Numerical and experimental analysis of multi-channel spiral twist extrusion processing of AA5083

Materials Science & Engineering A 764 (2019) 138216

Contents lists available at ScienceDirect

Materials Science & Engineering A journal homepage: www.elsevier.com/locate/msea

Numerical and experimental analysis of multi-channel spiral twist extrusion processing of AA5083

T

D.M. Fouada,*, A. Moataza, W.H. El-Garaihyb,c, H.G. Salema a

Mechanical Engineering Department, The American University in Cairo, Cairo, 11835, Egypt Mechanical Engineering Department, Unayzah College of Engineering, Qassim University, Unaizah, 51911, Kingdom of Saudi Arabia c Mechanical Engineering Department, Suez Canal University, Ismailia, 41522, Egypt b

A R T I C LE I N FO

A B S T R A C T

Keywords: Multi-channel spiral twist extrusion Ultrafine grained material Severe plastic deformation Finite element analysis Nano-indentation

In the current study, a comprehensive evaluation of AA5083 processed via the novel Multi-Channel Spiral Twist Extrusion (MCSTE) method was conducted. The induced stress-strain state and the deformation mechanism of multiple pass deformation via MCSTE dies with twist angle β (30°) and (40°) were analyzed using finite element analysis. Nanohardness measurements were carried out along the billet surfaces to validate the numerical model output. The micro-hardness, tensile-up-to-fracture, fracture behavior and microstructural properties were investigated. The numerical model and the empirical findings reveal that the increase in the mechanical properties of the billets processed via MCSTE die with a twist angle β (40°) was associated with a plastic strain of 0.9 (mm/ mm) compared to 1.2 (mm/mm) for conventional twist extrusion dies of β (60°). For the MCSTE die with angle β (30°), the hardness and tensile properties increased as a function of increasing the number of passes, with an insignificant reduction in ductility. Processing via a die angle β (40°) was limited to one pass due to excessive strain hardening, which resulted in shear localization during the second pass. The detailed analysis presented herein validates the effectiveness of MCSTE processing as a severe plastic deformation tool with a favorable potential for industrial applications.

1. Introduction Strengthening of polycrystalline materials is a perpetual interest of researchers and a recurring subject in mechanical metallurgy. A principle discovery has been the correlation between mechanical properties such as yield, flow, and fracture stress and their dependence on macroand microstructural features such as chemical composition, crystallographic defects, grain size and texture. The infamous Hall-Petch equation signifies the precise importance of grain size and its impact on the strength of polycrystalline materials. Along with the enhancement in strength are key improvements in mechanical and physical properties, which have led to the growing notion of fabricating ultrafine grained (UFG) materials [1]. Conventionally, various thermomechanical forming techniques are used to refine the microstructure of several polycrystalline metals and alloys. Different methods such as rolling and extrusion are the most common work hardening processes. However, conventional techniques are commonly associated with a major sacrifice in ductility as a result of strengthening [2]. Additionally, they are usually restricted by the amount of strain that can be induced without major reductions in the billet dimensions. Therefore, this

*

constraint has limited their capability to refine grains only on the order of ~ 1–2 μm [1]. Moreover, 5xxx series aluminum alloys constitute the highest strength Al alloys in the nonheat-treatable category. Alloys in this series are known for their good weldability and resistance to corrosion in harsh conditions such as those present in marine environments [3]. An appreciable increase in tensile properties can be attained as a function of increasing deformation; for example, appreciable 130% and 50% increases in yield strength (YS) and ultimate tensile strength (UTS) have been achieved after 90% cold rolling reduction of AA5052 [3]. However, it should be noted that such significant increases are coupled with a significant 95% decrease in ductility, which limits the amount of cold work that can be performed. An interesting resolution to such major constraints is the processing via severe plastic deformation (SPD) instead of traditional techniques, and several studies have reported on the synergy of strength and ductility in deformed billets [4]. Recent SPD techniques have been developed with the capability of refining the grain size to less than 1 μm without the adverse dimensional reductions associated with conventional metal forming techniques. However, the integration of grain refinement and SPD

Corresponding author. E-mail address: [email protected] (D.M. Fouad).

https://doi.org/10.1016/j.msea.2019.138216 Received 20 May 2019; Received in revised form 23 July 2019; Accepted 24 July 2019 Available online 25 July 2019 0921-5093/ © 2019 Published by Elsevier B.V.

Materials Science & Engineering A 764 (2019) 138216

D.M. Fouad, et al.

Fig. 1. Modeling of the MCSTE (a) die setup and (b) billet-disk assembly through the twist channel.

number of TE passes, which is associated with an improvement in the homogeneity of strain distribution across the simulated billet post multiple pass deformation [13]. Consistent results have been reported in Refs. [14,15]. Thus, FEM has become an indispensable tool that provides useful information regarding the effective stresses and strains induced as a result of deformation and that facilitates process optimization [11]. Although there have been numerous studies on TE, most of them have been mainly dedicated to and focused on evaluating the deformation characteristics of the process and introducing different variants designed for incrementing the effective strain induced, with disregard for crucial limiting aspects necessary for industrial integration such as die durability, tool cost, and material waste. However, recently, a TE-based technique, multichannel spiral twist extrusion (MCSTE), was developed and introduced in Ref. [16] as a means of providing a modified TE die with several cost-saving features such as a lower twist angle and the use of stacked discs compared to conventional TE. In this study, a comprehensive evaluation of the novel MCSTE technique is conducted through numerical modeling to study the effective induced stress-strain behaviors and their distribution after multiple passes of MCSTE (β 30° and 40°) processing of strain hardenable AA5083. Validation of the numerical model is carried out by investigating the mechanical properties and nano-indentation of the MCSTE-processed billets after deformation.

