimental analysis of concrete behavior under high confinement: Effect of the saturation ratio. Int J Solids Struct 46:1105–1120. https://doi.org/10.1016/j.ijsolstr.2008.10.015 Lima VN, Silva F, de Skadsem A, Beltrán-Jiménez HJ, Sunde K (2022) Effects of confinement pressure on the mechanical behavior of an oil well cement paste. J Petrol Sci Eng 208:109769. https://doi.org/10.1016/j.petrol.2021.109769 Sfer D, Carol I, Gettu R, Etse G (2002) Study of the Behavior of Concrete under Triaxial Compression. J Eng Mech 128:156–163. https://doi.org/10.1061/(ASCE)0733-9399(2002)128 Cite Share Download PDF Status: Published Journal Publication published 30 Oct, 2024 Read the published version in The International Journal of Advanced Manufacturing Technology → Version 1 posted Editorial decision: Major Revisions Needed 18 Aug, 2024 Reviewers agreed at journal 01 Jul, 2024 Reviewers invited by journal 30 Jun, 2024 Editor assigned by journal 30 Jun, 2024 First submitted to journal 27 Jun, 2024 You are reading this latest preprint version Research Square lets you share your work early, gain feedback from the community, and start making changes to your manuscript prior to peer review in a journal. As a division of Research Square Company, we’re committed to making research communication faster, fairer, and more useful. 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Also discoverable on Platform About Our Team In Review Editorial Policies Advisory Board Help Center Resources Author Services Accessibility API Access RSS feed Manage Cookie Preferences © Research Square 2026 | ISSN 2693-5015 (online) Privacy Policy Terms of Service Do Not Sell My Personal Information {"props":{"pageProps":{"initialData":{"identity":"rs-4652135","acceptedTermsAndConditions":true,"allowDirectSubmit":false,"archivedVersions":[],"articleType":"Research Article","associatedPublications":[],"authors":[{"id":320932988,"identity":"1753dcce-d641-4d48-ab44-16fff3474f80","order_by":0,"name":"Madhura Athale","email":"","orcid":"","institution":"","correspondingAuthor":false,"prefix":"","firstName":"Madhura","middleName":"","lastName":"Athale","suffix":""},{"id":320932989,"identity":"8ca50ab7-4b98-4119-a5ee-0af63ebb8c61","order_by":1,"name":"Taejoon Park","email":"data:image/png;base64,iVBORw0KGgoAAAANSUhEUgAAAZAAAAAyAQMAAABI0h/eAAAABlBMVEX///8AAABVwtN+AAAACXBIWXMAAA7EAAAOxAGVKw4bAAAAzUlEQVRIie3OsQqCUBTG8SMXrsuhVsXsGS4IiiD5KjeC09IDODY1RXNCD9EjXBBqMVrdBaeGoj2ztpaubQ33Pxw4ww8+AJPpP2MAGcf0uHx/1rIfKQcjKNUvxFqNE6hkTxLZp1oAD9HK6+KOkPh7pSHxeh5IQELmEXkIFGiJUMQUOAfk3iJkCMVUT85NR0SL6JZBN6ztQSpiEiRHx0HRDVN6Em8bJqTiKJDI3YlZkOtINCTm3B48FXZRXC/ZxN9oh72O/Hz7EJPJZDJ97Qk3bDpA6esa/gAAAABJRU5ErkJggg==","orcid":"https://orcid.org/0000-0003-0086-2674","institution":"The Ohio State University","correspondingAuthor":true,"prefix":"","firstName":"Taejoon","middleName":"","lastName":"Park","suffix":""},{"id":320932990,"identity":"2143a693-85ef-4bb1-9447-73a822b9face","order_by":2,"name":"Ryan Hahnlen","email":"","orcid":"","institution":"","correspondingAuthor":false,"prefix":"","firstName":"Ryan","middleName":"","lastName":"Hahnlen","suffix":""},{"id":320932991,"identity":"f627dd66-49da-4d66-8d42-2acd69fc6e5d","order_by":3,"name":"Farhang Pourboghrat","email":"","orcid":"","institution":"","correspondingAuthor":false,"prefix":"","firstName":"Farhang","middleName":"","lastName":"Pourboghrat","suffix":""}],"badges":[],"createdAt":"2024-06-28 05:08:40","currentVersionCode":1,"declarations":"","doi":"10.21203/rs.3.rs-4652135/v1","doiUrl":"https://doi.org/10.21203/rs.3.rs-4652135/v1","draftVersion":[],"editorialEvents":[{"content":"https://doi.org/10.1007/s00170-024-14705-2","type":"published","date":"2024-10-30T16:12:55+00:00"}],"editorialNote":"","failedWorkflow":false,"files":[{"id":60966015,"identity":"61f1f129-1a61-4e9f-a05d-001effc42b2a","added_by":"auto","created_at":"2024-07-24 06:01:20","extension":"png","order_by":1,"title":"Figure 1","display":"","copyAsset":false,"role":"figure","size":11858,"visible":true,"origin":"","legend":"\u003cp\u003eSchematic representation of shell and backfill tool design\u003c/p\u003e","description":"","filename":"1.png","url":"https://assets-eu.researchsquare.com/files/rs-4652135/v1/fd16bbf0454b90a3521fcca8.png"},{"id":60966016,"identity":"561ab4f1-9aea-4089-9992-deda8846a548","added_by":"auto","created_at":"2024-07-24 06:01:20","extension":"png","order_by":2,"title":"Figure 2","display":"","copyAsset":false,"role":"figure","size":213571,"visible":true,"origin":"","legend":"\u003cp\u003eUFT tool components: (a) Die Insert (b) Primary Die (c) Blank Holder (d) Punch (e) Detailed View: Primary Die with Insert (f) Detailed View: Punch [22]\u003c/p\u003e","description":"","filename":"2.png","url":"https://assets-eu.researchsquare.com/files/rs-4652135/v1/4d4564203716503033f07648.png"},{"id":60966404,"identity":"46d85925-90ca-442b-a351-03634b55405a","added_by":"auto","created_at":"2024-07-24 06:09:20","extension":"png","order_by":3,"title":"Figure 3","display":"","copyAsset":false,"role":"figure","size":36488,"visible":true,"origin":"","legend":"\u003cp\u003eCylindrical cup drawinggeometry for FE simulations of shell thickness optimization (unit: mm)\u003c/p\u003e","description":"","filename":"3.png","url":"https://assets-eu.researchsquare.com/files/rs-4652135/v1/d8ec56891bed52ba3e216148.png"},{"id":60967274,"identity":"91a053c5-b336-47cd-96f4-ec550c52509b","added_by":"auto","created_at":"2024-07-24 06:17:20","extension":"png","order_by":4,"title":"Figure 4","display":"","copyAsset":false,"role":"figure","size":97180,"visible":true,"origin":"","legend":"\u003cp\u003eFE simulation model of cup drawing tools with different shell thicknesses\u003c/p\u003e","description":"","filename":"4.png","url":"https://assets-eu.researchsquare.com/files/rs-4652135/v1/d210d2f4b8ecf2b220b6c3a0.png"},{"id":60966400,"identity":"3b3baa11-52aa-44c1-8783-5293c130d223","added_by":"auto","created_at":"2024-07-24 06:09:20","extension":"png","order_by":5,"title":"Figure 5","display":"","copyAsset":false,"role":"figure","size":186434,"visible":true,"origin":"","legend":"\u003cp\u003eEffective plastic strain on polymer punch shell during the stamping simulation for 6 tested cases\u003c/p\u003e","description":"","filename":"5.png","url":"https://assets-eu.researchsquare.com/files/rs-4652135/v1/5c32ce559760f8292856eaec.png"},{"id":60966407,"identity":"b010d693-d1f0-4196-8c74-a215886d690f","added_by":"auto","created_at":"2024-07-24 06:09:21","extension":"png","order_by":6,"title":"Figure 6","display":"","copyAsset":false,"role":"figure","size":167630,"visible":true,"origin":"","legend":"\u003cp\u003eEffective plastic strain (PEEQ) on concrete punch backfill during the stamping simulation for 6 tested cases\u003c/p\u003e","description":"","filename":"6.png","url":"https://assets-eu.researchsquare.com/files/rs-4652135/v1/062c83b9b769025b90036804.png"},{"id":60966019,"identity":"1f3024a7-245a-4f90-8c9f-c783f6d3c13d","added_by":"auto","created_at":"2024-07-24 06:01:20","extension":"png","order_by":7,"title":"Figure 7","display":"","copyAsset":false,"role":"figure","size":522998,"visible":true,"origin":"","legend":"\u003cp\u003e(a) FDM CF-Nylon punch shell (top) and backfilled with cement (bottom) (b) FDM CF-Nylon die shell (top) and backfilled with cement (bottom)\u003c/p\u003e","description":"","filename":"7.png","url":"https://assets-eu.researchsquare.com/files/rs-4652135/v1/423cdfbff87c6d9d3f18491f.png"},{"id":60966021,"identity":"5b55cab5-d155-416f-98f2-d6837f67fcf4","added_by":"auto","created_at":"2024-07-24 06:01:20","extension":"png","order_by":8,"title":"Figure 8","display":"","copyAsset":false,"role":"figure","size":102051,"visible":true,"origin":"","legend":"\u003cp\u003eSimulation model of stamping analysis with shell and backfill tooling\u003c/p\u003e","description":"","filename":"8.png","url":"https://assets-eu.researchsquare.com/files/rs-4652135/v1/e940378806ae03a0c2595778.png"},{"id":60967276,"identity":"e9d04d16-f281-404f-bef4-5b1ff68f6214","added_by":"auto","created_at":"2024-07-24 