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Original Article
DOI: 10.1016/j.jmrt.2019.01.010
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Available online 14 March 2019
Efficiency of dissimilar friction welded 1045 medium carbon steel and 316L austenitic stainless steel joints
Gawhar Ibraheem Khidhir
Corresponding author

Corresponding author.
, Sherko A. Baban
Department of Mechanical and Mechatronics Engineering, Salahaddin University, Erbil, Iraq
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Figures (9)
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Tables (4)
Table 1. Chemical compositions of base materials (wt.%).
Table 2. Mechanical properties of base materials.
Table 3. Process parameters used.
Table 4. Tensile test results for welded specimen.
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This study investigated the effects of frictional welding parameters on the microstructure and mechanical properties of dissimilar steel materials, namely, AISI 1045 medium carbon steel and AISI 316L austenitic stainless steel. The welded joints were produced by changing the forging pressure while the friction pressure, friction time, forging time and rotational speed were kept constant to achieve a constant range of temperature (780–800°C). Experimental results showed that when the forging pressure increases, the hardness value of the weld interface increases whereas the tensile strength decreases. The hardness profiles also indicated that the welds exhibited higher hardness numbers than the two base metals. The highest weld joint efficiency obtained was 90% while the lowest was 63%. The joints failed in the thermo mechanical affected zone on the 316L austenite stainless steel side. Scanning electron microscopy attached with energy dispersive spectroscopy was used to analyse the fracture surface in the tensile test.

Friction welding
AISI 1045
Tensile strength
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Iron-based alloys, including stainless steels with a significant chromium content of 12–30% chromium and 8–25% nickel, display characteristic resistance to both corrosion and high temperature [1]. The largest stainless-steel group produced and the most preferable are the austenitic stainless steel (ASS), due to their excellent mechanical properties and corrosion resistance [2,3]. ASSs have found a wide range of applications. For instance, they are extensively used in applications in which they are subjected to high temperatures, such as boilers, heat exchangers, and nuclear facilities [4]. The only way to increase the traction resistance of ASSs is through hardening by cold plastic deformation. In general, austenitic alloys are considered excessively weldable materials [1]. The 316L ASS is a chromium–nickel–molybdenum alloy. As a modified form of AISI 316, it contains minimal carbon (about 0.03%) and has limited sensitivity to carbide precipitation. The addition of molybdenum (approximately 2–3%) improves pitting corrosion resistance [5]. Friction welding (FW) is a solid-state welding process, in which the relative rotation produces friction between the two parts and heat is released at the contact surface. After reaching the welding temperature, the rotation seizes and an upsetting force is exerted on the plastically deformed material, leaving the joint to cool and consolidate. FW is considered a kind of forge welding technique because no melting occurs, and welding is done by the application of pressure [6]. Input parameters that control the joint include friction pressure, friction time, forging pressure, forging time and rotational speed. The interesting advantages of FW include substantial material saving, shorter production time and the capability of welding different metals or alloys; moreover, FW enables the joining of different configurations of circular or non-circular cross sections [1]. Nowadays, manufacturers and researchers focus on dissimilar metal welding, which moderately improves the performance of the welded parts and minimises the material costs [7]. Joining stainless steel to carbon steel through dissimilar joining is being used in the petrochemical and power generation industries. Significant challenges occur when joining materials of different chemical, mechanical, thermal or electrical properties. The incompatibility of these properties can create problems both the joining process itself and produces different residual stresses in the weld zone relative to similar metal welding [8]. Investigating the effect of the parameters of FW between AISI 1045 medium carbon steel (MCS) and AISI 316L austenite stainless steel has received minimal attention. However, some studies performed on 316L austenite stainless steel have been welded by linear FW, electron beam welding, friction stir welding and laser welding [9–12].

This study seeks to investigate the effect of forging pressure on the efficiency of dissimilar friction welded 1045 MCS and AISI 316L ASS joints at a constant range of welding temperature. This type of FW can be used in applications that require great strength and efficiency.

2Experimental procedures2.1Materials

The workpieces used in this study include cylindrical rods (16mm diameter and 50mm length) of AISI 1045 MCS and AISI 316L ASS. The chemical compositions and mechanical properties of the materials used are listed in Tables 1 and 2, respectively.

Table 1.

Chemical compositions of base materials (wt.%).

Elements  Si  Mn  Cr  Mo  Ni  Fe 
316L(ASS)  0.02  0.37  1.28  0.031  0.001  16.42  2.05  10.57  Bal 
1045(MCS)  0.497  0.262  0.735  0.0124  0.0249  0.0428  0.0044  0.0156  Bal 
Table 2.