strengthening metal forming technology in the industry has been hindered by several limitations associated with SPD processing such as upscaling, deformation uniformity, material waste associated with distortions in the deformed billets, and tool cost [5,6]. In quest of incorporating SPD methods in industry, extrusion-based SPD processes such as twist extrusion (TE) have been developed as a means of overcoming some of the aforementioned constraints [7]. In 1999, TE was introduced by Beygelzimer as a principle processing scheme that combines the advantages of both shear and torsional deformations. Inspired by the effectiveness of equal channel angular pressing (ECAP) and high-pressure torsion (HPT) in producing UFG materials, Beygelzimer presented TE as an alternative tool that aims to produce UFG bulk products characterized by long dimensions and intricate profiles that are not possible to attain via other SPD techniques [8]. Since the first publication on TE two decades ago, several empirical and numerical studies have been conducted to investigate the deformation characteristics of TE and its impact on the structural and mechanical properties of Cu, Al, and Ti alloys. Several advanced techniques such as transmission electron microscopy (TEM), electron backscattered diffraction (EBSD), and nano-indentation have been utilized to emphasize the microstructural features and ascertain the improvements in properties [9]. Moreover, efficient computer simulations have been replacing traditional experimental and analytical methods. The stress-strain states of materials subject to various bulk forming and SPD techniques have been commonly assessed with the aid of finite element modeling (FEM). The need arises from the difficulty of evaluating strains associated with deformation and their distribution along the deformed cross-section. Additionally, due to the large number of processing parameters and their impact on the flow behavior and deformation characteristics of any given system, experimentally evaluating their individual effects has been extremely challenging and costly [10]. In 2008, Mousavi and Shahab used the FE commercial code “ABAQUS 6.5” software to evaluate the magnitude of strain distribution after 3 passes of TE with a twist angle β (60°), imposing an effective plastic strain of 3.6 on AA1100 [12]. Because the microstructure evolution depends on the imposed strain, the model has been validated by metallographic images and grain size measurements corresponding with the strain distribution output from the FE simulations. Several simulations have revealed that the strain accumulated heterogeneously across the cross-section, where the highest magnitude of strain occurred at the corner and the lowest strain occurred at the center, with the strain varying from 1.8 to 0.6, after a single TE pass. Additionally, the numerical model depicted the variation of strain along the extrusion direction, where the bottom ends experienced higher strain values than the top [12]. Moreover, in 2011, Mousavi and Bahadori revealed through FE simulations of pure Al that the distribution of properties improved as a function of increasing the

2. Numerical modeling In this study, Simufact Forming 14 was employed to simulate the MCSTE novel design and to obtain the stress-strain behaviors associated with the current die setup. All geometries were imported from SolidWorks design software. The model consisted of a plunger, a twist die, and a billet-disk assembly, all of which were modeled as rigid objects, as shown in Fig. 1. The billet was modeled as a deformable object, and the AA5083 material was identified from the built-in library. Furthermore, a hexahedral mesh was employed, which is typically used in computational modeling of 3D regular shapes. Knowing that the accuracy of the model is directly related to the mesh size, a mesh refinement scheme was followed in which a preliminary coarse element size was chosen and then gradually decreased in several iterations until the solution converged. This was a key step in validating the model and gaining additional confidence in the results, which the model formulated. Finally, a mesh size of 0.5 mm, which yielded a total of 16,710 nodes solved during the simulation and consumed the least computational time, was chosen. In addition, as a result of deformation, a remeshing criterion was set to take into consideration any changes in the geometry and dimensions of the billet processing. The Johnson-Cook 2

Materials Science & Engineering A 764 (2019) 138216

D.M. Fouad, et al.

Table 1 The adopted Johnson-Cook model parameters for AA5083 [17].

Table 3 Comparative experiment matrix. Comparative Exp.

(1)

τ=mk

Material

A(MPa)

B (MPa)

N

C

M

Tm (oC)

AA5083

167

596

0.551

0.001

1

620

where A is the yield strength of the material at room temperature, B is the hardening parameter, C is the strain rate sensitivity, N is the hardening exponent, M is the thermal softening and Tm is the melting temperature. A constant temperature of 20 °C was maintained throughout the process to model deformation at room temperature. The default setting for the heat transfer coefficient between the two objects was used, and a constant value of 50 W/m2 K was applied. According to the Simufact software, the frictional shear stresses (τ) for all surfaces in contact were defined according to the equation below:

Process Parameters

Experiment #

Alloy

No. of Passes

Twist angle (β)

Exp. 1 Exp. 2

AA 5083 AA 5083

Up to 3 Up to 1

30° 40°

replacement of twist channels with different angles while keeping the remaining channels and parts the same. Moreover, to minimize friction between the disk-billet assembly and the inner die walls, the billets, inner disk cavity, and inner die channels were all well lubricated using a general-purpose graphite-based lubricant. Additionally, the support base at the outlet of the exit channel, connected to the backpressure springs, was used to create a minor backpressure of approximately 5 MPa against the disk motion during the extrusion process. Table 3 illustrates the experimental scheme followed using the 250kN Schenck Trebel testing machine. Moreover, it is important to note that strain saturation occurred in the AA5083, which led to shear localization and fracture after 3 passes for the billets processed via twist dies with a β (30°) and after 1-pass for the billets processed via dies with a β (40°). Further characterization was carried out to compare the mechanical and microstructural properties after deformation via MCSTE processing with β (30°) and (40°). Nano-indentation experiments were performed to validate the numerical model and simulation output. The test was conducted using the basic hardness technique. The indentation depth limit was 2000 nm and the peak hold time was 15 s to obtain the nanohardness values over an array of 3 × 5 indents. Thus, the test array represented points along the peripheral region of the billet-disk interface, as shown in Fig. 2, where the maximum strain is induced at the billet-disk interface (points 1 and 5) and the lower values are depicted in between (Points 2, 3, and 4). To evaluate the degree of uniformity of the mechanical properties, the extrudates were sectioned using a high-precision Isomet cutter to obtain both the longitudinal sections (parallel to the extrusion direction) (20-mm long) and the transverse sections (perpendicular to the extrusion direction). The two sections were then mounted, ground and polished to investigate the hardness, nano-indentation and microstructural evolution. Vickers microhardness mapping (Hv) of the

constitutive material model was used, and the parameters were set as indicated in Table 1.where m is the friction factor (ranging from zero for frictionless to 1 for fully sticking objects), and k is the shear flow of the material. According to Ref. [5] and the preset friction factors recommended by the Simufact Forming software for processes using lubricants, a friction factor of m = 0.12 was applied. Moreover, the contact definition of all surfaces was identified for the model to accurately simulate the actual interaction taking place during deformation. The contact between the disks and the billet was identified as inseparable and touching. Similarly, the contact between the assembly and the plunger was modeled. The boundary conditions were set as in the actual experimental procedure, for which all degrees of freedom are foreclosed expect for the translational and rotational motions (tabular motion) in the z-direction (extrusion direction). The results obtained from the simulation were then validated by the experimental results. The plastic strain values and distributions were correlated with the nano-indentation results along the examined cross-sections. Several simulations were then performed to estimate the equivalent stress-strain behavior that developed as a result of MCSTE processing under the various experimental conditions. This was a vital step for the evaluation of the MCSTE technique and its deformation characteristics.

