Currently, c omponents, processes and materials are mainly developed independently. However, to 12 exploit the full potential of modern materials in component design, integrative development work is necessary. 13 Component performance -based requirements and corresponding loc al material properties must be taken into 14 account. In this work, a component -driven approach and therefore an inverted process chain is presented and 15 demonstrated on the performance of a crash box, produced from DP600 steel. This is aimed to increase the 16 energy absorption of the crash box without losing progressive buckling behavior. The finite element 17 simulations were carried out on the crash box. It was shown that the crash box corners play an important role 18 in the crash box performance, and required material properties for improving the crash box performance were 19 derived based on the simulations. Heat treatment strategies were afterward developed and experimentally 20 validated to fulfill these requirements. Then different local heat -treatment processes were applied to the 21 corners of the crash boxes and tested experimentally. The experiments results validate the potential of the 22 inverted process chain to improve the components performance. Furthermore, in this pape r, optimal material 23 properties were extracted for crash box strengthening, which lead to high energy absorption a nd a low peak 24 force of it. 25 Keywords: Gradient microstructure , local heat treatment, square crash box , axial crushing, DP600 steel , peak 26 force 27 1. Introduction 28 Crash boxes are thin -walled tubes , which are used widely for energy -absorption purposes in collisions. 29 Hereby, the kinetic energy of a n impactor is converted by continuous folding into plastic deformation of the 30 crash box. Different aspects should be considered in a crash box design. In the last decades, a considerable 31 amount of research was conducted on the mechanisms of crash boxes . Several approaches have been introduced 32 for estimating the ir energy absorption [1–3]. Progressive buckling is a prerequisite for the crash box function . 33 Buckling of a tube is progressive when it produces several local folds (wrinkles ) along the structures under a xial 34 loading. The p rogressive buckling of crash boxes has been studied in several research works [4–8]. The c rash 35 box peak force is another important aspect of a crash box design . Axial crushing of thin -walled tubes leads to a 36 high initial peak force , which is related to the first fold cr eation on the tube. A high peak force endangers 37 protected structures and passengers’ safety [9]. Apart from the mentioned aspects , cracking in corners [10] and 38 global -stability failure [4,8] of the component should also be considered in the crash box design. C omprehensive 39 books and review papers have been published on crash box analysis [11–15]. Even though a considerable 40 amount of research has been devoted to crash boxes, they require constant improvements due to the demand for 41 increasing pas sengers’ safety and lightweight design concerns. 42 Different studies have been conducted to improve crash box performance . Materials optimization and also 43 foam filling effects were studied in [16–19]. It is reported that the foam filling can enhance the energy 44 absorption, even though the exact response of the c ellular material to dynamic loading is not yet 45 comprehensively understood. Wang et al. used a multi -objective optimization approach in combination with a 46 biomimetic -inspired lattice filling of the crash box to improve its crashworthiness [20]. Beside s the material 47 approaches, geometrical approaches are also widely used to improve the energy absorption. Jones et al. studied 48 the effect of stiffening elements on crash box efficiency [21]. The stiffeners were joined to the inner and outer 49 surface of the crash box to increase the crash box energy absorption. Even though the energy absorption was 50 increased, the stiffeners also increased the chance of crash box global buckling. External linea r-gradient grooves 51 2 effect on crashworthiness of the circular tubes was studied by Yao et al. [22]. The c rashworthiness design of 52 the windowed tubes under axial loading was studied by Nikkhah et al. [23]. Multi -cell tubes as a replacement 53 for typical tubes ha ve been studied in several research papers [13,18,24,25] . The studies revealed that increasing 54 the number of the corners results in energy -absorption raise. Thickness grading and corrugating effect on the 55 multi -cell tube performance have been studied by Fang et.al [26] and Li et al. [27]. Zhang et al. have studied 56 the thickness grading effect on energy absorption of hollow square tube s [28]. They reported that the grading 57 can increase the energy absorption up to 30% without increasing the crash box peak force . This thickness 58 grading can be achieved by different process strategies [29], e.g., flexible rolling , which is used to produce 59 tailor rolled blanks [30], or adapted welding processes to manufacture tailor welded blanks [31]. The crash box 60 manufacturing itself was investigated by Dimirci et al. who characterized the influence of spot welding in crash 61 box assembly regarding its performance [32]. 62 Additionally, local heat treatment has been introduced in the literature to improve crash box performance. 63 Bambach et al. [33] have used laser heat treatment to obtain spatially graded components to control folding 64 patterns of crash box es. By applying the same crush energy, the crushing distance of the crash box decreased 65 almost by 28% due to the strain -hardening engineering. Similar results were reported by Quadfasel et al. who 66 also used strain -hardening engineering to enhance the material properties and thus the crash box efficiency [34]. 67 The optimization resulted in a 35% weight reduction of the crash box. Conrads et al. [35] have used local heat 68 treatment for increasing the energy absorption capacity of the crash box and also increasing ductility at the 69 corners of the steel crash box to prevent crack initiation. Based on their results, the locally heat -treated crash 70 box in comparison to the fully -recrystallized crash box can absorb the same amount of energy with 17% less 71 crushing distance. The h igh effect of the crash box corners on e nergy absorption is reported by other researche rs 72 as well [26,28,36] . However, the local heat treatment s affect also other crash box features such as peak force 73 and progressive buckling, which need to be considered to design an optimized crash box . Contrary to the laser 74 heat source, which ha s been used by Bambach et al. [33], Hosseini et al. used an arc heat treatment technique 75 to produce graded microstructures through controlled temperature gradients [37]. A comprehensive review on 76 different approaches to improve crash -box performance is provided in [13]. 