Prediction of charge welds in hollow profiles extrusion by FEM simulations and experimental validation

Transcription

1 DOI /s ORIGINAL ARTICLE Prediction of charge welds in hollow profiles extrusion by FEM simulations and experimental validation Barbara Reggiani Antonio Segatori Lorenzo Donati Luca Tomesani Received: 7 August 2012 / Accepted: 17 June 2013 # Springer-Verlag London 2013 Abstract In direct extrusion of aluminum alloys, billets are discretely loaded into the press and joined by the high hydrostatic pressure field. The contamination of the billet-to-billet interface by oxides, dust, or lubricant produces a welded zone (charge weld) with reduced mechanical properties that requires profile discharge. For an efficient material scrapping, both the position of the transition zone and its extent in the profile must be accurately identified. In industrial practice, in relation to a lack of experimental and numerical studies on this specific matter, the determination of the zone to be discarded is still performed mainly by experience or labor-intensive analyses. The aim of the present study is to bridge this gap by investigating the evolution of the charge welds inside an industrial multi-profiles and determining their exact position and extension by experimental microstructural analyses coupled with comprehensive 3D FE simulations performed with the Arbitrary Lagrangian Eulerian code HyperXtrude. Skin and rest defects are also experimentally investigated and a numerical sensitivity study on the influence of the friction model selection is performed. Comparison between numerical and experimental results shows a good agreement both in terms of general trend and exhausting points of the charge welds. The results prove that the FE code is a reliable tool in assisting and driving the die and process design stages, not only for process optimization as reported in literature but also for the scrap length determination. Finally, a process efficiency index is defined and, for the specific case study, it is found to be increased from 82.6 %, as resulting from the actual industrial practice, to a 88.3 % as optimized by the performed coupled experimental and numerical activities. B. Reggiani : A. Segatori : L. Donati (*) : L. Tomesani DIN Department of Engineering for Industry, ALMA MATER STUDIORUM University of Bologna, Viale Risorgimento 2, Bologna, Italy l.donati@unibo.it Keywords Extrusion. Welding. Charge weld. Seam weld. Finite element method (FEM) 1 Introduction Nowadays, aluminum extruded profiles are used extensively in several types of applications in civil and industrial engineering, furniture design, and in transportation sectors (motorbikes and automotive frames, buses, trucks and trains, up to airplane primary structures). The main challenge of the extruders and die makers is to achieve the optimal compromise between tighter geometrical and mechanical tolerances and higher productivity that can be reached only by a deep and comprehensive understanding of the process itself. Aluminum extrusion is a complex forming process that involves severe deformations and complex friction and heat transfer conditions that, for the past two decades, has been merely optimized on the basis of experience, intuition, and trialsand-error experimental procedures. Although this method led to a progressive adjustment of the processing technique, nowadays it is not compatible with the requirement for high-speed and efficient process development. In this context, finite element modeling (FEM) has become a common adopted tool in supporting extruders and die makers to optimize the extrusion process saving expensive and time consuming experimental procedures. A number of dedicated codes are present on the market and a strong scientific and industrial interest on selection and validation of the FE code to be used remains mandatory so as to highlight specific strengths and weaknesses of each numerical approach in consideration of the complex phenomena involved and of the user goals [1 4]. In the field of extrusion, the Arbitrary Lagrangian Eulerian (ALE) technique, as used by the commercial package HyperXtrude (HX) [5], has been widely proven to be a suitable tool combining advantages of both Lagrangian and Eulerian methods. In the ALE method, mesh and material are detached, thus avoiding

