Clinching of High-Strength Steel Sheets with Local Preheating

Abstract

Clinching is a manufacturing method of mechanically joining two or more materials without the use of heat or additional components. This process relies on high plastic deformation to create a secure bond. Clinching technology is widely used for joining materials of various grades and thicknesses. Especially in the automotive industry, clinching is an alternative to resistance spot welding. However, the load-bearing capacity of clinched joints is comparatively lower when compared to resistance spot-welded joints. This research aimed to increase the load-carrying capacity of clinched joints. To enhance the load-bearing capacity of the clinched joints, localized modification of the microstructure was carried out, primarily focusing on the neck area of the joint. The alteration of the microstructure within the clinched joint was accomplished through the application of localized heating using the resistance spot welding method. The microstructure distribution in the clinched joint region was analyzed using light and scanning electron microscopy, as well as microhardness measurements. Two material grades, micro-alloyed steel HX420LAD+Z and dual-phase ferritic–martensitic steel HCT600X+Z, were tested. Each grade underwent five groups of ten samples, which were subjected to identical experimental conditions of local heating by resistance spot welding (RSW) and clinching. The utilization of RSW on the clinched joint region resulted in an average enhancement of 17% in the load-carrying capacity for material HCT600X+Z, and an average increase of 25% for material HX420LAD+Z.

1. Introduction

Steel sheets of various grades and thicknesses are widely employed in the automotive industry to produce lightweight car bodies. Joining various body parts is essential, and commonly used methods for steel parts include resistance spot welding, laser welding, brazing, soldering etc. [1,2]. The oldest and most widespread method of joining steel sheets in the automotive industry is resistance spot welding. Resistance spot welding (RSW) is a joining technique that eliminates the need for additional material during the process. It involves applying high-intensity current and pressure with welding electrodes. The process utilizes the heat generated by the current to achieve the necessary welding temperature for joining two or more materials. However, in welding processes, thermal after-effects and distortion are obtained [3,4,5]. RSW, as the predominant technique for metal joining, is commonly employed for similar metal materials. However, when dissimilar metals are welded together, there is a risk of welding defects due to substantial variations in their physical and chemical properties. This can compromise the overall quality of the joints [6,7]. The process of RSW depends on the types of used materials and the specific welding parameters. The parameters typical for RSW are the welding current, welding time, and electrode pressing force. It is important to note that welding parameters significantly affect the quality of the joint [8,9]. Ma et al. [10] conducted research on the resistance spot welding of high-strength (HSS) steel sheets. They investigated the formation process of the nugget under different welding conditions using experimental measurements and FEM simulation. The FEM simulation accurately predicted the nugget sizes and formation process. The effect of the welding parameters on the quality of spot welds of a combination of various high-strength steels was investigated by Kaščák et al. [11]. Increasing the welding current values led to higher carrying capacities in a linear dependence. H. Oikawa et al. [12] described in their paper that the suitable welding current range of HSS sheets during resistance spot welding shifted to the lower current side compared to that of mild steel sheets, and it was affected by the electrode force. Midawi et al. [13] studied the impact of the microstructures and mechanical properties of resistance spot welds in advanced high-strength steels (AHSS) on the overall failure response of structural assemblies in crash events. Strain localization occurred in regions with lower hardness, such as the subcritical heat-affected zone (HAZ), while limited deformation was observed in harder regions, like the upper-critical HAZ and fusion zone. Eftekharimilani et al. [14] presented a study on the effects of single- and double-pulse resistance spot welding on the microstructures of an advanced high-strength steel (AHSS). The focus was on how the double-pulse welding schemes affected the primary weld nugget, particularly by partially remelting it and annealing the fusion boundary area. The double-pulse welds demonstrated a favorable failure mode. Chung et al. [15] analyzed the failure of resistance spot-welded AHSS. The authors concluded that in the examined welds, there was a significant decrease in the ductility at the weld nugget such that it resulted in interfacial failure. X. Wan et al. [16] showed in their article that the current welding parameter influenced the mechanical properties of resistance spot-welded dual-phase steel welds. H. Moshayedi and I. Sattari-Far [17] declared, as part of their research, that the current welding parameter had a significant impact on nugget formation compared to the welding time, using finite element analysis. M. Eshraghi et al. [18], in their article, concluded that the welding current was the most influential and important factor of the weld properties, using SYSWELD simulation software. B. Varbai et al. [19] investigated the shear tension strength of resistance spot-welded ultra-high-strength steels. The authors concluded that the softening in the sub-critical heat-affected zone was found to be greater with a longer welding time.

Due to factors such as high thermal conductivity, the presence of a natural surface oxide layer, a zinc layer, and other challenges, connecting lightweight materials through welding techniques poses difficulties. Furthermore, the emission of sparks and smoke during the welding process raises environmental concerns. Hence, other assembly process methods can be utilized in joining high-strength steel sheets, such as adhesive bonding, fastening, riveting, clinching, clinch–riveting, or self-piercing riveting [20,21].

Over the past three decades, clinching has emerged as a fast-growing mechanical joining method in the automotive and aerospace sectors. Clinching joining technology has emerged as a promising solution for lightweight constructions that require high production rates, especially when joining hybrid metal–composite parts [22,23]. Clinching is a fast and localized mechanical fastening technique that involves deforming two or more sheet components through plastic deformation. The process utilizes a punch, a blank holder, and a specialized die assembly. By pressing the forming sheets, which are initially clamped between the blank holder and the die assembly, a friction-based interlocking connection is created between the sheet components [24,25,26,27]. The clinching process of materials with various thicknesses and mechanical property arrangements was investigated by Mucha et al. [28]. The sheet thickness arrangement in relation to the die is an important parameter influencing the load-bearing capacity. Abe et al. [29] studied mechanical clinching, using modified dies, and the resistance spot welding that is used to join ultra-high strength steel sheets with low ductility. The static joint loads for resistance spot welding were larger than those of clinching in the tension–shearing and cross-tension tests.

In the automotive industry, clinched joints are taken as an alternative to resistance spot welding. The main disadvantage of clinched joints is their lower load-bearing capacity compared to resistance spot welds [29,30,31]. This results in efforts to increase the load-bearing capacity of clinched joints. Efforts to increase the load-bearing capacity of clinching joints have led to research on the heat treatment of clinched joints and its combination with adhesive bonding.

Lei et al. [32] studied the application and evaluation of clinch–bonded hybrid joining technology in various manufacturing disciplines. The study focused on specimens made of similar and dissimilar sheets of H62 copper alloy, aluminum alloy, and galvanized steel. The shear strength of the specimens, primarily influenced by the adhesive, outperformed their peeling strength, which was closely associated with the clinched structures. Ma et al. [33] investigated the dynamic flow behavior of adhesive during the clinch–bonding process and its impact on joint formation. A lower clamping force should be employed in clinch–bonding to ensure the sufficient handling strength of the joint before the adhesive is cured. Balawender [34] dealt with two different manufacturing procedures: adhesive curing before clinching and clinching before adhesive curing. They found that clinching before adhesive curing was a more effective technology for achieving stronger joints.

Sia et al. [35] studied the behavior and durability of clinch joints at room temperature, 100 °C, and 250 °C. The ultimate tensile–shear strength remained relatively unchanged with a temperature increase, but the stiffness and energy absorption decreased at higher temperatures. Zeuner et al. [36] investigated the effect of heat treatment on the forming process and the mechanical properties of clinch joints made from a highly formable sheet material. The study examined the impact of heat treatment, which involves the formation of an oxidation layer on the sheet surface, on the forming process and resulting mechanical properties of the clinch joints. The findings demonstrated clear influences of heat treatment-induced surface roughness on the joint geometry and strength. Kaczyński et al. [37] studied a new method to enhance the strength of clinch joints, which involved heating the sheet from the side of the die prior to the joining process. The results demonstrated that increasing the temperature of the bottom component before deformation led to an increase in the shear strength as well as in the cross-tensile strength. Zhang et al. [38] evaluated a novel method of joining for aluminum alloy 5052. The method combined principles of mechanical clinching and resistance spot welding. However, the resulting joint shape was not a clinched joint with its characteristic shape. The results demonstrated that the resistance spot clinched joints exhibited superior load-bearing capacity compared to traditional resistance spot welding joints. Wang et al. [39] introduced a novel method called incremental laser shock clinching, which is used to create a round-end rectangular joint with a total length much larger than the diameter of the laser spot. The research verified the feasibility of this method to join copper, aluminum, and stainless steel. The development of the hot stamping clinching tool was investigated by Chen [40]. The stamping clinching tool incorporates a forming system, a heating system, and a cooling system. They found that the high cooling rate employed in the experiments increased the tensile strength of the sheet at the clinching point by 3 to 4 times. A novel joining method called electro-hydraulic clinching was introduced by Babalo et al. [41]. By utilizing electrical energy discharge to generate a shock wave in the fluid, a clinching joint is formed. The resulting joint efficiency was approximately twice as high as that achieved through the traditional clinching process. Abe et al. [42] studied the rectangular shear clinching process of ultra-high strength steel sheets. The study indicated that the rectangular shear clinching process, even without an airtight joint, was effective for joining ultra-high-strength steel sheets with low ductility and high flow stress. The same author [43] presented a clinching technique with the application of counter-pressure from a rubber ring for joining galvanized ultra-high-strength steel sheets. The interlock between the sheets was enhanced by promoting metal flow using the counter-pressure exerted by the rubber ring in the die cavity. It was observed that the appropriate shape of the rubber ring significantly improved the joinability, enabling the successful joining of the sheets with the aid of counter-pressure.

Quite extensive research has been devoted to mechanical joining methods using the rivet as a fastener. The methods using this fastener are mainly clinch–riveting and self-piercing riveting [44]. These techniques are employed in the joining of sheet materials, establishing a mechanical interlock between the sheets. This method has gained significant attention in the automotive industry due to its suitability for joining lightweight, high-strength, and dissimilar materials.

The biggest shortcoming of the clinching method is the low load-bearing capacity of the joints compared to resistance spot welding and other methods of mechanical joining, such as clinch–riveting and self-piercing riveting. This study deals with increasing the load-bearing capacity of clinched joints by locally modifying the microstructure of the joined materials HCT600X+Z and HX420LAD+Z, especially in the area of the neck of the joint. A local change in the microstructure of the clinched joint was achieved by using resistance spot welding. Numerical simulation software was used to predict the material flow of the preheated microstructure of the joined materials during the clinching process. The distribution of the thus-modified microstructure in the clinched joint was analyzed by light and electron microscopy techniques and by measuring the microhardness. The load-bearing capacity of clinched joints treated by local heating in this way was compared with the load-bearing capacity of joints without heat treatment. This method of joint modification can be applied to achieve higher strength of the joined units.

2. Materials and Methods

2.1. Experimental Materials

The selection of the experimental material was targeted at two-phase hot-dip galvanized steels 1.5 mm thick and fine-grained micro-alloyed low-carbon hot-dip galvanized steels 1.5 mm thick. The material of the first experimental group belongs to progressive high-strength steels. Its designation, according to EN 10346/09 [45], was HCT 600 X+Z. The HCT 600 X+Z steel sheet had a surface galvanized on both sides, and the thickness of the sheet was 1.5 mm. In the second experimental group, the steel sheet HX420LAD+Z was galvanized on both sides, and its thickness was 1.5 mm. The chemical composition of the steels was determined by optical emission spectroscopy (OES, Thermo ARL, Lausanne, Switzerland). The chemical elemental composition of the experimental materials was focused on the determination of the concentration of the main alloying elements (

Table 1. Chemical elemental composition of HCT600X+Z steel, No. 1.0941 (wt.%).

C	Mn	Si	P	S	Al	Cu	Ni	Cr	Ti	V	Nb	Mo
0.075	1.880	0.016	0.015	0.005	0.055	0.022	0.009	0.208	0.001	0.003	0.003	0.170

and

Table 2. Chemical elemental composition of HX420LAD+Z, No. 1.0935 (wt.%).

C	Mn	Si	P	S	Al	Cu	Ni	Cr	Ti	V	Nb	Mo
0.083	1.214	0.013	0.015	0.005	0.040	0.035	0.010	0.024	0.014	0.032	0.058	0.002

).

The mechanical properties of the experimental materials were determined by tensile testing at an ambient temperature according to ISO 6892-1:2019 [46] on flat samples 1.5 mm thick, with a measured length of L₀ = 80 mm and a cross-member movement rate of 0.05 mm/min. For the ambient temperature tensile tests, a TiraTest 2300 universal testing machine (TIRA GmbH, Schalkau, Germany) with a force transducer up to 100 kN was used. The basic mechanical properties of the joined materials are shown in

Table 3. Mechanical properties and thicknesses of joined materials.

Material	Thickness (mm)	Rp0.2 (MPa)	Rm (MPa)	A80
HX420LAD+Z	1.5	503	565	19.5
HCT600X+Z	1.5	343	593	27.5

.

HCT600X is a dual-phase (DP) steel with a multiphase structure consisting of hard martensite islands dispersed within a soft ferrite matrix. DP steel finds extensive application in the automotive sector. It is commonly employed in various structural components such as automobile chassis, beams, guide rails, and other parts. These parts experience intricate loads during practical usage, particularly in areas near geometric irregularities where stress tends to concentrate. Consequently, local plastic deformation may transpire when subjected to extreme loads [16,18,47].

HX420LAD is a high-strength low-alloy (HSLA) steel, a subset of high-strength steels (HSSs). HSLA steels have gained significant popularity due to their improved weldability and cost-effectiveness. HSLA steel offers several advantages, such as high specific strength and good low-temperature toughness. Compared to traditional steel, it enables reduced material consumption, enhanced reliability, and decreased costs when manufacturing components with an equivalent load-bearing capacity [48]. The combination of micro-alloying and controlled rolling allows for the attainment of high strength, good toughness, and improved weldability in low-carbon plate steels. These improvements in the mechanical properties primarily stem from the refinement of the ferrite grain size and the controlled precipitation strengthening. Micro-alloying elements like titanium, niobium, and vanadium aid in grain refinement by precipitating in the austenite phase, and they contribute to dispersion hardening by precipitating in ferrite during the austenite–ferrite transformation [49,50]. The effectiveness of titanium, niobium, and vanadium in these processes depends on the solubilities of their nitrides and carbides in austenite. HSLA steel can be subjected to severe dynamic and cyclic loading, which may lead to fatigue damage [51].

2.2. Numerical Simulation

Mechanical clinching is an inexpensive joining process that involves cold-joining two sheets through localized hemming using a punch and die. In the traditional mechanical clinching method, the punch and die are utilized to shape the sheets and create an interlock between the lower and upper sheets. During the clinching forming process, the punch applies pressure to the upper sheet, causing it to flow onto the lower sheet. The lower sheet is supported by a rigid die. As a result, both sheets undergo plastic deformation, leading to the creation of a mechanically interlocked structure known as the clinched point [52,53].

The typical steps of creating the clinched joint include clamping, pressing, bottom-forming, and interlock-forming (Figure 1). The pre-forming of the parts has a significant impact on the outcome of the joining process. Particularly, the pre-forming of the material on the punch side has a greater influence on the joining parameters compared to the material on the die side.

Figure 2a shows the thickness of the neck (t_N), and the joint bottom thickness (X) and interlock (t_u) are the main parameters that significantly affect the load-bearing capacity of the clinched joint [27,28,29]. The goal of the numerical simulation was to identify the zones in the sheet metal that are moved to the area of the neck of the clinched joint.

Ansys Workbench software was used for a numerical simulation of the clinching process. To simplify the task, a 2D model was employed due to the axial symmetry of the tools and the joint. Consequently, the anisotropy of the joined materials was not taken into account in the calculation. Figure 2b illustrates a schematic representation of the simulation model’s input geometry, including the basic dimensions and the limiting conditions of the simulation. The nodal points of the die were removed to prevent movement during the simulation, while the nodal points of the punch elements were allowed to move only along the Y axis. Another boundary condition involved applying a force of 400 N on the sheet holder to secure the joined materials during the simulation of the CL joining. Once the desired path was achieved, the punch moved backward to its starting position, simultaneously releasing the force from the sheet holder. In this simulation, the punch, die, and sheet holder were considered perfectly rigid and non-deformable bodies, whereas the joined sheets were modeled as deformable bodies. A multilinear material model was employed to describe the response of the joined material to the applied load. The material model input values for the joined materials can be found in Table 3. The plastic portion of the material model followed the Holomon hardening law, with the strain hardening exponent determined experimentally based on a standardized tensile test.

For the joined materials, a finite element mesh consisting of 2D quadrangular elements (PLANE 182) was used. These elements are recommended for simulating 2D axis-symmetric problems. Each node of these elements had two degrees of freedom, representing displacements along the X and Y axes. The element size of the mesh was set to 0.1 mm.

In processes like clinching, where both the sheet and the finite element mesh undergo significant deformation, it becomes necessary to remesh the elements at certain intervals [54]. This remeshing process, known as adaptive remeshing, is controlled by the deformation energy criterion.

At predefined intervals, the deformation energy of each element in the finite element mesh was monitored. If the energy exceeded the limits calculated by the program, adaptive remeshing was initiated. Frictional contacts were modeled between the tools and the joined materials, as well as between the joined materials themselves, using friction coefficients of 0.12 for the tool–sheet contact and 0.2 for the sheet–sheet contact.

2.3. Local Preheating

For the investigation of the clinching method with local preheating, overlapped samples with the same material combination were prepared: HX420LAD+Z with HX420LAD+Z and HCT600X+Z with HCT600X+Z.

Local intense heating (RSW) was applied to a pair of overlapping hot-dip galvanized sheets, each 1.5 mm thick. The local heating occurred at the location of the future clinched joint (Figure 3a). The local heating was performed with a BPK 20 spot pneumatic welder (VTS Elektro s.r.o. Bratislava, Slovakia) with adjustable welding parameters (electrode pressure, welding time, welding current), and with 30 different pre-set welding programs. The heating parameters were selected on the basis of the experience gained in the field of welding and then optimized for the implementation of local heating of the material. For local heating, Cu-Cr welding electrodes were used. The working surfaces of the welding electrodes were Ø8 mm with a conical finish.

The diameter of the welding electrode was selected based on a die diameter of Ø8 mm for clinching joining. The resistance spot welding parameters used in the preheat optimization are shown in

Table 4. Welding parameters for optimization of preheating process for both investigated materials.

Sample Series	Welding Current I (mm)	Welding Time T (per.)	Pressing Force Fz (kN)
I	3.5	12	6
II	3.5	18	6
III	3.5	24	6
IV	3.5	30	6
V	3.5	36	6
VI	3.9	12	6
VII	3.9	18	6
VIII	3.9	24	6
IX	3.9	30	6
X	3.9	36	6
XI	4.3	12	6
XII	4.3	18	6
XIII	4.3	24	6
XIV	4.3	30	6
XV	4.3	36	6

. Preheating optimization started from the lowest of the listed values. The aim was to achieve a change in the microstructure that reached the neck region of the joint (t_N) during the clinching process. However, a spot weld with a typical weld nugget was not allowed to occur. After the preheating phase ended, the samples were joined by the clinching method in place of the local change in the microstructure.

Based on the optimization of the preheating process, the following welding parameters were selected for further research:

• Welding current of I = 3.9 kA;
• Welding time of T = 36 periods (1 per. = 0.02 s);
• Pressing force of electrodes of Fz = 6 kN.

Preheating with the higher values already led to the formation of a welding lens, which was undesirable with respect to the subsequent clinching joining. The weld nugget prevented formation at the bottom of the clinched joint. After the local preheating using spot welding electrodes, the samples were subsequently joined at the preheating area by the clinching method (Figure 3b). A punch with a diameter of Ø5 mm and a die with a diameter of Ø8 mm were used. The force applied to the die was set to 80 kN.

A flow chart illustrating the research procedure applied in this study is shown in Figure 4.

Local preheated samples prepared by resistance spot welding and joined by clinching with local preheating are shown in Figure 5.

2.4. Tensile Test

Tensile–shear tests were performed for single-lap joint specimens in accordance with the guidelines from ISO 12996: 2013 [55]. Test samples measuring 40 × 90 mm were used to evaluate the load-carrying capacity of the clinched joints (see Figure 6). The TiraTest 2300 universal testing machine (TIRA GmbH, Schalkau, Germany) was used for the tensile tests.

After the tensile test, the method of failure of the clinched joint was also evaluated. All of the failure modes of clinched joints are illustrated in Figure 7. The typical failure modes of clinched joints, according to standard ISO 12996:2013, are: pull-out (Figure 7a), neck fracture (Figure 7b), neck fracture with plastic deformation—mixed mode failure (Figure 7c), and pull-out with neck fracture—mixed failure (Figure 7d).

The two most frequently occurring failure modes of clinched joints are [56] the pull-out mode and the neck fracture mode.

The pull-out mode is distinct from the neck fracture mode as it involves the complete separation of the upper and lower sheets. This mode is characterized by a lack of sufficient geometric interlocking in the mechanical clinched joint. When subjected to tensile–shear loading, the two sheets experience plastic deformation and eventually slide apart when the strength of the interlock becomes weaker than the applied tensile load.

The neck fracture mode is characterized by a fracture occurring in the thinnest part of the upper sheet’s neck. When subjected to tensile–shear loading, a force is applied to the neck of the upper sheet in a shearing motion. As the load increases, the shearing force on the thinnest part of the neck gradually increases. Eventually, when the neck becomes thinner, a ring fracture occurs once the shearing force reaches the strength of the thinnest neck.

2.5. Light and Scanning Electron Microscopy

The experimental materials and the fabricated clinched joints were sampled by electrospark machining for light microscope and scanning electron microscope analyses. Subsequently, the samples were prepared in conductive dentacryl, ground on 240, 400, 600, and 800 grit sandpaper moistened with water, polished with diamond paste 1µm on satin moistened with kerosene, washed and rinsed with alcohol. Samples were cleaned in methanol in ultrasonication before observation. The microstructure was developed by etching in Nital (2% nitric acid solution in ethyl alcohol).

An experimental light microscopy technique was used for the analyses, utilizing the Olympus CX71 inverted metallographic microscope with the Olympus DP12 camera (OLYMPUS Europa Holding GmbH, Hamburg, Germany). For the Olympus CX71 microscope, the accessories for observation in polarized light and by differential interference contrast were used. For detailed analyses of the microstructures, the JEOL JSM-7000F scanning electron microscope (SEM) (Jeol Ltd., Tokyo, Japan) with the INCA X-sight model 7557 energy-dispersive X-ray (EDX) analyzer (Oxford Instruments, Abingdon, Oxfordshire, UK) and the EVO MA15 environmental scanning electron microscope (Carl Zeiss Microscopy GmbH, Jena, Germany) with the X-Max 50 model 51-XMX1003 EDX analyzer (Oxford Instruments, Abingdon, Oxfordshire, UK). An accelerating voltage of 20 kV and a sample-to-surface distance of 10 mm was used for analyses in the secondary electron (SEI) mode and in the backscattered electron (BSE) mode for observations in chemical contrast.

2.6. Microhardness Measurement

The WILSON-WOLPERT Tukon 1102 microhardness tester (Buehler ITW Co., Lake Bluff, IL, USA) was used to measure the microhardness HV0.3 in the gradient microstructure zones and the clinching junction. The distance between the measurements was 0.15 mm. The individual hardness measurement zones for the welded joint and the clinched joint with the gradient microstructure are shown in Figure 8.

3. Results and Discussion

3.1. Material Flow in Clinched Joint

The use of numerical simulation for predicting the material flow of the studied steel sheets in the process of clinching is shown in Figure 9. The results of the numerical simulation were transformed into graphical representations using vectors. These graphical representations show the downward movement of the material towards the die cavity in the early phases of the joint-forming process, specifically during clamping and pressing. Once the material reached the die, it began to flow sideways into the die groove. The formation of the clinched joint concluded at the clinching stage, where both joined materials completely fulfilled the die groove, resulting in the formation of an interlock in the neck region of the joint.

From the point of view of increasing the load-bearing capacity of clinched joints, it was important that the structure of the preheated material reached the place of the neck of the joint during the clinching process.

3.2. Microstructure of Joined Materials

The microstructure of steel No. 1.0935 (HX420LAD+Z), h = 1.5 mm, in the rolling direction and perpendicular to the rolling direction was ferritic–perlitic, consisting of fine-grained ferrite and isolated islands of perlitic colonies. The average ferritic grain size in both directions was 4.6 ± 0.2 µm. The topography of the hot-dip galvanized coating was stochastic, formed by shallow depressions. Local qualitative EDX microanalysis detected zinc, aluminum, and oxygen lines on the surface (Figure 10). The thickness of the annealed zinc coating was 8 µm.

The microstructure of steel No. 1.0941 (HCT 600X+Z), h = 1.5 mm, in the rolling direction and perpendicular to the rolling direction was ferritic–martensitic. The average ferritic grain size in both directions was 7.2 ± 0.5 µm, and the area fraction of martensite was 13 ± 1.4%. The surface morphology of the hot-dip galvanized steel sheets was stochastic, formed by shallow depressions in the zinc coating of irregular shape. The thickness of the hot-dip galvanized coating on both surfaces was 10 μm. Local qualitative EDX microanalysis detected zinc, aluminum, and oxygen lines on the surface, as shown in Figure 11.

3.3. Clinched Joints

On the experimentally prepared clinched joints of hot-dip galvanized steel sheets of grade No. 1.0941 and grade No. 1.0935, there was an intense thinning of the thickness of the sheets in the area of the neck (Figure 12a,b).

The information about the intensity of hardening in the area of the clinching joint at the neck was obtained by measuring the microhardness HV0.3 at the distance between the centers of the indenters of 0.15 mm. The microhardness measurement was in the direction from the upper sheet (US) to the lower sheet (LS).

In the clinched joint of the sheets of material grade No. 1.0941 with the original ferritic–martensitic microstructure, the microhardness value in the measured directions was at the level of about 300 HV0.3, locally in the neck at the thinning point of the upper sheet (US-1.0941) the microhardness was 350 HV0.3 (Figure 13a,b). In the clinched joints of the No. 1.0935 grade sheets in the measurement directions, the microhardness HV0.3 was at the level of 250 HV0.3, locally in the neck at the thinning point of the top sheet (US-1.0935), the microhardness was 300 HV0.3 (Figure 13c,d).

3.4. Local Preheating of the Clinched Joint Zone

Local intense heating was applied to a pair of overlapping hot-dip galvanized sheets, each 1.5 mm thick. The heating was at the location of the future clinched joint. Tests were carried out to locally modify (change) the ferritic–martensitic microstructure to a ferritic–carbidic microstructure with a sufficient value of uniform plastic deformation suitable for the implementation of clinching. In the zones where the joints were subsequently made by the clinching method, a gradient microstructure was obtained by local intensive heating of both joining sheets, copying the heat flow from the heating zone to the surrounding material after the local heating phase was completed. Throughout the local heating zone, the ferritic–martensitic and ferritic–perlitic microstructures were transformed from the different temperature values reached in the respective heating zone due to the intense heat flux from the heating zone into other types of structure.

3.4.1. Zones with Gradient Microstructure Prior to Clinching, Steel No. 1.0941

Local heating (RSW) was applied to two overlapping galvanized steel sheets of grade 1.0941 in order to transform the microstructure from the original ferritic–martensitic microstructure to a gradient fine-grained ferritic–sorbitic microstructure. Different states of microstructure were observed on the upper sheet US-1.0941 (upper sheet = US) and lower sheet LS-1.0941 (lower sheet = LS) sheets at the Cu-Cr electrode contact sites and in their surroundings (Figure 14a).

In zone P1, there was a ferritic–martensitic microstructure (Figure 15a), which continuously transitioned to the original ferritic–martensitic microstructure of sheet grade 1.0941 and to a fine martensitic microstructure in the local heating zone (Figure 14b). The microstructure in Figure 14b was formed after rapid cooling. This area was heated to a temperature in the interval Ac1–Ac3, and a two-phase ferritic–martensitic structure was formed by rapid cooling.

The microstructure of zone P2 was transformed from a fine-grained ferritic–martensitic microstructure to austenite without an increase in the austenitic grain size as a result of local rapid intense heating to temperatures above Ac3 of material 1.0941. By the same intense heat dissipation to the surrounding of the heating zone, the austenite was transformed into a fine-martensitic microstructure (Figure 16). A martensitic microstructure was formed in the central region in zone C, characterized as a lens during the spot welding (Figure 17). There was intense heat dissipation into the electrode body at the contact points of the cooled Cu-Cr electrodes and the overlapping sheets. In zone E, there was then a region to a depth of about 0.3 mm in which no microstructure transformation occurred. This region was not heated above the temperature of Ac1 (Figure 18).

The microhardness of each zone with a gradient microstructure was measured in the region of the original base material (BM), zone P1, with a ferritic–martensitic microstructure, zone P2, with a fine martensitic microstructure, and the central region, which included zone C, with a martensitic microstructure, and zone E, in which no microstructural transformation occurred. The microhardness was measured on the upper sheet (Figure 19a) and the lower sheet (Figure 19b).

The microhardness values in the P1 zone were slightly increased with respect to the microhardness values of the original ferritic–martensitic microstructure. The microhardness values of the fine-grained sorbitic microstructure of the P2 zone were twice the microhardness values of the original ferritic–martensitic microstructure. In the central zone, the martensitic microstructure had almost twice the microhardness values of the original ferritic–martensitic microstructure. The microstructure gradient formed by the narrow ferritic–martensitic region at the point of contact of the electrodes with the sheets (zone E) caused by the intense heat dissipation into the water-cooled electrodes was manifested by the microhardness values at the level of the original matrix (Figure 20a,b).

3.4.2. Zones with Gradient Microstructure Prior to Clinching, Steel No. 1.0935

The original ferritic–perlitic microstructure of two overlapping galvanized steel sheets of grade 1.0935 was transformed by intensive local heating into a gradient fine-grained ferritic–sorbitic microstructure. Different states of microstructure were observed on the upper sheet US-1.0935 (upper sheet = US) and the lower sheet LS-1.0935 (lower sheet = LS) sheets at the Cu-Cr electrode contact points and in their surroundings (Figure 21).

In zone P1, there was a ferritic–martensitic microstructure (Figure 20), which continuously transitioned to the original ferritic–perlitic microstructure of sheet grade 1.0935 and to the fine sorbitic microstructure of zone P2 (Figure 22).

The microstructure of zone P2 was transformed from a fine-grained ferritic–perlitic microstructure to austenite without an increase in the austenitic grain size as a result of local rapid intense heating to temperatures above Ac3 of material 1.0935. By the same intensive heat dissipation to the surroundings of the heating zone, the austenite was transformed into a martensitic microstructure (Figure 23).

In the central region in zone C, there was a martensitic microstructure (Figure 24). There was intense heat dissipation into the electrode body at the contact points of the cooled Cu-Cr electrodes and the overlapping sheets. In zone E, there was then a region to a depth of about 0.3 mm in which no microstructure transformation occurred. This region was not heated above the temperature of Ac1 (Figure 25).

The microhardness of each zone with a gradient microstructure was measured in the region of the original base material (BM), zone P1, with a ferritic–martensitic microstructure, zone P2, with a fine martensitic microstructure, and the central region, which included zone C, with a martensitic microstructure, and zone E, in which no microstructural transformation occurred. The microhardness was measured on the upper sheet (Figure 26a) and the lower sheet (Figure 26b).

The microhardness values in the P1 zone were slightly increased with respect to the microhardness values of the original ferritic–pearlitic microstructure (BM).

The microhardness values of the fine-grained sorbitic microstructure of zone P2 were at the level of 250 to 300 HV0.3. In the central zone, there was a martensitic microstructure with almost twice the microhardness values compared to the original ferritic–perlitic microstructure. Analogous to the material of grade No. 1.0941, a microstructure gradient was observed, formed by a narrow ferritic–perlitic region at the point of contact of the electrodes with the sheets in zone E (Figure 27).