3. Experimental procedure Experimental work was carried out on strain hardenable AA5083, which was received in the form of rolled plates with a 10-mm thickness and an average grain size of 32 μm after annealing. The chemical composition is presented in Table 2. AA5083 billets were sectioned and machined using wire-cut electrical discharge machining (Wire-cut EDM) and a lathe to form 10*10*40-mm billets. Prior to processing, the billets were annealed at 673 K for 1hr as a homogenization and softening treatment, followed by air cooling. All processing was conducted via route-A (the disk-billet assembly was inserted in the inlet channel without rotation between successive passes) at a constant punch speed of 10 mm/min on a 250-kN electromechanical testing machine; Schenck Trebel. Additionally, two MCSTE dies having twist channels with β (30°) and (40) were used with the objective of analyzing and comparing the plastic strain induced during the processing of AA5083 as a function of multiple passes at room temperature (RT). In most of the literature on TE, the dies are split designs [11]. However, the MCSTE monolithic dies enable the Table 2 The chemical composition of AA5083. Material

Element

Fe

Mg

Mn

Cu

AA5083

Weight, %

0.4

4.5

0.4

0.1

Fig. 2. Nano-indentation layout - points 1–5 represent the nano-indentation locations selected. 3

Materials Science & Engineering A 764 (2019) 138216

D.M. Fouad, et al.

extrudates was conducted with the aid of a Mitutoyo HM112 hardness testing machine. To evaluate the hardness variation across the deformed cross-sections, indentations were taken on the longitudinal section (CS) over an equidistant grid with a spacing of 4 mm between each measurement taken across the peripheral region and at a distance of 1.25 mm from the peripheral region toward the center. Similarly, an investigation of the transverse cross-section was conducted with indentations taken along two diagonal lines with a spacing of 1 mm between each measurement, starting from the corner regions. Moreover, to examine the tensile properties of the extrudates, a uniaxial tensile test was performed at RT on a universal testing machine with a capacity of 100 kN at a constant strain rate of 10−3 s−1. Tensile test samples were prepared according to E8M/ASTM standards. Subsequently, the tensile test fracture surfaces were investigated by scanning electron microscopy (SEM) using a LEO Field emission operated at 8 kV with a 60-μm aperture using an SE2 lens. Fractographs for the 1-, 2- and 3-pass (β 30°) test specimens were taken and compared with the as-annealed (AA) and 1-pass (β 40°) specimens. Furthermore, microstructural evolution was characterized using a Leica optical microscope. The samples were chemically etched using Flick's etchant. Afterwards, micrographs were obtained along the peripheral and corresponding central regions of the longitudinal sections for each condition. Representative images of each extrudate taken at low and high magnifications are presented in the results section. Additionally, the average grain size was computed using the linear intercept method.

Fig. 4. Distribution of the plastic strain of AA1100 across the diagonal distance of the transverse cross-section for 1, 2, and 3 MCSTE (β 30°) passes.

Fig. 3(a–c), despite the enhancement in the homogeneity of the strain distribution with multiple passes, the upper section of the extrudates suffered from lower strain values compared to the lower ends. The authors of a previous study have explained such occurrences by the possibility that the residual stresses act as backpressure in the lower sections of the billet, thus increasing the stresses at the end of the sample compared to the top section [13]. Furthermore, it is evident from Fig. 4 that the strain was the highest at the corners of the transverse CS, similar to the results for the hardness contours in Fig. 3. The anisotropy gradually decreased as the number of passes increased, and after 2 passes via MSCTE, the plastic strain was 1.3 and 0.3 at the corner and center locations, respectively. Conversely, the third pass resulted in an increase of ε to 1.7 and 1 at the corner and center locations, respectively. Other FEM studies have suggested that increasing the number of passes and the use of rectangular cross-sections could result in a more uniform distribution of strain across the deformed cross-sections [18]. A similar pattern was observed after deformation via MCSTE processing (β 40°); however, an increase was observed in the εeff values. Table 4 lists the MSCTE-induced maximum strain values for die twisting angle β (30°) and (40°).

4. Results and discussion 4.1. Numerical analysis 4.1.1. Plastic strain distribution Color contour maps showing the distribution of the plastic strains throughout the longitudinal section of the simulated AA5083 extrudates after 1, 2, and 3 MCSTE (β 30°) passes are illustrated in Fig. 3. The distribution of the plastic strains was plotted against the diagonal distance of the transverse CS as a function of MCSTE passes, as shown in Fig. 4. As illustrated in Figs. 3(a) and Fig. 4, the plastic strain per MCSTE (β 30°) pass is approximately 0.5. The plastic strain progressively increased as the number of passes increased until a maximum of 1.7 was reached after 3 passes. Moreover, it is observed that the maximum strain occurred along the partition lines between the successively aligned disks as a result of their relative motion, which led to a heterogeneous strain distribution, as depicted in Fig. 3(a and b). However, the strain distribution became more uniform and homogenous as the number of passes increased, as shown in Fig. 3(c). As demonstrated in

4.2. Mechanical properties 4.2.1. Nano-indentation The load-displacement curves and the nanohardness variation as a function of the number of passes are displayed in Figs. 5 and 6, respectively. Increasing the number of passes resulted in increasing the load bearing capacity on the nanoscale (Fig. 5 (a)), which agrees with the simulation results (Fig. 3) for the locally induced strains. This was also reflected in the nanohardness results shown in Fig. 6. It is observed

Fig. 3. Plastic strain of AA5083 after MCSTE (β 30°) processing across the longitudinal cross-section after (a) 1-pass, (b) 2-passes and (c) 3-passes. 4

Materials Science & Engineering A 764 (2019) 138216

D.M. Fouad, et al.

lists the average Vickers microhardness (Hv) values measured at the peripheral and central regions of the AA5083 AA plate before and after MCSTE processing with 1, 2, and 3 passes. The average hardness for the AA plate was 80 Hv. From Table 5, it is clear that the hardness increased consistently as the number of passes increased, and gradual increases of 28% and 35% took place for the first 20 passes compared to the AA condition. A significant increase occurred for the billets processed for 3 passes, and the average measured hardness was 58% higher than that of the AA 5083 plate. Comparatively, a significant increase in the Hv values was attained after only 1 MCSTE pass with β (40°). As recorded in Table 5, the Hv values produced via 1-pass using a die angle β (40°) were almost equivalent to those after 3 MCSTE passes with a die angle β (30°). However, due to the severe strain hardening produced by the die with β (40°), the number of passes was limited to only one due to shear localization of AA5083. As illustrated in Fig. 7 (a), the contour hardness profile across the longitudinal cross-section of the extrudate from the 1-pass MCSTE processing with a β (30°) reveal the least homogeneity in the hardness distribution compared to the extrudates formed by 2 and 3 passes, as shown in Fig. 7 (b) and (c), respectively. As reported in Table 5, processing via 1-pass led to 25% and 31% increases in the Hv values at the central and peripheral regions, respectively, compared to the AA plate, while the heterogeneity decreased significantly after 3 passes (Fig. 7(c)). This was associated with sharp increases of 56% and 58% in Hv values at the central and peripheral regions, respectively, compared to the AA condition. These hardness results are consistent with the findings in Ref. [19], which reported a similar increase in Hv values after 3 passes of TE for AA6061. Moreover, Fig. 7 reveals a similar trend to that depicted in Fig. 3, where the hardness values are the highest at the bottom sections of the billet compared to the top ones. The hardness contour maps for the transverse CS presented in Fig. 8 reveal that the strain hardening was the highest at the corner and peripheral regions and decreased towards the central regions. Similarly, the contours clearly show the formation of a vortex flow associated with the material twisting within the twist extrusion channel, which is consistent with the results discussed in Ref. [20]. Moreover, the enhanced hardness homogeneity corresponded to the distribution of strain depicted in Fig. 3. Conversely, Fig. 9 (b) reveals a higher degree of heterogeneity in the 1-pass MCSTE β (40°) extrudate compared to the 3pass extrudate processed by an MCSTE die with a lower twist angle, as shown in Fig. 8 (c). In the case of higher twist angles, the magnitude of strain was higher and thus resulted in a higher increase in hardness accompanied by a slight enhancement in heterogeneity, and the hardness increased nonuniformly from the center (~112 Hv) to the periphery (~125 Hv) in the case of MCSTE processing with β (40°). In the case of the deformation of AA5083 through MCSTE processing with β (30°), the hardness values at the center (~121 Hv) varied slightly from those attained at the peripheral regions (~124 Hv) after conducting multiple passes. Similar conclusions were made in Refs. [6,20].

Table 4 Plastic strain across the partition lines between the aligned disks along the peripheral regions of the longitudinal cross-section after MCSTE processing. Twist Angle (β)

Number of Passes

Maximum Plastic Strain

30°

1-pass 2-pass 4-pass 1-pass

0.5 1 2 0.9

40°

Fig. 5. Load-displacement curves as a function of the number of MSCTE passes with β (30°).

that increasing the number of passes led to extrudates with higher strengths, and thus, less penetration under a constant load was recorded. Accordingly, as shown in Fig. 5, the AA condition displayed the highest depth of penetration (1800 nm), while 3 passes of MSCTE processing produced the lowest depth of penetration (1500 nm). Furthermore, Fig. 6 depicts a uniform distribution of the nanohardness values across the investigated section of the AA plate; an average value of 1.12 GPa ( ± 0.03) was recorded. Generally, there was a consistent increase in the hardness as the number of passes increased. This agrees with the findings of Mesbah et al. [9], who showed that the average nanohardness values for 1 pass and 3 passes were 1.36 and 1.53, respectively. Moreover, a general trend similar was depicted by the strain distributions displayed in Figs. 2 and 3; higher hardness values were recorded at points coinciding with the billet-disk interface, and lower values were recorded at the center of the disk. This validated the model output. 4.2.2. Microhardness Table 5 refers to the mechanical properties of the AA5083 processed via MCSTE processing (β 30° and 40°) compared to the AA plates and

Fig. 6. AA5083 nanohardness variation across the peripheral region of a disk-billet interface as a function of the number of MSCTE passes with β o(30°). 5

Materials Science & Engineering A 764 (2019) 138216

D.M. Fouad, et al.

Table 5 Mechanical properties of AA5083 processed via MCSTE processing (β 30° and 40°) compared to the AA plates. Twist Angle (β)

30°

40°

Processing Condition

As-annealed (AA) 1-Pass 2-Pass 3-Pass 1-pass

Hv Values

YS (MPa)

Center

Periphery

80 ± 2 99.6 ± 2 110.6 ± 5 125 ± 3 120 ± 3.5

80 ± 2 104.7 ± 4 106 ± 4.5 127 ± 3 126 ± 2

90 ± 2 190 ± 1 200 ± 1 215 ± 2 230 ± 2

UTS (MPa)

230 250 266 300 315

± ± ± ± ±

1 1 3 2 2

Elongation (%)

25 ± 0.5 19.8 ± 0.25 20 ± 2 19 ± 1.5 16 ± 1

Fig. 7. Color-coded contour for the microhardness values recorded on the longitudinal planes of the AA5083 billets processed via MCSTE processing (β (30°)) route A after (a) 1-pass, (b) 2-passes, and (c) 3-passes. (For interpretation of the references to colour in this figure legend, the reader is referred to the Web version of this article.)

Fig. 8. Color-coded contour for the microhardness values recorded on the transverse planes of the AA5083 billets processed via MCSTE processing (β 30°) route A after (a) 1-pass, (b) 2-passes, and (c) 3-passes. (For interpretation of the references to colour in this figure legend, the reader is referred to the Web version of this article.) 6

Materials Science & Engineering A 764 (2019) 138216

D.M. Fouad, et al.

Fig. 9. Color-coded contour for the microhardness values recorded after MCSTE processing for 1-pass. (a) Longitudinal plane, and (b) transverse plane for AA 5083 billets processed via MCSTE with β (40°). (For interpretation of the references to colour in this figure legend, the reader is referred to the Web version of this article.)

Fig. 10. Eng. σ-ε diagram for AA5083 billets processed by 1, 2, and 3-passes of MCSTE processing with β (30°) and 1-pass of MCSTE processing with β (40°)compared to the AA plates.

MCSTE processing resulted in strengthening of the AA plates, as reflected in the obvious increase in the yield strength after 1-pass MSCTE processing. From Table 5, it is clear that a two-fold increase in the Yeild Strength (YS) was achieved after the first pass, compared to the AA condition. This result is in good agreement with the findings of references [23,24]. Moreover, increasing the number of passes from 1 to 3 ensured a further 38% increase in the YS. The enhancement in the YS of AA5083 after MCSTE processing was comparable to the gains attained for similar high strength aluminum alloys processed by TE. Iqbal et al. reported a significant increase in the YS of AA6061-T6 from 60 to 150 MPa after 3 TE passes [25]. This led to a 150% increase in the alloy YS compared to a 140% increase for the AA5083 alloy processed via MCSTE. It is important to note that the incremental increase in the ultimate tensile strength (UTS) was slightly lower compared to the YS increase. The UTS increased by 9% and 30% after 1 and 3-passes, respectively. A similar trend was reported by Kumar and Iqbal, who showed 3% and 15% increases in the UTS of AA6082 after 1 and 3-passes of TE, respectively [19]. This phenomenon could be related to the strain hardening effect, which led to the onset of necking soon after yielding and

4.2.3. Tensile properties The engineering stress-strain (σ-ε) curves for the AA plates compared to the processed billets after 1, 2 and 3-passes via MCSTE processing with β (30°) and after 1-pass via MCSTE processing with β (40°) are shown in Fig. 10. Moreover, the variation in yield, ultimate tensile strength and elongation percentage as a function of the number of passes compared to the AA condition are listed in Table 5. As shown in Fig. 10, the σ-ε curves reveal an irregular flow pattern that was manifested in a continuous serrated pattern, which is commonly associated with the portevin- Le Chatelier (PLC) effect [21]. The PLC phenomenon is a typical behavior of Al-Mg alloys that occurs as a result of the dynamic interaction between the diffusing solutes/precipitates and the pinning of mobile dislocations [22]. Under specific strain rate and temperature conditions, the solute atoms agglomerate by diffusion to form solute clouds that effectively impede dislocation motion. Conversely, the unpinning of dislocations under applied stress may cause a drop in the stress values, and thus, the serrated pattern represents the continuous pinning and unpinning processes between the mobile dislocations and the solute atoms [23]. Moreover, the σ-ε curves in Fig. 9 reveal that deforming AA5083 via 7

Materials Science & Engineering A 764 (2019) 138216

D.M. Fouad, et al.

due to its brittle nature and the stress concentration created by its needle-like structure [31]. In addition, cleaved planes with minimum deformation were observed at facets with aligned needle-like particles, as depicted by the red circle in the high magnification image of Fig. 11(b). Multiple pass processing via MCSTE (β 30° and 40°) resulted in the formation of a combination of equiaxed dimples and couplets slightly oriented in the direction of shear induced by the MCSTE processing, as shown in Fig. 11 (b) and (d). These results are consistent with the findings discussed in Ref. [31]. However, increasing the number of passes up to 3 (Fig. (d)) resulted in shallower and finer dimples, indicative of a relatively lower ductility prior to the separation of the two surfaces. Moreover, couplet shearing in the direction of shear increased as the number of passes increased and their depth decreased, as depicted in Fig. 11 (c) and (d), respectively. The formation of a uniform dimple size (diameter and depth) across the fractured surfaces as the number of passes increases is indicative of enhancing the deformation uniformity across the cross-sections of the 2 and 3-pass billets, respectively, compared to the 1-pass billets, as shown in Fig. 11 (b). Furthermore, Fig. 11 (d) shows evidence for crack opening after void coalescence in the direction of shear (indicated by a black circle). Sedighi et al. showed that as a result of the higher strain hardening rate exhibited by AA5083 compared to 1xxx Al Alloys, the ductility of the material is reduced, which limits the degree of microvoid edge stretching prior to fracture within the necked region and thus results in shallower couplets associated with shear stresses [32]. Additionally, evidence of second-phase particle cracks is indicated by the blue circle in Fig. 11 (d), which depicts the typical coalescence of microvoids at the particle-matrix decohesion sites [33]. The presence of coarse secondphase particles in AA5083 was also reported in Ref. [34], which explained the shearing and consequent fracture of these particles into finer particles as a result of severe imposed deformations. The size of these coarse intermetallic phases -Al6 (Fe, Mg) - is in the range of 10 μm, as reported by Ref. [35]. In contrast, for AA5083, after one MCSTE pass (β 40°), the fracture surface of the billet at low magnification reveals a ductile failure mode manifested by the formation of coarse nonuniformly distributed dimples and ultrafine dimples that dominated the fractured surfaces, compared to the high fraction of deep dimples dominating the fractured surfaces of AA5083 after 3 MCSTE passes with β (30°).

thus decreased the nonuniform region of the tensile curve [25]. Moreover, the relatively small increase in the UTS associated with a relatively low increase in strain hardening after yielding was positively reflected in the minimum reduction in ductility of the billets processed by 1, 2, and 3-passes, as displayed in Fig. 10. Thus, increasing the number of passes via route A using MCSTE (β 30°) significantly increased the YS of the AA5083 alloy and relatively retained the ductility, which resulted in a stronger and tougher material. The reduction in ductility after MCSTE processing relative to the increase in the YS is considered insignificant, and Zendehdel and Hassani reported a 129% increase in the YS of AA6083 processed with 2 passes of TE at RT, with an accompanying 10% reduction in ductility [26]. Conversely, processing with 3 passes via MCSTE (β 30°) increased the YS by 140% and decreased the ductility only by 6%. Moreover, Fig. 10 reveals a notable strengthening effect. A significant increase in the YS of the processed billets was achieved after only one pass of MCSTE processing with β (40°) and was higher than the values obtained after three MSCTE passes with β (30°). From Table 5, it is clear that the YS increased by more than two and half times, compared to the AA condition, from 90 MPa to 230 MPa after 1pass. The increase in YS after 1 pass of MCSTE processing with β (40°) was analogous to that achieved after 3-passes with a lower twist angle. Similarly, the UTS increased by 40% after 1-pass compared to the AA condition. No studies have reported on the fracture of extrudates after 1-pass of TE. Kumar and Iqbal emphasized the necessity of conducting TE at elevated temperatures for high-strength aluminum alloys such as AA5083 [19]. The researchers were able to conduct 3 TE passes at 350 °C on AA6082, which yielded only a 17% increase in the UTS. Another complementary study was conducted on AA5083 by deforming it via 4-passes of a variant of ECAP at 240 °C. The authors reported a 15% increase in the UTS after 4passes, coupled with a 7% loss in ductility [27]. Furthermore, processing AA5083 by MCSTE with a higher twist angle led to a 9% reduction in ductility compared to the 6% decrease in AA5083 extrudates deformed via MCSTE processing (β 30°). The results suggest that strain saturation occurred in AA5083 after only one pass of MCSTE processing, which led to substantial grain refinement and strengthening in a single process [28]. These findings indicate that MCSTE processing (β 40°) was able to considerably enhance the mechanical properties of AA5083 and similar Al-alloys with less deformation loads and energy compared to similar SPD processes at elevated temperatures. Additionally, the fact that only one pass was needed to achieve this considerable increase in strength is considered highly favorable for the uptake of MCSTE processing in industry. Accordingly, the displayed results are indicative of the ability of the MCSTE process as an SPD technique to effectively deform high-strength materials to increase the YS and produce tougher materials.

4.2.5. Microstructural evolution The AA, 1-pass, 2-pass and 3-pass billet microstructures were assessed using OM after MCSTE processing (β 30°). The 1-pass billet microstructures were also assessed after MSCTE processing with β (40°). Fig. 12 demonstrates the microstructure at the peripheral and central regions of the AA billets before and after processing by MCSTE (β 30°) for 1, 2, and 3-passes. The microstructure of the AA plates reveals relatively uniform grains with an average size of 32 μm both at the peripheries (Fig. 12 (a)) and the center (Fig. 12(b)). Table 6 lists the grain size variation for the samples after 1, 2, and 3-passes, measured both at the center and the peripheries of the AA5083 MCSTE (β 30° and 40°) billets. Increasing the number of MCSTE passes resulted in significant grain refinement at the periphery (Fig. 12 (c)) and at the center (Fig. 12 (d)) for the billets processed via a single MCSTE pass. From the displayed results, the grain size of the AA5083 AA plates at the central regions were refined by 43%, 68% and 90% after 1, 2, and 3passes, respectively. Moreover, the grain size at the peripheries of the billets was refined by 50%, 85% and 95%, for the billets processed with 1, 2 and 3passes, respectively. As demonstrated, the average grain size at the center of the billet tends to be coarser compared to that at the peripheries, which agrees with the results obtained for AA1100 [36] and the distribution of hardness values across the cross-section for AA5083 (Figs. 7–9). Augmenting the intensity of plastic straining through MCSTE processing with up to 3-passes led to a significant grain

4.2.4. Fracture behavior Fig. 11 illustrates the SEM fractographs of tensile tested specimens after successive MCSTE passes. The fracture surfaces of AA5083 billets processed via MCSTE (β 30°) with and without processing and 1-pass MCSTE processing (β 40°) extrudates were examined at low magnification; Fig. 11 (a) reveals a ductile failure mode manifested by the formation of relatively shallow and fine dimples. Similar findings were reported in a study investigating the impact of accumulative rolling bonding (ARB) conducted on both commercially pure aluminum AA1050 and strain hardenable AA5083 [29]. For the AA5083 alloy fracture surfaces, Fig. 11 depicts second-phase particles within the deep dimples, suggesting that failure was initiated at those sites. It is suggested that those particles are mainly Fe-rich intermetallic compounds that react with Mg in the AA5083 to form intermetallic Al6 (Fe, Mg) [30]. This agrees with the chemical composition (Table 2). Additionally, Fe is the most common impurity present in aluminum alloys due to its ability to add strength [30]. However, the Fe-intermetallic compounds display a platelet-like morphology, which is known for having the most detrimental effect on the tensile properties of AA5083 8

Materials Science & Engineering A 764 (2019) 138216

D.M. Fouad, et al.

Fig. 11. SEM images of the AA5083 fractured tensile specimens illustrating the fracture topography of the (a) AA, the MCSTE extrudates for β (30°) after (b) 1-pass, (c) 2-passes, and (d) 3-passes, and the MCSTE extrudates for β (40°) after (e) 1-pass. *The blue arrow represents the shear direction, and the white arrow indicates at a magnified image. The red circle indicates a cleaved plane, while the black circle represents a crack opening. The blue circle shows fracture of second-phase particles. (For interpretation of the references to colour in this figure legend, the reader is referred to the Web version of this article.) Table 6 Average grain size measured at the central and peripheral regions of AA5083 MCSTE extrudates (β 30° and 40°) compared to the AA plate. Twist angle (β)

30° 40°

Location

Center Periphery Center Periphery

Mean Grain Size (μm) AA

1-Pass

2-Pass

3-Pass

20 20 – –

14 ± 2 10 ± 3.5 2.8 ± 2.25 3 ± 2.5

6.5 ± 2.5 3 ± 1.5 – –

2±1 1.75 + 0.5 – –

refinement, and the average grain size measured at the peripheral and central regions was 2 μm and 1 μm, respectively. Such grain refinement occurred simultaneously with the increase in the induced strain for the 3-pass processed billets, as reported in Table 4. Consequently, increasing the number of passes resulted in enhancing the structural uniformity across the billet cross-section, as shown in Fig. 12(g and h), which also supports the observations made for the fractured surfaces in Fig. 11. Zendehdel and Hassani reported similar findings after performing 4 TE (β 30°) passes on AA6063 [25]. The average grain size increased by 92% and 94% at the central and peripheral regions, respectively. Comparing the average grain size measured for AA5083 to that of AA1100 processed under the same conditions (number of passes and twist angle), presented in Ref. [37], it was evident that the grain refinement was higher for AA5083 compared to AA1100. Furthermore, it has been shown that materials exhibiting a lower stacking fault energy (SFE), in comparison to materials exhibiting higher SFE values, display lower rates of dynamic recovery, and hence,

Fig. 12. OM micrographs of A5083 at peripheral and central regions along the longitudinal cross-section for AA5083 (a–b) AA and MCSTE extrudates with β (30°) after (c–d) 1-pass, (e–f) 2-passes and (g–h) 3-passes, at high magnification X.

9

Materials Science & Engineering A 764 (2019) 138216

D.M. Fouad, et al.

Fig. 13. OM micrographs of AA5083 at peripheral regions across the longitudinal cross-section of (a) 1pass extrudate processed via MCSTE (β 30°) compared to (b) 1-pass extrudate processed via MCSTE (β 40°). * The white arrows depict the structural orientation.

and shearing of the dimples into couplets that dominated the fractured surfaces. Finally, MCSTE processing via a twisting angle β (40°) resulted in a higher shear deformation intensity, which was manifested by the formation of couplets after 1-pass processing, while an angle β (30°) showed mixed equiaxed dimples and couplets after 2-pass processing. From the displayed results, the MCSTE novel die design for SPD offers several advantages over its counterparts, which makes it useful for industrial applications.

the continuous dynamic recrystallization (cDRX) process is more pronounced, especially if severe deformation is imposed [37]. Accordingly, the microstructure of the SPD-deformed billets manifest finer grain sizes coupled with an isotropic distribution. It has been suggested that because the 5083 Al alloy is a solid-solution-strengthened material, the alloying elements reduce the SFE of the Al alloy, reduce its tendency towards dynamic recovery and aid in achieving finer grains with the application of SPD [26]. Similar to the previous investigations, the microstructure of 1-pass AA5083 deformed by MCSTE processing (β 40°) has been assessed through OM. Fig. 13 demonstrates the microstructure at the peripheral regions for the AA5083 billets processed via MCSTE with β (40°) after 1pass compared to MCSTE processing with β (30°). The grain size of the AA plate at the central region was refined by 87.5%, while the peripheries were refined by 97.5%. The grain size analysis (Table 6) and the microstructure presented in Fig. 13 (b) reveal substantial grain refinement after processing via MCSTE (β 40°) for only 1-pass. These findings agree with the enhancement in mechanical properties reported in this study. Comparable reductions in the grain size were reported in a previous study on the impact of the TE deformation of AA6083 on the grain size, and a 98% reduction in grain size was achieved after 16 TE passes [26]. Accordingly, the strain distribution in the deformation zones is strongly dependent on the die geometry (path and twist angle); in this case, the MCSTE die (β 40°) was capable of inducing adequate stresses and strain that led to the formation of fine equiaxed grains, possibly with HAGBs [11]. Moreover, the OM micrographs of the AA5083 1-pass MCSTE (β 40°) extrudate demonstrated in Fig. 13(b) reveal intensive shear compared to that of 1-pass MCSTE (β 30°), as depicted in Fig. 13(a) by the white arrow. This agrees with observations made for the fracture surfaces after 1-pass for both angles. Couplets were evident after 1-pass MCSTE processing (β 40°), while the fracture surface of the 1-pass MCSTE extrudate was mainly comprised of equiaxed dimples. This can be attributed to increases in the plastic strain per pass, from 0.5 to 0.9, associated with processing via MCSTE with β (30°) and (40°), respectively.

Data availability The raw data required to reproduce these findings are available to download from [https://doi.org/10.17632/7zp7x32v53.1]. Acknowledgments The authors of this work would like to acknowledge the American University in Cairo for granting the research funds for this project. Gratitude is also extended to the MSC software company for offering a trial version of Simufact Forming. Appendix A. Supplementary data Supplementary data to this article can be found online at https:// doi.org/10.1016/j.msea.2019.138216. References [1] T.G. Langdon, Twenty-five years of ultrafine-grained materials: achieving exceptional properties through grain refinement, Acta Mater. 61 (2013) 7035–7059. [2] A. Azushima, R. Kopp, A. Korhonen, D.Y. Yang, F. Micari, G.D. Lahoti, P. Groche, J. Yanagimoto, N. Tsuji, A. Rosochoski, A. Yanagida, Severe plastic deformation (SPD) processes for metals, CIRP Ann. - Manuf. Technol. 57 (2008) 716–735. [3] J.R. Davis, Alloying: Understanding the Basics, ASM International, Newyork, 2001. [4] Y.T. Zhu, T.G. Langdon, The fundamentals of nanostructured materials processed by severe plastic deformation, JOM 56 (2004) 58–63. [5] J.G. Kim, M. Latypov, N. Pardis, Y. Beygelzimer, H.S. Kim, Finite element analysis of the plastic deformation in tandem process of simple shear extrusion and twist extrusion, Mater. Des. 83 (2015) 858–865. [6] M. Latypov, E.Y. Yoon, D.J. Lee, R. Kulagin, Y. Beygelzimer, M.S. Salehi, H.S. Kim, Microstructure and mechanical properties of copper processed by twist extrusion with a reduced twist-line slope, Metall. Mater. Trans. A 45A (2014) 2232–2241. [7] Y. Beygelzimer, V. Varyukhin, D. Orlov, B. Efros, V. Stolyarov, H. Salimgareyev, Microstructural evolution of titanium under titanium, in: Y.T. Zhu, T.G. Langdon, R.S. Mishra, S.L. Setniatin, M.J. Saran, T.C. Lowe (Eds.), Ultrafine Grained Materials II, Wiley online Library, 2013, pp. 43–46. [8] V.V. Stolyrov, Y. Beygelzimer, D. Orlov, R. Valiev, Refinement of microstructure and mechanical properties of titanium processed by twist extrusion and subsequent rolling, Phys. Met. Metallogr. 99 (2005) 204–211. [9] M. Mesbah, F. Fadaifard, A. Karmzadeh, B. Tabrizi, A. Rafieerad, G. Faraji, A. Bushoa, Nano-mechanical properties and microstructure of UFG brass tubes processed by parallel tubular channel angular pressing, Met. Mater. Int. 22 (2016) 1098–1107. [10] H. Bakhtiari, M. Karimi, S. Razazadeh, Modeling, analysis and multi-objective optimization of twist extrusion process using predictive models and meta-heuristic approaches, based on finite element results, J. Intell. Manuf. 27 (2016) 463–473. [11] Y. Beygelizmer, R. Kulagin, Y. Estrin, L. Toth, H. Kim, M. Latypov, Twist extrusion as a potent tool for obtaining advanced engineering materials: a review, Adv. Eng. Mater. 19 (2017) 1–24. [12] S.A. Mousavi, A.R. Shahab, Influence of strain accumulation on microstructure of aluminum 1100 in the twist extrusion, Int. J. Mod. Phys. B 22 (2008) 2858–2865. [13] S.A.A.A. Mousavi, ShR. Bahadori, Examination of an aluminum alloy behavior under different routes of twist extrusion processing, Mater. Sci. Eng. A 528 (2011) 6527–6534. [14] U.M. Iqbal, V.S. Kumar, An analysis on effect of multipass twist extrusion process of AA6061 alloy, Mater. Des. 50 (2013) 946–953.

5. Conclusion MCSTE processing is a rather simple and effective solution that makes TE a viable method for commercialization purposes. With MCSTE processing, customized stacked disks are used to host a noncircular cross-sectional billet that is processed through a TE die with a reduced twist angle (β) 30° and 40°, compared to the typical 60° angle used in conventional TE dies. The introduced modifications to the conventional TE die provide a rather simple and effective solution that might translate into processing cost savings and higher die durability. The FE analysis revealed that MCSTE processing is associated with a plastic strain that varies with the twist channel angle and is independent of the material type. TE with a twisting angle of β = 30° produces an average plastic strain of ~0.5 per pass compared to ~0.9 for MCSTE processing with β = 40°, for which the strain approached that induced by conventional TE. Moreover, MCSTE processing showed the ability for grain refinement of AA5083 alloys, which can be attributed to increased hardness and tensile properties as the number of passes increases. Failure of the processed AA5083 followed the typical ductile failure mode of AA alloys, which is augmented with refinement 10

Materials Science & Engineering A 764 (2019) 138216

D.M. Fouad, et al.

and mechanical properties of 6063 aluminum alloy, Mater. Des. 37 (2012) 13–18. [27] N. Fakhar, F. Fereshteh-Saniee, R. Mahmudi, Significant improvements in mechanical properties of AA5083 aluminum alloy using dual equal channel lateral extrusion, Trans. Nonferrous Metals Soc. China 26 (2016) 3081–3090. [28] S.W. Cheng, C.F. Wu, Factor screening and response surface exploration, Stat. Sin. 11 (2001) 553–604. [29] M. Sedighi, P. Fahadipour, M.H. Vini, Mechanical properties and microstructural evolution of bimetal 1050/Al203/5083 composites fabricated by warm accumulative roll bonding, JOM 68 (2016) 3193–3200. [30] S.S. Kumari, R.M. Pillai, B.C. Pai, Role of calcium in aluminum based alloys and composites, Int. Mater. Rev. 50 (2005) 216–238. [31] Y. Liu, Y. Sun, L. Zhang, Y. Zhao, J. Wang, C. Liu, Microstructure and mechanical properties of Al-5Mg-0.8Mn alloys with various contents of Fe and Si cast under near- rapid cooling, Metals 7 (2017) 1–12. [32] D. Singh, P.N. Rao, R. Jayaganhan, Microstructures and impact toughness behavior of Al 5083 alloy processed by cryorolling and afterwards annealing, Int. J. Miner Metall.Mat 20 (2013) 759–769. [33] N. Fakhar, F. Saniee, R. Mahmudi, High strain-rate superplasticity of fine- and ultrafine-grained AA5083aluminum alloy at intermediate temperatures, Mater. Des. 85 (2015) 342–348. [34] J.S. Washfold, I.R. Dover, I.J. Polmear, The thermomechanical processing of an AlMg alloy, Met. For. 8 (1985) 56–59. [35] D.M. Fouad, W.H. ElGaraihy, M.M.Z. ahmed, M.M. EL Sayed Selman, H.G. Salem, Influence of Multi-Channel spiral twist extrusion (MCSTE) processing on structural evolution crystolographic texture and mechanical properties of AA1100, Mater. Sci. Eng. A 737 (2018) 166–175. [36] E.V. Kozlov, A.N. Zhdanov, N.A. Koneva, Deformation mechanisms and mechanical properties of NanoCrystalline materials, Phys. Mesomech. 11 (2008) 42–50. [37] M. Bacca, D. Hayurst, R. Mceeking, Continuous dynamic recrystallization during severe plastic deformation, Mech. Mater. 90 (2015) 148–156.

[15] M.I. Latypov, I.V. Alexdrov, Y.E. Beygelzimer, S. Lee, H.S. Kim, Finite Element analysis of plastic deformation in twist extrusion, Comp.Mat.Sci. 60 (2012) 194–200. [16] W. ElGaraihy, D.M. Fouad, H.G. Salem, Multi-Channel spiral twist extrusion (MCSTE): a novel severe plastic deformation technique for grain refinement, Metall. Mater. Trans. A 49 (2018) 2854–2864. [17] H. Yalavarthy, Friction Stir Welding Process and Material Microstructure Evolution in 2000 and 5000 Series of Aluminum Alloys, All theses, 2009, p. 745. [18] V.S. Kumar, U.M. Iqbal, Effect of process parameters on the microstructure homogeneity of AA6082 aluminum alloy deformed by twist extrusion, ASME 2013: Adv. Manuf. 2A (2013) 1–6. [19] Yu Beygelzimer, R. Kulagin, O. Davyenko, Vortex formation under twist extrusion, NanoMat 14 (2016) 553–560. [20] S.S. Faregh, A. Hassani, Stress and strain distribution in twist extrusion of AA6063 aluminum alloy, Int. J. Material Form. 11 (2018) 175–184. [21] Q. Hu, Q. Zhang, S. Fu, P. Cao, M. Cong, Influence of precipitation on the portevinLeChatelier effect in Al-Mg alloys, Theor.App.Mech.Lett. 1 (2011) 1–4. [22] H. Dierke, F. Krawehl, S. Graff, S. Forest, J. Sachl, H. Neuhauser, PortevinLeChatelier effect in Al-Mg alloys: influence of Obstacles-experiments and modelling, Comput. Mater. Sci. 39 (2007) 106–112. [23] D. Orlov, Y. Beygelzimer, S. synkov, V. Varyukhin, N. Tsuji, Z. Horita, Microstructure evolution in pure Al processed with twist extrusion, Mater. Trans. 50 (2009) 96–100. [24] S.A. Mousavi, ShR. Bahadori, A.R. Shahab, Numerical and experimental studies of the plastic strains distribution using subsequent direct extrusion after three twist extrusion passes, Mater. Sci. Eng. A 527 (2010) 3967–3974. [25] U.M. Iqbal, V.S. Kumar, S. Gopalakannan, Application of response surface methodology in optimizing the process parameters of twist extrusion process for AA6061T6 aluminum alloy, Measurement 94 (2016) 126–138. [26] H. Zendehdel, A. Hassani, Influence of twist extrusion process on microstructure

11