2 inevitable mesh distortion and regeneration as seen with Lagrangian codes with large deformation. In addition, mesh is not fixed in space thus being suitable for the description of movement of free surfaces. In recent years, many works have been performed by means of the ALE code HyperXtrude. As examples, Guana et al. [6] studied the effects of number and layout of portholes on extrusion process by comparing the metal flow, temperature distribution, welding pressure, extrusion load, and die stress for different die designs. Zhang et al. [7] investigated the effects of stem speed on metal flow behavior at die exit, temperature distribution, extrusion force, and welding pressure. Bastani et al.[8] discussed the material flow and temperature evolution inside the container and die, aiming at minimizing the change in shape and mechanical properties of the extruded material. Kumar compared the HX code predictions by comparing processing load at different processing temperatures and different alloys [9]; Lof described the resistance in an equivalent bearing without using a large number of elements [10], while Zhuang simulated the sheet metal extrusion process [11]. Liu et al. analyzed the optimization of different die design parameters in the case of a large, non-symmetric, and complex thin-walled multi-cavity aluminum profile, aiming at a uniform flow velocity [12]. Zhao et al. [13] used the FE simulation to detect flow unbalances in an extruded complex thin profile and compare the outcome defects with experimental data. All these works validated the ALE formulation capability, in particular the HyperXtrude code capability, to cope the industrial extrusion complexity on a wide range of process parameters, aluminum alloys, and materials and die designs. Moving to process-related issues, the requirements on profiles quality strongly depend on the final application: if for aesthetical applications (i.e., window frames or interior design structures), the surface aspect is of primary importance, while in more severe loaded conditions (like automotive or aerospace), the mechanical properties of the profiles remain of critical importance. Several authors [14, 15] have classified and described the most common extrusion defects. Pick-up, blistering, streaks, and back-end defects play a primary role on the surface aspects, while, in terms of mechanical properties, the most critical problems are incorrect heat treatment, seam weld failure, and the presence of charge welds inside the profile. In particular, seam and charge welds defects often mutually interact, but the studies available in literature usually analyze each imperfection individually. Concerning charge weld phenomenon, transportation standards (EN :2003) require 100 % production testing at the beginning, middle, and end of each extruded bar. These tests are necessary because both extremities of the bar may be corrupted by such defects that strongly reduce the final mechanical properties [16] thus requiring defective segments to be discarded. For an accurate understanding of the extension of these segments, it is important to focus attention on three phenomena: back-end effect, seam welding, and charge welding itself. Back-end defect It is also often called contamination by billet skin because it consists in the presence of billet skin inside the final profile (Fig. 1). In the direct hot extrusion of aluminum profiles, a complex material flow is generated in the container and in the die in relation to the high friction conditions generated at the billet container interface. The friction produces an imbalanced material flow that on one side positively affects the process (reduction of contamination by billet skin) but on the other side generates dead metal zones and complex material flow inside the die. The skin of the billet is usually characterized by a different chemical composition or microstructure than the inner material. This difference can be either the result of an altered chemical composition (as a consequence of cooling from casting) or of a contamination due to impurities (oxides, dust, oil, etc.) during billet shipping, pre-heating, and loading in the press. The key parameter used to avoid billet skin contamination is the length of the discarded rest ; a longer rest allows holding more billet skin and consequently it guarantees a profile free from contaminants, while a shorter rest reduces material losses. Figure 1 shows the evolution of the billet skin during a typical unlubricated hot extrusion processing of an aluminum alloy. In relation to the high friction factor with the container, the billet skin is accumulated in the proximity of the ram and, if an adequate rest length is set, it is entirely discarded (Fig. 1 (a)), while if a too short rest is used, the billet skin can flow inside the die up to the welding chambers and eventually also into the profile (Fig. 1 (b)). In order to understand if the rest length is adequate, microstructure analysis of the profiles nearby the profile end must be performed. Specifically, when the ram stops at the end of the stroke, the profile sticks to the bearing zones, thus creating a very evident mark on the profile surface called stop mark. The stop mark is a permanent shape marker on the profile that indicates the end of a billet stroke. If the billet skin overcomes the stop mark, the contaminated part of the profile has to be cut and discarded, while if the skin remains inside the die, the profile will be skin free. In an industrial environment, this type of analysis is actually performed by cutting several slices of profile on the right and left side of a stop mark and grinding and etching them to verify the complete absence of skin contamination. The phenomenon of the billet skin contamination has been investigated in literature mainly from an experimental point of view or on simplified 2D round bar simulations [17 19]. Seam weld While open shape profiles can be easily produced by a flat die, hollow profiles require a porthole die that is split into two parts: the mandrel and the die (Fig. 2). The mandrel produces the shape of the inner cavities while the die the shape of the outer cavities. The billet is initially divided around the mandrel legs, then it rejoins in the welding chamber producing the one called seam or longitudinal weld. These welding lines

3 Fig. 1 Billet skin contamination in the extrusion of a tube (schematic): (a) long rest (no profile contamination); (b) short rest (profile contamination) are generated in the profile section all along the profile length [20, 21], thus making this phenomenon not limited to a portion of the profile and motivating the wide literature on the subject. Donati et al. [22] and Nanninga et al. [23] showed that seam welds, when the die design and the extrusion process are optimized, can exhibit a resistance equal to that of the base material. Donati and Tomesani [24, 25] investigated the influence of die design and process parameters on the effectiveness of welding, demonstrating that a significant improvement in product deformability can be obtained with larger welding chambers. In a number of studies, the prediction of seam welds quality has been investigated by means of numerical models. Liu et al. [26] analyzed the weld formation parameters in a thinwalled square tube made of a wrought magnesium alloy. Finally, Donati and Tomesani [22, 27] presented a comprehensive review of the literature focused on FE analysis of AA 6082 aluminum alloy with the aim of evaluating the fields of pressure, flow stress, and velocity on the welding surface, comparing numerical and experimental results for different welding criteria Charge weld When multiple billets are consecutively loaded into the press, as happens during continuous extrusion, at the end of a process stroke, the die remains completely filled by the material (Fig. 3a). This condition can be easily identified on the profile by the stop mark. When a new billet is loaded into the press, the old and new billets start to interact and a long transition zone is produced in the profile extruded after the stop mark. This is the so-called charge or transverse weld often called carrot, in relation to the shape of the new material into the old one. As is visible in Fig. 3c, d, this transition extends along the profile to a variable length depending on die design and processing parameters [16, 28]. As for the seam weld, the charge weld is generated at high hydrostatic pressure. However, unlike the former, the charge weld is usually contaminated by oxides, dust, or by lubricant collected during loading into the press [29]. As a consequence, the profile length affected by the charge weld must be discarded due to its inferior mechanical properties. In order to effectively remove all the critical parts, two parameters are of critical importance: the position of the transition zone on the profile with respect to the stop mark (which is usually different form hole to hole in multipleholes dies) and its extent. To date, many experimental investigations have been undertaken into the analyses of charge welds and metal flow in complex dies, but all of them have been mainly focused on the influence of the charge welds on the quality and strength of the produced profile and on interaction with seam welds. A good review of pioneering works up to 1995 on extrusion welding may be found in Valberg[19]: the works illustrated and analyzed both charge and seam welding by means of experimental trials and analytical formulations. In more recent years, Loukus et al. [30] studied the extrusion weld of AA6082- Fig. 2 Schematic representation of porthole die and seam weld localization