06:17:20","extension":"png","order_by":9,"title":"Figure 9","display":"","copyAsset":false,"role":"figure","size":353718,"visible":true,"origin":"","legend":"\u003cp\u003eBolt locations on shell and backfill die (left) and corresponding boundary conditions on simulation model (right)\u003c/p\u003e","description":"","filename":"9.png","url":"https://assets-eu.researchsquare.com/files/rs-4652135/v1/2bd82cccc13cf9317e83b3b9.png"},{"id":60966027,"identity":"22762c8d-9dc0-4618-b71e-b1de488995eb","added_by":"auto","created_at":"2024-07-24 06:01:20","extension":"png","order_by":10,"title":"Figure 10","display":"","copyAsset":false,"role":"figure","size":522963,"visible":true,"origin":"","legend":"\u003cp\u003eCF-Nylon shell and cement backfill die after stamping 39 parts, showing failure in flange area\u003c/p\u003e","description":"","filename":"10.png","url":"https://assets-eu.researchsquare.com/files/rs-4652135/v1/bb7c60dbc07cf31bd9b23ffa.png"},{"id":60966402,"identity":"6d0aab99-ca53-4ba6-87d8-0e041f79d5b9","added_by":"auto","created_at":"2024-07-24 06:09:20","extension":"png","order_by":11,"title":"Figure 11","display":"","copyAsset":false,"role":"figure","size":777637,"visible":true,"origin":"","legend":"\u003cp\u003eComparison of wrinkling in part #1 vs. #39 stamped with CF-Nylon + cement shell and backfill tools\u003c/p\u003e","description":"","filename":"11.png","url":"https://assets-eu.researchsquare.com/files/rs-4652135/v1/57c8ed34f0091c07a018671f.png"},{"id":60966029,"identity":"640f8c42-82a2-4b35-93eb-50d039de1568","added_by":"auto","created_at":"2024-07-24 06:01:20","extension":"png","order_by":12,"title":"Figure 12","display":"","copyAsset":false,"role":"figure","size":249339,"visible":true,"origin":"","legend":"\u003cp\u003eATOS 3D scan comparison of CF-Nylon die before and after stamping trials to show outward expansion on the outer walls\u003c/p\u003e","description":"","filename":"12.png","url":"https://assets-eu.researchsquare.com/files/rs-4652135/v1/a02c0011e465ed03e1815eee.png"},{"id":60966406,"identity":"2ebfdaf9-2069-4f61-a15e-742e538b3251","added_by":"auto","created_at":"2024-07-24 06:09:20","extension":"png","order_by":13,"title":"Figure 13","display":"","copyAsset":false,"role":"figure","size":428765,"visible":true,"origin":"","legend":"\u003cp\u003eATOS 3D scan comparison of CF-Nylon punch before and after stamping trials to show outward expansion on the outer walls\u003c/p\u003e","description":"","filename":"13.png","url":"https://assets-eu.researchsquare.com/files/rs-4652135/v1/1840ad93b68c5a5e930dc2db.png"},{"id":60966024,"identity":"af447e9e-a32f-4586-b56f-3d1eb26d57c8","added_by":"auto","created_at":"2024-07-24 06:01:20","extension":"png","order_by":14,"title":"Figure 14","display":"","copyAsset":false,"role":"figure","size":81551,"visible":true,"origin":"","legend":"\u003cp\u003eCross section comparison of part stamped with CF-Nylon + cement tools and all steel tooling from ATOS\u003c/p\u003e","description":"","filename":"14.png","url":"https://assets-eu.researchsquare.com/files/rs-4652135/v1/b8fbb5862de605572bc7709d.png"},{"id":60966031,"identity":"91eb44dc-ef69-4f78-9fbc-b9bd07364dfc","added_by":"auto","created_at":"2024-07-24 06:01:21","extension":"png","order_by":15,"title":"Figure 15","display":"","copyAsset":false,"role":"figure","size":265836,"visible":true,"origin":"","legend":"\u003cp\u003ein-plane major strain distribution on stamped part from ARGUS strain measurements and simulation results\u003c/p\u003e","description":"","filename":"15.png","url":"https://assets-eu.researchsquare.com/files/rs-4652135/v1/1eedf876d1515c5d254e4a13.png"},{"id":60966034,"identity":"67c983ea-578b-40fe-8cfb-2466eaeb1f2a","added_by":"auto","created_at":"2024-07-24 06:01:21","extension":"png","order_by":16,"title":"Figure 16","display":"","copyAsset":false,"role":"figure","size":156747,"visible":true,"origin":"","legend":"\u003cp\u003eOutward displacement of CF-Nylon die shell walls from FE simulation of stamping with \"no bonding\" interface condition\u003c/p\u003e","description":"","filename":"16.png","url":"https://assets-eu.researchsquare.com/files/rs-4652135/v1/1f9f63a4aaa3dd714dbc3e9f.png"},{"id":60966033,"identity":"09cafbd8-4ef2-44c2-a752-0cca8560f3b6","added_by":"auto","created_at":"2024-07-24 06:01:21","extension":"png","order_by":17,"title":"Figure 17","display":"","copyAsset":false,"role":"figure","size":243981,"visible":true,"origin":"","legend":"\u003cp\u003eOut of plane displacement of CF-Nylon die shell flange area for two interface conditions\u003c/p\u003e","description":"","filename":"17.png","url":"https://assets-eu.researchsquare.com/files/rs-4652135/v1/0c19696074658973c3279533.png"},{"id":60967847,"identity":"825e6628-10d7-4ddb-902d-fc35fd870477","added_by":"auto","created_at":"2024-07-24 06:25:20","extension":"png","order_by":18,"title":"Figure 18","display":"","copyAsset":false,"role":"figure","size":347225,"visible":true,"origin":"","legend":"\u003cp\u003eComparison of permanent deformation of CF-Nylon die shell insert area from experimental and simulation results\u003c/p\u003e","description":"","filename":"18.png","url":"https://assets-eu.researchsquare.com/files/rs-4652135/v1/cc5f65ef322d549a994936b0.png"},{"id":60966030,"identity":"a5976562-7202-4c63-bb53-44c91a79ae85","added_by":"auto","created_at":"2024-07-24 06:01:20","extension":"png","order_by":19,"title":"Figure 19","display":"","copyAsset":false,"role":"figure","size":254571,"visible":true,"origin":"","legend":"\u003cp\u003eContour plots of concrete backfill in the die showing hydrostatic pressure (MPa) (left) and the effective plastic strain (right) at the end of stamping\u003c/p\u003e","description":"","filename":"19.png","url":"https://assets-eu.researchsquare.com/files/rs-4652135/v1/ba84f61d5ca65181bcd2cc6c.png"},{"id":68206408,"identity":"42cbcca1-a0c2-4d37-b5a8-4e9b2816a259","added_by":"auto","created_at":"2024-11-04 16:32:19","extension":"pdf","order_by":0,"title":"","display":"","copyAsset":false,"role":"manuscript-pdf","size":6167610,"visible":true,"origin":"","legend":"","description":"","filename":"manuscript.pdf","url":"https://assets-eu.researchsquare.com/files/rs-4652135/v1/9b053e3d-f9ce-415b-a615-d2be59c70611.pdf"}],"financialInterests":"","formattedTitle":"Novel Design of Low-Cost Composite Shell and Backfill Tool for Stamping of HSS 590 Sheet Metal","fulltext":[{"header":"1. Introduction","content":"\u003cp\u003eSheet metal forming is a popular manufacturing technique used in various industries as it is an economical way of mass production. Typical mass production forming toolsets are machined from tool steel and designed to last long enough to produce large volumes of parts, sometimes up to hundreds of thousands. Thus the cost of the machined steel dies can be easily justified in this scenario [\u003cspan citationid=\"CR1\" class=\"CitationRef\"\u003e1\u003c/span\u003e]. However, there are certain situations where these conventional forming tools fall short in the economics aspect \u0026ndash; such as for low or medium volume production of custom parts or prototype parts [\u003cspan citationid=\"CR2\" class=\"CitationRef\"\u003e2\u003c/span\u003e]. The need to manufacture inexpensive yet effective dies quickly for prototyping has paved way for the use of polymer AM technologies for forming tool production [\u003cspan citationid=\"CR3\" class=\"CitationRef\"\u003e3\u003c/span\u003e]. Fused Deposition Modeling (FDM) or alternatively, Fused Filament Fabrication (FFF) is the most commonly used polymer AM technology for tool fabrication.