Micro-roll forming of stainless steel bipolar plates for fuel cells

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Micro-roll forming of stainless steel bipolar plates for fuel cells Buddhika Abeyrathna a,*, Peng Zhang a, Michael P. Pereira b, Daniel Wilkosz c, Matthias Weiss a a

Deakin University, Geelong, Australia, Institute for Frontier Materials, Waurn Ponds, Pigdons Rd., VIC. 3216 Deakin University, Geelong, Australia, School of Engineering, Waurn Ponds, Pigdons Rd., VIC. 3216 c Ford Motor Company, Research and Innovation Center, Dearborn, 2101 Village Rd., MI 48121, United States b

article info

abstract

Article history:

Stainless steel bipolar plates for use in proton exchange membrane (PEM) fuel cells have

Received 17 September 2018

been identified as a lighter and cheaper alternative to graphite plates. Current

Received in revised form

manufacturing of metal bipolar plates by hydroforming or micro-stamping leads to

23 November 2018

excessive stretching of the material and therefore limits the channel depths that can be

Accepted 3 December 2018

formed. Low channel depths for the bipolar plates will result in low overall fuel cell effi-

Available online 8 January 2019

ciency. In comparison, the bending-dominated deformation mode present in roll forming provides the potential to form metal bipolar plates with less thinning and to greater

Keywords:

channel depths. In this work, the roll forming process is employed for the first time to form

Roll forming

thin stainless steel sheets to micro-scale channel sections of the kind required for bipolar

Micro-roll forming

plates. This paper describes the process and machine design as well as the establishment

Longitudinal bow

of the forming methodology. Experimental trials are performed and the final part quality is

Thinning

evaluated in terms of material thinning, longitudinal bow and cross-sectional shape. The

Bipolar plates

process was numerically analysed to understand the causes of the forming problems and

PEM fuel cells

shape defects observed in the experimental trials. The results of this work show that roll forming of micro-scale corrugated bipolar sheets is feasible. Furthermore, the findings provide a summary of both the practical difficulties and the possible advantages of using micro-roll forming to manufacture improved thin metal micro-corrugations for bipolar plates. © 2018 Hydrogen Energy Publications LLC. Published by Elsevier Ltd. All rights reserved.

Introduction To reduce the anthropogenic emission of CO2, automotive manufacturers such as Ford have invested 11 billion US dollars on electric cars with the plan to introduce 40 electric cars by 2022 [1]. Recent research estimated that the number of electric vehicles sold in 2040 will account for 35% of the total

car sales [2]. Proton exchange membrane (PEM) fuel cells have attracted interest as a potential power source for electric vehicles due to their high power density, compact size and rapid start up [3]. The bipolar plate (BPP) is one of the main components of a PEM fuel cell and accounts for more than 80% of the weight [4] and 45% of the total cost in a fuel cell stack [5]. Three common types of bipolar plate materials are currently being

* Corresponding author. E-mail address: [email protected] (B. Abeyrathna). https://doi.org/10.1016/j.ijhydene.2018.12.013 0360-3199/© 2018 Hydrogen Energy Publications LLC. Published by Elsevier Ltd. All rights reserved.

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considered: graphite, polymerecarbon composites and thin metal sheet [6]. Conventional bipolar plates are manufactured from graphite, but high cost, brittleness, thickness and high overall weight pose challenges to their use in mass production [7,8]. Compression or injection moulded polymerecarbon composites BPPs are cheaper and faster to manufacture compared to graphite BPPs. However, they have lower conductivity, combined with a limited operating temperature range compared to graphite BPPs [3]. One alternative is thin metal bipolar plates, which promise superior conductivity, lower weight, high strength and reduced manufacturing cost combined with excellent mechanical properties [3,9]. A number of methods have been employed for manufacturing metal BPPs, such as micro electrical discharge machining (mEDM), lithography galvanic moulding (LIGA), electrochemical micro-machining (EMM), hydroforming, micro-stamping and the recently developed tube wringer process [10e15]. While the first three methods do not appear to be economical, micro-stamping, hydroforming and the tube wringer process are cost efficient and suitable for the mass production of BPPs. In micro-stamping, a flat sheet is formed between male and female dies in a single operation. This allows a large number of corrugations to be formed in a short time [13,16,17]. However, the major deformation mode in micro-stamping is stretching and excessive thinning of material can occur. This limits the channel depths that can be formed. Recently Bong et al. [18] showed that two stage micro-stamping reduces the amount of thinning at the bending corners and allows forming of higher channel depths. However, other problems such as springback, wrinkling and warping of material are currently limiting the widespread application of micro-stamping of thin foils of 0.1 mm thickness and below [9,19]. Additionally, it has been identified that high volume manufacturing is difficult in micro-stamping due to the high precision material handling processes required [20]. The process of hydroforming uses one tooling surface and fluids under high pressure to force the workpiece into the cavities of the tooling [13]. Lower surface roughness and less variation of channel dimensions in this process have led to higher corrosion resistance than in micro-stamped plates [14,21]. Reduced springback and a higher limiting drawing ratio are further advantages of the hydroforming process [13]. On the other hand, high pressure equipment and accompanying sealing problems make this process expensive [13]. Both micro-stamping and hydroforming, according to the literature, can only produce corrugations with a maximum aspect ratio (ratio between channel height and width) of less than 1. The power density of a PEM fuel cell improves with increasing aspect ratio (AR) and this increases its performance [22]. In the tube wringer process, a thin metal strip is moved between two rolls of corrugated shape to form long corrugated channels in the sheet width direction. Even though the process has been mainly used so far to form micro-channels for honeycomb core production [23], Nikam and Reddy [15] recently showed that the longitudinal channels produced this way can also be applied for bipolar plates. The tube wringer process allows the forming of AR as high as 1 with less than 19% of material thinning, but only achieves low