4 Fig. 3 Sequence of a charge weld formation in longitudinal and transverse views T4 alloy by means of tensile testing and microstructural analyses: several specimens differently oriented with respect to the extrusion direction were tested thus evidencing different mechanical proprieties to be related to local grain pattern and precipitation of Mg 2 Si particles in seam-welded zones. Zhu and Young [31] investigated the effects of charge welds on the strength of 6063-T5 and 6061-T6 heat-treated aluminum alloy columns with respect to the section slenderness on simple square and round profiles and determined the weld material properties by means of longitudinal tensile test. Bingöl and Keskin [32] discussed the effect of billet temperature and process speed on the structure and discrimination of charge and seam welds, using a profile obtained from a real manufacturing process; however, the analyses were limited to a single cross section for each process conditions and to a macrostructural level. Zhang et al. [29] reported oxide distribution and microstructure analyses in seam and charge welding zones obtained by SEM and optical microscopy but for a single cross section of the profile. In addition, to the best of the authors knowledge, only three experimental works have dealt with the matter of analysis and estimation of charge welds extension in the profile for a practical prediction of the segment to be discarded. Duplanĉić and Prgin [33] carried out a Fig. 4 a The die used in the experimental design with the profile labeling (press exit view); b profile shape a b D A C B

5 Table 1 Initial temperatures and geometries of the billets and tools involved in the experimental setup Temperature [ C] Length [mm] Diameter [mm] Billets Die 468 Container 410 1, Ram comprehensive investigation and a regression analysis on the effect of the temperature, degree of deformation, and height of the welding chamber on the charge weld length, but on a simple round tube with a fixed shape. Akeret [16] investigated the effect of alloy and die design parameters (like ports dimensions and positions) on charge weld length and mechanical properties at variable distance from the stop mark, but, again, the study was mainly based on data regression from experimental trials so that results cannot be extended to different die design. Saha [34] proposed the following empirical equation in order to estimate charge weld extension for different die geometries: d e ¼ V ð1þ A s n with d e is length of the segment to be discarded from the stop mark, V total volume of ports and welding chambers, A s cross section area of a single profile, and n number of profiles simultaneously extruded by the die. Main limit of such equation is the hypothesis of a constant material flow in all die ports and chambers, while it is commonly known that friction deeply affect material flow thus generating fast and dead metal zones [17, 18, 21]. The number of studies on charge welds that employ numerical modeling tools is also limited. Qiang et al. [28] and Zhang et al. [29] performed FE simulations of charge welds phenomenon in an AA6061 alloy by means of Lagrangian codes. The former investigated the formation and evolution of the interface between the new billet and the old billet at a very early stage of the process (less than 10 mm stroke); the latter focused the study on the oxide distribution at the interface with particular attention to oxides breakings during port filling and seam joining in welding chambers. Both investigations, however, utilized only 2D FE models and very simplified profile shapes due to the high computational cost of the Lagrangian approach, making a difficult task the extension of the results to more complex geometries. In addition, no experimental validation was performed in terms of charge weld formation and evolution. As a result, no accurate information is available today for the prediction of charge weld extension and evolution in complex industrial hollow profiles. Currently, due to this limited literature, determination of the zone to be discarded in industrial everyday practice for weld contamination is still performed by intuition, analogy with similar profile shapes, and experience. For example, it is a common industrial practice to discard 1,000 mm of profile before the stop mark; another common applied rule is to scrap a profile length after the stop mark in relation to the extrusion ratio R. As guideline, 1,000 mm for R<30, 2,000 mm for 30<R<40, and 3,000 mm for R>40 are scrapped. In more critical applications, or when imposed by standards requirements, labor-intensive microstructure investigations must be undertaken at the beginning and end of each extruded lot. This results in considerable process inefficiency, as either a large portion of profiles must be discarded or a great amount of time must be spent for scrap reduction. Finally, the determination of charge weld extent at the press is extremely critical also because it follows the die manufacturing. The opportunity to check the effects of the die design in an early stage of concept through FEM tools Fig. 5 a The four extruded profiles; b stop mark evidence in one of the extruded profiles and the experimental sample labeling with d distance from stop mark: d s =start of charge weld, d e =end of charge weld