\u003c/p\u003e \u003cp\u003eOver the last two decades, several research works have been carried out that look into various aspects of this topic. It has been experimentally demonstrated that polymer AM tooling can successfully form Aluminum and low-strength steels for small production volumes [\u003cspan additionalcitationids=\"CR5\" citationid=\"CR4\" class=\"CitationRef\"\u003e4\u003c/span\u003e\u0026ndash;\u003cspan citationid=\"CR6\" class=\"CitationRef\"\u003e6\u003c/span\u003e]. Good friction behavior of polymers makes them attractive as tooling materials [\u003cspan citationid=\"CR7\" class=\"CitationRef\"\u003e7\u003c/span\u003e]. Tribological properties of polymer composite tools for contact with metals were studied by Park and Colton [\u003cspan citationid=\"CR8\" class=\"CitationRef\"\u003e8\u003c/span\u003e] as well as Liewald and de Souza [\u003cspan citationid=\"CR9\" class=\"CitationRef\"\u003e9\u003c/span\u003e]. Park and Colton also investigated the failure modes of polymer composite deep drawing dies, including fracture, wear and plastic deformation [\u003cspan citationid=\"CR10\" class=\"CitationRef\"\u003e10\u003c/span\u003e]. Kuo and Li reported that addition of Zirconia particles improved the wear performance of epoxy resin dies [\u003cspan citationid=\"CR11\" class=\"CitationRef\"\u003e11\u003c/span\u003e]. Many studies have shown that FDM parts have significant anisotropy in mechanical properties [\u003cspan additionalcitationids=\"CR13\" citationid=\"CR12\" class=\"CitationRef\"\u003e12\u003c/span\u003e\u0026ndash;\u003cspan citationid=\"CR14\" class=\"CitationRef\"\u003e14\u003c/span\u003e] with the raster direction being significantly stronger than the transverse direction [\u003cspan citationid=\"CR15\" class=\"CitationRef\"\u003e15\u003c/span\u003e]. Several factors such as layer-wise fabrication, fiber reinforcements, internal porosity etc. may cause this anisotropy [\u003cspan citationid=\"CR16\" class=\"CitationRef\"\u003e16\u003c/span\u003e]. However, most numerical investigations into the use of polymer AM tooling for sheet metal forming have used simple isotropic material models for the polymer tooling [\u003cspan citationid=\"CR9\" class=\"CitationRef\"\u003e9\u003c/span\u003e, \u003cspan citationid=\"CR10\" class=\"CitationRef\"\u003e10\u003c/span\u003e, \u003cspan citationid=\"CR17\" class=\"CitationRef\"\u003e17\u003c/span\u003e]. The authors previously demonstrated the importance of using anisotropic material models for accurately predicting the elastic-plastic deformation of FDM CF-Nylon dies during stamping of HSS sheets using finite element analysis [\u003cspan citationid=\"CR18\" class=\"CitationRef\"\u003e18\u003c/span\u003e].\u003c/p\u003e \u003cp\u003eAn important issue with polymer AM tooling that is reported in published research is their elastic and plastic deformation during forming and resultant part deviations. Nakamura et al. performed V bending of steel sheets using FDM PLA tools and reported a higher geometrical deviation in final parts as compared to steel tools [\u003cspan citationid=\"CR19\" class=\"CitationRef\"\u003e19\u003c/span\u003e]. Durgun performed stamping of steel sheets with FDM Polycarbonate dies and reported that MC355 parts were outside the dimensional tolerance range due to tool deformations [\u003cspan citationid=\"CR20\" class=\"CitationRef\"\u003e20\u003c/span\u003e]. Similarly, Tondini et al. reported higher springback in V bending with polymer tools due to their low stiffness and elastic deformation [\u003cspan citationid=\"CR21\" class=\"CitationRef\"\u003e21\u003c/span\u003e]. In a previously published work [\u003cspan citationid=\"CR22\" class=\"CitationRef\"\u003e22\u003c/span\u003e], the authors showed that stamping 100 HSS590 sheets with glass fiber reinforced Polycarbonate dies resulted in a consistent reduced draw depth of 2mm across all parts as compared to steel tooling. Finite element analysis revealed the elastic deformation of the dies to be a potential cause of this [\u003cspan citationid=\"CR22\" class=\"CitationRef\"\u003e22\u003c/span\u003e]. Similar observations of lower resultant draw depth of up to 1.7 mm for steel blanks in case of CF-Nylon tooling compared to steel tooling were reported by Giorleo and Ceretti [\u003cspan citationid=\"CR23\" class=\"CitationRef\"\u003e23\u003c/span\u003e]. In contrast to elastic deformation of dies, most of the plastic deformation takes place in the first few stamping passes and reaches a steady state value [\u003cspan citationid=\"CR17\" class=\"CitationRef\"\u003e17\u003c/span\u003e, \u003cspan citationid=\"CR18\" class=\"CitationRef\"\u003e18\u003c/span\u003e, \u003cspan citationid=\"CR21\" class=\"CitationRef\"\u003e21\u003c/span\u003e]. Deviation in the final part dimension when using AM polymer tooling is an important aspect that needs to be further investigated.\u003c/p\u003e \u003cp\u003eIn this study, the authors have proposed a novel \u0026ldquo;shell and backfill\u0026rdquo; die design concept to mitigate some of the issues with using fully dense or solid AM tooling for sheet metal stamping reported in the previous work [\u003cspan citationid=\"CR22\" class=\"CitationRef\"\u003e22\u003c/span\u003e]. A \u0026ldquo;shell and backfill\u0026rdquo; type tool consists of an outer thin shell of the stamping die or punch produced by 3D printing of polymer composite material and backfilled with a less expensive material. This is expected to achieve higher cost savings than solid AM tooling. Geueke et al. used geometry optimization of solid AM polymer tooling for cup drawing by removing material strategically from low load areas in the tool for cost reduction [\u003cspan citationid=\"CR24\" class=\"CitationRef\"\u003e24\u003c/span\u003e]. However, such a design approach may only be applicable to relatively simple geometries and extension of it to more complex geometries would incur high computational costs during design. On the other hand, the shell and backfill design is also highly adaptable to any tool geometry. Additionally, using a backfill material that is stiffer than the polymer shell material, limits the elastic deformation of the tools during stamping. This reduces the possibility of insufficient draw depth or other geometric deviations in final parts as seen with solid AM polymer tooling.\u003c/p\u003e \u003cp\u003eThis paper is organized as follows: In Section 2, the shell and backfill tool design concept is explained in further detail, as well as material testing and characterization information is provided. Details of the finite element models and the experimental stamping setup are also given. Section 3 is dedicated to a case study of the shell and backfill tooling including tool design, finite element analysis, experimental stamping trials and results. Finally, Section 4 provides conclusions and future work.\u003c/p\u003e"},{"header":"2. Materials and Methods","content":"\u003cdiv id=\"Sec3\" class=\"Section2\"\u003e \u003ch2\u003e2.1. Shell and backfill tooling\u003c/h2\u003e \u003cp\u003eIn the previous research, glass fiber polycarbonate (GFPC) solid tools showed only 0.25 mm deformation in the highest stress area after stamping 100 parts from 1.539 mm HSS 590 sheets [\u003cspan citationid=\"CR22\" class=\"CitationRef\"\u003e22\u003c/span\u003e]. This indicates that the GFPC tools could possibly be used for stamping more parts, but it also implies that their potential may not be fully leveraged in a typical prototype or low volume production scenario. Moreover, the solid GFPC die set was able to achieve 11% cost savings compared to a conventional steel tool. In the innovative \u0026ldquo;shell and backfill\u0026rdquo; type tooling approach, an outer thin shell of the stamping die or punch is produced by 3D printing of polymer composite material, which is then backfilled with a less expensive material. Figure\u0026nbsp;\u003cspan refid=\"Fig1\" class=\"InternalRef\"\u003e1\u003c/span\u003e shows schematic representation of the shell and backfill tool design. As most of the die volume is made up of the inexpensive backfill material, there is potential for additional cost savings of up to 50% as compared to the conventional steel tool cost, depending on the specific backfill material used. This approach takes advantage of the most important feature of 3D printing, that is to fabricate complex shapes rapidly, by fabricating the forming surfaces of the tools, while also reducing the printing time and material as well as machine cost by removing a large volume from the 3D printed part. Backfilling operation is very fast and requires minimal operator skill.