Effect of aluminum pre-straining on strength of clinched galvanized SAE1004 steel-to-AA6111-T4 aluminum

Journal of Materials Processing Technology 215 (2015) 193–204

Contents lists available at ScienceDirect

Journal of Materials Processing Technology journal homepage: www.elsevier.com/locate/jmatprotec

Effect of aluminum pre-straining on strength of clinched galvanized SAE1004 steel-to-AA6111-T4 aluminum Teng Jiang a , Zhong-Xia Liu a,∗ , Pei-Chung Wang b a b

School of Physics and Engineering, Key Lab of Materials Physics, Zhengzhou University, Zhengzhou 450052, PR China Manufacturing Systems Research Lab, General Motors Global Research and Development Center, 30500 Mound Road, Warren, MI 48090, USA

a r t i c l e

i n f o

Article history: Received 13 December 2013 Received in revised form 9 August 2014 Accepted 16 August 2014 Available online 29 August 2014 Keywords: Aluminum Pre-strain Clinching joining Section parameters Joint strength

a b s t r a c t Normally, the sheet metal may be pressed or stamped prior to mechanical clinching. The pre-strain from the pressing or stamping may have an influence on the quality of the clinched joints. To understand how the pre-strained sheets materials affect the joint quality, we conducted the study to understand the effect of pre-straining aluminum on the static strength of the clinched aluminum-to-steel joint. The section parameters (i.e. undercut, neck thickness, and bottom thickness) and joint strength were measured. It was found that the work hardening resulting from pre-straining decreased the ductility of aluminum AA6111-T4 and induced some ductile damage on the clinched aluminum workpieces. A 5% pre-strain on AA6111-T4 caused a significant decrease (about 20%) in joint strength. Though the bottom thickness is a good indicator for detecting the strength variation for the clinched as-received and pre-strained aluminum–steel joints, it barely detected the effect of ductile damage in the pre-strained aluminum on the joint strength. The application of the similar “X” parameter (i.e., bottom thickness of the clinched joint) from the as-received aluminum to monitor the quality of the clinched steel-pre-strained aluminum joint resulted in overestimates of the joint strength. Therefore, the optimum clinching tools and process variables for clinching of the pre-strained AA6111-T4-to-steel cannot be derived from the experiments using the as-received aluminum workpieces. To obtain the desired strength of the joints made with dissimilar materials, it is important to achieve a good balance between the undercut and neck thickness by optimizing the clinching tools and process parameters. The variation of electrical resistance of the joints during the clinching process can differentiate the difference of the damage between the as-received and pre-strained clinched joints. Therefore, a method of monitoring the variation of electrical resistance is proposed to inspect the quality of the clinched steel-to-pre-strained aluminum joints. © 2014 Elsevier B.V. All rights reserved.

1. Introduction In current lightweight design, mass reduction of body-in-white assemblies is realized by the systematic application of lightweight materials, such as aluminum or magnesium alloys, or a multimaterial mix, such as the combination of steel and aluminum alloys, steel and magnesium alloys and steel and polymer material. In particular, the combination of steel and aluminum alloys is currently applied most extensively in the automotive industry which describes the state of the technology in automotive engineering (Miller et al., 2000). Because of the different melting points and thermal conductivity and intermetallic compounds formed at the welding interface of steel/aluminum alloys, it is difficult to join

∗ Corresponding author. Tel.: +86 37167767776; fax: +86 37167767776. E-mail addresses: [email protected], [email protected] (Z.-X. Liu). http://dx.doi.org/10.1016/j.jmatprotec.2014.08.016 0924-0136/© 2014 Elsevier B.V. All rights reserved.

them by conventional fusion welding techniques such as resistance welding (Qiu et al., 2010). An alternative method for joining the sheet metal is the mechanical clinching which geometrically constrains two workpieces by local deformation of the workpiece metals using a punch and die. It has been shown that clinching is environmentally friendly due to low energy requirements, low-noise output, no fume emissions and typically involves no heating compared with resistance welding (Michalos et al., 2010), more efficient in mass production compared to friction stir welding (Briskham et al., 2006), and needs lower production costs if tool service life is guaranteed compared to self-pierce riveting (Varis, 2006a). Clinching consists of producing an appropriate material flow of two sheets by means of a punch and a die designed to produce a mechanical interlock. It is a mechanical method involving severe local plastic deformation of two or more workpieces. The mechanical strengths of the clinched joints depend primarily on the joint

194

T. Jiang et al. / Journal of Materials Processing Technology 215 (2015) 193–204

Fig. 1. Schematic model of TOX Clinch Setup (TOX@ , 2014a).

profile which mainly is determined by the clinch die, mechanical behavior and process-induced strain hardening of materials. To obtain suitable joint profile, two different schemes are (i.e., fixed die and extensible die) available. Fig. 1 schematically shows the extensible SKB die and TOX clinch model (TOX@ , 2014a). As shown, the punch moves the die and exerts a press force on the workpieces to be joined. The material of the workpieces is penetrated down to the die bottom. Then, the material is upset between the punch and solid segment of the die such that a radial material flow occurs toward the outside. The flexible segment of SKB die opening during the upsetting phase controls the radial material flow such that an undercut is created between the materials to be joined on the punch and die side (Lambiase, 2013). The radial motion of the flexible segments increases the die cavity volume, and thus decreases the forming force during the clinching process (Lambiase, 2013). As a result, greater undercut may be obtained with a smaller forming force than that of the fixed die. Therefore, the extensible die can guarantee the joint symmetry by controlling the material flow in the die to ensure the joint having sufficient undercut and desired joint strength (Lambiase and Di Ilio, 2013). Furthermore, a series of sheets having a wide range of thicknesses can be clinched with a single set of clinching tools, which is especially beneficial for vehicle body assembly. The main section parameters describing a clinched joint are the undercut between the upper and lower workpieces, neck thickness of the upper workpieces, and bottom thickness (Varis and Lepisto, 2003) as shown in Fig. 1. The undercut and neck thickness can determine the strength of the clinched joint while the bottom thickness is often used to monitor the joint quality (Varis and Lepisto, 2003). Therefore, the mechanical strengths of the joint highly depend on the final joint geometry after clinching. Various factors influence the joint geometry, and consequently the strength of the clinched joints. Sufficient rigidity of the support C-frame and suitable tool parameters are important to obtain a quality joint. Markowski et al. (2013) discovered that C-frame with low rigidity resulted in an increase of tool axis deviation and this in turn resulted in a distorted, non-circular joint and the decrease of the bottom thickness of the clinched joints. Therefore, the frame used for forming the joint should be sufficiently rigid. In addition,