6 numerical analyses when compared to the actual industrial value. 2 Experimental investigation 2.1 Experimental setup Fig. 6 Extrusion load allows the definition of optimal geometries in order to minimize the extent of the profile to be discarded. In consideration of these limitations, the aim of the present study is to investigate the evolution of charge welds inside an industrial multi-profiles die, in order to determine their exact position and extent, and thus determine the minimal profile length to be discarded. For the first time, this investigation is carried out systematically by experimental microstructure analyses coupled with 3D FE simulations. The numerical computation of charge weld position and extent is performed with the ALE code HyperXtrude by means of transient moving boundary simulations. A sensitivity study on the influence of the selected friction model on charge weld prediction is also carried out. In addition, the selected multi-hole die allows some conclusions to be attained with regard to the influence of design parameters on charge weld evolution. Finally, a process efficiency index is defined in relation to billet-to-billet defects and the value optimized on the basis of the performed experimental and The experiment was performed with an industrial 1,600- ton extrusion press at the Alutitan plant. The porthole die consisted of eight cavities producing four hollow profiles (Fig. 4a). The profile geometry was composed by a massive part and a thin long tongue (Fig. 4b). The inner cavity was produced by a mandrel supported by two bridges, so that seam welding occurred. A nonsymmetrical position of the profiles, with respect to the horizontal axis of the die, allowed investigation of the influence of holes placement and orientation on metal flow (profiles A-D have the massive part of the section toward the center of the billet, unlike profiles C- B) while the vertical symmetry allowed FEM computation of half the process. The minimum profile thickness was 1.4 mm, the section area of each profile 135 mm 2, and the maximum bearing length 11 mm. Four billets made of an AA6060 alloy were extruded consecutively during the experimental investigation at a ram speed of 6.2 mm/s. The temperature of the container was held at 410 C and that of the ram at 450 C. The die and the billet were preheated to 468 and 478 C, respectively. Table 1 reports the initial thermal and geometrical conditions of the experimental setup. The temperature of profile A was continuously monitored by means of a pyrometer (accuracy of ±5 C) pointed at the middle of the flat profile portion, 1,000 mm after the die exit. Extrusion load, ram speed, and profile velocity were also Fig. 7 Profile temperature acquired by pyrometer

7 the die. After a distance d e from the stop mark, the transition was completely finished and the profile was made of 100 % new billet material. In Fig. 5, the four extruded profiles are shown (Fig. 5a), together with the labeling method used to track the sections of the profiles: d is the distance from the stop mark in millimeter, positive if after the stop mark with respect to the extrusion direction, negative before. The transition between the first two billets was chosen to explore the welding phenomena interaction, thus limiting interference from previous billet material deposits. The four extruded profiles were analyzed by cutting slices every 50 mm before and after the stop mark (Fig. 5b). Each specimen was ground, polished and etched by diluted hydrofluoric acid (HF 10 %), and observed under an optical microscope. For each slice, the percentage area of the new billet was computed by means of image analysis software (AxioVision). The analysis was carried out until complete exhaustion of the charge welds. Analogous activity was performed on the right side of the profile from the stop mark to check for the presence of backend defects. Fig. 8 a 3D view of metal flow throughout the die. b Position of the bridge (dotted line) and experimental seam welds (dashed lines) for profile A at 200 mm from the tip registered in order to validate numerical simulations. The four profiles were sequentially marked in order to track profile correspondence with each die hole (A to D) and billet number. The profiles clearly showed stop marks (Fig. 5b), thus allowing stable position referencing in each profile independently of the process stroke. The charge welds appeared at a distance d s (Fig. 5b) from the stop mark, depending on different positions of the profile openings in 2.2 Experimental results In Figs. 6 and 7, the process parameters monitored during the experiment for the four extruded billets are reported. The extrusion load curve for the first billet presents the typical three-peaks trend. At the start of the extrusion, the billet was upset and then offloaded in order to exhaust entrapped air: a small peak with sudden unload is visible. This is present for all the billets. The second peak presents a smooth increase due to the low starting process speed and the filling of the empty die by the material. The extrusion was interrupted at this point for about 70 s to allow engagement of the puller. The last peak presents instead an abrupt increase as the die Fig. 9 Composition of micrographs of profile section: details of charge weld contour

8 Fig. 10 Charge weld evolution on profile B: image analysis of the profile at increasing distance d from the stop mark in millimeter was already completely filled. This variation is also visible in the profile temperature graph, where the movement of the profile for the puller engagement brought the pyrometer to focus the spot on a different measurement point; this effect was adjusted at around the middle of the billet stroke. A similar problem occurred during the second billet, where an interruption of signal is evident. The position of the seam welds in the early portions of the profiles extruded by the first billet (profiles noses) was investigated. The profiles were sliced and analyzed at a distance of 50 mm from the profile tip. A seam weld occurs as the material in deformation is divided into two flows by the mandrel bridges (in the present study Fig. 8a, flow Y fills the thinner part of profile and flow X fills the more massive one) and, after streaming through the ports, joins in the welding chamber. Welds are therefore expected to be in the proximity of the bridge symmetry axis (Fig. 8b, dotted lines). Conversely, they emerged as dark curved lines (Fig. 8b, dashed lines) shifted toward the thinner part of the profile with respect to the position of the two bridges. This effect is associated with a faster flow in the thicker part, in agreement with the results of charge welds analysis and FEM outputs that are presented later. Thereafter, the transition from the first to the second billet was investigated for presence of back-end defect. Analyses were performed on profile section starting from 50 mm before the stop mark. No section of the four profiles revealed any contamination by billet skin, which can be identified by the typical variation in microstructure color associated with oxide or carbide contamination. Then, the charge welds transition between first and second billet was investigated. The analysis was performed by combining multiple micrographs to generate a complete view of the weld within the profile, as shown in Fig. 9. It was then possible to evaluate the area enclosed by the weld line and therefore the percentage of new material at that particular position of the profile. This procedure was repeated every 50 mm all along the charge weld transition length for all four profiles. The exact positions of the start and end points of the charge welds were also determined. Figure 10 shows four steps of the sequence required to assess the evolution of the transition within the profile section. Fig. 11 Profile A, end of charge weld in flow X and begin in flow Y