77 Most of the existing design and optimization approaches of the thin-walled crash box can be categorized 78 into geometrical -based, material -based and process -based approaches. Usually , if a novel steel alloy is 79 developed its application on the crash box is investigated regarding the energy absorption. The s ame applies if 80 a heat treatment method or forming process is refined. Whereas in this paper, a n integrative method is 81 established, which can be seen as a n improvement of the current component design : From a material -focused 82 structural design approach to a component performance -based mat erial and process development. This allows 83 the identification of component -derived local property requirement s, which are converted into tailored material 84 properties. Similar con cepts can be found in the literature . Ituarte et al. [38] developed a functional graded 85 material design and manufacturing workflow which includes problem specification, graded s tructural design, 86 the grading realization and product fabrication. This work focuses on additive manufactured structures of so - 87 called digital materials. Similar approaches of fully integrated design workflows can be found in the field of 88 aminated continuou s fiber -reinforced composites [39]. In this context, local property grading for nodular ca st 89 iron has been studied by Gebhardt et al. [40] and Ripplinger et al [41]. Both works showed a benefit in terms 90 of increased fatigue strength in the case of high silicone -alloyed no dular cast iron and ferrite graded nodular 91 cast iron. Even though the approach shows high efficiency, it has not been applied to dual-phase steel, nor to 92 crash relevant s tructural automotive components , to the best of the authors ' knowledge. 93 In this paper , the above -mentioned approach of a component performance -based material and process 94 development is demonstrated on property -graded crash boxes. Here , energy absorption is taken as the main 95 measure for the assessment of the crash box performance . The aim is to increase the energy absorption without 96 endangering the progressive buckling of the crash box by local adjustment of the material properties . Using 97 finite element models , highly stressed regions in the crash box were identified and artificia l material cards were 98 assigned locally in order to optimize the energy absorption of the crash box . With this approach , local property 99 requirements were defined . Isothermal -heat-treatment investigations were afterward used to determine 100 information regarding holding time and temperature that allows fulfilling the previously identified mechanical 101 property requirements . Using a process -adapted heat treatment approach, process parameters leading to the 102 desired holding time and temperature were then determined and applied on cras h boxes, which were tested 103 quasi -statically afterward. Furthermore , finite element analysis was used to investigate the crushing behavior 104 of the crash box es. Thereby, the effects of different material behaviors on crash box features such as peak force 105 3 were studied. Furthermore, finite element analysis was used to identify the advantages of local grading over the 106 homogenous heat treatment of the entire crash box regarding the progressive -buckling behavior of it. In 107 comparison to previous studies, current work contains the following novelties: 108  implementing a new design approach, from a material focused structural design approach to a 109 component performance -based ma terial and process development ; 110  identification of component -derived local property requirements, which can be converted into 111 tailored material properties ; 112  compari son of different local heat treatment strategies on enhancing the crash box energy - 113 absorption; 114  definition of advantages and limitations of property -graded crash boxes. 115 2. Materials and Methods 116 2.1 Material 117 In this study , a cold -rolled DP600 steel with a material thickness of 1.5 mm was investigated. The chemical 118 composition of the material is displayed in Table 1. As part of the 2nd Gen . Advanced High Strength Steels 119 (AHSS ), dual -phase (DP) steels are characterized by their good combination of strength and ductility, which 120 makes them attractive for crash components in t he automotive industry. This mechanical propert y profile of DP 121 steels is attributed to the ferritic -martensitic microstructure . Depending on the hard to soft phase ratio the 122 mechanical prop erties can be designed. The investigated DP600 steel features a microstructure containing 9.1 123 % martensite and 88.9 % ferrite. 124 Table 1: Chemical composition of DP600 steel, in mass %. 125 C Si Mn P Cu Cr Mo Ni N 0.113 0.390 1.38 0.0167 0.0152 0.179 0.0515 0.0198 0.00001 126 Tensile properties of the DP600 steel were determined following the standard DIN EN ISO 6892 -1 using 127 a tensile specimen with a gauge length of 20 mm. According to the tensile test, a yield strength (YS) of 391 128 MPa, ultimate tensile strength (UTS) of 705 MPa, and total elongation (TE) of 31.3 % w ere extracted . The 129 stress -strain curve and the microstructure of the DP600 are shown in Figure 1. 130 (a) (b) Figure 1: (a) Stress -strain curve; (b) Microstructure of DP600. 131 2.2 Identification of local property requirements using a c omponent -driven approach 132 In the present work, a novel methodical approach for integrative material, process , and component design 133 are presented using the example of a thin -walled crash box . This approach aim s to demonstrate the potential of 134 a component -driven material and process design in order to increase the energy absorption potential of a crash 135 box. Unlike conventional component design, where global homogenous material properties are defined, this 136 approach first derives required local material properties on the component level by a preliminary analysis that 137 combines FEM simulations with artificial material cards and analytical progressive buckling condition s (1). 138 The experimental fulfil lment o f these material requirements within the constraints of the applied process 139 4 technology ( e.g., short hol ding times) is investigated by selected, isothermal heat treatments of the base material 140 DP600 in a salt bath. This type of heat treatment is further cal led isothermal heat treatment. The following 141 material characterization, including microstructural and mecha nical investigations, enables the identif ication of 142 functional relationships between heat treatment process parameters and resulting mechanical prope rties (2). In 143 order to transfer the outcomes of isothermal heat treatment s to a locally ap plicable heat treatment process, 144 further experiments to define cooling rates and process -specific boundary conditions ( e.g., critical cooling rates) 145 are performed . Based on this information , the crash box is locally heat-treated with a moving heat source to 146 adjust the material properties in the required regions, identified in step (1) of this approach. In the following , 147 this type of heat treatment is called process -adapted heat treatment. This processing enables high flexibility 148 considering the location and size of the heat-treated region and thus on the local properties. In the following 149 heat treatment patterns are defi ned either as line heat treatment or combinations of single spot heat treatments 150 using different spot densities and orientations. To assess the different heat treatment strategies , tensile test 151 specimens were processed by different heat treatment pattern s and subsequently tested under quasi -static 152 conditions . The same patterns were applied locally to the corners of the crash boxes (3). The locally heat -treated 153 and reference component s were then experimentally tested under axial compression . Based on the force - 154 displacement curves, the e nergy absorption can be derived . Further FE simulation s were carried out to assess 155 the performance of isothermal heat-treated crash boxes (4). A Schematic process map of the component -driven 156 approach is shown in Figure 2. 157 158 159 Figure 2: Schematic process map of the component -driven approach . 160 2.3 Crash box es under axial loading 161 To carry out and validate the abovementioned method, both experimental and numerical compression tests 162 on crash boxes are requir ed. These tests were used a s a basis for formulating local -material -property 163 requirements to improve the component performance. The crash box specimen s were manufactured from 164 DP600 steel, w hich was characterized in s ection 2.1 . The crash box es are manufactured by arc welding of two 165 u-shape profiles , which were bent from the DP600 steel sheets . Thus, t he weld seams are located in the middle 166 of the side faces . Dimensions of t he manufactured crash box es are 40 mm x 40 mm x 200 mm and the thickness 167 is 2 mm (Figure 4 (a)). 