Mechanical properties of base materials.

Material  Hardness (HV)  Yield strength (MPa)  Ultimate tensile strength (MPa)  %Elongation L0=25mm  %Reduction in area 
316L(ASS)  180  255  570  54  70 
1045(MCS)  225  450  750  15  62 
2.2Welding process

Continuous drive FW between 1045 MCS and 316L ASS was performed using a lathe machine type (TSL, SER.No.NN94298). Different ranges of rotation were present between 30–1600rpm. The part of 1045 MCS was fixed on the rotating chuck and the 316L ASS was on the stationary clamp on the tail stock. A load cell was arranged on the side of the tail stock to read the friction and forging pressures during the welding process. Then, a stop watch was adjusted to obtain the friction and forging times. During the application of friction pressure, the temperature of the weld zone was measured by using a thermocouple device K-type (ProsKiteMT-1280), that can read temperatures between −20 and 1000°C. Once the system reached a certain temperature, a brake instantaneously stopped the rotation, and a specified forging pressure was applied which compresses the samples together. In this research, the following parameters were held constant at the corresponding specified values to obtain a constant temperature range of 780 to 800°C: friction pressure at 75MPa, friction time at 38s, forging time at 2s and rotational speed at 560rpm. Table 3 shows the forging pressure as the only varying parameter, and Fig. 1 presents the welded samples 1045 MCS and 316L ASS. The goal was to adjust the friction pressure and friction time to reach a constant range of temperature and determine the effect of forging pressure on the microstructure of the specimens and their mechanical properties.

Table 3.

Process parameters used.

Sample No.  Forging pressure (MPa) 
Sample (1)  75 
Sample (2)  95 
Sample (3)  115 
Sample (4)  135 
Sample (5)  155 
Fig. 1.

Welded samples.

2.3Microstructure analysis

Microstructure analysis of the welded joints of 1045 MCS and 316L ASS were conducted by using a computerised METKAN IMM901 microscope and a scanning electron microscopy (SEM) model (TESCAN, MIRA 3) equipped with an energy dispersive spectroscopy (EDS) for analysing the chemical composition of the weld fracture after the tensile test. The specimens were prepared for microstructure analyses according to ASTM E3-11. Etching for 1045 MCS was done by Nital, which is 98% water and 2% HNO3, and the etching time was fifteen seconds. Aqua regia was used to reveal the microstructures of the 316L ASS specimen [13].

2.4Mechanical tests

The mechanical properties were determined through the hardness and tensile tests. The hardness profile was taken across the weld centreline, according to the ASTM E92-03 standard test [14]. Hardness measurement was carried out with a load of 50kg for 15s by a testing machine AKASHI type model AVA. Tensile test samples were prepared according to the ASTM A370 standard, with a gage length (G) of 25mm and diameter (D) of 6.25mm. The weld interface was located in the centre of the gage length (Fig. 2). The samples were tested in tension by using Hualong type WAW 600. Joint efficiency (η) was calculated in terms of the tensile strength of the welded metal divided by the tensile strength of the softer metal as shown in Eq. (1).

Fig. 2.

(a) and (b) The geometry and the welded specimen prepared for tensile test.

3Results and discussion3.1Visual examinations

The visual test confirmed that no significant defects or surface discontinuities were present in the welded joints, which indicates satisfactory metallurgical joint in the weld zone. Moreover, the flash of the 1045 MCS side is greater than that of the 316L ASS side for all cases as in Fig. 3(a). This result is due to the presence of alloying elements, especially Cr in stainless steel, and the higher hardness of 316L ASS than 1045 MCS at higher temperatures. These results are identical to those obtained by [15].

Fig. 3.

(a) Macro structure of the welded Sample (1), (b) microstructure of base 1045 MCS, (c) microstructure of base 316L ASS.