\u003c/p\u003e \u003cp\u003e \u003c/p\u003e \u003c/div\u003e \u003cdiv id=\"Sec4\" class=\"Section2\"\u003e \u003ch2\u003e2.2. Materials\u003c/h2\u003e \u003cp\u003eIn this study, the AM shell material was carbon fiber reinforced nylon (CF-Nylon) printed with the FDM technique. FDM CF-Nylon was chosen as the die shell material based on its excellent mechanical properties and demonstrated success as a tooling material for sheet metal stamping [\u003cspan citationid=\"CR18\" class=\"CitationRef\"\u003e18\u003c/span\u003e]. CF-Nylon mechanical characterization was performed using coupons machined from FDM printed blocks along four unique directions to capture the anisotropy in their mechanical properties as detailed in [\u003cspan citationid=\"CR18\" class=\"CitationRef\"\u003e18\u003c/span\u003e]. High strength steel sheets (HSS 590) with a grade of 590 MPa and a thickness of 1.539 mm were used throughout this study, as it is a common material grade and gauge used in the automotive industry [\u003cspan citationid=\"CR1\" class=\"CitationRef\"\u003e1\u003c/span\u003e]. HSS 590 material properties, including Young\u0026rsquo;s modulus, Possion\u0026rsquo;s ratio, 0.2%-offset initial yield stress, and hardening curve, were obtained from uniaxial tension tests performed along the rolling direction (RD) [\u003cspan citationid=\"CR18\" class=\"CitationRef\"\u003e18\u003c/span\u003e].\u003c/p\u003e \u003cp\u003eConcrete was considered as the backfill material in this study for its high stiffness compared to polymers, and its low cost, and ease of use. Mechanical properties of concrete, including average values for Young\u0026rsquo;s modulus and Poisson\u0026rsquo;s ratio listed in Table\u0026nbsp;\u003cspan refid=\"Tab1\" class=\"InternalRef\"\u003e1\u003c/span\u003e, were obtained from various online sources [\u003cspan citationid=\"CR25\" class=\"CitationRef\"\u003e25\u003c/span\u003e]. Compressive strength for concrete was obtained from the manufacturer\u0026rsquo;s specification. Commercially available high-strength concrete mix with a cured strength of 27.5 MPa was used for this study [\u003cspan citationid=\"CR26\" class=\"CitationRef\"\u003e26\u003c/span\u003e]. Mechanical properties of materials used in this study are given in Table\u0026nbsp;\u003cspan refid=\"Tab1\" class=\"InternalRef\"\u003e1\u003c/span\u003e and Table\u0026nbsp;\u003cspan refid=\"Tab2\" class=\"InternalRef\"\u003e2\u003c/span\u003e.\u003c/p\u003e \u003cp\u003e \u003cdiv class=\"gridtable\"\u003e\u003ctable float=\"Yes\" id=\"Tab1\" border=\"1\"\u003e \u003ccaption language=\"En\"\u003e \u003cdiv class=\"CaptionNumber\"\u003eTable 1\u003c/div\u003e \u003cdiv class=\"CaptionContent\"\u003e \u003cp\u003eMechanical properties of materials.\u003c/p\u003e \u003c/div\u003e \u003c/caption\u003e \u003ccolgroup cols=\"3\"\u003e \u003cdiv align=\"left\" class=\"colspec\" colname=\"c1\" colnum=\"1\"\u003e\u003c/div\u003e \u003cdiv align=\"left\" class=\"colspec\" colname=\"c2\" colnum=\"2\"\u003e\u003c/div\u003e \u003cdiv align=\"char\" char=\".\" class=\"colspec\" colname=\"c3\" colnum=\"3\"\u003e\u003c/div\u003e \u003cthead\u003e \u003ctr\u003e \u003cth align=\"left\" colname=\"c1\" morerows=\"1\" rowspan=\"2\"\u003e \u003cp\u003eProperty\u003c/p\u003e \u003c/th\u003e \u003cth align=\"left\" colspan=\"2\" nameend=\"c3\" namest=\"c2\"\u003e \u003cp\u003eMaterial\u003c/p\u003e \u003c/th\u003e \u003c/tr\u003e \u003ctr\u003e \u003cth align=\"left\" colname=\"c2\"\u003e \u003cp\u003e\u003cem\u003eConcrete\u003c/em\u003e\u003c/p\u003e \u003c/th\u003e \u003cth align=\"left\" colname=\"c3\"\u003e \u003cp\u003e\u003cem\u003eHSS 590 RD\u003c/em\u003e\u003csup\u003e\u003cem\u003e*\u003c/em\u003e\u003c/sup\u003e\u003c/p\u003e \u003c/th\u003e \u003c/tr\u003e \u003c/thead\u003e \u003ctbody\u003e \u003ctr\u003e \u003ctd align=\"left\" colname=\"c1\"\u003e \u003cp\u003eYoung\u0026rsquo;s Modulus (\u003cem\u003eE\u003c/em\u003e) (GPa)\u003c/p\u003e \u003c/td\u003e \u003ctd align=\"left\" colname=\"c2\"\u003e \u003cp\u003e24\u003c/p\u003e \u003c/td\u003e \u003ctd align=\"char\" char=\".\" colname=\"c3\"\u003e \u003cp\u003e206.5\u003c/p\u003e \u003c/td\u003e \u003c/tr\u003e \u003ctr\u003e \u003ctd align=\"left\" colname=\"c1\"\u003e \u003cp\u003ePoisson\u0026rsquo;s ratio (\u003cem\u003eν\u003c/em\u003e)\u003c/p\u003e \u003c/td\u003e \u003ctd align=\"left\" colname=\"c2\"\u003e \u003cp\u003e0.21\u003c/p\u003e \u003c/td\u003e \u003ctd align=\"char\" char=\".\" colname=\"c3\"\u003e \u003cp\u003e0.31\u003c/p\u003e \u003c/td\u003e \u003c/tr\u003e \u003ctr\u003e \u003ctd align=\"left\" colname=\"c1\"\u003e \u003cp\u003eYield stress (\u003cem\u003eσ\u003c/em\u003e\u003csub\u003e\u003cem\u003ey\u003c/em\u003e\u003c/sub\u003e) (MPa)\u003c/p\u003e \u003c/td\u003e \u003ctd align=\"left\" colname=\"c2\"\u003e \u003cp\u003e--\u003c/p\u003e \u003c/td\u003e \u003ctd align=\"char\" char=\".\" colname=\"c3\"\u003e \u003cp\u003e435.2\u003c/p\u003e \u003c/td\u003e \u003c/tr\u003e \u003ctr\u003e \u003ctd align=\"left\" colname=\"c1\"\u003e \u003cp\u003eStrength (MPa)\u003c/p\u003e \u003c/td\u003e \u003ctd align=\"left\" colname=\"c2\"\u003e \u003cp\u003e27.5\u003c/p\u003e \u003c/td\u003e \u003ctd align=\"char\" char=\".\" colname=\"c3\"\u003e \u003cp\u003e709.8\u003c/p\u003e \u003c/td\u003e \u003c/tr\u003e \u003c/tbody\u003e \u003c/colgroup\u003e \u003ctfoot\u003e \u003ctr\u003e\u003ctd colspan=\"3\"\u003e\u003csup\u003e*\u003c/sup\u003e Rolling direction\u003c/td\u003e\u003c/tr\u003e \u003c/tfoot\u003e \u003c/table\u003e\u003c/div\u003e \u003c/p\u003e \u003cp\u003e \u003cdiv class=\"gridtable\"\u003e\u003ctable float=\"Yes\" id=\"Tab2\" border=\"1\"\u003e \u003ccaption language=\"En\"\u003e \u003cdiv class=\"CaptionNumber\"\u003eTable 2\u003c/div\u003e \u003cdiv class=\"CaptionContent\"\u003e \u003cp\u003eAveraged uniaxial compression results with deviations for CF-Nylon [\u003cspan citationid=\"CR18\" class=\"CitationRef\"\u003e18\u003c/span\u003e].\u003c/p\u003e \u003c/div\u003e \u003c/caption\u003e \u003ccolgroup cols=\"5\"\u003e \u003cdiv align=\"left\" class=\"colspec\" colname=\"c1\" colnum=\"1\"\u003e\u003c/div\u003e \u003cdiv align=\"left\" class=\"colspec\" colname=\"c2\" colnum=\"2\"\u003e\u003c/div\u003e \u003cdiv align=\"left\" class=\"colspec\" colname=\"c3\" colnum=\"3\"\u003e\u003c/div\u003e \u003cdiv align=\"left\" class=\"colspec\" colname=\"c4\" colnum=\"4\"\u003e\u003c/div\u003e \u003cdiv align=\"left\" class=\"colspec\" colname=\"c5\" colnum=\"5\"\u003e\u003c/div\u003e \u003cthead\u003e \u003ctr\u003e \u003cth align=\"left\" colname=\"c1\" morerows=\"1\" rowspan=\"2\"\u003e \u003cp\u003eProperty\u003c/p\u003e \u003c/th\u003e \u003cth align=\"left\" colspan=\"4\" nameend=\"c5\" namest=\"c2\"\u003e \u003cp\u003eOrientation\u003c/p\u003e \u003c/th\u003e \u003c/tr\u003e \u003ctr\u003e \u003cth align=\"left\" colname=\"c2\"\u003e \u003cp\u003e\u003cem\u003eX\u003c/em\u003e\u003c/p\u003e \u003c/th\u003e \u003cth align=\"left\" colname=\"c3\"\u003e \u003cp\u003e\u003cem\u003eZ\u003c/em\u003e\u003c/p\u003e \u003c/th\u003e \u003cth align=\"left\" colname=\"c4\"\u003e \u003cp\u003e\u003cem\u003eXY\u003c/em\u003e\u003c/p\u003e \u003c/th\u003e \u003cth align=\"left\" colname=\"c5\"\u003e \u003cp\u003e\u003cem\u003eXZ\u003c/em\u003e\u003c/p\u003e \u003c/th\u003e \u003c/tr\u003e \u003c/thead\u003e \u003ctbody\u003e \u003ctr\u003e \u003ctd align=\"left\" colname=\"c1\"\u003e \u003cp\u003eYoung\u0026rsquo;s Modulus (\u003cem\u003eE\u003c/em\u003e) (GPa)\u003c/p\u003e \u003c/td\u003e \u003ctd align=\"left\" colname=\"c2\"\u003e \u003cp\u003e2.20\u0026thinsp;\u0026plusmn;\u0026thinsp;0.20\u003c/p\u003e \u003c/td\u003e \u003ctd align=\"left\" colname=\"c3\"\u003e \u003cp\u003e1.85\u0026thinsp;\u0026plusmn;\u0026thinsp;0.08\u003c/p\u003e \u003c/td\u003e \u003ctd align=\"left\" colname=\"c4\"\u003e \u003cp\u003e3.95\u0026thinsp;\u0026plusmn;\u0026thinsp;0.52\u003c/p\u003e \u003c/td\u003e \u003ctd align=\"left\" colname=\"c5\"\u003e \u003cp\u003e1.6\u0026thinsp;\u0026plusmn;\u0026thinsp;0.12\u003c/p\u003e \u003c/td\u003e \u003c/tr\u003e \u003ctr\u003e \u003ctd align=\"left\" colname=\"c1\" morerows=\"1\" rowspan=\"2\"\u003e \u003cp\u003ePoisson\u0026rsquo;s ratio (\u003cem\u003eν\u003c/em\u003e)\u003c/p\u003e \u003c/td\u003e \u003ctd align=\"left\" colname=\"c2\"\u003e \u003cp\u003eν\u003csub\u003exy\u003c/sub\u003e\u0026thinsp;=\u0026thinsp;0.55\u0026thinsp;\u0026plusmn;\u0026thinsp;0.05\u003c/p\u003e \u003c/td\u003e \u003ctd align=\"left\" colname=\"c3\"\u003e \u003cp\u003eν\u003csub\u003exy\u003c/sub\u003e\u0026thinsp;=\u0026thinsp;0.14\u0026thinsp;\u0026plusmn;\u0026thinsp;0.03\u003c/p\u003e \u003c/td\u003e \u003ctd align=\"left\" colname=\"c4\" morerows=\"1\" rowspan=\"2\"\u003e \u003cp\u003eν\u003csub\u003eplanar\u003c/sub\u003e\u0026thinsp;=\u0026thinsp;0.09\u0026thinsp;\u0026plusmn;\u0026thinsp;0.01\u003c/p\u003e \u003c/td\u003e \u003ctd align=\"left\" colname=\"c5\" morerows=\"1\" rowspan=\"2\"\u003e \u003cp\u003eν\u003csub\u003eplanar\u003c/sub\u003e\u0026thinsp;=\u0026thinsp;0.31\u0026thinsp;\u0026plusmn;\u0026thinsp;0.02\u003c/p\u003e \u003c/td\u003e \u003c/tr\u003e \u003ctr\u003e \u003ctd align=\"left\" colname=\"c2\"\u003e \u003cp\u003eν\u003csub\u003exz\u003c/sub\u003e\u0026thinsp;=\u0026thinsp;0.25\u0026thinsp;\u0026plusmn;\u0026thinsp;0.04\u003c/p\u003e \u003c/td\u003e \u003ctd align=\"left\" colname=\"c3\"\u003e \u003cp\u003eν\u003csub\u003exy\u003c/sub\u003e\u0026thinsp;=\u0026thinsp;0.12\u0026thinsp;\u0026plusmn;\u0026thinsp;0.04\u003c/p\u003e \u003c/td\u003e \u003c/tr\u003e \u003ctr\u003e \u003ctd align=\"left\" colname=\"c1\"\u003e \u003cp\u003eYield stress (\u003cem\u003eσ\u003c/em\u003e\u003csub\u003e\u003cem\u003ey\u003c/em\u003e\u003c/sub\u003e) (MPa)\u003c/p\u003e \u003c/td\u003e \u003ctd align=\"left\" colname=\"c2\"\u003e \u003cp\u003e26.4\u0026thinsp;\u0026plusmn;\u0026thinsp;2.51\u003c/p\u003e \u003c/td\u003e \u003ctd align=\"left\" colname=\"c3\"\u003e \u003cp\u003e29.2\u0026thinsp;\u0026plusmn;\u0026thinsp;0.76\u003c/p\u003e \u003c/td\u003e \u003ctd align=\"left\" colname=\"c4\"\u003e \u003cp\u003e45.5\u0026thinsp;\u0026plusmn;\u0026thinsp;1.92\u003c/p\u003e \u003c/td\u003e \u003ctd align=\"left\" colname=\"c5\"\u003e \u003cp\u003e25.8\u0026thinsp;\u0026plusmn;\u0026thinsp;2.81\u003c/p\u003e \u003c/td\u003e \u003c/tr\u003e \u003c/tbody\u003e \u003c/colgroup\u003e \u003c/table\u003e\u003c/div\u003e \u003c/p\u003e \u003c/div\u003e \u003cdiv id=\"Sec5\" class=\"Section2\"\u003e \u003ch2\u003e2.3. Numerical analysis\u003c/h2\u003e \u003cp\u003eFinite element (FE) analysis was performed using the commercial solver ABAQUS/Explicit throughout this study for stamping simulations. Tool components including the AM polymer shell and backfill were modeled as deformable bodies and meshed with linear tetrahedral elements (C3D4). HSS 590 blanks were meshed with linear quadrilateral shell elements. In each case, a steel blank holder was modeled as a rigid body.\u003c/p\u003e \u003cp\u003eHSS 590 was modeled with von Mises yield criterion since the difference in the measured stress-strain curves was negligible in uniaxial tensile tests along 0, 45, and 90 degrees to the rolling direction. Isotropic hardening rule was assumed based on the stress-strain curve measured in the uniaxial tension test. The hardening curve was extrapolated beyond the uniform elongation limit using a power-law-type hardening equation. To account for the severe anisotropy of the CF-Nylon FDM shell, orthotropic elasticity and Hill 1948 yield criterion were assumed with the tetragonal symmetry constraints. Additionally, an isotropic hardening rule was assumed to represent the anisotropic expansion of the yield stress preserving the initial anisotropy. Material model parameters for the FDM CF-Nylon were calibrated based on experimental data as detailed in [\u003cspan citationid=\"CR18\" class=\"CitationRef\"\u003e18\u003c/span\u003e]. Although concrete is typically brittle under the uniaxial compression, it demonstrates brittle-to-ductile transition, increased strength, strain hardening, and increased failure strain under compressive hydrostatic pressure [\u003cspan citationid=\"CR27\" class=\"CitationRef\"\u003e27\u003c/span\u003e, \u003cspan citationid=\"CR28\" class=\"CitationRef\"\u003e28\u003c/span\u003e]. In other words, under confinement, when the hydrostatic pressure in the concrete is high, it exhibits higher load-carrying capacity than its nominal compressive strength. Studies have shown that applying hydrostatic pressure of 100% or 200% of the nominal compressive strength, increased the compressive strength of concrete significantly \u0026ndash; by 200\u0026ndash;400% [\u003cspan citationid=\"CR29\" class=\"CitationRef\"\u003e29\u003c/span\u003e]. In sheet metal stamping, the tools experience mostly compressive stresses. As the concrete backfill is surrounded by the polymer composite shell, the compressive stresses arising during the stamping process are unlikely to lead to catastrophic failure of the concrete. For simplification of the material modeling, the concrete backfill was assumed as isotropic elastic \u0026ndash; perfectly plastic material. As the perfect plasticity model does not consider the increased yield stress and hardening due to the hydrostatic pressure, this concrete backfill modeling can be seen as a type of worst-case scenario assumption. In other words, if the simulation results show high hydrostatic pressure comparable to nominal stress in the critical regions, it can be expected that in reality, the tool performance will be better due to higher yield stress and hardening behavior of concrete. Additionally, depending on the confining pressure, high strains in concrete even up to 5% may not result in failure [\u003cspan citationid=\"CR27\" class=\"CitationRef\"\u003e27\u003c/span\u003e, \u003cspan citationid=\"CR29\" class=\"CitationRef\"\u003e29\u003c/span\u003e].\u003c/p\u003e \u003c/div\u003e \u003cdiv id=\"Sec6\" class=\"Section2\"\u003e \u003ch2\u003e2.4. Experimental setup\u003c/h2\u003e \u003cp\u003eThe same stamping tool geometry used in the previous study with solid AM tooling was also used in this study to maintain consistency. Figure\u0026nbsp;\u003cspan refid=\"Fig2\" class=\"InternalRef\"\u003e2\u003c/span\u003e shows the components of the Universal Formability tool (UFT) \u0026ndash; punch, primary die with die inserts, and blank holder \u0026ndash; and their dimensions. Square sheets of HSS 590 with 470 mm side length and 1.539 mm thickness were used for stamping.\u003c/p\u003e \u003cp\u003eA \u0026ldquo;baseline\u0026rdquo; stamping trial was first performed with HSS 590 blanks using conventional steel tooling for comparing the performance of shell and backfill tools against a reference. A steel blank holder was used during all the stamping trials. All forming trials utilized a 300-ton servo press programmed with a crank slide motion. The blank holder force was maintained at 200 kN for the duration of the draw phase. Total draw depth was set at 65 mm with a speed of 14 strokes per minute. These parameters were chosen based on prior experience with the sheet metal material and relevance to mass-production manufacturing. For all trials, the HSS 590 blanks were lubricated with the same mass-production relevant water-based lubricant and selected blanks were etched for post-forming strain analysis. All process parameters were kept consistent between the steel tools and composite tools. ATOS 3D scans of the shell and backfill tools were performed before and after the stamping trial to get a quantitative measure of the dimensional change of tools due to plastic deformation, damage, and wear. ATOS and ARGUS scans of the AM-stamped parts were also performed and compared against those from steel-stamped parts to evaluate the accuracy and consistency of the tools.