168 Even t hough the normal function of the crash boxe s is under dynamic loading s, such as collisions, it is 169 enlightening to first investigate the crash box behav ior under a quasi -static load. Moreover, most of the vehicle 170 collisions are in low velocities, about 40 m/s at most, and the influence of the stress -wave speed on the crash 171 box behavior during an impact weakens as the impact velocity decreases [42]. As a reasonable simplification, 172 the a xial impact behavior of tubes with velocities up to tens of meters per second, for metal tubes, is taken as 173 quasi -static in the lit erature [12]. Therefore , in this study quasi-static tests were carried out as well for 174 experimental investigation of the energy absorption of the heat -treated crash boxes and also to validate the 175 numerical simulation of the crash box . The specimens were painted with a stochastic pattern to investigate the 176 local strain s by digital image correlation (DIC) , Aramis GOM. The mechanical tests of the investigated crash 177 boxes were performed by a servo -hydraulic testing machine, Instron 8804. The sampling frequency was set at 178 100 Hz. Force signal was co rrelated with DIC data (sampling frequency of 4Hz) to extract local strain s at certain 179 load points of interest during the test. The compression load was applied via a load stamp with a displacement 180 rate of 30 mm/min until the crash box is bottomed out. The experimental setup is shown in Figure 3. 181 182 Process -adapted heat treatment (3) Investigation of material propertiesInvestigation of material propertiesComponent LevelInvestigation of component properties SimulationMaterial RequirementsHeat treatment strategyComponent processingInvestigation of component properties Simulation Experiment Specimen LevelSalt bathHeat treatment windowLocally heat treated component Isothermal material propertiesProcess -affected material propertiesLocally moving heat sourceIsothermal heat treatment (2)Preliminary Analysis (1)Component assessment (4) 5 183 Figure 3: The e xperimental setup for quasi -static compression of the crash box. DIC method was used for 184 investigating the local strains. 185 The numerical simulations of the crash boxes were performed using Abaqus commercial software. By 186 considering the symmetry of the crash box and also the symmetry of its crushed form, half of the crash box was 187 simulated. 4 -node doubly -curved shell elements w ith full integration were used for meshing the crash box. By 188 performing a mesh convergence study, a 1.25 mm mesh size was chosen for the meshing. On the tested crash 189 boxes, there was not any imperfection made deliberately; however, the production phase cou ld not be perfect. 190 Therefore implementing a small imperfection on the FE model can result in more realistic results [43]. Thus, a 191 small imperfection with 0.1 mm deflection was applied on the second fold region of the crash box (see Figure 192 4 (b)). The fold region can be located by calculating the fold length of the crash box, as a progressively buckled 193 crash box has a specific fold length. The fold length means the length of the crash box that contributes to a local 194 fold creation and it can be extracted analytically or experimentally , by dividing the crash box length by the 195 numbe r of the folds produced along it [17,44,45] . The material card of the DP600 steel was applied to the model. 196 The flow curve was derived from the previously discussed quasi -static tensile tests of the DP600 steel. The 197 equivalent plastic strain value in the simulated crash box can go beyond the strain at the necking point extracted 198 by the one -dimensional tensile test [46]. Thus, the extracted flow -curve was extrapolated using Voce hardening 199 function, 200 𝜎y(𝜀p)=𝐾0+𝑄(1−𝑒−𝑏𝜀p), (1) where σy(εp) donates the stress as a function of plastic strain, K 0 and Q are material constants and b is the strain 201 hardening exponent, which were extracted experimentally, based on the tensile test results (see Table 3). 202 The explicit method can be used for low-speed quasi -static simulations [47]. To reduce the calculation 203 time, it is often useful to use mass -scaling or reduce the analysis time. Both approaches yield similar results for 204 the rate -independent models [48]. The analysis time of 0.3 s and mass -scaling factor of 4 were considered for 205 the analysis. Even though the m ass scaling and the analysis -time reduction are common approach es to decrease 206 computation time, attention should be taken as they can affect the results accuracy [43,49] . Therefore the FE 207 results should be analyzed to be sure about the mode l accuracy. The calculated kinetic energy was insignificant 208 compared to the energy absorbed by the specimen and the load-displacement diagram was not affected 209 remarkably . Furthermore, i n order to ensure the simula tion accuracy, the model is validated with the 210 experiment al results . 211 6 Figure 4: (a) Dimensions of the tested crash boxes, the thickness of the crash boxes is 2 mm. (b) imperfection 212 on the simulated crash box at the second -fold region (exaggerated); (c) force -displacement diagrams of the 213 experimental test and numerical simulation of DP600 crash box (none heat -treated). 214 Validating the FE model is crucial in a crash box simulation. FE and experiment al results of the 215 compression test on the none heat -treated crash box are presented in Figure 4 (c). Furthermore, Figure 5 show s 216 the comparison of the out -of-plane displacement , in the x-direction , between the FE analysis ( a) and the 217 experiment ( b) at the peak force moment, captured with DIC method. The extracted peak force is 116 KN (see 218 Figure 4 (c)), t his peak force is related to the first fold creation on the tube. The first fold is located at the bottom 219 end of the crash box , both in the FE and experiment. Furthermore, t wo pre-folds are emerging as well , in the 220 middle and at the top edge of the crash box . Considering the crash box crushed shape and force -displacement 221 diagrams, o verall a good agreement could be obtained between the FE model and the experiment. As can be 222 seen in Figure 6, the corners of the crash box face the highest strain. Therefore , increasing strength and ductility 223 of the corners region can improve the energy absorption , which is going to be discussed more in detail in the 224 result section . 225 The energy absorption of the crash box is defined as the area under the force -displacement curv e derived 226 from a component test; whereas , the material’s energy d issipation in homogeneous uni axial deformation 227 (dissipation parameter ) is defined as the area under its stress -strain curve extracted by tensile tests and can be 228 seen as the absorbed energy per unit volume . The energy absorption E (kJ) and the so -called dissipation 229 parameter , D (MPa), describing the combined strength and ductility effect up to the UTS can be calculated as: 230 231 𝐸=∫𝐹(𝛿)𝑑𝛿, (2) 232 𝐷=∫𝜎(𝜀)𝑑𝜀, (3) 233 (a) (b) (c) 7 where F represents the force, 𝛿 the displacement, σ the stress and ε the plastic strain , respectively . In some 234 resources [50], the dissipation parameter is named modulus of toughness when the upper bound of the 235 integration is the fracture strain of the material . 236 237 238 Figure 5: Comparison of the out -of-plane displacement in the x -direction (mm) for peak force between FE 239 analysis (left) and experiment (right) captured with DIC method. 