dimensional accuracy. Furthermore, studies that have used the tube wringer process to date have been limited to the forming of metal sheet thicknesses and channel dimensions substantially larger than those required for bipolar plate manufacture [24]. Roll forming is a well-established sheet forming process in which a flat sheet is formed into the desired shape incrementally by passing it through a number of roll stations [25]. Since transverse bending is the main deformation mode in roll forming, stretching of the material is low compared with other methods. This leads to low material thinning and the potential to produce corrugations with high AR. The process further combines low cost and high production rates with the ability to compensate for defects such as springback. Roll forming is already currently used for wide panel manufacturing in industry and for many other applications, such as roof sheeting, automotive structures and structural steel [26,27]. Micro-roll forming may therefore represent a cost-effective alternative for producing micro-scale bipolar plates for fuel cell applications. The micro-roll forming of 0.1 mm thick SS304L to produce three micro channels over three forming stations was first introduced by Zhang et al. [28]. Even though promising results were shown, the study mainly focused on the development and experimental validation of a computationally efficient model routine and was limited to the forming of micro channels with large radii that are unsuitable for bipolar plate production. This paper investigates, for the first time, the micro-roll forming of micro channel shapes that are representative of those currently used in metal bipolar plate production. The authors are aware that a common bipolar plate consists of micro channels surrounded by flat regions of undeformed material. Manufacturing such panels from micro-roll formed channel sections will require a secondary manufacturing operation to achieve the flat regions or would use the approach already successfully applied by Nikam and Reddy [15]. However, this study is limited to the forming of only the micro channels by micro-roll forming. The material deformation and common defects such as material thinning, final cross-sectional shape and longitudinal bow are discussed. In addition, practical difficulties and process inaccuracies are identified and further improvements are proposed to improve this new micro-forming technique. To the authors’ knowledge, this is the first time that micro-roll forming has been successfully used to produce micro channel geometry that is suitable for the application to bipolar plates.

Experimental procedure Material Stainless steel SS304L foil of 0.1 mm thickness was used and the chemical composition obtained from spectroscopy is given in Table 1. Tensile tests were conducted according to ASTM E8/E8M [29] in an Instron 5967 machine with a 30 kN load cell and bone-shaped samples cut along the rolling direction, 45
to the rolling direction and 90
to the rolling direction. The samples were prepared by Wire Cut Electrical Discharge Machining

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Table 1 e Chemical composition of SS304L foil obtained from spectroscopy measurement. Percent by weight (max) C

Mn

Si

Cr

Ni

Mo

P

S

N

Cu

0.049 1.49 0.291 19.19 7.74 0.369 0.032 0.003 0.114 0.502

(WCEDM) to achieve high precision and keep cutting forces low. A test speed of 0.025mm s1 was used giving a strain rate of 0.001 s1 . Three samples from each condition were tested and the averaged engineering stress strain curve is shown in Fig. 1a and the corresponding material parameters are given in Table 2. The anisotropy ratio was determined for each sample direction which indicated near isotropic material behaviour. The microstructure of the material was analysed by Scanning Electron Microscopy (SEM) with an angular selective backscattered electrons (AsB) detector in the in-plane direction and an example micrograph is shown in Fig. 1b. The grain size was analysed with the line intersection method and based on the standard ASTM E112-13 [30]. This gave an average grain diameter of d ¼ 14:7 mm and a ratio between material thickness to grain diameter below 10, suggesting that material behaviour is influenced by size effects [31]. Despite this, size effects were neglected in the finite element analysis below.