3.1.1Macrostructure and microstructure

Macro analysis shows different zones in Fig. 3(a), such as the weld interface (WI), thermo-mechanically affected zone (TMAZ) and heat affected zone (HAZ) on both sides. The widths of TMAZ and HAZ of 1045 MCS are thicker than those in 316L ASS due to the higher thermal conductivity of the former. These characteristic zones are discussed in detail in the microstructure field. The microstructure of 1045 MCS base metal is shown in Fig. 3(b) which consists of ferrite (light) and perlite (dark) areas. The ferrites are uniformly distributed along the pearlite grain boundaries and have grain size of approximately 28μm. By contrast, Fig. 3(c) shows the microstructure of 316L ASS base metal with grain size of nearly 40μm. The chromium and nickel equivalents (Creq, Nieq) were determined using Eqs. (2) and (3) according to the Schaeffler diagram [16]. For 316L ASS base metal, Creq=19.05 and Nieq=11.8; thus, the phase structure of 316L ASS consists of austenite with approximately 8% ferrite:

The microstructures from the welded joint to the base metal 1045 MCS for Sample (1) are shown in Fig. 4(a)–(c). The microstructure of the weld interface (WI) is obvious, as it is the line between 1045 MCS and 316L ASS. Close to the weld interface (WI), a thin proeutectoid ferrite layer appeared due to element diffusion, especially of concentrated carbon. In TMAZ, rapid cooling after high temperature led to the formation of coarse pearlite grains and less ferrite precipitates along the pearlite grain boundary as shown in Fig. 4(a) and (b). The microstructure in HAZ shows refined perlite grains with higher amount of ferrite content due to a lower process temperature compared with the TMAZ, as shown in Fig. 4(c). The same phenomena are shown in Fig. 4(d)–(f), except for Sample (5), whose WI, TMAZ and HAZ showed finer structure those for Sample (1) because of severe plastic deformation. Additionally, Fig. 4(d) indicates that due to the overheating of TMAZ, some ferrites exist in the needle form. Such structure is called acicular ferrite. Acicular ferrite nucleates were found on prior-austenite grain boundaries. Fig. 5(a) reveals the grain refinement occurring at WI for the 316L ASS side of Sample (1) from the dynamic recrystallisation of austenite grains, which results in improvement of the hardness properties. The decarburised layer consists of a thin ferrite layer emerging along the weld interface towards the 316L ASS side because of carbon diffusion from the 1045 MCS to the 316L ASS. Fig. 5(b) shows that the deformed austenite grains and twins appear in TMAZ, while the austenite grains become coarser and appear to have no deformation in HAZ. Sample (5) reveals more refined grains in the WI and TMAZ due to more severe plastic deformation. High temperature phase delta ferrite (δ) does not appear because of the low welding temperature, which is approximately 800°C. In Fig. 5(d) and (e) chromium carbides appear given the migration of carbon, which may be responsible for higher hardness and lower tensile strength and may lead to the fracture at the TMAZ of 316L ASS. This type of diffusion is also discussed by Ma H [17].

Fig. 4.

Microstructure on the 1045 MCS side, Sample (1): (a) WI and TMAZ, (b) TMAZ and (c) HAZ. Sample (5): (d) WI and TMAZ, (e) TMAZ and (f) HAZ.

Fig. 5.

Microstructure of different regions on 316L (ASS) side, Sample (1): (a) WI and TMAZ; (b) TMAZ and (c) HAZ. Sample (5): (d) WI and TMAZ (e) TMAZ and (f) HAZ.

3.2Mechanical properties3.2.1Hardness tests

Micro-hardness was measured on both sides of the interface at a distance of 1mm apart, as depicted in Fig. 6. All samples showed similar hardness profile and the weld interface of all samples exhibit higher hardness values than TMAZ, HAZ and the base metal of both sides, which may be attributed to brittle intermetallic formation. In addition, as forging pressure increased, the temperature remained constant, but plastic deformation increased and the hardness value in WI increased from 251 HV for 75MPa to 316 HV for 155MPa. This difference may be due to chemical composition and chromium content [2]. Moreover, a decreasing trend from weld interface on either side was observed. In the 316L ASS side, the hardness value decreased sharply in TMAZ because of coarse austenite grains and then increased slightly as one moved along the HAZ to the parent material. Conversely, in the 1045 MCS side, the hardness value decreased regularly in the HAZ as a result of the increasing content of ferrite until the hardness value of the parent material was reached.

Fig. 6.

Vickers hardness distributions for welded joint.