240 (a) (b) Figure 6: (a) Validated FE mode; (b) experimental test result of the crash box with DP600 steel material 241 behavior. 242 2.4 Progressive buckling and corner -strengthen ing of a crash box 243 Progressive buckling is a prerequisite for the crash box function . A progressive buckling results in local 244 folds production along the crash box length. Plastic deformation is the main energy -absorption mechanism of a 245 crash box and just by a progressive buc kling considerable portion of the crash box deforms plastically. 246 Progressive buckling is also named stable deformation under axial load ing. A crash box has a progressive 247 behavior when it s side plates buckle plastically [6,51] . A crash box can be considered as four connected plates , 248 and c ritical buckling stress of a plate can be calculated as 249 250 𝜎𝑐𝑟=𝑘𝜋2𝐸 12(1−𝜐2)(𝑡 𝑏)2 , (4) 251 where E is Young modulus, 𝜐 Poisson’s ratio, k buckling coefficient, b width and t thickness of the plate 252 respectively [52,53] . Based on the dimensions of the plates, the boundary condition of the crash box plates was 253 considered simply supported [54]. Besides the plastic bucking necessity, it was shown in [6] that a crash box 254 Mises 8 cannot buckle progressively when the length of the local folds is higher than the crash box width. Therefore , 255 the maximum length of a plate that contributes to a local fold production is the same as the crash box width, see 256 Figure 7. In order to have a plastic buckling, t he critical buckling stress of this plate should be beyond the yield 257 stress of the material, 𝜎𝑦𝑖𝑒𝑙𝑑<𝜎𝑐𝑟, which is a necessary condition for the crash box to buckle progressively 258 [6,51] . 259 Figure 7: Crash box under compression load F with the plate region engaged in a local fold. 260 Crash box corners face the highest deformation by the local folds production. The m echanism of a fold 261 production on a crash box starts with the local buckling of the crash box plates . Plates have a special capability 262 to withstand the axial compression load even after buckling [55] By initiation of the buckling , the mid-side of 263 the crash box plates cannot bear the compression load anymore , and the stress decreases at these regions . 264 However the compressive stress can still increase at the crash box corners owing to the higher rigidity at these 265 regions . Therefore by initiation of the buckling , the stress distribution is not uniformon on the cross section any 266 more . By further loading, crash box corners buckles as well, which is known as crippling, and local folds are 267 produced along the crash box. Production of t hese l ocal folds result s in the creation of the sharp curves and high 268 plastic deformation at the corners , which results c onsequently in high energy absorption at these regions . Thus, 269 the corners play an important role in the energy absorption of a crash box . High -stress concentration at the crash 270 box corners can be seen in the FE results as well, see Figure 6. 271 Due to the high importance of the crash box corners in energy absorption, corner strengthening can 272 enhance energy absorption significantly. In this paper the effect of corner strengthening on crash box energy 273 absorption was studied experimentally and numerically, using the validated FE model. A parametric study was 274 conducted to study the effect of different corner strengths . The higher strength material properties were applied 275 on crash box corners, as depicted in Figure 4 (a). Results of these simulations were compared regarding the 276 energy absorption and also peak force of the crash box in the result section . Furthermore, heat -treatment 277 processes on DP600 steel were studied in order to increase the strength at the crash box corners as indicated by 278 the computa tions . 279 2.5 Heat treatments and characterization methods 280 The heat treatments to examine suitable strategies to fulfill propert y requirements were carried out in a 281 salt bath. The material was isothermally heat-treated for one minute at a specific holding temperature, which 282 was ranged from 750 °C to 950 °C, before quenching in oil . These heat treatments lead to partial or complete 283 austenitization before quenching, so that microstructural configuration is significantly different from the initial 284 state. For microstructural characterization of the different heat treatment states, the samples were etched with 285 Nital after metallographic preparation. The mechanical properties were determined by Vickers hardness test 286 and tensile test according to DIN EN ISO 6892 -1. 287 To transfer the conditions from the isothermal heat treatment to a process -adapted heat treatment that uses 288 a moving heat source, the phase -transformation behavior of the material is studied by dilatometric 289 investigations. Therefore, continuous cooling curves with different cooling rates that range between 1.6 - 204 290 K/s are measured. The combination of this dilatometric data yields a continuous cooling transformation (CCT) 291 diagram ( see Figure 13), which can be used to identify critical cooling rates for the DP600 steel . To minimize 292 9 radiation effects on the measurement , the temperature of the specimen is determined on the inside at a point of 293 heat treatment. The crash box specimens were heat-treated by using a tungsten inert gas arc welding torch . 294 Different process -adapted heat -treatment s trategies were applied to identify a suitable pattern for increasing the 295 energy absorption of the crash box . Conrads et al. have shown that the energy absorption of crash boxes can be 296 increased by implementing a wavelike pat tern using laser heat treatment [35]. However , the arc welding 297 technique has not been investigated for such application. The application of an arc provides a fast and flexible 298 way of heat treatment without the necessity of additional equipment other than a relatively small power source 299 and a welding torch. Line and spot heat treatment patterns have been chosen, which are easy to implement in 300 an industrial context. Figure 8 depicts the different heat treatment patterns . In the following , it can be 301 distinguished between non -heat-treated (N), line heat -treated (L), spot heat -treated (S) and spot double -dense 302 heat-treated (SDD) patterns. In addition to crash boxes , tensile test samples with similar heat treatment patterns 303 were manufactured to compare material properties resulting from process -adapted heat treatment and isothermal 304 heat treatment. The shape DIN 5 0125 – H 20 x 120 wa s chosen for the tensile test samples . In accordance with 305 DIN 6892 -1, an elongation speed of 15mm/min wa s used for these tests. 306 307 308 Figure 8: (a) Heat -treatment pattern on the crash box ; (b) Heat treatment pattern on the tensile specimen; 309 (c) schemat ic of heat treatment setup ; (d) Indication of the thermoelement positioning. 310 3. Results and Discussion 311 3.1 Energy absorption of the corner -strengthened crash box 312 The FE-analysis results of the corner -strengthened crash box are provided in this section. As it was 313 discussed, strengthening the crash box corners can improve the crash box energy absorption. A parametric study 314 was conducted and different material cards were applied on the crash box corners in order to study the edge 315 strengthening effect . The DP600 steel has true ultimate stress of 8 41 MPa. Three artificial materials were 316 defined with increased ultimate stress of 1000 MPa , 1200 MPa and 1400 MPa, named U1000, U1200 and 317 U1400 respectively. Except the plasticity parameters, all the other parameters, e.g., Young’s modulus and yield 318 stress, of the artificial materials were kept the same as the DP600 steel . The f low curve s of the materials were 319 extrapolated using Voce -law. The extrapolated stress-strain curve s of the artificial materials and DP600 steel 320 are shown in Figure 10 (a). The extrapolated curves were used for the numerical simulations of the crash box . 