Experimental micro-roll forming trials The active area of a bipolar plate generally consists of approximately 50e100 corrugations. This initial investigation was limited to the forming of three corrugations with the profile shape shown in Fig. 2a. The profile shape requires the forming of very small radii, with an internal radius of just 0.15 mm at each of the corners shown in Fig. 2a. These very

small radii would be very difficult to achieve with the forming sequence that is typically used in roll forming process design; where each set of rollers at every roll station would be required to have these very small radii manufactured. To avoid this significant issue, in this work we have formulated a new forming approach, where the sheet is first roll formed into a preform shape with larger than the targeted profile radii, as depicted in Fig. 2b. Preforming is followed by a coining/pressing type forming step in order to achieve the final geometry. The depth of the preform is chosen to be higher than that of the final shape to ensure that during the final forming step, the material is compressively deformed into the corners to reduce material thinning (see Fig. 2b). One half of the corrugated profile and the proposed preform shape are shown in Fig. 2c and d respectively. The preform consists of two arcs with radius, R, and the included angle, a. Both the preform and the final profile must have the same base length, l, to allow a symmetric material flow. Eq. (1) is used to determine l, where the value of 1.16 mm in the equation refers to the total width of the desired profile shape shown in Fig. 2a. The arc length, L; of both the preform and the final profile is the same to avoid unwanted thinning or thickening. Therefore, the arc length, L; can be calculated with the radius, R, and the included angle, a, given in Eq. (2). By considering the triangle OAB in Fig. 2c, Eq. (3) can be developed. Solving Eqs (2) and (3), the values for R and a are obtained, namely, 0.305 mm and 71.8
respectively. This provides the required parameters for the desired geometry of the preform. 2l ¼ 1:16

(1)

2  R  a ¼ 2L ¼ 1:53

(2)

sinðaÞ ¼ l=2R

(3)

Fig. 1 e (a) Average true stress strain curves measured with respect to the rolling direction for SS304L, (b) microstructure of the SS304L material determined in the sheet plane.

Table 2 e Mechanical properties of the SS304L material, measured with respect to the rolling direction. Orientation of the sample Yield Strength (MPa) Ultimate tensile strength (MPa) Uniform elongation Anisotropy ratio (r) RD 45
to RD 90
to RD

364 365 346

751 751 716

0.53 0.54 0.69

0.9324 1.0102 0.8719

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Fig. 2 e (a) Geometry profile of the corrugations to be formed (dimensions in mm), (b) the preform and the final shape for the full corrugation, (c) half of the corrugation of the final part, (d) centreline of preform for half of the corrugation.

Based on the preform geometry determined, the forming sequence and the corresponding tooling was developed using the commercial software COPRA RF and is given in Fig. 3a and b. To achieve smooth material flow, the three corrugations were formed to a preform shape in the first three forming stations, followed by the forming of the final channel geometry in the fourth station. For the tool design, the conventional approach of creating the tool around the given profile geometry was employed. A schematic of the experimental micro-roll forming equipment is given in Fig. 4a. As shown, this consists of the four forming stations and a feeding station. The top and bottom roll diameters are 30 mm and 15 mm, respectively, resulting in a top to bottom roll diameter ratio of 2:1. A crank handle is used to manually drive a chain that drives the bottom rolls at each of the stations. Each of the top rolls are driven from the bottom rolls by a pair of spur gears with the same gear ratio as the roll diameter ratio (2:1). In this way, a constant line speed is achieved and slippage between the forming rolls and the strip minimised. Roller gaps are adjusted with shimming foil positioned between the top and bottom station blocks as shown in Fig. 4b. The inter station distance (SD) was set to be 37 mm and the stainless steel foil strips that were formed were 27 mm wide  100 mm long in all experiments. A high level of care was taken to align the rolls as accurately as possible. This required the development of a specific alignment procedure, as follows: (i) Firstly, the existing roll gap was examined optically using a Canon EOS single-lens reflex camera equipped

(ii)

(iii)

(iv)

(v)

(vi)

with a macro lens. This gave high resolution images of the roll gap. Based on these images, the transverse alignment of the rolls was assessed and the position of the rolls was adjusted to improve the alignment with the lock nuts shown in Fig. 4b. The roll gap was subsequently re-checked using the camera images and steps (i) and (ii) were repeated until the roll gap was visually observed to be consistent across the entire tool profile. Roll forming trials were performed and the formed profile shapes were measured after each forming station by optical microscopy. (The procedure used for optical microscopy is explained below.) These microscopy images were used to assess if the material thinning was consistent across the entire formed profile at each of the stations. An uneven thinning distribution (i.e. from one side of the profile to the other) indicated tool misalignment in the transverse direction. If the thinning distribution of the formed profile showed inconsistency at any station, steps (i) to (iv) were repeated until a final satisfactory alignment of the top forming rolls was achieved for that roll station. Steps (i) to (v) were repeated for each of the four forming stations.

In some cases, this alignment process required the adjustment of the position of the top rolls by less than approximately 10 mm, which was very difficult to achieve. Therefore, the alignment procedure often required many

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Fig. 3 e (a) New forming sequence proposed, including the coining/pressing step at station 4, (b) profile geometries for the top and bottom roll for each of the tooling stations.

Fig. 4 e (a) Schematic of the micro-roll forming experimental setup, (b) roll gap setting procedure.

attempts in order to achieve a satisfactory result for each roll station.