\u003c/p\u003e \u003cp\u003e \u003c/p\u003e \u003c/div\u003e"},{"header":"3. Case study: CF-Nylon + Concrete tool","content":"\u003cdiv id=\"Sec8\" class=\"Section2\"\u003e \u003ch2\u003e3.1. Tool design: shell thickness optimization\u003c/h2\u003e \u003cp\u003e \u003cem\u003eSimulation conditions\u003c/em\u003e \u003c/p\u003e \u003cp\u003eTo minimize computational cost, the optimum thickness of the outer shell die was determined by performing FE simulations for cup drawing with a simple geometry shown in Fig.\u0026nbsp;\u003cspan refid=\"Fig3\" class=\"InternalRef\"\u003e3\u003c/span\u003e. The conditions for cup drawing were designed to closely mirror those of UFT forming. The punch and die radii for cup drawing were selected to be similar to the tightest tool radius in the UFT die. Additionally, the 3.4 mm gap between punch and die was maintained. For the initial circular blank, the same 1.539 mm thick HSS 590 sheet was employed. As for the shell and backfill materials, FDM-produced CF-Nylon and concrete were used, respectively. To further save computational cost, a quarter model was utilized for the FE simulations. Three different shell thicknesses were considered for the optimization simulations: 2.5 mm, 5 mm, and 10 mm. Figure\u0026nbsp;\u003cspan refid=\"Fig4\" class=\"InternalRef\"\u003e4\u003c/span\u003e shows the simulation model for different shell thicknesses.\u003c/p\u003e \u003cp\u003eThe contact condition at the interface between the shell and backfill has uncertainties in terms of friction and bonding. The surface of the CF-Nylon shell, produced by the FDM process, may exhibit irregularities, and the interfacial bonding can become extremely weak due to dry shrinkage and capillary water evaporation during the concrete curing process. Therefore, the simulations need to account for any possible interface condition and select the most suitable shell thickness. Two extreme scenarios were considered for the FE simulations to include all possible friction conditions: 1) smooth contact to simulate no friction or bonding/adhesion between the shell and backfill, and 2) bonded interface to simulate infinite friction or perfect bonding between the shell and backfill. The bonded interface condition was simulated using tie constraints between nodes on the surfaces of shell and backfill that are in contact with each other. With 3 different shell thicknesses and 2 different interface conditions, a total of 6 unique simulation conditions were examined.\u003c/p\u003e \u003cp\u003e \u003c/p\u003e \u003cp\u003e \u003c/p\u003e \u003cp\u003e \u003cem\u003eSimulation results\u003c/em\u003e \u003c/p\u003e \u003cp\u003eFE simulations of cup drawing using the shell and backfill tools showed that the bonding between the shell and backfill at the interface has a significant effect on the tool performance in terms of resultant stress, strain, and plastic deformation. Critical areas in the tools such as the punch and die corner radii as well as the corner radii at the shell and backfill interface were examined for different shell thicknesses and two extreme interface conditions as described previously. Figure\u0026nbsp;\u003cspan refid=\"Fig5\" class=\"InternalRef\"\u003e5\u003c/span\u003e shows a comparison of the effective plastic strain (PEEQ) on the CF-Nylon punch shell for all 6 simulation cases. The two interface conditions, namely perfect bonding and no bonding, generated reverse trends in the amount of plastic strain as a function of the shell thickness. Increasing shell thickness led to higher accumulated plastic strain near the punch corner of the CF-Nylon shell in the perfect bonding interface condition. Conversely, in the no bonding case, as the shell thickness increased, the amount of accumulated plastic strain in the CF-Nylon shell decreased. As the interface condition is uncertain, the optimal shell thickness should be chosen based on the lowest accumulated plastic strain in both extreme interface conditions. In this case, 5 mm thickness was found to be optimal. For the 5 mm shell thickness case, the no bonding interface condition was found to have maximum plastic strains of 2\u0026ndash;3% in the shell. As forming tools mostly experience compressive loads, the higher value of 2\u0026ndash;3% plastic strain is still below failure strain [\u003cspan citationid=\"CR18\" class=\"CitationRef\"\u003e18\u003c/span\u003e].\u003c/p\u003e \u003cp\u003e \u003c/p\u003e \u003cp\u003eHowever, the effect of shell thickness and interface condition on the backfill material should be confirmed before finalizing the optimal shell thickness. Figure\u0026nbsp;\u003cspan refid=\"Fig6\" class=\"InternalRef\"\u003e6\u003c/span\u003e shows the effective plastic strain in the concrete punch backfill for the 6 simulation cases. The backfill concrete demonstrated an increase in the accumulated plastic strain as the shell thickness increased in both interface conditions. Unlike the CF-Nylon shell shown in Fig.\u0026nbsp;\u003cspan refid=\"Fig5\" class=\"InternalRef\"\u003e5\u003c/span\u003e, the concrete backfill exhibited higher plastic strain in the perfect bonding interface case. This is the result of a more effective transfer of load from the shell to the backfill when there is bonding between the two. In the no bonding interface condition, shear load and plastic deformation remains isolated within the shell. In general, some bonding between the two materials is desired so that the shell does not experience very high deformations which may lead to part deviations or tool failure. Considering the resultant plastic strain on both shell and backfill materials under the two interface conditions, 5 mm shell thickness was found to be optimal. However, if there is good bonding between AM material and concrete, the concrete material is likely to develop internal cracks. As the backfill is completely surrounded by the polymer shell, risk of catastrophic failure due to concrete was considered to be low. A similar analysis was performed for the cup drawing die, considering FDM Nylon shell and concrete backfill to obtain the optimized die shell thickness. Shell thickness of 5 mm was again found to be the most suitable.\u003c/p\u003e \u003cp\u003e \u003c/p\u003e \u003c/div\u003e \u003cdiv id=\"Sec9\" class=\"Section2\"\u003e \u003ch2\u003e3.2. Experimental stamping trial\u003c/h2\u003e \u003cp\u003eBased on the shell thickness optimization presented in the previous section, tool shells with 5 mm wall thickness were fabricated with CF-Nylon using FDM technique for the UFT forming. A track width of 0.5 mm in the build plane (X-Y) and a layer height of 0.254 mm along the build direction (Z) were used to make all tools. In this case, the punch travel is in Z direction as well. Each layer was printed using a single perimeter pass and filled with an alternating raster orientation of \u0026plusmn;\u0026thinsp;45 degrees from the X axis. Printed tools were then backfilled with concrete and allowed to cure for the specified amount of time. Figure\u0026nbsp;\u003cspan refid=\"Fig7\" class=\"InternalRef\"\u003e7\u003c/span\u003e shows the tools before and after the backfilling operation. The original steel toolset for the UFT geometry cost \u003cspan\u003e$\u003c/span\u003e20,500. The FDM CF-Nylon shell for the punch and die combined cost \u003cspan\u003e$\u003c/span\u003e7,100 and the concrete backfill cost \u003cspan\u003e$\u003c/span\u003e10. The CF-Nylon shell and concrete backfill toolset achieved 65% cost reduction compared to the conventional toolset. As the same steel blank holder was used for both toolsets, it was left out of the cost comparison.\u003c/p\u003e \u003cp\u003eStamping of HSS 590 blanks was performed with a steel blank holder using process parameters described in Section 2.4. A total of 39 parts were stamped with the shell and backfill toolset.\u003c/p\u003e \u003cp\u003e \u003c/p\u003e \u003c/div\u003e \u003cdiv id=\"Sec10\" class=\"Section2\"\u003e \u003ch2\u003e3.3. Numerical analysis of stamping\u003c/h2\u003e \u003cp\u003eFE simulations of the stamping of HSS 590 blanks with shell and backfill tooling were performed using the Abaqus/Explicit commercial solver. Relevant simulation conditions and material model details are given in Section 2.3. Figure\u0026nbsp;\u003cspan refid=\"Fig8\" class=\"InternalRef\"\u003e8\u003c/span\u003e shows a sectional view of the Abaqus simulation model used for the FE analysis. Boundary conditions for controlling the displacement of punch and die were determined considering the physical attachment points where tools were attached to the press plates during the stamping trials indicated by red circles in Fig.\u0026nbsp;\u003cspan refid=\"Fig9\" class=\"InternalRef\"\u003e9\u003c/span\u003e (a). Specifically, displacement boundary conditions, including punch displacement in Z direction and fixed displacements in X and Y directions (\u003cspan class=\"InlineEquation\"\u003e\u003cspan class=\"mathinline\"\u003e\\({U}_{X}={U}_{Y}=0\\)\u003c/span\u003e\u003c/span\u003e), were applied to the corresponding elements highlighted in red in Fig.\u0026nbsp;\u003cspan refid=\"Fig9\" class=\"InternalRef\"\u003e9\u003c/span\u003e (b). Additionally, similar to the shell thickness optimization simulations in Section 3.1, two interface conditions corresponding to perfect bonding and no bonding were considered to evaluate the effect of the interface between shell and backfill.\u003c/p\u003e \u003cp\u003e \u003c/p\u003e \u003cp\u003e \u003c/p\u003e \u003c/div\u003e \u003cdiv id=\"Sec11\" class=\"Section2\"\u003e \u003ch2\u003e3.4. Results\u003c/h2\u003e \u003cp\u003eDuring the experimental stamping trial with shell and backfill tools, 39 parts were formed before stopping due to failure of the tool shell that accumulated over the course of the trial as shown in Fig.\u0026nbsp;\u003cspan refid=\"Fig10\" class=\"InternalRef\"\u003e10\u003c/span\u003e. Progressively increasing wrinkling was observed in the flange region of the stamped parts, which finally caused die failure at the location of the largest wrinkle due to the penetration of the wrinkle into the polymer die shell as shown in Fig.\u0026nbsp;\u003cspan refid=\"Fig11\" class=\"InternalRef\"\u003e11\u003c/span\u003e.\u003c/p\u003e \u003cp\u003e \u003c/p\u003e \u003cp\u003e \u003c/p\u003e \u003cp\u003eBoth the punch and die were observed to be bulging outwardly on the outer vertical walls. This indicates insufficient adhesion or bonding between the shell and backfill. Due to the outward bulging, the punch sustained visible damage on the outer walls rubbing against the blank holder during each stamping pass. ATOS 3D scans were conducted to measure the deformation of the tools after the stamping trials. Figure\u0026nbsp;\u003cspan refid=\"Fig12\" class=\"InternalRef\"\u003e12\u003c/span\u003e shows a maximum bulging of 1.18 mm on the die wall, while Fig.\u0026nbsp;\u003cspan refid=\"Fig13\" class=\"InternalRef\"\u003e13\u003c/span\u003e shows bulging of 1.11 mm to 4.38 mm at different locations on the punch wall. ATOS 3D scans in both figures show the comparison of the deformed tool surfaces (after stamping 39 parts) with respect to the original tool surfaces (before the stamping trial). Here the outward direction from the tool surface is assigned positive sign. In other words, positive values indicate the tool surface has expanded outward compared to the original tool surface.\u003c/p\u003e \u003cp\u003e \u003c/p\u003e \u003cp\u003e \u003c/p\u003e \u003cp\u003eOne of the motivations behind using the shell and backfill die design was to mitigate the issue of the shallow draw as seen with the solid AM polymer tooling [\u003cspan citationid=\"CR22\" class=\"CitationRef\"\u003e22\u003c/span\u003e]. By using a backfill material with a higher elastic modulus, the elastic deformation of the tools is reduced, resulting in stamping deeper parts as intended. Figure\u0026nbsp;\u003cspan refid=\"Fig14\" class=\"InternalRef\"\u003e14\u003c/span\u003e shows a cross-section of the formed part after #1 trial using this tooling and a comparison with that of a steel-stamped part as obtained from ATOS 3D scanning. No draw deficit was observed with the new CF-Nylon shell and concrete backfill tooling. However, a small amount of flattening of the stamped part at the die insert nose region is observed as compared to the steel-stamped part due to the deformation of the polymer die at that location.\u003c/p\u003e \u003cp\u003e \u003c/p\u003e \u003cp\u003eFinal part strains compared between experimental values obtained through ARGUS strain analysis and those from simulation results show good agreement, providing a basic validation of simulation results as shown in Fig.\u0026nbsp;\u003cspan refid=\"Fig15\" class=\"InternalRef\"\u003e15\u003c/span\u003e. The ARGUS grid pattern was washed out from some regions on the part during stamping where strain measurements could not be obtained. Differences between simulations and experiments are attributed to different averaging methods and averaging areas used between ARGUS and simulations.\u003c/p\u003e \u003cp\u003e \u003c/p\u003e \u003cp\u003e \u003c/p\u003e \u003cp\u003eFigure\u0026nbsp;\u003cspan refid=\"Fig16\" class=\"InternalRef\"\u003e16\u003c/span\u003e shows the deformation of the CF-Nylon die shell walls at the end of the stamping stroke with no bonding interface condition between the shell and backfill. Like the experimental observations, the die shell walls showed outward deformation or bulging. The excessive wrinkling of stamped parts seen in the experimental trials (refer to Fig.\u0026nbsp;\u003cspan refid=\"Fig11\" class=\"InternalRef\"\u003e11\u003c/span\u003e) is thought to result from the uneven bulging of the shell and the consequent uneven distribution of blank holding force. This claim is supported by simulation results showing non-uniform deformation of the top surface of the die shell in the stamping direction. Figure\u0026nbsp;\u003cspan refid=\"Fig17\" class=\"InternalRef\"\u003e17\u003c/span\u003e shows a comparison of die top surface deformation in the Z direction (stamping direction) between the two simulation cases of perfect bonding and no bonding. The perfect bonding case shows negligible deformation of the die surface near the flange area. However, the no bonding case shows highly non-uniform deformation, which could create uneven gaps in between the die and blank holder, resulting in non-uniform blank holder force in different areas of the flange, promoting wrinkling. Actual location and height of wrinkles on the formed part could not be accurately predicted with the simple material model used for the blank, as wrinkling and formability of the formed part were not the focus of this study.\u003c/p\u003e \u003cp\u003e \u003c/p\u003e \u003cp\u003eAdditionally, the simulation was able to qualitatively predict the deformation of the die insert area, which is the highest-stress area in the tool. Figure\u0026nbsp;\u003cspan refid=\"Fig18\" class=\"InternalRef\"\u003e18\u003c/span\u003e shows a comparison of experimental die scan results after 39 stamping and simulation results of the same at the end of the first stamping pass. Accurate quantitative prediction of plastic deformation depends on the material model for AM polymers which should account for various effects such as anisotropy, tension-compression asymmetry, and strain rate sensitivity. Successful qualitative and quantitative prediction of FDM CF-Nylon tool deformation was demonstrated through the use of appropriate material models previously [\u003cspan citationid=\"CR18\" class=\"CitationRef\"\u003e18\u003c/span\u003e]. In the case of shell and backfill die design, the plastic deformation is also dependent on the material model used for the backfill and the unloading behavior of the backfill.\u003c/p\u003e \u003cp\u003e \u003c/p\u003e \u003cp\u003eTo assess the damage to the concrete backfill, Fig.\u0026nbsp;\u003cspan refid=\"Fig19\" class=\"InternalRef\"\u003e19\u003c/span\u003e shows the hydrostatic pressure and effective plastic strain (PEEQ) in the insert region of the die at the end of the stamping simulation. The maximum hydrostatic pressure is 196 MPa, which is more than 7 times the nominal compressive strength of concrete. As mentioned in Section 2.3, when hydrostatic pressure is twice the nominal compression strength, the compressive strength can increase up to four times compared to scenarios without hydrostatic pressure. Additionally, the maximum PEEQ in the concrete backfill of the die was 2.1%, which is well within the safe zone for such high confining pressure.\u003c/p\u003e \u003cp\u003e \u003c/p\u003e \u003c/div\u003e"},{"header":"4. Conclusions and future work","content":"\u003cp\u003eThe shell and backfill die design is a novel approach for further lowering the cost of polymer AM tooling for sheet metal forming applications. Solid AM GF-PC toolset achieved 11% cost reduction, while the new CF-Nylon\u0026thinsp;+\u0026thinsp;Concrete shell and backfill toolset achieved 65% lower cost compared to steel tools. Inexpensive backfill materials such as cement can achieve high cost-savings but may not be the most ideal backfill material to meet all performance criteria and recycling feasibility. The contact and bonding condition at the interface between the shell and backfill has a significant impact on the tool performance in terms of tool longevity as well as the final part geometry. The presence of bonding between the shell and backfill is essential for preventing buckling or bulging of the shell away from the backfill, which may lead to failure of the shell. Additionally, shell bulging may result in non-uniform distribution of the blank holding force on the flange region, promoting wrinkling in the part.\u003c/p\u003e \u003cp\u003ePolymer materials for the shell often do not have sufficient bond strength with the various backfill materials under consideration. To address this issue, various mechanical interlocking mechanisms such as bolts, dovetails, etc. to attach the shell and backfill together need to be explored. This is also expected to improve part accuracy and tool life by eliminating she shell and backfill separation. Additionally, recyclable and energy-friendly materials such as metal alloys with low melting temperatures will be considered as potential backfill materials for future work.\u003c/p\u003e"},{"header":"Declarations","content":"\u003cp\u003e \u003ch2\u003eCompeting Interests\u003c/h2\u003e \u003cp\u003eThe authors have no relevant financial or non-financial interests to disclose.\u003c/p\u003e \u003c/p\u003e\u003ch2\u003eFunding\u003c/h2\u003e \u003cp\u003eThis work was supported by Honda Development and Manufacturing of America, LLC (AWD-115703).\u003c/p\u003e\u003ch2\u003eAuthor contributions\u003c/h2\u003e \u003cp\u003e \u003cb\u003eMadhura Athale\u003c/b\u003e: Methodology, Software, Validation, Formal analysis, Investigation, Writing \u0026ndash; original draft, Visualization; \u003cb\u003eTaejoon Park\u003c/b\u003e: Conceptualization, Methodology, Writing \u0026ndash; Review \u0026amp; editing; \u003cb\u003eRyan Hahnlen\u003c/b\u003e: Supervision; \u003cb\u003eFarhang Pourboghrat\u003c/b\u003e: Writing \u0026ndash; review \u0026amp; editing, Supervision, Project administration, Funding acquisition\u003c/p\u003e\u003ch2\u003eAcknowledgements\u003c/h2\u003e \u003cp\u003eThe authors wish to thank Honda Development and Manufacturing of America, LLC. for their support of this project through the award AWD-115703: An inverse design methodology to fabricate low-cost agile tools for manufacturing lightweight automotive components.\u003c/p\u003e"},{"header":"References","content":"\u003col\u003e\u003cli\u003e\u003cspan\u003eHahnlen R, Pourboghrat F, Park T, Hoffman B, Athale M (2021) accessed October 27,. 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However, these tools have limitations when it comes to prototype production volumes, due to the high cost and time required for tool fabrication. Forming tools fabricated with polymers using additive manufacturing (AM) offer an inexpensive alternative suitable for low volume production and prototyping. For successful implementation of polymer AM tooling in sheet metal forming, tool cost, tool life, and part accuracy are important metrics. A novel composite tool design concept consisting of two distinct components \u0026ndash; an outer polymer AM shell, and inner backfill \u0026ndash; to make up the composite tool is proposed. Experimental and numerical investigation of stamping of high strength steel sheets with the new tool design is presented. It is shown that the new composite tool design concept improves performance and is more economical compared with fully dense or solid AM polymer tools.\u003c/p\u003e","manuscriptTitle":"Novel Design of Low-Cost Composite Shell and Backfill Tool for Stamping of HSS 590 Sheet Metal","msid":"","msnumber":"","nonDraftVersions":[{"code":1,"date":"2024-07-24 06:01:15","doi":"10.21203/rs.3.rs-4652135/v1","editorialEvents":[{"type":"communityComments","content":0},{"type":"decision","content":"Major Revisions Needed","date":"2024-08-18T11:20:42+00:00","index":"","fulltext":""},{"type":"reviewerAgreed","content":"","date":"2024-07-01T07:51:31+00:00","index":0,"fulltext":""},{"type":"reviewersInvited","content":"","date":"2024-07-01T02:27:04+00:00","index":"","fulltext":""},{"type":"editorAssigned","content":"","date":"2024-06-30T22:34:04+00:00","index":"","fulltext":""},{"type":"submitted","content":"The International Journal of Advanced Manufacturing Technology","date":"2024-06-28T01:08:29+00:00","index":"","fulltext":""}],"status":"published","journal":{"display":true,"email":"[email protected]","identity":"the-international-journal-of-advanced-manufacturing-technology","isNatureJournal":false,"hasQc":true,"allowDirectSubmit":false,"externalIdentity":"jamt","sideBox":"Learn more about [The International Journal of Advanced Manufacturing Technology](https://www.springer.com/journal/170)","snPcode":"170","submissionUrl":"https://submission.nature.com/new-submission/170/3","title":"The International Journal of Advanced Manufacturing Technology","twitterHandle":"","acdcEnabled":true,"dfaEnabled":true,"editorialSystem":"em","reportingPortfolio":"Springer Hybrid","inReviewEnabled":true,"inReviewRevisionsEnabled":false}}],"origin":"","ownerIdentity":"61adb5f1-9519-4fbf-bb0a-df789f30ddef","owner":[],"postedDate":"July 24th, 2024","published":true,"recentEditorialEvents":[],"rejectedJournal":[],"revision":"","amendment":"","status":"published-in-journal","subjectAreas":[],"tags":[],"updatedAt":"2024-11-04T16:21:45+00:00","versionOfRecord":{"articleIdentity":"rs-4652135","link":"https://doi.org/10.1007/s00170-024-14705-2","journal":{"identity":"the-international-journal-of-advanced-manufacturing-technology","isVorOnly":false,"title":"The International Journal of Advanced Manufacturing Technology"},"publishedOn":"2024-10-30 16:12:55","publishedOnDateReadable":"October 30th, 2024"},"versionCreatedAt":"2024-07-24 06:01:15","video":"","vorDoi":"10.1007/s00170-024-14705-2","vorDoiUrl":"https://doi.org/10.1007/s00170-024-14705-2","workflowStages":[]},"version":"v1","identity":"rs-4652135","journalConfig":"researchsquare"},"__N_SSP":true},"page":"/article/[identity]/[[...version]]","query":{"redirect":"/article/rs-4652135","identity":"rs-4652135","version":["v1"]},"buildId":"qtupq5eGEP_6zYnWcrvyt","isFallback":false,"isExperimentalCompile":false,"dynamicIds":[84888],"gssp":true,"scriptLoader":[]}