321 The a rtificial materials were applied to 5 mm wide strips along the crash box corners ( Figure 4 (a)). Voce -law 322 parameters , the energy absorption and the peak force of the corner -strengthened crash boxes are shown in Table 323 3. The force-displacement diagram of the U1400 corner -strengthened crash box is compared with the DP600 324 10 crash box in Figure 10 (b). The force-displacement diagram of the other strengthened crash boxes are located 325 between the shown graphs. As it is obvious in the diagram s; even though the energy absorption of the 326 stren gthened crash box is increased considerably, the peak force is almost the same as the non-heat-treated DP 327 600 crash box . According to the FE results, by applying U1400 locally to the corner of the crash box, the energy 328 absorption is increase d up to 32% without any change in structural and folding behavior. This increase is 329 correlated to the improved balance of strength and ductility caused by the increased UTS. Thus , a component 330 motivated local -material requirement can be defined as follows: 331 332 The corners of a crash box should have a hi gher strength level compared to the base material 333 334 However , in common steels an increase in strength goes along with a decrease in ductility which 335 subsequentially leads to an earlier failure event. Therefore, a local strength increase of a crash box might be 336 limited by the possibility of an early local fracture. Even though large cracks during the fold creation should be 337 avoided, Liu et al. showed that the influenc e of local fracture appearance after folding has a negligible effect on 338 the energy absorption of a crash box [10]. 339 As it was discussed, a n improved performance can be achieved by grading the material properties on the 340 crash box . A low yield stress at mid -plate region results in progressive buckling of the crash box and a high 341 stiffness at the corners increases its energy absorption. However, a minimum required ductility at the corners 342 should also be considered in the design to prevent large cracks formation, as the highest deformation takes 343 places at the se regions . This minimum ductility acts as a boundary condition for the design. The radar chart of 344 Figure 9 depicts the design requirements for corner -strengthened crash boxes . 345 346 Figure 9: Requirement definition of the corner strengthened crash box (depicted as a radar chart) and 347 corresponding microstructures (schematic). 348 3.2 Isothermal heat treatments 349 The component simulations showed that an increase of strength at the corners of the crash box leads to an 350 improvement of the overall crash box pe rformance, see Figure 10. In order to fulfill the proposed requirement 351 and achieve similar improvements in the experimental component tests, the local microstructure, and thus the 352 mechanical property profile of DP600 steel in the crash box corners needs to be adapted. The suitabl e 353 microstructure adjustment can be achieved by isothermal heat treatment. During this heat treatment, the material 354 is held at a temperature within the austenite -ferrite or fully austenitic range before it is quenched to room 355 temperature. The rapid cooling transforms the austenite into martensite, resulting in a ferritic -martensitic or 356 fully -martensitic microstructure whose phase fraction composition, morphology, grain size, and local chemical 357 composition depend on the isothermal heat treatment temperature. The choice of a suitable isothermal holding 358 temperature is therefore important to fulfill the requirements from the component simulations. To propose a 359 suitable process window for the process -adapted heat treatment, isothermal heat treatments at different 360 temperatures were carried out. In Figure 11 (b) it can be seen that the martensite (M) phase fraction, determined 361 in light optical measurements, increas es between 775 °C and 850 °C from 26 % to 98 %. At the same time, the 362 amount of ferrite (F) in the microstructure decreases proportionally. At higher holding temperatures, the 363 11 microstructure remains similar and the martensite fraction shows a low variation of 1-2 %. A similar trend can 364 be found for the hardness profile. The hardness increases with raising the temperature from 235 HV up to 413 365 HV where a plateau is reached. The constant hardness and the high martensite content above a holding 366 temperature of 850°C indicate that Ac3 temperature has been reached. This hypothesis was verified by 367 dilatometry studies. According to the investigations, the Ac3 temperature lies at 837 °C. 368 369 (a) (b) (c) Figure 10: (a) Extrapolated DP600 steel and the artificial flow curves defined for corner -strengthening of the 370 crash box ; including maximum PEEQ values experienced in the simulation (stars), (b) FE results of DP600 371 crash box and U1400 -stren gthened crash box . Force -displacement diagram s of other strengthened crash boxes 372 are located between the shown graphs ; (c) Absolute and proportional (compared to DP600) energy absorption 373 for each artificial material 374 (a) (b) Figure 11: (a) Influence of the isothermal holding temperature on the hardness and martensite phase fraction of 375 heat-treated DP600 ; (b) L ight optical microscopy images of the heat -treated DP600, natal etching. 376 Based on the previous results , tensile test specimens were processed at a holding temperature ranging from 377 800 °C to 875 °C. Figure 12 shows the stress -strain curves of the heat-treated DP600 specimens . It can be seen 378 that y ield strength and ultimate tensile strength increase with higher holding temperature, while the un iform 379 12 and total elongation decrease. However, the stress -strain curves at a holding temperature of 850 °C and 875 °C 380 show similar behavior , which correlates to the microstructural and hardness findings. 381 382 383 Figure 12: Stress -strain curves of heat -treated DP600 samples. 384 From the tensile test curves, it is clear to see that the strength level required by the component simulation , 385 i.e., U1400, cannot be achieved completely, but at least the desired UTS level can be approximated by an 386 isothermal heat treatment at 850 -875 °C. Thus, a reasonable process window could be defined for the process - 387 adapted heat treatment : 388 389 According to isothermal heat treatment , the lo cal property requirements can best be met by isothermal 390 heat treatment of DP600 steel for 60 s at a holding temperature of 850 °C. 391 3.3 Process -adapted heat treatments 392 The transfer of the process window determined by isothermal heat treatment in salt baths to process -adapted 393 heat treatment using a moving heat source requires additional investigations. To investigate the phase - 394 transformation behavior of the DP600 steel , dilatometric measurement s under continuous cooling conditions 395 were conducted. The results are shown in Figure 13. Based on the CCT diagram, a feasible process window to 396 achieve the desired mechanical properties can be determined . In accorda nce with this study , local heat treatment 397 should achieve temperat ures above Ac3 for the material to be fully austenitic. The cooling rate after the heating 398 process determines the distribution of austenite, ferrite and martensite. Higher cooling rates lower the 399 transformation temperatures to around 350 °C and 290 °C and therefore increase the austenite ferrite ratio . 