3.2.2Tensile properties

Table 4 lists the results of the tensile tests of welded samples, including ultimate tensile strength, elongation % and efficiency %. From Fig. 7(a) and (b), note that the ultimate tensile strength value of the welded samples is less than that of the parent material. Furthermore, the maximum tensile strength of 513MPa was achieved for Sample (1), the minimum tensile strength of 361MPa is obtained for Sample (5), and the efficiency decreased from 90% to 63%. Therefore, when the forging pressure is 75MPa, the two metals are strongly welded because the plastic deformation is sufficient and temperature range of 780–800°C is optimal for this welding process; moreover, at low temperatures, twins leads to high flow stresses and increases dynamic recrystallization, and this result is in good agreement with the equation of the Zener–Holloman parameter (Z) discussed by H. Mirzadeh in [18]. A very thin and discontinuous carbide layer was also obtained. In the intervening period, weld defects such as cracks and inclusion do not appear in the weld interface. Conversely, when the forging pressure was increased to 155MPa, plastic deformation increased and more diffusion occurred, leading to formation of thicker carbide layers in the TMAZ. For this reason, all welded joints failed close to the weld interface or in the TMAZ of the 316L ASS side, Fig. 8(a) and (b).

Table 4.

Tensile test results for welded specimen.

No. of sample  Ultimate tensile strength (MPa)  Elongation %  Efficiency of joints % 
S1  513  21  90 
S2  473.1  19  83 
S3  440  18  77 
S4  397  15  69 
S5  361  11  63 
Fig. 7.

(a) Ultimate tensile strength, (b) elongation % and efficiency of joints %.

Fig. 8.

Fracture appearance and SEM morphology after the tensile test: (a) for Sample (1), (b) for Sample (5), (c) SEM for Sample (1) and (d) SEM for Sample (5).


Concerning the elongation in Fig. 7(b), the elongation decreased from 21% for Sample (1) to 11% for Sample (5), and they were lower than the elongation of the 316L base metal. These decrements were caused by brittle intermetallic phases and severe plastic deformation. Forging pressure increment exhibits a fine microstructure, but also leads to more alloy elements extruded out, which resulted in descending tensile strength. Moreover, the microstructure at the TMAZ of the 316L ASS side is heterogeneous, which caused a quasi-cleavage fracture for Sample (5). In sum, when the temperature range is fixed, as in our case, exaggerated forging pressure is unnecessary because it degrades the resistance of the weaker side.

Sample (1) exhibits a ductile mode fracture, shown in Fig. 8(c). This mode consists of a spherical dimples region, which creates micro voids that initiate crack formation. Coalescence of these micro voids leads to crack formation. For Sample (5), Fig. 8(d) shows quasi-cleavage mode, mean dimple with flat surface expounding the brittle fracture. This feature results from the greater chromium carbide formation on the grain boundary due to the severe diffusion process. This result agrees well with the tensile results discussed previously. The diffusion is further supported by EDS results, in Fig. 9(a) and (b), and in Table (5), which represent the quantitative results for the major element content in wt. % on the stainless steel side for Samples (1) and (5). Sample (1) involves low forging pressure and contains low carbon, chromium, and nickel content. When the forging pressure increased, intensive plastic deformation and greater element diffusion occurred for Sample (5), in which the amount of C, Cr, Ni increased, thereby leading to lower tensile strength. This result coincides with the same feature for welding 304 ASS with 1045 MCS in [17].

Fig. 9.

EDS analysis after the tensile test for the 316L ASS side: (a) for Sample (1) and (b) for Sample (5).


From the EDS result, the chromium and nickel equivalents were determined using Eqs. (2) and (3), respectively, which proved that the fracture occurred on the 316L ASS side. An increase in the C and Cr content was found to decrease the flow stress due to formation of chromium carbide and according to Zener–Holloman parameter the apparent activation energy for self-diffusion decreased as in Sample (5). While in Sample 1, lesser amounts of C and Cr leading to higher flow stress, thus activation energy increased. This result coincides with those obtained by Menapace C [19], who stated that activation energy decreased as a result of increasing the Fe self-diffusion due to high C content. While, almost all other alloying elements like Mn, Si, Mo, Ti and Nb reduced Fe self-diffusivity, thus increased the activation energy.


The effect of forging pressure on the microstructure and mechanical properties of dissimilar metal joint was evaluated and the conclusions are summarised as follows:

  • 1.

    FW is successfully carried out to joint 316L ASS and 1045 MCS in a constant range of welding temperatures (780–800°C), with a maximum efficiency of 90% was achieved at the minimum forging force of 75MPa. Thus, exaggerated forging pressure is unnecessary.

  • 2.

    The microstructure reveals that the widths of the TMAZ and HAZ of 1045 MCS were larger than those of 316L ASS.

  • 3.

    As the forging pressure increased, the hardness increased at the interface of the weld joint.

  • 4.