3.5. Analysis of Clinched Joints with Local Preheating

After applying preheating by resistance spot welding to the No. 1.0941 and No. 1.0935 grade sheets and making the clinched joints, the fine martensitic region, which was marked in Figure 14a and Figure 21a, was relocated to the neck region in the upper sheet and to the locking region in the lower sheet (Figure 28a and Figure 29a). Zone C, with a martensitic microstructure, was in the bottom region after the formation of the clinched joint. The narrow ferritic–martensitic and ferritic–perlitic zone E, which was formed at the point of contact of the electrodes with the sheets, was in the bottom region on the outer surfaces of the joined sheets after the formation of the clinched joint.

Figure 30 shows a comparison of the load-bearing capacity Fmax [N] of the pure clinched joints (CL) and the preheated clinched joints (RSW + CL). Five groups of specimens were tested for each of the jointed steel sheets. Within one group, seven specimens were prepared. All of the samples shown in Figure 30 were preheated by resistance spot welding, with optimized values of the welding parameters: a welding current of I = 3.9 kA, a welding time of T = 36 periods (1 per. = 0.02 s), and a pressing force of the electrodes of Fz = 6 kN. For both types of tested steel sheets, a higher load-bearing capacity was obtained for the preheated clinched joints. The load-bearing capacity of the pure clinched joints for the 1.0941 (HCT600X+Z) sheets ranged from 5398 N to 5540 N. On the other hand, the load-bearing capacity of the preheated clinched joints for these steels ranged from 6227 N to 6525 N. The measured load-bearing capacity values for the preheated clinched joints of the 1.0941 steels were higher by an average of 17%. Regarding the joining of steel sheets 1.0935 (HX420LAD+Z), the measured load-bearing capacity values of the pure clinched joints ranged from 3712 N to 3831 N. For the preheated clinched joints for these steels, the load-bearing capacity values ranged from 4224 N to 5200 N. On average, the load-bearing capacity of the preheated clinched joints for the 1.0935 steel was higher by 25%.

The increase in the load-bearing capacity of the preheated clinched joints is related to the change in the microstructure of the observed parts of the joint (P1, P2, center), which was proven by measuring the microhardness (Figure 13, Figure 28 and Figure 29). The microhardness values in the observed parts of the joint were higher in the preheated clinched joints in both types of steel sheets.

Figure 31 shows the failure mode of the clinched joints after the tensile test. It is obvious that, with the clinching method, the mode of failure “crack in the neck with plastic deformation” occurred in both examined steel sheets. This is typical when the strength of the interlock is slightly greater than that of the clinched neck (as was described by Lei et al. [56]). This mode is also known as the hybrid neck fracture mode. After the application of local heating using resistance spot welding, the failure mode of the clinching joints was changed to a “neck fracture”. An increase in the hardness of the neck zone of the preheated clinched joints resulted in an increase in the strength of the material, which affected the joint failure mode.

4. Conclusions

The aim of this research was to find out how to increase the load-carrying capacity of clinched joints. Based on experimental clinching with local preheating achieved by resistance spot welding, the following conclusions can be stated:

• Trial tests of clinched joints with a locally modified microstructure from ferritic–martensitic to fine-grained ferritic–sorbitic and fine-grained sorbitic–martensitic microstructures and increased strength values were carried out in order to localize them in particular regions of the compression joint.
• The microstructure distribution of the clinched joint region was analyzed by light and scanning electron microscopy and microhardness measurement techniques in the clinched joint.
• In the entire clinched joint, no discontinuities were observed in the material of the clinched joint by light and scanning electron microscopy.
• Fine-grained ferritic–sorbitic and fine-grained sorbitic microstructures were observed in the clinched joint at the neck when local heating technology was used.
• Two grades of materials were tested with a chemical concept corresponding to micro-alloyed steels and two-phase ferritic–martensitic steels. For each grade, five groups of seven samples were tested under the same experimental conditions of local heating (RSW) and clinching. The application of RSW to the clinched joint area increased the load-carrying capacity by an average of 17% for material No. 1.0941, and by an average of 25% for material No. 1.0935.

We consider the limitation of this study to be the joining of hot-dip galvanized steel sheets with greater thickness than those investigated. It is necessary to take into account higher values of the welding parameters, which can negatively affect the zinc layer on the surface. At elevated temperatures, degradation of the coating may occur, and thus, may significantly affect the corrosion resistance of the joints. Another significant limit was joining a combination of ferrous and non-ferrous metals, such as joining steel sheets and aluminum alloy sheets. Due to the different melting temperatures of steel and aluminum, it is very difficult to achieve a local change in the microstructure.

Further investigation into clinching joining with a local microstructure modification could be directed to the area of joining different combinations of materials, with different thicknesses and different positioning of the materials in relation to the punch and die.