9 x y Stop mark distance [mm] Profile D Percentage of new billet [%] Y New billet X Stop Mark Extrusion direction Old billet x y Stop mark distance [mm] Profile A Percentage of new billet [%] New billet Y X Stop Mark Extrusion direction Old billet x y Profile B Percentage of new billet [%] New billet Y X Stop Mark Extrusion direction Old billet Stop mark distance [mm] x y Profile C Percentage of new billet [%] Y X Stop Mark Extrusion direction Stop mark distance [mm] New billet Old billet Fig. 12 Percentage of new billet material, divided for X and Y flows, in the profiles over stop mark distance and corresponding schematic representation In particular, the different sections analyzed for profile B at increasing distances d from the stop mark are reported with the welding line highlighted. Figure 10a shows the starting point of the charge weld that always appears as a small closed loop on the section. Figure 10b indicates how the new material replaces the old one within the profile section. It is well apparent in Fig. 10c how the new material has nearly filled the thicker part of the section (flow X) and, at the same time, has just appeared in the thinner part of the profile (flow Y). Both flows X and Y converge onto the two seam welds (Fig. 8). In Fig. 10d, there is no residual trace of charge weld in part X of the profile, while the flow Y is not yet complete. The way the profile section is filled depends on the particular die design: aluminum flows faster in part X due to the bigger

10 Table 2 Charge weld starting and ending points Profile Start stop distance from stop mark [mm] X [mm] Y [mm] A 565 1, ,000 1,090 1,950 B 500 1, , ,600 C 400 1, , ,550 D 500 2, ,000 1,080 2,100 die feeder and profile section, thus producing a faster charge weld transition with respect to the bottom part. The filling of the part Y ends only at 1,900 mm after the stop mark. For profile A (Fig. 11), two results are evident in comparison with profile B. Firstly, the filling of part X by the new billet material is completed before any charge weld appearing in part Y: the charge weld appears therefore disjointed in the two flows thus increasing the total extent of it (Fig. 12, profile A). Secondly, while the charge weld in flow X presents a shorter extent than profile B (Fig. 12), in flow Y the effect it is the opposite, thus revealing a higher unbalance in metal flow. Figure 12 reports the percentage of the new billet over the stop mark distance for the X and Y flows for the four profiles. Graphical representation allows an immediate overview of the charge weld over a single profile, as well as between profiles. It may be noted how the charge welds start at different distances d s for the different profiles (in the range mm) and always at the more massive part (X) of each section. On the other hand, the complete transition length of the phenomenon strongly varies from 1,550 mm on profile C to the 2,100 mm on profile D. Moreover, for profiles A and D, the filling of parts X and Y are completely disjointed and in particular the profile D shows about 100 mm distance between the end of the upper weld and the onset of the lower one, thus producing a greater transition distance. On the other hand, in profiles B and C, the two transitions partially overlap. This is particularly relevant in profile C, where the shortest transition is seen. The die design, in particular the feeding port dimension or the localization of the profiles on the die, together with their orientation, has a great effect on the velocity unbalance. Such transition is too complex to be modeled analytically with rough calculations as with volume constancy relationship or by fixed percentage of extruded profile in relation to extrusion ratio, so that, for scrap optimization, massive microstructure analysis would be required for every new die design. Profiles A and D have the greatest transition length because their thinner sections are located at a greater distance from the die center (which causes a slower material flow). In order to minimize such a transition, the different flows should emerge at the same time inside every feeder for all profiles and, if possible, at a minimum distance from the stop mark, which represents a scrap itself. Table 2 reports the precise starting and ending points of the charge weld. Data are also reported for each of the two flows (massive, X, and thin, Y) Greater differences emerge comparing the bottom profiles (B,C)withthetopones(A,D).Thisisdeterminedbyprofile orientation in the die: top profiles present the thin flow (Y) on the outer part of the billet, which presents a slower material flow, while the bottom one present Y flow toward the center of the billet. Accordingly, the weld extension for the Y profile is characterized by later onset and longer extent for the top profiles. Noting the die symmetry, same results should be found for symmetric profile, that is between A and D on the top and between B and C on the bottom. Differences between the two sides are indeed small and have a maximum value of 100 mm between the start of the weld in B and C profiles. 3 Numerical analysis 3.1 FE model The evolution of the charge welds in the investigated extruded profiles was simulated by means of Altair HX [5]. HX is a commercial FE simulation code for the analysis and optimization of the extrusion process and dies, based on a fluid dynamic approach for modeling incompressible flows, including non-newtonian fluid behavior. Fig. 13 a The 3D CAD model of the billet as required in input in HX; b magnification of the bearing curves