Measurement of thinning Material thinning can be identified as one of the main defects observed in the manufacture of stainless steel bipolar plates [32]. Material thinning was measured optically following the steps shown in Fig. 5. Most samples showed some degree of longitudinal bow after forming. Therefore, all samples were

clamped to stiff aluminium backings using binder clips (as shown in Fig. 5b), to keep the samples flat during the thinning measurement process steps. The assembly of the roll formed component, aluminium backing plate and binder clips (shown in Fig. 5b) was then mounted in epoxy resin (see Fig. 5c) and section cuts were made with a Struers Accutom 100 [33]. To avoid thermal deformation, cold mounting was employed. In this procedure, the mounting material was poured around the sample and allowed to solidify in a container filled with epoxy resin at room temperature for 24 h. Three cross-sections were

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Fig. 5 e Schematic of the process used to prepare samples for cross-sectional measurements and analysis: (a) formed part, (b) straightening with aluminium backing plates, (c) mounting in epoxy resin, (d) cutting and grinding.

obtained from each sample. After cutting, the cutting plane was ground using SiC grinding paper with 1200 grit size (see Fig. 5d). An optical 3D surface measurement system, Alicona InfiniteFocus [34], was used to capture images of the cross-

section surfaces that were prepared. An example image of a cross-section after the second roll stand is presented in Fig. 6a. The measurements were taken with 20 magnification lens and a lateral resolution of 1 mm. The thickness of the sheet

Fig. 6 e (a) Image field of a cross-section after the second roll stand and the measurement of t0 and t1 , (b) an image from the 3D optical profile measurement of the corrugations of the formed sheet, and (c) the 2D cross-section profile obtained from this measurement.

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was then measured using the Alicona InfiniteFocus measurement. To ensure that the thickness was measured perpendicular to the surface, three points were obtained as shown in Fig. 6a. The first two points define a short line on one surface, while the third point gives a point on the opposite surface to which the distance representing the thickness is calculated. A straight line is drawn between the first and the third point and the angle between the two defined lines is displayed; this needs to be set close to 90
. To determine the percentage material thinning, Eq. (4) was used, where t0 and t1 are the initial sheet thickness measured in the non-deformed material region of the cross-section and the measured thickness in the deformed region, respectively (see Fig. 6a). To plot the material thinning with respect to the profile position, the coordinates of point P1 in each measurement were used to determine the position of the thinning measurement.  Thinning ¼

1

t1 t0

  100%

(4)

Measurement of the cross-section profile of the formed parts and the tools The profile of a cross-section after each forming station was measured using the same optical 3D surface measurement system (Alicona InfiniteFocus). A roll formed part was directly placed on the sample table so that the corrugations were aligned with the axis of the table movement. Since the roll formed parts had some degree of longitudinal bow, the sample table was moved until the highest point of the sample was in focus. Then, the 3D measurement and image of the profile was generated as shown in Fig. 6b. The geometry of the crosssection profile was determined by calculating the average

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height of the surface, determined over a 10 mm wide area across the corrugations (see Fig. 6b). An example of the output graph of the cross-section geometry is shown in Fig. 6c. The measurements were conducted with a 10 magnification lens, and the lateral and vertical resolution was set to approximately 3.5 mm and 400 nm, respectively. The same procedure was followed to measure the tool profile and to analyse any profile inaccuracy that may exists on the forming rolls.

Measurement of the tool roughness The surface roughness of the rolls was measured with the Alicona InfiniteFocus following the specifications given in EN ISO 4287 [35]. The roll surfaces were scanned with a 10 magnification lens and the profile roughness measured outside the profiled portions of the rolls over a 10  500 mm area, as shown in Fig. 7a. An example of the roughness profile is presented in Fig. 7b. The cut-off wavelength was set to 80 mm, i.e., wavelengths higher than 80 mm were not part of the roughness profile and not taken into account for the roughness calculation. The arithmetic mean roughness value, Ra (the absolute mean of the roughness profile), and the maximum height of the roughness profile, Rz (the absolute vertical distance between the maximum peak and the maximum valley of the roughness profile), were determined.

Measurement of longitudinal bow Longitudinal bow was measured after forming stations 3 and 4. A Kreon Baces arm scanner [36] was used to measure the geometry of the surface of the bowed sample. Due to the high reflectivity of the material, the surface was sprayed with a non-reflective matt spray. Based on the measured surface, the

Fig. 7 e (a) Roughness scan path, (b) resulting roughness profile, (c) section cut considered for bow measurement, and (d) longitudinal bow distribution, showing maximum bow height measurement.

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bow evaluation procedure is shown in Fig. 7c and d. The surface generated was first aligned with a defined reference plane and then a longitudinal cross-section of the scanned sample was considered for bow evaluation (see Fig. 7c). The x-y-coordinates of the cross-section along the longitudinal direction of the sample were determined and used to display the shape of the bow distribution shown in Fig. 7d. The magnitude of the bow was then determined using Eq. (5). Longitudinal bow ¼

Maximum bow height Strip length

(5)

Finite element model development The roll forming process was numerically analysed with the commercial software package COPRA® FEA RF, which utilises the MSC Marc implicit nonlinear finite element analysis solver. Due to symmetry, half of the sheet (13.5 mm wide and 100 mm long), was modelled. The rolls were modelled as rigid bodies while the stainless steel foil strip was discretised with three-dimensional, eightnode, solid shell elements with 11 integration points through the thickness. A refined mesh was used in the area where the corrugations were formed and a friction coefficient of m ¼ 0:5 was assumed due to the high surface roughness measured on the micro-roll forming tooling. A Poisson's ratio of 0.3 and a Young's modulus of 210 GPa were used to define the elastic properties of the material [37]. The von Mises criterion was used together with the plastic component of the true stress strain curve in the rolling direction (shown in Fig. 1a), to define the plastic material behaviour. The boundary conditions that were applied on the strip are summarised in Fig. 8. As shown, an X e lock was applied on all top and bottom nodes on the plane of symmetry. A displacement boundary condition in the Z-direction was also applied at the first row of nodes at the front of the strip and also at the end of the strip (Z e move) to pull the sheet into the first roll stand. This boundary condition is only applied until the contact between the strip and the rolls is established. In the same period, three nodes at the end of the strip are fixed with a Y e lock boundary condition to

prevent the strip from moving downwards (in the negative ydirection) before it becomes supported by the rolls. The rotational speed of the top roll was set to be 5.72 RPM, resulting in a translational speed of 9 mm/s in the positive z-direction. The same parameters that were investigated experimentally e such as material thinning, longitudinal bow and crosssection measurements after each pass e were also evaluated with FEA. COPRA® FEA RF provides an inbuilt function to determine the thinning, where the distance between the top and the bottom nodes of the strip before and after deformation is determined. The software calculates the percentage thinning similar to the experiments by using Eq. (4). For the cross-section measurements and the longitudinal bow measurements, the coordinates of the nodes across the width and the length were considered respectively. In addition the distribution of longitudinal strain (z-direction) over the strip width (x-direction) was determined after RS3. This was used to understand the origin of longitudinal bow.

Results The results from the experimental trials show that, using the proposed micro-roll forming method, it is possible to successfully form the corrugated profile shape. This was achieved without failure of the 0.1 mm thick stainless steel foil and with reasonable accuracy of the final part shape. This Results section will describe the measured thinning, cross-section shape and longitudinal bow of the experimentally produced samples and compare these results to the FEA predictions. Finally, the tool profile shape and roughness will be presented. These results, the causes for inaccuracies and areas for improvements will be discussed in further detail in the subsequent section (Discussion).

Material thinning Fig. 9a depicts the cross-sections of the formed strip after each forming station, which were used to analyse thinning given in Fig. 9bee. Fig. 9b shows the experimental and numerical

Fig. 8 e Boundary conditions applied on the strip.

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No additional material thinning occurs in the centre corrugation after station 2, but thinning increases in the neighbouring corrugations after station 3 to values of up to 30% (compare Fig. 9c and d). Similarly, in the FEA, thinning in the centre corrugation only slightly increases from approximately 14%e18% from the second to the third forming station, but shows a large increase from 4% to approximately 15% in the neighbouring corrugations. Both the FEA model and the experimental results suggest that no further material thinning occurs in the coining step (station 4). The cross-sections of the formed strip shown in Fig. 9a confirm that no necking occurred, even though the maximum thinning reaches up to 30% in the side walls according to Fig. 9e.

Cross-section shape

Fig. 9 e (a) Cross-section images of the formed profile after each station, (a) experimental and numerical thinning distribution after the first, (b) second, (c) third, (d) fourth forming stations.

The cross-sectional shape after the first forming station is given in Fig. 10a, showing the comparison between the experimental measurement (EXP), the finite element model predictions (FEA) and the ideal profile shape. Note that the “ideal” profile refers to the desired profile shape of the top of the sheet surface for station 1 in Fig. 3b, which is based on the CAD model. It can be seen that there is a good agreement between the FEA and the ideal profile shape. However, the experiments show a 0.05 mm lower corrugation depth compared to the ideal profile. There is also a reasonable agreement between the ideal part shape and the FEA prediction after station 2 (see Fig. 10b). The depth of the experimental cross-section shape is approximately 0.04 mm shallower than that of the ideal profile at the inner region of the part and 0.07 mm shallower at the outer region of the part. After three forming stations, both the experiments and the numerical results show good agreement with the ideal profile shape for the centre corrugation, while at the outer region of the part there is a difference in profile depth of 0.07 mm that remains between the ideal and experimental shape (see Fig. 10c). Comparison of the final profile shapes after station 4 suggests that there is an insufficient amount of material forming especially in the corner regions, as shown in Fig. 10d. This leads to a rounded profile shape, rather than the desired sharp corner radii for the experiments and the FEA. In this outer part of the cross-section, the profile shape that is experimentally achieved is approximately 0.07 mm deeper compared to the ideal profile shape.

Longitudinal bow thinning distribution over the strip width after the first roll station together with the ideal profile shape. In the experiments, material thinning is non-symmetric with a maximum thinning of approximately 15% on the right and 12% on the left. Maximum thinning occurs in the side walls of the corrugations in both the experiments and the FEA, but thinning is much lower in the FEA. Thinning is the highest in the centre corrugation and after station 2 (Fig. 9c) reaches values of up to 30% in the experiments. In contrast, the maximum thinning predicted by the FEA is only 14%. Both the experimental and the numerical results show a similar distribution of thinning over the strip width. It can be seen that after station 2 that maximum material thinning in the centre corrugation remains constant at approximately 30% (compare Fig. 9b and c).

The comparison between the experimental and the numerical bow measurements is shown in Fig. 11a. Two measurements were taken, one after roll station 3 and the other after roll station 4. Good correlation between the experimental and numerical results was achieved for the longitudinal bow measurement after the third forming station. However, bow is significantly underestimated by the FEA after the fourth forming station. The longitudinal bow is concave downwards as shown in Fig. 7c. The distribution of longitudinal strain across the width of the profile after RS3 is shown in Fig. 11b). After RS3, it can be seen that higher levels of longitudinal tensile strain occur at the top regions of the profile. Conversely, lower levels of longitudinal tensile strain and

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Fig. 10 e Comparison of the part shape after the (a) first, (b) second, (c) third and (d) fourth forming stations.

Fig. 11 e (a) Experimental and numerical longitudinal bow after the third and the fourth forming stations (RS3 and RS4), (b) longitudinal strain distribution along the width of the part after station RS3.

some small amount of longitudinal compressive strain exist at the bottom regions of the profile. This distribution of longitudinal strain is likely to create a downward moment along the length of the formed part and result in concave downward bow.

Investigation of the tool profile The ideal top and bottom tool profiles are compared with those measured with the Alicona system in Fig. 12a for forming station 1. In this figure, the profiles of the top and bottom tools have been aligned to give a centre tool gap of 0.1 mm as specified in the process design. Some deviation between the designed and the actual tool profile can be observed especially in the side wall regions. If a roll gap of

0.1 mm is set in the centre, the roll gap in the side walls is reduced to 0.067 mm (see Fig. 12a). The same situation can be observed in forming stations 2 and 3, where the minimum roll gaps in the side walls are estimated to be 0.82 and 0.68 mm respectively (see Fig. 12b and c). Also in the final forming step (station 4), the roll gaps in the side wall regions are lower than the initial material thickness. In addition, the actual diameter of the bottom roll is less than that required, which in some regions of the centre corrugation leads to roll gaps that are up to 0.026 mm larger than required (Fig. 12d).

Tool roughness The measured values for tool roughness are given in Table 3. While Ra was in the range between 0.37 and 0.76 mm, Rz varied

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Fig. 12 e Measured and ideal tool profile of the (a) first, (b) second, (c) third and (d) fourth forming stations.

between 1.92 and 4.02 mm. This is significantly higher compared with the surface roughness generally observed in stamping dies, where Ra is usually up to three times lower [38].

Discussion Material thinning Both the experimental results and the FEA show similar distributions for material thinning for all forming stations, with the highest level of thinning observed in the corrugation side walls. Both the FEA and the experimental results show that material thinning is the highest in the centre and remains

constant after the centre corrugation is fully formed in station 2. There is also a general trend of reduced thinning towards the side corrugations. This may suggest that in a micro-roll forming process e where the corrugations are progressively formed from the inside to the outside of the strip e thinning in the centre may be mostly unaffected by the forming of the neighbouring corrugations and independent of the overall number of corrugations that are formed. In addition, both experimental and FEA results show no increase in maximum thinning between stations 3 and 4. This may be related to the small material movement in the transverse direction from station 3 to 4, given that the arc lengths of both the preform and the final profile are designed to be the same.

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Table 3 e Tool roughness measurements for the four forming stations. Roll Bottom Top 1 Bottom Top 2 Bottom Top 3 Bottom Top 4

1 2 3 4

Ra [mm]

Rz [mm]