Lambiase (2013) investigated the influence of process parameters in mechanical clinching with extensible dies and obtained the relation of undercut and neck thickness to the tool parameters (e.g., tools corner radii, die diameter, die depth and punch diameter). For the given clinching workpieces and clinching facilities, the initial state of the clinched workpieces may affect the quality of the clinched joints. Based on a wide range of published results, Varis (1998) established the limiting values of the initial mechanical properties of the materials prior to clinching to assess the suitability of clinching (i.e., elongation at failure >10% and 0.2% proof stress <550 MPa). Mucha et al. (2011) found that when joining the similar materials, the joints made with sheets with high initial strength had higher lap-shear strength. Varis (2002) discovered that the joints with round shape had the maximum lap-shear strength for high strength steel. For low ductility material (e.g., magnesium alloys), pre-heating may be necessary during the clinching process to avoid the cracking (Neugebauer et al., 2008). For joining dissimilar workpieces having significantly different strengths and ductility, the material having higher formability should be located on the punch side while the higher strength material should be positioned on the die side (Kato et al., 2008). Saberi et al. (2008) discovered that the anisotropic mechanical properties of the material may affect the joint performance, and resulted in higher lap-shear strength than that of the isotropic materials. However, Coppieters et al. (2011a) realized that the anisotropic material behavior hardly induced large geometric differences in the interlock regime. By nature, clinching is a plastic connection technology and involves severe local plastic deformation of two or more workpieces created by the clinching tools. The strain hardening during the clinch forming should be accounted for accurately simulating the clinch forming process and predicting the strengths of the clinched joint. As the local strain hardening behavior of the sheet metal during the clinching process beyond the point of maximum uniform strain, which is a post-necking hardening behavior, it is difficult to evaluate the stress-strain relation of sheet metal using standard tensile tests. Coppieters et al. (2011b) demonstrated that the post-necking hardening behavior of sheet metal can be identified by minimization of the discrepancy between the internal and external work in the necking zone during a tensile test. When standard tensile specimens cannot be properly produced, multi-layered upsetting test (MLUT) was an alternative method for the identification of the local hardening behavior of sheet metal (Coppieters et al., 2010). Furthermore, larger local plastic deformation may cause ductile damage in the clinched regime and thus has an influence on the joint strength. Under the uniaxial (Coppieters et al., 2011a) and multi-axial quasi-static (Coppieters et al., 2013) loadings, the pull-out strength was less relevant while the lap-shear strength of the clinched joints was sensitive to the material state. If the ductile damage caused by clinching was not taken into account, the failure mode of the clinched joint was barely predicted both in the experimental study of uniaxial (Coppieters et al., 2011a) and multi-axial (Coppieters et al., 2013) quasi-static strengths of the clinched joints. Roux and Bouchard (2013) found that the mechanical strength and failure mode of the clinched 5774 aluminum joint can be accurately estimated with taking into account of the hardening behavior and ductile damage in the clinched regime. In actual production, the workpieces may be deformed (e.g., stamping) prior to clinching. The pre-strain of the wokpieces not only increases the deformation resistance, and thereby limiting further deformation but increases the ductile damage propensity in the clinched regime. Furthermore, a change in dimensions of the workpieces may be developed. Hahn et al. (2001) observed that the joint locations often were placed in areas (e.g., sheet flange) that had been deformed in previous operation (e.g., stamping), which would result in the change of characteristic features. Because the specimens from the pre-strained workpieces were often

T. Jiang et al. / Journal of Materials Processing Technology 215 (2015) 193–204

unavailable in the product development process, as-received workpieces are often used for the process and tool optimization development. The clinching conditions established using the as-received materials may be inappropriate for clinching the prestrained materials. A new parameters optimization strategy of the clinching tools and clinching process may be required by considering the new state of the pre-strained materials. As the pre-strain prior to clinching varies with the deformation process, it may require a new set of the clinching tools and process parameters for various deformation processes. If so, a frequent replacement and adjustment of the tools and process parameters may cause a decrease in production efficiency and an increase in the production cost, which is undesirable for automotive production. It is desirable that few clinching tools and process parameters can tolerate various pre-strains of the clinched materials to some extent. How the pre-strained materials influence the quality of the clinched workpieces is the main concern in this study. Though various studies studied the influence of the initial mechanical properties of asreceived materials on the clinching performance, there were few studies reported on the influence of pre-strained materials. As the bottom thickness provides a good indication of the joint quality, it has been used to monitor the joint quality in actual production, which named as “X” parameters (Varis, 2006b). Realtime monitoring method based “X” parameter had been studied and applied in clinching process (Tan et al., 2005). Force sensors measure the punch force while a position sensor monitors the cylinder travel to achieve the quality control dimension “X”. The force–displacement reference curve can be obtained and a tolerance range can be set for one clinching process. Then, the generated force–displacement curve during the clinching process is compared with a reference curve. Once the actual process curve is within the pre-specified tolerance range, a good clinched joint is obtained. To detect and show the fault of the process variables such as workpiece thickness and properties, damage to punch and die, various new real-time monitor methods based windows technology have been developed (HBM.com, 2014; TOX@ , 2014b). If the process variables cause a significant deviation of the force–displacement curve from the reference curve, the clinching process would halt. As discussed in previous section, the plastic deformation prior to the clinching causes a material strain hardening and ductile damage in the clinched region, which may induce a deviation of process curve from the reference curve to some extent. The key issue is whether the “X” parameters can reflect the influence of pre-strain on the clinched workpieces. The electrical resistance of metallic materials is determined by the scattering effect of lattice vibration and crystal defect on the free electron (Brown, 1977). An increase of impurity atom content, dislocation density (Brown, 1977), grain boundary (Karolik and Luhvich, 1994), precipitation of secondary phase (Gibiansky and Torquato, 1996) and micro-cracking (Sevostianov, 2003) may cause an increase in electrical resistance. Plastic deformation of metal causes the grain distortion and introduces the imperfection in the crystal structure, which would have an influence on the electrical property of the alloy. The electrical resistance has been used to investigate the plastic deformation and work-hardening behavior of steel (Dominguez and Sevostianov, 2011) and aluminum materials (Adeosun et al., 2011) and the cracking in materials (Sevostianov, 2003). Since the plastic deformation of the workpieces prior to clinching induced higher density dislocation and dislocation pile-ups in the pre-strained workpices than that of the as-received aluminum in the clinched region, more ductile damage may be developed in the pre-strained workpieces than that of the as-received workpieces. Therefore, the difference in ductile damage in the clinched region can be reflected by the difference of the electrical resistance between the as-received and pre-strained clinched joint.

195

Table 1 Chemical composition of AA6111 (mass %). Mn

Si

Mg

Cu

Cr

Zn

Ti

Fe

Al

0.15–0.45

0.7–1.1

0.5–1.0

0.5–0.9

Max 0.10

Max 0.15

Max 0.10

Max 0.04

Bal.

Table 2 Chemical composition of SAE1004 steel (mass %). C

Mn

P

S

Cu

Ni

Cr

Mo

Fe

0.02–0.08

Max 0.35

Max 0.03

Max 0.03

Max 0.20

Max 0.20

Max 0.15

Max 0.06

Bal.

Table 3 Mechanical properties of SAE1004 steel and AA6111-T4. Material

Yield strength (MPa)

Tensile strength (MPa)

Elongation (%)

SAE1004 steel AA 6111-T4

151.12 151.27

288.78 290.79

43.84 23.45

To address this concern, the effect of pre-strained AA 6111-T4 aluminum on the quality and strength of clinched 0.7 mm thick galvanized SAE1004 steel and 1.0 mm thick aluminum AA 6111-T4 was investigated. The aluminum AA6111-T4 was selected because it has lower work hardening rate and formability than steel (Vincze et al., 2005) and has been widely applied in automotive industry. There are three main parts in this study; the first presents the experimental procedure including material, pre-strain aluminum workpiece, sample fabrication, process variable development, joint characterization, mechanical testing and measurement of joint electrical resistance online. The next section shows the effects of process parameters and pre-stained aluminum on the quality of the clinched joints. The effect of process parameters including punch force and die depth is also investigated in order to identify the key variables associated with the clinching technologies. The attributes (i.e., undercut, neck thickness and bottom thickness) and the shear strength of the clinched joints are presented where the strengths of the clinched joints are evaluated by the bottom thickness of the joints. The variation of electrical resistance of the clinched joints was used to evaluate the effect of pre-strained aluminum on the strength of the clinched joints. Finally, a discussion is presented concerning the use of the bottom thickness and variation of electrical resistance to detect the quality of clinched pre-strained aluminum-to-steel joints. 2. Experimental procedure 2.1. Material 0.7 mm thick galvanized SAE1004 steel and 1.0 mm thick aluminum AA 6111-T4 were selected in this study. To simulate the aluminum sheets transport from aluminum mill to assembly plant, all AA6111 aluminum workpieces were held 15 days at room temperature after solution treatment (T4 condition) prior to clinching. The chemical compositions and mechanical properties of these materials were measured and the results are listed in Tables 1–3. 2.2. Pre-straining of aluminum AA6111-T4 To study the effect of pre-strained AA6111-T4 on the joint strength, the as-received aluminum AA6111-T4 was subject to various amounts of pre-strain. Though the stress and strain fields in the clinched regime are very complex in actual applications while the standard uniaxial tensile test may not precisely

196

T. Jiang et al. / Journal of Materials Processing Technology 215 (2015) 193–204

Fig. 3. Configuration of a lap-shear specimen (dimension in mm).

2.3. Joint fabrication

Fig. 2. Configuration of (a) specimen for tensile pre-strain, (b) coupon for pre-strain and (c) pre-strained coupon for clinching (dimension in mm).

simulate the real stress and strain state, limited to the existing facilities, uniaxial tensile pre-straining was still applied in this study. Standard tensile specimens as shown in Fig. 2(a) machined from the 300 mm × 300 mm × 1.0 mm aluminum AA6111-T4 workpieces were used to measure the mechanical properties of the pre-strained aluminum. 2.5, 5, 7.5, and 10% plastic strain were obtained with a MTS 810 tensile tester at a stroke rate of 2 mm/min. To precisely control the pre-strain, the extensometer shown in Fig. 2(a) was used. After the pre-strain, the specimens were aged for 24 h in ambient environment prior to tensile testing. The 24 h aging time is used to simulate the time interval between the part stamping and clinching for automotive applications. Three replicates were performed for each pre-strain condition, and the average mechanical properties were reported. To prepare the pre-stained aluminum for joint clinching, coupons with a size of 38.1 mm × 176 mm as shown in Fig. 2(b) were machined from 300 mm × 300 mm × 1.0 mm aluminum AA6111 workpieces. To ensure the pre-strained aluminum coupons have the precise pre-strain within the area of the clinched joint, the extensometer was mounted at the clinching site of the aluminum workpieces as shown Fig. 2(b). The coupons were first pre-strained to various amounts of plastic strain along axial direction with MTS 810 tester at a stroke rate of 2 mm/min. The grip areas of the coupons were then removed and the remaining pre-strained coupon shown in Fig. 2(c) with a gauge area of 38.1 mm × 126 mm was used for clinching.

The effect of aluminum pre-straining on the joint strength is determined by constructing and testing a set of three samples. The samples were fabricated by forming the mechanical joints in a combination of steel and pre-strained aluminum with prestraining levels of 0, 5%, and 10%, respectively with the same process parameters as that of the as-received aluminum. The coupons had the dimensions of 38.1 mm × 126 mm were prepared from 0.7 mm thick galvanized SAE1004 steel workpieces and 1.0 mm thick aluminum AA 6111-T4 workpieces. The rolling direction is parallel to the longer axis of the coupons. The lap-shear specimen configuration having an overlap of 15 mm which simulates the vehicle body flange presented in Fig. 3 was selected in this study. Because most of the joints made with the pre-strained aluminum were cracked even though the aluminum had only a 5% pre-strain in the available clinching tools, a TOX CEB15 Hydraulic and Pneumatic Clinch System with TOX@ SKB extensive die as shown in Fig. 1 was adopted. As shown, the SKB die has both solid and flexible segments. The clinched workpieces and the punch are centered by the fixed segments, and thereby guaranteeing that the joint formation is perfectly concentric. The mobile elements between the solid segments allow an undercut of the materials in the joint (TOX@ , 2014a). To examine the joint quality, the X parameter which is the bottom thickness at the center of the clinched joint shown in Fig. 1 was measured. Before fabricating a test sample, a proper set of joining process parameters were determined by a trial and error process. Two kinds of joints were fabricated. One is the as-received SAE1004 steel-as-received aluminum AA6111-T4 joints, and the other is the as-received SAE1004 steel-pre-strained aluminum AA6111-T4 joints. Utilizing existing guidelines for joint quality, satisfactory process parameters were determined by a combination of visual inspection and static joint strength. Table 4 lists the process variables and tools used in this study. Both the as-received and pre-strained workpieces were fabricated with the same clinching parameters as shown in Table 4. Once the reference X parameter and corresponding punch force were determined, all joints for the same clinched workpieces were clinched with the same punch force. In order to reduce the scatter in the data, three duplicates were made for each pre-strain condition.

Table 4 Process variables for clinching of 0.7 mm thick galvanized SAE1004 and 1.0 mm thick AA6111-T4. Punch

Die

Punch force (kN)

Type

Diameter (mm)

Type

Diameter (mm)

Depth (mm)

TOX 10.240.341995

5.6

TOX SKB14.180.226579

8.2

1.0, 1.2, 1.4

32, 40, 48, 56

T. Jiang et al. / Journal of Materials Processing Technology 215 (2015) 193–204

197

Fig. 4. Configuration of electrical resistance test.

2.4. Characterization of clinched joints To investigate the quality of the clinched joint, the joints were characterized by three section parameters, i.e., undercut, neck thickness and bottom thickness refer to Fig. 1. The joints were sectioned and examined using an optical microscope (OEM). Three replicates were performed for each condition and the average value was reported for each parameter. 2.5. Quasi-static testing Quasi-static tests were performed by loading each specimen to failure in a MTS 810 tester. To minimize the bending stresses inherent in the testing of lap shear specimens, filler plates were attached to both ends of the sample using masking tape to accommodate the sample offset as shown in Fig. 3. Load vs. displacement curves were obtained as the specimens were loaded at a stroke rate of 2 mm/min. Three replicates were performed, and the average peak loads were reported. Post-failure analysis was performed with optical microscopy to study the failure mechanisms. 2.6. Electrical resistance test To evaluate the ductile damage in the clinched joints, the change of electrical resistance of the clinched joint during the clinching process was measured to indirectly monitor the effect of pre-strain of the clinched workpieces on the ductile damage. The electrical resistance of the joints was measured with a STR-B DCR (i.e., direct current resistance) test instrument during the clinching process. Four-probe method was introduced to eliminate the effects of contact resistance and conductor resistance. It had been applied to measure the electrical resistance for pure aluminum and aluminum alloys (Park and Niewczas, 2008) as well as porous metal (Zhou et al., 2012). Fig. 4 presents the schematics of electrical resistance test method. As shown, to ensure the electric current passes through the joint section, the positive voltage and current electrodes were fixed on the side of the upper workpiece while the negative voltage and current electrodes were symmetrically fixed on the opposite side of the lower workpiece. The voltage and current electrodes should be closely contacted with the upper and lower workpieces to ensure the measurement accuracy. The distance between the voltage electrodes was kept at 5 mm from the end of overlap and the distance between the voltage electrode and

current electrode was kept at 10 mm. During the measurement, once the punch was lowered to clinch the workpieces, DCR test instrument supplied a direct current of 10 ampere passing through the clinched joints along the positive and negative current electrodes. And then the instantaneous electrical voltage between the positive and negative voltage electrodes in the clinched joints was measured online. After eliminating the internal resistance of the instrument and wire using the build-in program in the DCR test instrument, the value of the electrical resistance of the clinched joints was obtained in real time. Two kinds of electrical resistance were obtained. One is the initial electrical resistance R0 of the upper and lower workpieces, and was measured after the workpieces clamped by the holder as shown in Fig. 1 prior to clinching. To obtain the consistent measurement while avoiding the plastic deformation of the workpieces, 8 kN pre-load was exerted to ensure the close contract between two clinched workpieces during the measurement of R0 . The other is the finial electrical resistance R of the clinched joint and was measured after the joint was clinched. The values of both R0 and R vary with the pre-straining and punch force. The variation of electrical resistance R, which is the difference of the electrical resistance before clinching (R0 ) and after clinching (R), is defined as: R =

R − R0 × 100% R0

(1)

R is used to evaluate the effect of the pre-strained aluminum on the quality of the clinched joints. In order to reduce the measurement scatter, three duplicates were made for each case. 3. Results and discussion 3.1. Effect of process parameters on clinched steel-to-as-received aluminum joint quality To understand the effect of process parameters on the quality of the clinched joints, 0.7 mm thick galvanized SAE1004 steel and 1.0 mm thick as-received aluminum AA6111-T4 were clinched with various die depths and punch forces, and the cross-section and joint strength were examined. Fig. 5 presents the typical cross-section of the clinched joints under a punch force of 40 kN. As shown, steel and aluminum workpieces were interlocked intimately by the clinching. To characterize the joint strength, the section parameters and strength of the joints as a function of the punch force

198

T. Jiang et al. / Journal of Materials Processing Technology 215 (2015) 193–204

Fig. 5. Typical section of clinched 0.7 mm thick galvanized SAE1004 steel and 1.0 mm thick as-received aluminum AA6111-T4 joints with a punch force of 40 kN.

Fig. 6. Relation between the section parameters and joint strength of clinched 0.7 mm thick. SAE1004 and 1.0 mm thick as-received AA6111-T4 joints fabricated with a die depth of 1.2 mm under various punch forces.

and die depth were measured and the results are presented in Figs. 6 and 7, respectively. It can be seen that increasing the punch force and decreasing the die depth increased the joint strength. The joint strength increased initially and reached a maximum at a punch force of 48 kN. The analyses of the test results revealed that increasing the punch force resulted in an increase in undercut which provides the interlock between the workpieces and consequently, is beneficial to the joint strength. However, as shown in Fig. 6, as the punch force was raised to 56 kN the undercut increased and the neck thickness remained unchanged but the joint strength

decreased. These results revealed that the joint strength was barely related to the undercut and neck thickness. We speculated that this was likely attributed to the excessive punch forces which caused significant strain hardening, and consequently ductile damage in the clinched materials. As the joint strength was related to the failure mode, the influence of punch force and die depth on the joint strength can be identified from the failure mode. Fig. 8(a–c) presents the three different failure modes for the clinched joint, that are, button pull-out, neck cracking and a mixture of button pull-out and neck cracking failure modes, respectively. Test results showed that a low punch force (e.g., 32 kN) resulted in insufficient deformation of the steel and aluminum during the clinching process, and thus led to a minor undercut shown in Fig. 6. The small undercut produced a weak interlock between the steel and aluminum workpieces, which resulted in the joints failing in a button pull-out mode as shown in Fig. 8(a) and poor joint strength. On the other hand, a large punch force (e.g., 56 kN) and deeper die depth (e.g., 1.4 mm) increased the plastic deformation in the clinched region. This likely not only caused severe localized strain hardening in the clinched region, and thus caused a severe strain hardening but induced more ductile damage in the neck regime of the clinched joint shown in Fig. 8(b). This may overrule the increase in undercut, and thereby lead to a decrease in joint strength. A good balance of die depth and punch force (e.g., a punch force of 48 kN and a die depth of 1.2 mm as shown in Figs. 6 and 7) led to a mixed button pull-out and neck cracking failure mode as shown in Fig. 8(c), which corresponded to the optimum joint strength. It is noted from Fig. 6 that though the joint strength increased with a decrease in bottom, the bottom thickness barely reflected the adverse effect of ductile damage on the joint strength. As shown in Fig. 6, although the bottom thickness continuously decreased at the 56 kN punch force, the joint strength decreased. Based on the above discussion, the reduction in joint strength can be attributed to the ductile damage caused by the excessive punch force applied to the clinched aluminum workpieces. Therefore, the bottom thickness may be an indicator for the quality of the clinched joints. But it barely detected the effect of ductile damage induced by severe plastic deformation of the clinched materials on the joint strength. 3.2. Relationship between the section parameters and joint strength To further understand the effects of section parameters on the joint strength, the results shown in Fig. 6 were replotted in Fig. 9. As shown, iso-shear strength map for the different combinations of the undercut and neck thickness was derived in the range of experimental data. The combined effects of undercut and neck thickness on the joint strength can be approximately evaluated. It can be seen that although various combinations of the undercut and neck thickness produced the same joint strength, there is an optimum combination of bottom thickness and undercut in terms of tool life for a given joint strength. Experimental observations showed that while a large punch force (∼48 kN) is beneficial to the joint strength, it may damage the punch and die, and consequently shorten the tool life. Therefore, the determination of punch force should consider the tool life of the punch and die as well as the desired joint strength. To obtain the desired joint strength, it is important to achieve a good balance between the undercut and neck thickness by selecting the proper tools and clinching process parameters. 3.3. Effect of aluminum pre-straining on joint quality

Fig. 7. Relation between the section parameters and joint strength of clinched 0.7 mm thick galvanized SAE1004 and 1.0 mm thick as-received AA6111-T4 joints fabricated with various die depths under a punch force of 40 kN.

The aforementioned results are for the clinched joint made with the as-received aluminum AA6111-T4 and galvanized SAE1004 steel sheets. In reality, the sheet workpieces are pre-strained (e.g. stamped) before clinching. The strain hardening resulting from

T. Jiang et al. / Journal of Materials Processing Technology 215 (2015) 193–204

199

Fig. 8. Fractograph of clinched 0.7 mm thick galvanized SAE1004-to-1.0 mm thick as-received AA6111-T4 joints: (a) button pull-out (b) neck cracking, and (c) a mixed button pull-out and neck cracking.

pre-straining and induced ductile damage may limit the ability of the material to be clinched, and consequently degrade the joint quality. In this study, only aluminum workpiece was prestrained. 1.0 mm thick aluminum AA6111-T4 was pre-strained to various amounts prior to clinching. Clinched joints were made

of 0.7 mm thick galvanized SAE1004 steel and 1.0 mm thick prestrained aluminum AA6111-T4 with the same process parameters as that as-received aluminum shown in Table 4. Table 5 and Fig. 10 present the effect of pre-strained aluminum AA611-T4 on the discrepancy propensity (e.g., cracking) and joint appearance, respectively. Experimental observations indicated that while the clinched AA6111-T4-to-SAE1004 joints made with a 10% pre-strain aluminum had the severe discrepancies at the locations normal and parallel to the rolling directions (Fig. 10(b)), the joints made with less pre-strain aluminum (<10%) had most of the discrepancies at the locations normal to the rolling direction. Furthermore, as shown in Table 5 while few discrepancies were observed for the joints with no pre-strain under the die depths of 1.0 mm to 1.4 mm, the joints made with only a 5% pre-strain aluminum were seriously cracked at the corners of the aluminum button under a die depth of 1.4 mm. These results implied that pre-straining of the aluminum decreased the clinchability of aluminum AA6111-T4 and the initial plastic anisotropy of the pre-strained aluminum prior to pre-straining had a significant effect on the discrepancies in the clinched joints. As the Table 5 Effect of the pre-strained aluminum and die depth on the quality of clinched 0.7 mm thick galvanized SAE1004 and 1.0 mm thick AA6111-T4. Die depth (mm)

1.0 Fig. 9. Combined effects of the undercut and neck thickness on the strength of clinched 0.7 mm thick galvanized SAE1004 and 1.0 mm thick AA6111-T4 joints.

1.2 1.4

Pre-strain (%)

0

2.5

5

7.5

10

No apparent cracking (NAC) NAC NAC

NAC

NAC

NAC

NAC

NAC NAC

NAC Cracking

NAC Cracking

NAC Cracking

200

T. Jiang et al. / Journal of Materials Processing Technology 215 (2015) 193–204

Fig. 10. (a) cracking site, (b) appearance, and (c) cross-section of clinched 0.7 mm thick galvanized SAE1004 and 1.0 mm thick AA6111-T4 with a 5% pre-strain under a die depth of 1.4 mm.

work hardening resulting from pre-straining aluminum decreased the ductility of aluminum and induced some ductile damage prior to clinching, the localized ductility loss in the clinched regime not only resisted the plastic flow of the materials in the die but also introduced additional ductile damage. The existing damage in the pre-strained aluminum may be further compounded the ductile damage under exertion of the multi-axial stresses. The anisotropic mechanical properties may further decrease the strength and ductility of the strained aluminum, and consequently result in more discrepancies. It is noted from Fig. 10 that the discrepancies of the clinched joints appeared at the outer radius of the protrusion into the die with respect to the plane parallel to the specimens in the middle of the section, which correspond to the flexible segment of the SKB die. As discussed by Lambiase and Di Ilio (2013), the cavity volume was modified during the clinching process. The radial displacement of the flexible segment, and in turn, the die cavity volume, depended on the developing system of forces acting on the flexible segment surfaces. As shown in Fig. 10(a), as the material was deformed into the die, the sheet contacted the solid segments of the die which acted to lock the sheet materials. With the punch stroke increased, the horizontal component of contact pressure on the flexible segment surface increased and triggered the flexible segment motion, and in turn the punch-die cavity volume was enlarged. The material between two adjacent solid segments continued to deform locally because of the flexible die segment. Therefore, the quality of the clinched joint is closely related to the materials flow in the extensible die. To alleviate this problem, a decrease in die depth may reduce the plastic deformation of the clinched workpieces, and thus decrease the ductile damage in the clinched regime, which

may be helpful to decrease the discrepancies of the pre-strained workpieces. Test results listed in Table 5 showed that while the discrepancy propensity of aluminum workpieces increased with an increase in pre-strain, it decreased with a decrease in die depth. The effect of discrepancies of aluminum workpieces on the joint strength will be discussed in the next section. 3.4. Effect of aluminum pre-straining on joint strength To understand the effect of pre-strain levels in the aluminum on the joint strength of 0.7 mm thick galvanized SAE1004 steel and 1.0 mm thick pre-strained aluminum AA6111-T4 clinch joints, the aluminum was pre-strained to various levels and the dissimilar metal workpieces were clinched together and the results are shown in Fig. 11. As shown, the level of aluminum pre-straining had a significant effect upon the joint strength. A 5% pre-strain caused an about 20% decrease in joint strength. The joint strength with a prestrain of 10% exhibited little difference as compared to that with a pre-strain of 5%. To demonstrate the influence of die depth on the section parameters and strength of the clinched joints made with pre-strained aluminum, the relations between the die depth, section parameters and joint strength are replotted in Fig. 12. Similar to the results shown in Fig. 7, an increase in die depth decreased the joint strength, especially for the joints fabricated with a die depth of 1.4 mm. Associated with the discrepancies in the clinched joints as shown in Table 5, the joint strength was decreased by as much as 20% compared to the joints made with the as-received aluminum AA6111-T4. It suggests that the discrepancies at the corner of

T. Jiang et al. / Journal of Materials Processing Technology 215 (2015) 193–204

201

stress-strain measurements. As shown, the pre-strain had little influence upon the tensile strength but increased the yield strength and decreased the elongation of the as-received AA6111-T4. The increase in yield strength indicated the pre-straining produced high density dislocations in the pre-strained aluminum. These pre-existed dislocations may have interacted with the additional dislocations produced during the clinching forming. Severe stress concentration in the pre-strained aluminum was likely produced, and consequently resulted in additional damage in the clinched region. Furthermore, the decrease in elongation of the pre-strained aluminum meant that pre-straining decreased the formability of aluminum workpieces during the clinching process. Therefore, the pre-strained aluminum is not only more susceptible to the discrepancies than the as-received AA6111-T4 but resulted in a reduced amount of undercut in the clinched joints as shown in Fig. 11. These appeared to be the reasons for the reduction of the joint strength shown in Figs. 11 and 12. It is noted from Fig. 11 that the pre-strained aluminum only induced a minor decrease in undercut and neck thickness while the joint strength was significantly degraded. Careful examinations of the test results suggested that the decrease in joint strength was likely attributed to the pre-strain-induced ductile damage in the clinched region. As the electrical resistance is sensitive to the variation of the dislocation density (Brown, 1977), it may be used as an indirect indicator of damage caused by pre-strain prior to clinching. To examine if ductile damage introduced by pre-straining, the electrical resistance of the clinched pre-strained joints was measured. Fig. 14 presents the relation among the pre-strain, variation of electrical resistance (R) and joint strength. It can be seen from Fig. 14 that though high initial electrical resistance R0 of the prestrained aluminums resulted in a decrease in variation of electrical resistance R for clinched pre-strained aluminum joints. Similar trend was observed for the joint strength. According to the cross-property connection between electrical resistance and strain hardening coefficient of the metal in the process of plastic deformation (Dominguez and Sevostianov, 2011), the increase in R caused by pre-strain prior to clinching is primarily attributed to high dislocation density in the pre-strained aluminum. When the pre-strained aluminum was clinched, the severe plastic deformation created new dislocations. The dense dislocations existed in the pre-strained aluminum would block the multiplication of new dislocation in the clinched regime, and consequently, would create more dislocation pile-up and stress concentration than that of the as-received aluminum. These would induce the occurrence of ductile damage and decrease the load capacity of the clinched joint. Combined Figs. 11 and 14, it can be seen that there is significant correlation between the joint strength and pre-strain of aluminum instead of the section parameters. Therefore, the ductile damage induced by pre-strain may be more relevant than the section parameters in terms of relating to the joint strength. Fig. 11. Effect of pre-strained aluminum on the strength of clinched 0.7 mm thick galvanized SAE1004 and 1.0 mm thick AA6111-T4 joints fabricated with a die depth of (a) 1.0, (b) 1.2 and (c) 1.4 mm.

button decreased the constraint of the material and it fully moved outward filling the die and resulting in an increased amount of undercut. This further caused the neck discrepancy of steel workpieces, which resulted in the joints failing in neck cracking mode shown in Fig. 8(b) thereby reducing the joint strength. Again, the decrease in joint strength is likely attributed to the strain hardening resulting from the pre-strained aluminum prior to the clinching. To validate this, experiments were performed to assess the effect of pre-strain on the mechanical properties of the as-received AA6111-T4 and the results are shown in Fig. 13 in which the strength and elongation were obtained from the engineering

3.5. Relation between bottom thickness and strength of clinched pre-strained aluminum As shown in Fig. 6, the bottom thickness can be used to monitor the quality of clinched galvanized SAE1004 steel – as received aluminum AA6111-T4 joint. To investigate if it can be used to monitor the quality of the joints made with the pre-strained aluminum, aluminum AA6111-T4 was pre-strained to 10% and then clinchedto-0.7 mm thick galvanized SAE1004 steel with a die depth of 1.2 mm under various punch forces. The strength and bottom thickness of the clinched joints were measured and the results are shown in Fig. 15. As shown, the bottom thickness correlated well with the strength for the clinched joints made with and without the pre-strained aluminum.

202

T. Jiang et al. / Journal of Materials Processing Technology 215 (2015) 193–204

Fig. 12. Effect of die depth on the lap-shear strength (a), bottom thickness (b), undercut (c) and neck thickness (d) of clinched 0.7 mm thick galvanized SAE1004 and 1.0 mm thick pre-strained AA6111-T4 joints.

To further validate the use of bottom thickness for monitoring the pre-strained aluminum joint quality, the correlations among the punch force, bottom thickness and strength of the clinched joints are presented in Fig. 16. For comparison purpose, the variation of electrical resistance of the clinched joints is included in Fig. 16. It can be seen that in general an increase in punch force

Fig. 13. Effect of pre-straining on the mechanical properties of aluminum AA6111T4.

decreased the bottom thickness. While there was an apparent difference in joint strength between the joints made with and without the pre-strained AA6111-T4 for a given punch force, there was little difference in bottom thickness. Especially for the punch force

Fig. 14. Effect of pre-straining on the variation of electrical resistance and joint strength of the clinched 0.7 mm thick galvanized SAE1004 and 1.0 mm thick AA6111-T4 fabricated with a punch force of 40 kN.

T. Jiang et al. / Journal of Materials Processing Technology 215 (2015) 193–204

203

the R to perceive the initial state of the clinched workpieces and the difference in strength between the joints made with the as-received and pre-strained aluminum. Therefore, the optimum clinching tools and process cannot be derived from the experiments using the as-received aluminum workpieces. To obtain a quality clinched pre-strained aluminum joint, it needs to optimize the clinching tools and process parameters by considering the ductile damage in the pre-strained aluminum. A damage-sensitive method, such as the variation of electrical resistance of the clinched joint, is needed to monitor the quality of the clinched steel-to-prestrained aluminum. 4. Conclusions

Fig. 15. Correlation between the bottom thickness and peak load for the joints made with 0% and 10% aluminum pre-straining.

over 48 kN, the joints made with and without the pre-strained aluminum virtually had the similar bottom thickness even though the pre-strained aluminum decreased the joint strength. As discussed earlier, the joint strength degradation caused by the pre-strain and excessive punch force was related to the ductile damage in the clinched aluminum. The results in Figs. 15 and 16 show that though the bottom thickness was a good indicator for detecting the strength variation for the clinched as-received and pre-strained aluminum–steel joints, it barely detected the effect of ductile damage in the pre-strained aluminum the joint strength. In production, the quality of the clinched joint is often monitored with the “X” parameter ( Varis, 2006b) which is the bottom thickness of the clinched joints. With the use of this approach, because the as-received aluminum resulted in a thinner bottom thickness than the pre-strained aluminum under the given clinching process parameters shown in Fig. 16, a larger punch force is required to obtain the similar “X” parameter as that of the as-received aluminum. As a result, it would overestimate the strengths of the joints made with the pre-strained aluminum. Different from the bottom thickness, the variation of electrical resistance R was proposed to monitor the joint quality. The electrical resistance of the joints was measured on-line during the clinching process and the results are included in Fig. 16. As shown, the corresponding relations between the joint strength and R are significant. It was apparent from

(1) The work hardening resulting from pre-strained aluminum AA6111-T4 likely induced the ductile damage in the clinched workpieces during the clinching process. To obtain sound clinched pre-strained aluminum–steel workpieces, the use of a shallow die to accommodate the pre-strain in aluminum workpiece is preferred. (2) Pre-straining of aluminum prior to joining had a significant effect on the strength of the clinched aluminum–steel joints. A 5% pre-strain resulted in as much as 20% reduction in strength of the clinched 0.7 mm thick galvanized SAE1004 and 1.0 mm thick AA6111-T4 joint. The significant reduction of joint strength was likely attributed to the combined effects of strain hardening and ductile damage in the pre-strained aluminum. (3) The bottom thickness barely detected the effect of ductile damage in the pre-strained aluminum the joint strength. The application of the similar “X” parameter from the as-received aluminum to monitor the quality of the clinched steel-prestrained aluminum joint resulted in overestimates of the joint strength. As the electrical resistance of the clinched joint can indirectly indicate the ductile damage of the pre-strained aluminum, it may be an alternative method to monitor the quality of the clinched steel-to-pre-strained aluminum. (4) To obtain the desired strength of the clinched 0.7 mm thick galvanized SAE1004 steel and 1.0 mm thick pre-strained aluminum AA6111-T4 joint, it is important to balance between undercut and neck thickness by selecting proper punch and die and process parameters (i.e., die depth and punch force) through considering the pre-strain and ductile damage in the pre-strained aluminum. Acknowledgement The authors gratefully acknowledge the financial (Grant number: TCS49382) and technical support provided by GM Global Research and Development Center to carry out the present work. References

Fig. 16. Effect of pre-strained aluminum on the bottom thickness and strength for the joints made with a die depth of 1.2 mm under various punch forces.

Adeosun, O., Sekunowo, O.I., Balogun, S.A., Osoba, L.O., 2011. Effect of deformation on the mechanical and electrical properties of aluminum–magnesium alloy. J. Miner. Mater. Charact. Eng. 10, 553–560. Briskham, P., Blundell, N., Han, L., Hewitt, R., Young, K., Boomer D., 2006. Comparison of self-pierce riveting, resistance spot welding and spot friction joining for aluminum automotive sheet http://www.sae.org/ SAE 2006. In: World Congress & Exhibition Technical Papers, SAE paper 2006-01-0774. Brown, R.A., 1977. Electrical resistivity of dislocations in metals. J. Phys. F: Metal Phys. 7 (7), 1283–1295. Coppieters, S., Lava, P., Sol, H., Van Bael, A., Van Houtte, P., Debruyne, D., 2010. Determination of the flow stress and contact friction of sheet metal in a multi-layered upsetting test. J. Mater. Process. Technol. 210, 1290–1296. Coppieters, S., Cooreman, S., Lava, P., Sol, H., Van Houtte, P., Debruyne, D., 2011a. Reproducing the experimental pull-out and shear strength of clinched sheet metal connections using FEA. Int. J. Mater. Form. 4, 429–440. Coppieters, S., Cooreman, S., Sol, H., Van Houtte, P., Debruyne, D., 2011b. Identification of the post-necking hardening behavior of sheet metal by comparison of

204

T. Jiang et al. / Journal of Materials Processing Technology 215 (2015) 193–204

the internal and external work in the necking zone. J. Mater. Process. Technol. 211, 545–552. Coppieters, S., Lava, P., Van Hecke, R., Cooreman, S., Sol, H., Van Houtte, P., Debruyne, D., 2013. Numerical and experimental study of the multi-axial quasi-static strength of clinched connections International. J. Mater. Form. 6, 437–451. Dominguez, D., Sevostianov, I., 2011. Cross-property connection between workhardening coefficient and electrical resistivity of stainless steel during plastic deformation. Int. J. Fract. 167, 281–287. Gibiansky, L.V., Torquato, S., 1996. Connection between the conductivity and bulk modulus of isotropic composite materials. J. Mech. Phys. Solids 44 (A452), 253–283. Hahn, O., Kurzok, J.R., Dolle, N., 2001. DaimlerChrysler – Technologiekolloquium: Simulation der Fugetechniken Potentiale und Grenzen (Simulation of mechanical joining technologies, Stuttgart, Germany, Januar)., pp. p35–p38. Hot note, 2014. Clinching-clinch monitoring with MP85 makes an innovative technique even safer, Available from: http://www.hbm.com/fileadmin/mediapool/ hotline/hl051 note2 clinching int.pdf Karolik, A.S., Luhvich, A.A., 1994. Calculation of electrical resistivity produced by dislocations and grain boundaries in metals. J. Phys.: Condens. Matter 6, 873–886. Kato, T., Abe, Y., Mori, K., 2008. Plastic joining of dissimilar material of high tensile strength steel and aluminum alloy sheets. Weld. Int. 22, 575–581. Lambiase, F., Di Ilio, A., 2013. Finite element analysis of material flow in mechanical clinching with extensible dies. J. Mater. Eng. Perform. 22, 1629. Lambiase, Francesco, 2013. Influence of process parameters in mechanical clinching with extensible dies. Int. J. Adv. Manuf. Technol. 66, 2123–2131. Markowski, T., Mucha, J., Witkowski, W., 2013. Eksploatacja i Niezawodnosc (FEM analysis of clinching joint machine’s C-frame rigidity). Maint. Reliab. 15, 51–57. Michalos, G., Makris, S., Papakostas, N., Mourtzis, D., Chryssolouris, G., 2010. Automotive assembly technologies review: challenges and outlook for a flexible and adaptive approach. CIRP J. Manuf. Sci. Technol. 2, 81–91. Miller, W.S., Zhuang, L., Bottema, J., Wittebrood, A.J., De Smet, P., Haszler, A., Vieregge, A., 2000. Recent development in aluminum alloys for the automotive industry. Mater. Sci. Eng. A 280, 37–49. Mucha, J., Kascak, L., Spisak, E., 2011. Joining the car-body sheets using clinching process with various thickness and mechanical property arrangements. Arch. Civil Mech. Eng. 1, 135–148. Neugebauer, R., Kraus, C., Dietrich, S., 2008. Advances in mechanical joining of magnesium. CIRP Ann. – Manuf. Technol. 57, 283–286. Park, D.-Y., Niewczas, M., 2008. Plastic deformation of Al and AA5754 between 4.2 K and 295 K. Mater. Sci. Eng. A 491, 88–102.

Qiu, R., Shi, H., Zhang, K., Tu, Y., Iwamoto, C., Satonaka, S., 2010. Interfacial characterization of joint between mild steel and aluminum alloy welded by resistance spot welding. Mater. Charact. 61, 684–688. Roux, E., Bouchard, P.-O., 2013. Kriging metamodel global optimization of clinching joining processes accounting for ductile damage. J. Mater. Process. Technol. 213, 1038–1047. Saberi, S., Enzinger, N., Vallant, R., Cerjak, H., Hinterdorfer, J., Rauch, R., 2008. Influence of plastic anisotropy on the mechanical behavior of clinching joint of different coated thin steel sheets International. J. Mater. Form. (Suppl. 1), 273–276. Sevostianov, I., 2003. Explicit relations between elastic and conductive properties of a material containing annular cracks. Philos. Trans. R. Soc. (Lond.) A-361, 987–999. Tan, Y., HAHN, Q., Du, F., 2005. Process monitoring method with window technique for clinch joining. ISIJ Int. 45, 723–729. Overview of the TOX@ Clinching Technology, 2014a, Available from: http://www. tox-us.com/fileadmin/pdf/us/joining systems/TB 80100 0508 US 10.pdf Process Monitoring CEP 400, 2014b, Available from: http://www.tox-us.com/ us/products/joining-systems/accessories/process-monitoring-cep-400.html Varis, J.P., (Licentiate Thesis) 1998. Clinching of zinc-coated high strength structural steel. Department of Mechanical Engineering, Lappeenranta University of Technology, Lappeenranta, Finland (in Finnish). Varis, J., 2002. The suitability of round clinching tools for high strength structural steel. Thin-Walled Struct. 40, 225–238. Varis, J.P., Lepisto, J., 2003. A simple testing-based procedure and simulation of the clinching process using finite element analysis for establishing clinching parameters. Thin-Walled Struct. 41, 691–709. Varis, J., 2006a. Economics of clinching joint compared to riveted joint and example of applying calculations to a volume product. J. Mater. Process. Technol. 172, 130–138. Varis, J., 2006b. Ensuring the integrity in clinching process. J. Mater. Process. Technol. 174, 277–285. Vincze, G., Rauch, E.F., Gracio, J.J., Barlat, F., Lopes, A.B., 2005. A comparison of the mechanical behavior of an AA1050 and a low carbon steel deformed upon strain reversal. Acta Mater. 53, 1005–1013. Zhou, W., Tangb, Y., Song, R., Jianga, L., Hui, K.S., Hui, K.N., 2012. Characterization of electrical conductivity of porous metal fiber sintered sheet using four-point probe method. Mater. Des. 37, 161–165.