11 Fig. 14 FE model of the billet with the defined components According to this type of approach, the model of the aluminum billet must copy the final shape of the extruded profile. It was therefore derived as the Boolean subtraction between a cylinder and the 3D CAD model of the die (Fig. 13a). The bearing curves were extracted by the 3D CAD model of the die and then projected on the billet surfaces (Fig. 13b). As explained before, due to the symmetry of the profile, only one half of the model was simulated focusing on profiles A and B, thus reducing the number of mesh elements and computational time. The resulting model was divided in four components due to the different boundary conditions involved: billet, porthole that included the welding chamber, bearing, and profile (Fig. 14). The computation of the charge welds evolution did not require the simulation of the tools that, for this reason, were not included in the model. However, the heat exchange at the tool billet interfaces was accounted for by setting proper values of the convective coefficients and reference temperatures for a third kind of boundary condition for the heat diffusion equation at those surfaces. The billet and porthole components were free meshed with 3D tetrahedral 4-noded elements. Due to the higher accuracy required in the bearing and profile regions, 3D prismatic sixnoded elements were used for these regions. The final model consisted of 1,323,583 elements and 538,026 nodes. The same billet temperature and ram speed of the experimental campaign were set as boundary conditions in the FE model; a temperature of 478 C was imposed to the billet as experimental recorded before the loading into the press and a ram speed of 6.2 mm/s set, producing a mean profile speed of 19.2 m/min. At the two exits, zero normal stress boundary condition was specified in order to simulate the steady state condition. A full sticking condition was initially set all around the billet surfaces, except on the bearings, according to what commonly observed in aluminum extrusion practice and validated by several papers [21, 35 39]. However, in order to evaluate the influence of different friction models on the FE model predictions in the porthole region, a sensitivity study was performed by considering other models available in the code (viscoplastic, Coulomb, and slip velocity) and then numerical predictions were compared to experimental data in terms of extrusion load, profile exit temperature, and velocity and charge welds evolution. The viscoplastic model is based on the following governing equation: t ¼ km p σe ffiffi 1 exp 1:25 σ n ð2þ 3 σ e where σ e is the effective Von Mises stress, σ n is the normal stress, and m is friction factor (0 m 1). The term k is a correction parameter respect to the original formulation of the viscoplastic model [40] that allows the material to slip if

12 Fig. 15 The boundary conditions applied to the 3D FE model of the extruded billet the workpiece temperature reaches liquidus temperature (k=1 if T<T solidus, k=0 if T>T liquidus ). In the Coulomb friction model, the critical sliding shear stress τ crit is correlated to the normal compressive stress σ p n by the friction coefficient μ (0 μ 1= ffiffiffi 3 ): t crit ¼ μσ n ð3þ For C <C crit, there will be sticking, while for C C crit sliding friction can occur. Lastly, the slip velocity friction model specifies the frictional shear stress C in at the wall as a linear function of the slip velocity: t in ¼ η C v i v BC i ð4þ where C is a friction coefficient that for aluminum extrusion has to be set to 0.3 [5], v i the velocity on the i th direction, v i BC the velocity of the boundary surface in the i th direction, and η the material viscosity. In the bearing lands, the viscoplastic friction model was used accounting for the complex interfacial conditions occurring at the bearing interfaces ranging from full sticking at the die entrance to slipping at the end of the bearing [37, 41]. The friction factor m was assumed 0.3 as report by Zhang et al. [7] and Chanda et al. [42]. A convective coefficient of 3,000 W/m 2 C was set between aluminum and the tools system (container, die, and ram) as reported in [7, 13, 39] while the experimental temperatures of the tools at the beginning of the process were used as input temperatures. The ram speed was imposed by applying the prescribed velocity of 6.2 mm/s at the billet back. The scheme of the applied boundary conditions is reported in Fig. 15. The flow of aluminum at elevated temperature exhibits a viscoplastic behavior and is thus non-newtonian. The effective flow stress of the AA6060 alloy was expressed by the Table 3 Physical, thermal, and mechanical properties of the AA6060 aluminum alloy used in the present study Density (kg/m 3 ) 2,685 Specific heat (J/kg K) 878 Conductivity (W/m K) 200 Coefficient of thermal expansion (1/K) 1E 5

13 Sellars Tegart inverse sine hyperbolic model [43] to yield the steady state effective deviatoric flow stress: " σ ¼ 1 # 1 Z n α sinh 1 A " ¼ 1 # 1 1 α sinh 1 A Q n ε exp ð5þ RT where Z is the Zener-Hollomon parameter [44]. Torsion tests were carried out at different temperatures and strain rates and experimental data interpolated in order to derive the following input values for (5) [38]: n=3.515, Q=144,000 J/mol, A= E9s 1, R=8.314 J/(K mol), α=3.464e 8 m 2 /N, and T is expressed in Kelvin. The physical, thermal, and mechanical properties used for the AA6060 alloy are reported in Table 3 as validated by Bastani et al. [8]. The weld length calculation was performed by means of a transient analysis with moving boundaries. In this type of problem, the boundary conditions for the flow and heat transfer equations are treated as time-dependent and the position of the billet ram and of the billet container interfaces are tracked during the simulation time. The mesh in the profile, bearing, porthole, and welding chamber remain fixed, but in the billet region, the elements scale down linearly in the extrusion direction at each time step. HX uses an Arbitrary Lagrange Eulerian description of motion in problems with moving boundaries, since it exhibits a good numerical stability and convergence properties with a robust mesh moving technique. Fig. 16 Evolution of the charge welds at various subsequent time steps