400 401 402 Figure 13: CCT -diagram measured with cooling rates between 200 K/s and 2 K/s . 403 The parameters for the welding power source were derived by these measurements and chosen as high as 404 possible without creating a weld pool on the surface of the material . For the process -adapted heat treatment of 405 tensile specimens and crash boxes, a weld ing current of 8 A and a voltage of 12 -15 V were proven to be suitable. 406 The temperature me asured during heat treatment is displayed in Figure 14. These results were obtained from 407 the opposite side of the heat source (see Figure 8) to minimize errors introduced by radiation emitted from the 408 arc. Clearly visible is the difference of time exposed to temperatures above 400 °C for the two different 409 13 configurations “S” (spot) and “L” (line) . To achieve the targeted effect of higher UTS th e measured 410 temperatures are insufficient because the desired holding temperatures have not been reached . However, d ue to 411 the highly inhomogeneous nature of the temperature distribution around the arc, it is possible areas even closer 412 to the heat source may have reached higher temperatures than those that were measured on the opposit e side of 413 the heat source, so it is l ikely that the desired temperature was reached. Therefore, the samples treated by a 414 movable heat source might perf orm similar to those that were heat -treated in salt baths. 415 416 417 Figure 14: Local t emperature experienced during the heat treatment for the two configurations “S” (spot) and 418 “L” (line) . 419 In order to assess the different pattern s applied to the crash box depicted in Figure 8, tensile specimens 420 with identical process -adapted heat treatment strategies were produced and tested. These tests we re 421 accompanied by a full -field strain measurement using DIC to obtain the difference in local material properties 422 at the location of heat t reatment. The results obtained from the analysis of local strains are displayed in Figure 423 15. There are clear distinctions at the location of local heat treatment in comparison to the untreated areas. The 424 positions are chosen in order to demonstrate the effect of local heat treatment as they display the highest strain 425 in the sample . The point of failure is not evaluated because of the shortcomings of DIC. The speckle pattern is 426 not stable if there is a crack in the sample, therefore the positio n with the next highest strain wa s chosen 427 throughout. The spot heat -treated (S) tensile test sample displays a difference of local strain by the factor of two 428 compare d with line heat -treated (L) at UTS. The projected raising of the UTS from around 800 MPa to over 429 1400 MPa is not visible. However, a notable deviation of the balance between strength and ductility up to UTS 430 is visible between the different heat -treated sp ecimens. 431 432 433 Figure 15: Engineering strain field of tensile specimen moments before failure with spot heat treatment ( a), line 434 heat treatment ( b) and spot double dense heat treatment ( c). 435 14 (a) (b) Figure 16: (a) Global true stress and true strain curve; (b) Local true stress -strain curve of heat -treated tensile 436 specimen at the depicted positions. Dissipation parameter s (D) were calculated up to the UTS. 437 The tensile test result s of the different heat -treatment strategies are shown in Figure 16. As plastic 438 deformation is the prevailing mechanism of the energy absorption in the crash boxes, the flow cu rves were 439 extracted and compared with the flow curve of the base material up to UTS. The material behavior after UTS 440 is highly dependent on the stress state, which in itself depends on factors like geometry and the heat treatment 441 pattern. Therefore, the ex tent of inference on crash boxes is very limited . The tensile test results show the 442 applied heat treatment process es do not result in a higher UTS. On the contrary , the results show a decrease in 443 UTS for “L” and S. The reduction of UTS is probably due to the huge temperature gradient induced by the arc. 444 However, the resulting property gradient displays an improved ductility and dissipation parameter until UTS 445 for several heat treatments ranging from T=62.62 MPa with “L” heat treatment to T=135.25 MPa with “S” heat 446 treatment , which affec t the energy absorption of the crash box es. 447 3.4 Compon ent assessment 448 Performance of the process -adapted heat treatment 449 The heat-treated crash boxes were teste d under quasi -static compression , see Figure 17. The experiment results 450 show that t he process -adapted heat treatment s do not affect the folding behavior of the crash box considerably . 451 In quasi -static compression of a crash box , the first fold is produced randomly along the crash box , based on its 452 imperfections. Due to the regular folding pattern, crash boxes have a specific fold length , thus based on the first 453 fold location a small devi ation in the number of the folds can be seen in the tested crash boxes . In the conducted 454 tests, “S” and “SDD” crash boxes have eight folds, whereas “N” and “L” crash box es have seven folds ; however, 455 some small -size unfolded regions remain at both ends of “N” and “L” crash boxes . Therefore it can be concluded 456 that the applied local heat treatments do not affect the fold length of the crash boxes. 457 458 459 Figure 17: Compression test result of the crash boxes with different heat treatments: non-heat-treated (N) , line 460 heat-treated (L) spot heat -treated (S) and spot double -dense heat -treated (SDD ). 461 The resulting force -displacement curves of “N”, “L”, “S” and “SDD” heat-treated crash boxes and overview of 462 the extracted results are shown in Figure 18 (a) and Table 2 respecti vely. “N” crash box represents the reference 463 configuration for the comparison . Due to the deviation in the number of the produced folds, absorbed energy 464 over crushing distance diagrams of the crash boxes were considered in extracting energy absorption change of 465 50 mmN L S SDD 15 them. In order to extract the effect of the local heat treatment on energy absorption of the crash boxes, first 466 energy abs orption o ver crushing distance wa s extracted using equation (2) for all the crash boxes . Subsequently, 467 the abs orbed energy of each crash box wa s correlated with the reference configuration and a regression line wa s 468 fitted to the curve. The slope of the fitted line shows the level of the energy -absorption change due to the local 469 heat treatment . An exemplar is shown for “S” crash box in Figure 18. The s lope of the fitted line is 1.02 8, which 470 shows a 2.8 % energy absorption increase compared to the “N” crash box . The energy absorption change s of 471 all the crash boxes were extracted using the same approach . A similar effect to “S” crash box can be seen for 472 the “SDD” crash box as well , which has an energy absorption increase of 2.2 % . In contrast, the “L” crash box 473 has an energy absorption decrease of 4 %. Obviously, a dditional factors have resulted in a negative effect on 474 the component performance. The particul ar metallurgical reasons for this behavior have to be the subject of 475 further studies. 476 (a) (b) Figure 18: (a) Force -displacement diagrams of all tested crash box es; (b) E nergy absorption of “S” crash box 477 over energy absorption of “N” crash box with a linear fit. 478 Table 2: Compression test results of the locally heat -treated crash boxes . 479 Heat treatment pattern N L S SDD Dissipation parameter (MPa) of corresponding tensile test 108.02 65.62 135.3 124.3 Peak force (kN) 116.3 115.4 112.8 115.4 Energy absorption experiment (kJ) 6.37 6.11 6.55 6.51 Normalized Energy Absorption (%) - -4.0 2.8 2.2 480 In this part, the energy absorption difference of “S” and “N” crash boxes is compared to have a more robust 481 data evaluation. The difference in energy absorption, ΔE, between the crash boxes was extracted over the 482 crushing distance, see 483 Figure 19. Unsynchronized local buckling of the crash boxes results in fluctuation of the ΔE curve. Using a 484 linear fit a smoothing of ΔE was achieved, and the energy absorption increase, owing to the corner strengthening 485 of the crash box, is obvious during the crushing. It can be seen that the “S” crash box continuously absorbs 486 more energy than the “N” crash box. 487 Different process -adapted heat treatment s lead to specif ic energy absorption potentials. According to the 488 global flow curves (see Figure 16) of the process -adapted heat-treated specimen s, the highest energy absorption 489 potential could be expected for “N” crash box. Nevertheless , “S” crash box and “SDD” crash box show a higher 490 energy absorption. Thus , the corresponding local material properties that are related to the individual heat 491 treatment pattern s should be taken into account. Therefore , the balance of strength and ductility (dissipation 492 parameter in equation 2) is defined as the reference material property for comparing each process -adapted heat 493 treatment . Figure 20 displays the crash box energy absorption over the dissipation parameter for each process - 494 adapted heat treatment. The graph indicates a linear correlation between the dissipation parameter and energy 495 absorption. Starting from the reference , “N” crash box, a lower dissipation parameter leads to an energy - 496 absorption decrease, while a higher dissipation parameter results in a n increase in energy absorption. The reason 497 for this correlation can be found in the local material properties in the corners of the crash box . Crash boxes “S” 498 16 and “SDD” show increased early strain hardening without significant reduction of UTS (Figure 16) so that their 499 dissipation parameter is higher compared to the reference crash box “N” . In contrast , crash box “L” shows a 500 slightly improved strain hardening at the early stage of yielding , but due to the different thermal conditions 501 compared to “S” and “SDD” crash boxes, UTS is significantly lowered compared to the reference crash box 502 “N”. In total, this results in a decreased dissipation parameter , accordingly , the energy absorption is decreased 503 compared to the reference configuration “N”. Comparing “S” and “SDD” crash boxes , owing to the interaction 504 of the spots in “SDD” heat-treatment pattern, the dissipation parameter of “SDD” is below “S” pattern, which 505 results in lower energy absorption as well . Likewise, the effect of the dissipation parameter on the crash box 506 energy absorption can be approximated as shown in Figure 20. The presented results serve as a good example 507 to demonstrate that materials design and component design show pronounced interactions that should be 508 considered to establish new lightweight eng ineering potential. 509 510 Figure 19. Energy absorption difference between “S” crash box and “N” crash box (reference) over crushing 511 distance. The positive slope of the fitted line shows enhanced energy absorption of the “S” crash box during the 512 crushing. 513 514 Figure 20 Energy absorption change of the crash boxes due to the dissipation parameter change at the crash box 515 corners. The linear trend indicates the strong influence of corner material properties on overall performance. 516 Performance of the isothermal heat treatment 517 Since the isothermal heat treatment in salt bath was not applied to the crash box , owing to the limitation 518 of the applied process -adapted heat treatments, the component assessment was carried out based on FE analysis. 519 Recommended isothermal heat -treated (I SO-HT) material behavior , discussed in Section 3.2, was applied to 520 the corner of the crash box as described in Figure 7. The s imulated crash box show s a 53% energy -absorption 521 increase compared to the DP600 reference model , which is much more than that of the U1400 strengthened 522 crash box , i.e., 32%, see Table 3. Comparing the energy absorption of the U1400 and ISO-HT strengthened 523 crash boxes , even though the ISO-HT steel has lower ultimate stress than U1400, its energy absorption is higher. 524 The higher energy absorption is related to the higher strain -hardening in the early phase of yielding in ISO -HT. 525 The high energy absorption can be explained by considering that the significant portion of the crash box is 526 deformed plastically at relatively small equivalent -plastic strains , and the ISO -HT affects this portion much 527 more than the U1400 . However, the strong hardening at the beginning of the ISO -HT flow curve results in an 528 unfavorable sharp increase of the peak force as well. In contrast to the high peak force of ISO-HT strengthened 529 17 crash box, peak force of U1000, U1200 and U1400 strengthen ed crash boxes are almost the same as the DP600 530 reference model (Table 3). The sharp increase of the peak force can be explained by analyzing the fold -creation 531 mechanism. In a plastically buckled crash box , by increasing the stress beyond the y ield point, material stiffness, 532 E, decreases significantly, which results in local buckling and fold creation on the crash box , see equation (4) . 533 Therefore t he fold is created at relatively small plastic strain values . Thus, both the yield stress and hardening 534 behavior of the material in the low strain region affect the peak force . For a better investigation, a new m aterial 535 card, named U1400Y952, wa s defined with the same yield stress as ISO-HT and higher ultimate stress of 1400 536 MPa . The new material h as a lower hardening behavior than ISO -HT at low strain values , see Figure 21 (a). 537 Comparing the crash boxes strengthened with ISO-HT and U1400Y952 material cards; even though both absorb 538 almost the same amount of energy, U1400Y952 has a 5.3% lower peak force due to a lower hardening behavior , 539 see Table 3 and Figure 21 (b). In sum, corner strengthening of crash boxes using a material behavior with a low 540 yield stress, low hardening and high ultimate stress enhances the crash box energy absorption and prevents the 541 destructive increase of the peak force . 542 (a) (b) Figure 21: (a) Material cards used for FE simulation ; (b) FE results of DP600 , ISO -HT strengthened and 543 U1400Y952 strengthened crash boxes . 544 Table 3: Extracted parameters of the Voce hardening functions for extrapolating the materials and energy 545 absorption value extracted using the validated FE model . 546 Material K0 (MPa) Q (MPa) b (-) UTS (MPa) Fpeak (kN) Energy Absorption (kJ) Normalized Energy - Absorption Change (%) DP600 388.7 447.4 17.9 841 95 6.13 - U1000 388.7 621 19.1 1000 95 6.2 1.14 U1200 388.7 829.3 20.3 1200 97.5 7.23 17.94 U1400 388.7 1023 20.6 1400 98.5 8.08 31.81 ISO-HT 952 387.1 146.1 1359 137 9.36 52.69 U1400Y952 952 453.9 20 1400 129.4 9.42 53.67 Benefit s of corner -strengthening on crash box progressive -buckling behavior 547 Enhancing the energy absorption capacity of a crash box should not endanger its progressive buckling. 548 Using equation (4) elastic buckling stress o f the experimentally tested crash box can be calculated as 1888 MPa . 549 The calculated buckling stress is higher than the yield stress of the DP600 steel, i.e. , 389 MPa , and it is expected 550 that the crash box buckles progressively , which is confirmed by the experimental results (Figure 17). In order 551 to explore the advantages of the corner -strengthening of the crash box over homogenous heat treatment of it , 552 three simulations were conducted using the validated FE model. By decrea sing the crash box thickness to 1 mm, 553 the critical buckling stress of the crash box decreases to 451 MPa. The buckling stress is below the yield stress 554 of the I SO-HT DP600 steel, i.e. , 952 MPa. However, the buckling stress is still beyond the yield stress of the 555 DP600 steel . Two simulations were conducted on crash boxes using the DP600 steel and ISO-HT material cards. 556 The results are shown in Figure 22 (a) and (c) . As it was predicted, the crash box made by the ISO-HT steel 557 buckles unprogressively, but the crash box made by DP600 steel buckles progressively. For longer crash boxes , 558 18 this unprogressiv e behavior can result in global bending of the crash box , which prevents the energy absorption 559 process . A corner -strengthen ed crash box was simulated in addition to the previous simulations. Material 560 properties of the DP600 steel were used for the mid-plate region and ISO -HT material card was applied to the 561 corners (see Figure 21 (a)). The corner -strengthened crash box shows progressive behavior and produces the 562 same number of the folds as DP 600 crash box, see Figure 22 (b). Thus, by corner -strengthening the crash boxes , 563 the energy absorption can be increased without any critical change in the crushing behavior . 564 (a) (b) (c) Figure 22: Crushed crash boxes (1mm thickness) of (a) DP600 steel ; (b) Homogenous heat-treated using I SO- 565 HT; (c) ISO-HT-strengthened crash box . 566 567 Evaluation of component -driven appr oach in the context of other methodologies 568 An increase in component performance can be achieved through either material, process or design techniques 569 [23,26,34] . In most cases, material -technical approaches to enhance the component performance first obtain 570 improved mechanical material properties by means of heat treatment strategies, which then positively influence 571 the component behavior [34]. The mentioned approach would possibly have led to a closer match of the 572 experimental and FEM results of the crash box performance. However, the focus of this investigation was to 573 postulate an inverse material -process -component design method, the component -driven approach, which 574 reverses the process chain perspective. Although this approach did not lead to an increase in energy absorpt ion 575 of 53 % as c alculated in the simulations owing to the points already discussed, it may in the future enable a 576 comprehensive, integrative material, process and component design. Thus, resulting in strong lightweight and 577 sustainable construction potentia l in the future. In addition to the mechanical properties addressed in this paper, 578 this approach makes it possible on the one hand to extend the range of properties and requirements to include 579 for example corrosive, welding and functional properties and on the other hand to fully exploit local property 580 profiles. The authors see further potential to increase the integrated component performance through a 581 numerical consideration of the process control and the associated microstructure evolution. Both the pote ntial 582 transferability of the methodology to other scientific and technical challenges and the flexibility of the approach 583 offer further potential for the presented approach. 584 5. Conclusions 585 In this paper , a component -driven approach is proposed and demonstrated on a crash box. It was shown 586 that a requirement profile for local material properties derived from the component can lead to energy absorption 587 increase . Starting from a calibrated and validated FE model of the crash box, a corridor of desired material 588 properties could be specified with simulations using artificial material cards and corner -strengthenin g. 589 Subsequently, isothermal heat treatment strategies and material characterization methods, including light - 590 optical microstructure, hardness and tensile property measurements , were used to achieve the required property 591 profile and to propose a process window for the process -adapted heat treatment. In order to transfer the given 592 process window to a flexible process with a moving heat source, differ ent heat treatment patterns were 593 characterized using thermal and mechanical investigations , The heat treatment patterns were applied to the crash 594 box corners, to strengthen the corners and enhance the energy absorption. Then, the crash boxes were tested 595 under axial quasi -static compression. The results show an increase in energy absorption by 2. 8 % compared to 596 the reference un -heat-treated crash box. Correlating the energy absorption to the dissipation parameter of the 597 heat treatment s patterns, a linear trend is observed, which indicates the strong influence of local material 598 properties at the crash box corners on the overall component performance. 599 The experiment results validate the potential of the inverted process chain; h owever, the numerical results 600 show a considerably higher energy -absorption increase for the strengthened crash box using the material card 601 19 extracted by the isothermal heat treatment. This difference is related to the limitations of the applied process for 602 local heat -treatment of the crash box; using this process the defined temperatures by the isothermal heat 603 treatment could no t be achieved. This process has to be studied and improved further in order to enable the 604 application of this technique in the future. Fast -moving modified welding processes have a good prospect of 605 success in this domain. 606 Even though the full potential of corner strengthening of crash boxes could not be achieved by the chosen 607 means of process -adapted heat treatment, the inverted process chain showed high potential for integrated design 608 approaches. However, further research is needed to implement additional property requirements and methods 609 such as microstructure evolution and process simulations into the inverted process chain description 610 Considering the structural behavior, this research shows that local heat treatment has advantages over 611 homogenous heat t reatment of a crash box. An increase in the whole crash box stiffness can result in non - 612 progressive buckling of the crash box. But, the local heat treatment can increase the energy absorption without 613 endangering the progressive -buckling behavior of the crash box. 614 The e ffect of corner strengthe ning on crash box features was studied by conducting parametric studies on 615 the material properties of the corners and the optimal material properties were extracted. The results show high 616 ultimate stress besides a low yield point and low hardening behavior at the beginning of the flow curve not only 617 enhances the crash box energy absorption but also prevents destructive increase of the peak force. 618 619 Author Contributions: Conceptualization , N. A. , Y. S., J. Z. , K. M. , R. S. , S. M. and K. S. ; methodology, N.A., U. R., K. 620 S. and S. M.; software, N. A. ; validation, N. A. , Y. S., J. Z. and K. M. ; formal analysis N. A. , Y. S., J. Z. and K. M .; 621 investigation, N. A. , Y. S., J. Z. and K. M .; resources, U. R., K. S. and S. M .; data curation, N. A. , Y. S., J. Z. and K. M .; 622 writing —original draft preparation, N. A. , Y. S., J. Z. and K. M .; writing —review and editing, N. A. ; visualization, K. M .; 623 supervision, R. S., U. R., K. S. and S. M .; project administration, U. R., K. S. and S. M .; funding acquisition, U. R., K. S . 624 and S. M . All authors have read and agreed to the published version of the manuscript. 625 Funding: This research received no external funding . 626 Conflicts of Interest: The authors declare no conflict of interest . 627 Data availability: The raw/processed data required to reproduce these findings cannot be shared at this time due 628 to time limitation s. 629 6.