0.37 0.76 0.54 0.53 0.56 0.67 0.67 0.65

1.92 4.02 2.42 2.88 3.13 3.23 3.21 3.21

The Aspect Ratio (AR) formed in the experimental trials of this study was 0.6 and this is higher than most of the ARs achieved by micro-stamping [17,21,39]. Maximum material thinning was 30% and significantly lower than the minimum uniform elongation determined in the tensile tests (Table 2). This, in combination with the findings presented below, suggests that even higher ARs are achievable especially when improved high precision tooling is used. The material thinning predicted by the FEA is up to 40% lower compared to that experimentally observed. The experimental analysis performed as part of this study suggests that this may be due to three major reasons: (i) tool shape inaccuracy, (ii) tool surface roughness and (iii) misalignment of the top and bottom rolls. These will be described in further detail below: i) Tool shape inaccuracy: The forming tool/roll surface profile measurements shown in Fig. 12 suggests that significant tool shape inaccuracies exist in the experimental tooling. By aligning the measured tool profiles and assuming the designed roll gap setting of 0.1 mm in the middle corrugation; these inaccuracies lead to locally reduced roll gaps (by up to 33%), as shown in Fig. 12. These locally reduced roll gaps are most evident in the corrugation side walls, which correlates well with regions of the highest level of material thinning in the experiments. This is also the location were the deviation between the experimental results and the numerical predictions for material thinning are the largest. The numerical model assumes the “ideal” roll profiles (i.e. corresponding to the CAD model) and a roll gap of 0.1 mm over the entire tool profile. This reduced roll gap in the experimental setup, caused by the tool profile inaccuracies, is likely to be the main reason why the experimental measurements show significantly higher material thinning compared to the numerical predictions. The authors have conducted a separate study to account for the inaccuracy of the real tool profile in the numerical analysis model. This lead to improved correlation with the experimental results for material thinning. However, additional experimental and numerical work is required to fully understand the effect tool inaccuracy and misalignment; this will be part of a future publication. ii) Tool surface roughness: According to Table 3, the roughness of the forming rolls is up to three times higher

compared to that normally observed in conventional stamping tools. Note that this higher roughness on the rolls occurred due to the compromise associated with manufacturing the highly accurate tool profile geometry on such as small scale and the need to achieve a high quality polished surface finish. This increased roughness may have restricted the material movement in the forming rolls, resulting in an increased transverse tension in the strip and therefore higher thinning in the strip. It has been shown that there is a significant influence on the surface roughness on the coefficient of friction [40]. Indeed it is well known that, especially when roll forming stainless steel alloys, a high roll surface roughness can lead to extensive material thinning in the profile radii [41]. In the numerical model of the micro-roll forming process, the higher tool surface roughness was accounted for by applying a friction coefficient of m ¼ 0:5. This is approximately three times higher than the friction values commonly used for sheet forming simulations, which range between 0.1 and 0.2 [42,43]. Despite this, the numerical model still underestimated material thinning and this suggests that the effect of tool surface roughness on friction may have been higher for the investigated micro-roll forming case. Previous studies have shown that in microforming, due to the smaller size of the forming tools, the impact of tool surface roughness on friction is more severe compared to conventional sheet forming applications [44]. To investigate this, additional numerical analysis with increased friction coefficients was performed. This additional study was limited to forming station 3 and showed a near linear relationship between the friction coefficient and material thinning. A friction coefficient of m ¼ 1 led to 25% maximum material thinning, which is close to the experimental results (Fig. 9d). Given the high surface roughness of the forming rolls observed (Table 3) such a high coefficient of friction may be justified. This suggests that much lower surface roughness will be required to achieve acceptable material flow in micro-forming processes. It further indicates that, despite the higher friction coefficient that was applied in the numerical model of this study, the friction forces on the sheet and the resulting material thinning were underestimated. iii) Tool misalignment: In all four roll forming stations, material thinning was asymmetric (see Fig. 9). While in forming stations 1, 2 and 4 thinning was higher on the right hand side of each of the corrugations; in station 3 the opposite trend can be observed. This is most likely related to a small misalignment in the transverse direction between the top and the bottom rolls. A better tool alignment would lead to a more homogeneous material flow and reduced material thinning that is also symmetric across each corrugation. However, it should be noted that in order to improve the tool alignment, very small adjustments of the position of the rolls in the lateral direction (i.e. of approximately 5 mm) would be required. This very small level of adjustment would require very precise control of the lateral position of the tools.

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Profile shape Fig. 10 shows that the ideal profile depth is not achieved in every station. This may be related to an enlarged roll gap, either due to an incorrect setting of the roll gap or the defection of the roll shaft. Distinguishing between both effects is not possible, but the experimental results of this study give clear evidence that the roll gap was larger than designed. For example, as can be seen in Fig. 12a, a roll gap of 0.1 mm set at the middle corrugation in forming station 1 would result in a reduced roll gap of 0.067 mm at the sidewalls, due to the roll profile inaccuracies. This reduced roll gap should result in material thinning of approximately 33% in the corrugation side walls. However, as can be seen in Fig. 9b the material thinning that was experimentally observed after station 1 was significantly lower, at only 15%. This suggests that the general roll gap during forming was much larger than the required 0.1 mm, which would explain the 0.05 mm deviation in profile depth between the formed profile and the ideal profile (see Fig. 10a). The same effect may have led to the deviation in cross-section shape between the formed and the ideal profile in the part regions next to the outer profile. In station 4, the experimentally formed part shows a rounded profile shape without the tight corner radii of the desired profile. This is especially evident in the corrugation peaks, as shown in Fig. 10d. This is likely due to the tool inaccuracies observed in forming station 4 which indicate an enlarged gap between the top and the bottom rolls in the peak corrugation regions (as shown in Fig. 12d). The resulting lack of tool contact and pressure could have led to insufficient material flow into the corner regions. Tool deflection and tool profile inaccuracies were not accounted for in the numerical model and therefore the FEA results represent the ideal forming case with a roll gap of 0.1 mm and forming rolls with perfect profile shapes. As can be seen in Fig. 10, the cross-section shape predicted by the numerical model is very close to the desired profile. This includes the final shape achieved in station 4 and suggests that the desired channel profile is achievable with an ideal microroll forming tool.

Bow Bow in roll forming is generally the result of an imbalance of permanent longitudinal strain between the top and bottom regions of a roll formed section [45]. Both the experimental results and the numerical prediction indicate significant downwards bow after forming station 3 (see Fig. 11). This suggests that the tensile longitudinal strain after station 3 was higher in the top profile region and generated a downwards bending moment on the profile, as shown in Fig. 11b. While a good correlation for bow is achieved between the experiment and the FEA after forming station 3, the numerical model significantly underestimates the level of bow after station 4. This may be related to tool/shaft deflection. According to the way the process was designed, in station 4 the material should be coined/pressed into the tight corner radii of the roll tooling, while in station 3 the major deformation mode is bending. Therefore, the forming forces on the tooling are likely to be much higher in station 4 compared to station 3

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and this would lead to higher tool deflection in station 4. This correlates with the results shown in Fig. 10, which indicates that the cross-section shape is almost perfectly formed in station 3, while the desired part shape is not achieved in station 4. This indicates high tool deflection and reduced material deformation in station 4 and may explain the lower reduction in bow observed experimentally, compared to the FEA results (which assumed a perfect roll gap of 0.1 mm and no tool deflection). Overall the results suggest that bow in a micro-roll formed parts can be controlled by the coining station (station 4) and may be completely removed with a micro-roll forming tool that has improved geometric accuracy of the tool profile shapes, more consistent roll gap and less shaft deflection.

Conclusion In this study, a proposed methodology for the roll forming of micro-scale corrugated sheets, relevant to fuel cell bipolar sheets, is examined. The process of micro-roll forming is experimentally and numerically investigated to evaluate the effectiveness of the proposed tool and process design, the tool robustness in the equipment and the effect of these on the accuracy of the final parts produced. To improve the understanding, defects in the cross-sections of the formed parts e in terms of thinning, profile shape accuracy and bow e at the intermediate and final stages of the process were examined. Limitations with the initial experimental setup were identified and improved tool settings implemented. Techniques to further reduce these issues and defects and improve the accuracy of the micro-roll forming process are presented. The following conclusions can be made based on the above investigation.  Roll forming of micro-scale corrugated bipolar sheets is shown to be feasible. The proposed forming process design has been shown to be successful. A significant feature of this process design is that the first forming stations involve micro-roll forming a preform shape via “traditional” roll forming stations and then subsequently coining/pressing the very tight radii into shape in the last station.  It was found that some material thinning does occur and its reduction is discussed. Local material thinning is generally greater than that predicted by FEA, but follows similar trends. Thinning is sensitive to practical difficulties such as incorrect roll gap, tool misalignment, tool inaccuracy and poor surface finish of the tooling.  Potential improvements in the design of the micro-roll forming machine have been identified from the results of these experiments and may lead to improved results. It must be emphasised, however, that the physical dimensions of the stainless steel foil and tooling are at least one order of magnitude smaller than in conventional roll forming. Therefore, changes become expensive and will involve precision manufacturing techniques to achieve extremely high manufacturing tolerances on the tool profile shapes and well-designed equipment to achieve accurate and adjustable roll gaps and tool alignment.  There is a slight deviation of the profile shape of the parts produced in this study compared to the ideal shape. It was

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identified that this is related to an incorrect roll gap as a result of shaft deflection and this might be minimised by increasing the effective shaft diameter.  Longitudinal bow can be significantly reduced by the process design proposed in this study, where a micro-roll formed preform shape is subsequently coined/pressed to the desired shape in the last forming station.  Even though the results of this study are promising, it needs to be emphasised that the presented micro-roll forming process is limited to the manufacture of micro channels that expand over the full length of the sheet. However, in most bipolar plate designs, the micro channels are surrounded by regions of undeformed flat sheet to enable plate assembly. Such shapes may be achieved via a secondary forming process subsequent to the micro-roll forming and will be part of future work.

[10]

[11]

[12]

[13]

[14]

[15]

Acknowledgement The authors would like to acknowledge the financial support by the Australian Research Council (LP150100059) and Ford Global. The authors would like to acknowledge the software and technical support provided by dataM Sheet Metal Solutions and Australian Rollforming Manufacturers. The authors further would like to thank Emeritus Professor J.L. Duncan for his assistance in writing this paper and appreciate the support from the RWTH master and internship students Christoph Muller, Marius Kaiser and Jens Muller, which was vital for the success of this project.

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