14 A variable decreasing number of time steps was defined, starting from the first 10 s, in which the transverse welds were supposed to evolve and the velocity and thermal fields to reach a steady state condition, to the remaining 70 s of simulation, used to evaluate the load stroke history. The maximum number of nonlinear iterations was set equal to 25 with a convergence tolerance on velocities taken to be 10 3.A 90 % of the plastic work was considered to be converted in heat [45, 46]. The total simulation time was 132 h (on a Dell T7400 with 4 Intel Xeon processors at 3 GHz, 20 GB of ram). 3.2 Numerical results and comparison with experimental data In the following, numerical results as obtained by adopting the most commonly used sticking condition in the porthole regions are firstly reported. Then, numerical predictions will be presented for the different friction models tested (viscoplastic, Coulomb, slip velocity) and compared to experimental data. In Fig. 16, the evolution of the transition between the new (gray) and the old (black) billets for the two profiles is shown. By qualitatively comparing Figs. 16 and 10, the analogies are extremely clear. The FE code correctly predicted the replacement of the billet that first affected the massive part (X) of the profile (steps 12 to 14) and then the thinner part (Y) (from step 15). Also, the asymmetric shape of the new material flowing inside the old one (step 13, profile B) is really consistent with the one experimentally obtained (Fig. 10b). In addition, the code was found able to capture the faster flow of new material in the thicker part of the section in profile A than in B, as experimentally observed. A more quantitative comparison of the charge evolution with experimental data will be described in the following. Figure 17 shows the percentage of the new billet as a function of the stop mark distance with the details of the replacement of the old billet by the new one at four simulation steps. In Fig. 18, numerical results are reported for the different friction models tested in the sensitivity study for the porthole region, including the sticking condition, together with the corresponding experimental data. By considering the simulation performed with the sticking condition, a fine agreement was found between experimental results and numerical prediction in terms of profile exit velocity with a peak error at the steady state less than 3 % (Fig. 18a). A similar accordance was observed for the profile exit temperature when the pyrometer was positioned in the planned location of the profile (after 40 s process time), with a peak difference of 7.8 % (Fig. 18b). This difference is consistent with the effect of heat exchange (which was not simulated) between the profile and the surrounding air in the travel from the die exit to the temperature monitoring point. A general slight overestimation in terms of extrusion load was finally found (Fig. 18c), but limited to 14 %. The detailed analysis of the charge weld evolution showed a reasonable agreement between the experimental and numerical trends of profile filling by the new material. Figure 18d, e reports a quantitative comparison of numerical and experimental results for profiles A and B. The discrepancy between the experimental and numerical onset of the charge welds curves is to be related to the upsetting of the billet in the container that cannot be computed by the FE code (billet diameter must be set as container diameter). In addition, as shown in Fig. 16 and experimentally observed, the simulation clearly evidenced the overlapping of the charge weld transition for the two flows X and Y in profile B (step 15) and the corresponding Fig. 17 Simulated percentage of new material as a function of the distance from the stop mark 2200 Profile B Distance from the stop mark [mm] Percentage of new billet [%] 100% OLD BILLET 100% NEW BILLET

15 Fig. 18 Comparison of the experimental vs. numerical results for different friction models in terms of exit velocity (a), temperature (b), extrusion load (c), percentage of new material over the stop mark distance for profile A (d) and B (e) separation in profile A (step 16 where the flow Y appears when the flow X has been completely filled by the new material). Concerning the sensitivity study on the different friction models, no significant differences were found among the coulomb (0.3), viscoplastic (0.3), and slip velocity friction models (0.3) in terms of loads, profile exit velocity, and temperature. Moreover, also the evolution of charge welds with respect to the distance from the stop mark (Fig. 18d, e) provided just the same pattern for all the friction models. If the effect of billet upsetting is neglected, it may appear that the aforementioned friction models provided a better accuracy with respect to experimental data. Indeed, the onset of the charge weld phenomenon occurred at a simulated distance d s nearest to the experimental value, in particular for profile A. Moreover, the comparison in terms of global pattern of charge

16 Fig. 19 The origin of process inefficiency in the extrusion production as quantified in Table 4 weldevolutionshowedanaverageerrorof13and7%for profile A and B with Coulomb, viscoplastic, and slip velocity models while values of 24 and 19 % were obtained with sticking model, respectively. However, it is also evident that the pattern of the curves did not copy the experimental behavior as the sticking model did. If the shifting related to the billet upsetting is considered (100 mm), an improved accordance is achieved between sticking model predictions and experimental values by producing a global error of 8.2 and 12.7 % for profile A and B, respectively. If the same correction is applied to Coulomb, viscoplastic, and slip velocity models, then a significant discrepancy with experimental data would emerge. Finally, if the empirical Eq. (1) proposed by Saha is applied, a constant value of 1,252 mm is found for all the profiles thus providing a clear underestimation of the profile extent to be discarded. 4 Process efficiency With the results showed above, it is clear that die design deeply influences the charge weld extent and evolution. It is now possible to perform an evaluation of the global efficiency of a specific die design. Such evaluation has two uses: first to show to process managers the efficiency status of their actual production and, second, to provide an index representative of the efficiency of different die designs directly at the die concept stage. The scrap index, which is the complement to 100 % of the efficiency index, is the sum of the different phenomena previously discussed (Fig. 19): the billet rest length b, the profile discharged before the stop mark related to billet skin contamination (d c ), and the profile discharged after the stop mark in relation to charge weld evolution (d e ). Charge weld d e From experimental as well as numerical analyses, a discharge of 2,100 mm of extruded profile is the minimal amount required in order to prevent charge weld contamination on profiles to be supplied. This value corresponds to the longest charge weld of the four profiles as the discharge is performed, for logistic constrains, simultaneously on all the profiles in the same cut. Before the study, the company was scrapping 2,000 mm for charge weld contamination; such value was selected on the basis of internal standards in relation to the specific extrusion ratio. In order to define an efficiency index, the discharged lengths have to be related to billet volume: the two values (before and after optimization) account respectively for 8.7 and 9.1 % of total billet volume. It is evident that the optimized solution produce, on this particular factor, an increase of the discharged material but the previous solution did not assure charge weld-free profiles. Skin contamination d c The d c length is deeply related to billet rest length and die volume. If a too short billet rest is used, skin contamination flows in the profile before the stop mark. For this reason, the company was scrapping 1,000 mm before the stop mark, which accounted for 4.3 % of billet volume. In an optimal choice of rest length, no skin contamination is present in the profile before the stop mark and profile scrap can then be kept at 0. Experimentally, no trace of skin contamination was found in front of the stop mark, profile scrapping was therefore unnecessary with the selected rest length. Table 4 Scrap and scrap index contributions for industrial practice and experimental-fe optimization Industrial practice Experimental-FE optimized Length [mm] Volume [mm 3 ] % billet Length [mm] Volume [mm 3 ] % billet Charge weld (d e ) 2,000 1,080, ,100 1,134, Skin contamination (d c ) 1, , Billet rest (b) , , Scrap index Efficiency index