    The SEM tests show a ductile fracture appearance in Sample (1) and quasi-cleavage fracture in Sample (5). This result was confirmed by the EDS test.

  • 5.

    Forging pressure has a positive effect on the ultimate tensile strength of the weld joint because fracture occurs in the TMAZ of the 316L ASS side.

Conflicts of interest

There is no conflict of interest to declare.


This study is funded and sponsored by Salahaddin University-Erbil.


The authors are grateful to the Mechanical and Mechatronics Department in the College of Engineering at Salahadin University for facilitating this research. The authors would like to thank Mr. Sardar A. for his help in performing the SEM and EDS test.

I. Kirik, N. Özdemir.
Effect of process parameters on the microstructure and mechanical properties of friction-welded joints of AISI 1040/AISI 304L steels.
Mater Technol, 49 (2015), pp. 825-832
A. Moteshakker, I. Danaee, S. Moeinifar, A. Ashrafi.
Hardness and tensile properties of dissimilar welds joints between SAF 2205 and AISI 316L.
Sci Technol Weld Join, 21 (2016), pp. 1-10
S. Shashi Kumar, N. Murugan, K.K. Ramachandran.
Influence of tool material on mechanical and microstructural properties of friction stir welded 316L austenitic stainless steel butt joints.
Int J Refract Met Hard Mater, 58 (2016), pp. 196-205
S.S. Rezaei-Nejad, A. Abdollah-zadeh, M. Hajian, F. Kargar, R. Seraj.
Formation of nanostructure in AISI 316L austenitic stainless steel by friction stir processing.
Procedia Mater Sci, 11 (2015), pp. 397-402
N. Ghosh, P.K. Pal, G. Nandi.
Parametric optimization of MIG welding on 316L austenitic stainless steel by grey-based taguchi method.
Procedia Technol, 25 (2016), pp. 1038-1048
V. Akhil, A.P. Nellore.
An investigation of mechanical and metallurgical properties of friction welded steel joint.
Int J Innov Sci Res Technol, 2 (2017), pp. 265-271
J. Verma, R.V. Taiwade, R.K. Khatirkar, A. Kumar.
A comparative study on the effect of electrode on microstructure and mechanical properties of dissimilar welds of 2205 austeno-ferritic and 316L austenitic stainless steel.
Mater Trans, 57 (2016), pp. 494-500
H. Eisazadeh, J. Bunn, A. Achuthan, J. Goldak, D.K. Aidun.
A residual stress study in similar and dissimilar welds.
Weld J Res Suppl, 95 (2016),
111–s to 119–s
I. Bhamji, M. Preuss, P.L. Threadgill, R.J. Moat, A.C. Addison, M.J. Peel.
Linear friction welding of AISI 316L stainless steel.
Mater Sci Eng A, 528 (2010), pp. 680-690
B. Joseph, D. Katherasan, P. Sathiya, C.V. Srinivasa Murthy.
Weld metal characterization of 316L(N) austenitic stainless steel by electron beam welding process.
Multicr Int J Eng Sci Technol, 4 (2012), pp. 169-176
A. Yazdipour, A. Heidarzadeh.
Effect of friction stir welding on microstructure and mechanical properties of dissimilar Al 5083-H321 and 316L stainless steel alloy joints.
J Alloys Compd, 680 (2016), pp. 595-603
K. Ks.
Investigation on transient thermal responses on 316L austenitic stainless steel and low carbon ferritic steel welding using pulsed Nd:YAG laser.
G.F. Voort Vander.
ASM International, (2004),
D.M. Howard Kuhn.
ASM International, (2000),
H. Ma, G. Qin, P. Geng, F. Li, B. Fu, X. Meng.
Microstructure characterization and properties of carbon steel to stainless steel dissimilar metal joint made by friction welding.
Mater Des, 86 (2015), pp. 587-597
H. Ma, G. Qin, P. Geng, F. Li, X. Meng, B. Fu.
Effect of post-weld heat treatment on friction welded joint of carbon steel to stainless steel.
J Mater Process Technol, 227 (2016), pp. 24-33
M. H Mirzadeh, M.H. Roostaei, R.M. Parsa.
Rate controlling mechanisms during hot deformation of Mg–3Gd–1Zn magnesium alloy: dislocation glide and climb, dynamic recrystallization, and mechanical twinning.
C. Menapace, N. Sartori, M. Pellizzari, G. Straffelini.
Hot deformation behavior of four steels: a comparative study.
J Eng Mater Technol, (2017),
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