17 Billet rest length b It is the portion of a billet not extruded and sheared at the end of each extrusion; as explained it should be calibrated to avoid contamination before the stop mark. A flow of skin can nevertheless be accepted in the die with no harmful effect as the extent of contaminated profile will remain within the length discharged for charge weld defect (Figs. 1 and 19). The company sets as standard a billet rest of 20 mm that accounted for 4.4 % of billet material. Considering that no trace of skin contamination was found in the profiles also after the stop mark (i.e., all the skin was retained in the billet rest), the rest length could be reduced to 12 mm. Such length allows to exhaust any skin contamination in the profile within the charge weld extent, accounting for a waste 2.6 % of billet material. Table 4 reports a summary of such contribution for the notoptimized production based on industrial practice and the optimal values that can be obtained either through FEM outputs or labor-extensive microstructure analysis. Experimental activities and consistent numerical analysis identified a maximum of efficiency index for optimized process settings of 88.3 % respect to the 82.6 % of the actual industrial conditions. This can be compared with other solutions of die geometries (e.g., by performing FE analysis) if a further optimization level would be reached without any trial and error activities. A comparison of the optimized condition with the company usual practice reveals that an excessive scrap level was used: it is to be noticed that although a small increase in the scrap for charge weld was necessary, a substantial reduction ( 6 % of billet material) could be performed regarding the scraps due to skin contamination. 5 Conclusions As outcomes of the work, the following observational conclusions are obtained: For the first time, the phenomenon of the charge welds evolution has been systematically investigated by means of experimental image analyses, collecting information section by section, on a complex industrial multi-holes hollow profile. The extent, onset, and evolution behavior of the charge welds for each profile have been determined also accounting for the interaction with seam welds. Aside from welds investigation, skin and rest defects were also investigated revealing any contamination of the billet skin in the profile. The charge welds evolution on the complex multi-holes profile has been also analytically modeled by means a comprehensive 3D FE model using the commercial code HyperXrude, the only actually available for this type of computation. In the present work, the code, already validated in literature against a wide range of process conditions, was further verified by the comparison with experimental collected data in terms of extrusion load, profiles exit velocity, and temperature, resulting in a good accordance. The ALE formulation allowed to achieve the required information on the transient evolution and extension of the charge welds in a low computational time if compared to the concurrent Lagrangian formulation for the same complex profile. Comparison between experimental and numerical results of the new material charge weld over the stop mark distance showed a good agreement both in terms of general trend and exhausting points. The numerical experimental comparison in terms of charge weld evolution was tested against different friction models in the porthole region (Coulomb, viscoplastic, and slip velocity models) showing that, when the billet upsetting is taken into account, the sticking friction model provided the best accordance with experimental data. The experimental analyses showed a significant influence of the feeding port dimension and of the localization and orientation of the profiles in the die on the charge weld evolution. These effects were also captured by the FE code. The FE code, with the selected set of boundary conditions, was proved to be a valuable tool to assist and drive the die and process design stages in order to select the optimal porthole channels shape and location and to set key process parameters such as rest and scraps length. The advantages of a simulation are both at the die design stage, where it can be used directly to maximize the overlapping of different material flows within the same section, thus minimizing the whole charge weld extent and on the plant, where it can be adopted to discard the minimum amount of material while avoiding expensive optical analyses over the profile sections. A process efficiency index was defined in relation to billet-to-billet defects; an optimized value of 88.6 % resulted from the performed coupling between experimental and numerical activities, thus achieving a significant improvement (scrap reduction) if compared to the actual industrial practice (82.6 %). Acknowledgments The authors would like to thank Eng. Floriano Bagagli of Alutitan S.p.A. San Marino, RSM, for the interest, support, and for his help in experimental trials. The support team of Altair is also deeply acknowledged for the details provided on HyperXtrude simulation settings. References 1. Reddy M, Mazzoni L (2007) Benchmark results. In the Proc. of the International Conference on Extrusion and Benchmark. ISBN: : Los Rios G, Mayavaram R, Reddy M (2009) Benchmark results. In the Proc. of the International Conference on Extrusion and Benchmark: