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Vol. 8. Issue 3.
Pages 2481-3388 (May - June 2019)
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Vol. 8. Issue 3.
Pages 2481-3388 (May - June 2019)
Original Article
DOI: 10.1016/j.jmrt.2019.02.011
Open Access
Residual stress characterization in friction stir welds of alloy 625
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Guilherme Vieira Braga Lemosa,b, Alexandre Bellegard Farinac, Rafael Menezes Nunesb, Pedro Henrique Costa Pereira da Cunhad,
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pedrohcunha@hotmail.com

Corresponding author.
, Luciano Bergmanna, Jorge Fernandez dos Santosa, Afonso Regulyb
a Helmholtz-Zentrum Geesthacht (HZG), Institute of Materials Research, Materials Mechanics, Solid State Joining Processes (WMP), Geesthacht, Germany
b Laboratório de Metalurgia Física (LAMEF), Programa de Pós Graduação em Engenharia de Minas, Metalúrgica e de Materiais (PPGE3M), Universidade Federal do Rio Grande do Sul (UFRGS), Porto Alegre, Brazil
c Centro de Pesquisa e Desenvolvimento, Villares Metals S/A, Sumaré, São Paulo, Brazil
d Pós-Graduação em Engenharia Mecânica – PPMec, Escola de Engenharia, Universidade Federal do Rio Grande (FURG), Rio Grande, RS, Brazil
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Table 1. Chemical composition of alloy 625 base material (% by weight).
Table 2. Overview of FSW process parameters and welded depth achieved.
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Abstract

Alloy 625 (UNS N06625) welded sheets were evaluated in the present study. Friction stir welding (FSW) was performed in a range of tool rotational speed from 1200 to 200rpm, welding speed from 1.0 to 1.5mm/s and constant axial force. Residual stress states were investigated by X-ray diffraction (XRD). Besides, microstructural features were analyzed by optical microscopy (OM), and scanning electron microscopy (SEM). The FSW application was effective in the Ni-based alloy, promoting different levels of grain refinement, microstructural characteristics, and enhanced microhardness. Results also showed that distinct process parameters led to changes in the joints and distinguishable residual stress distributions. In general, as the tool rotational speed decreased, the grain refinement increased, more homogeneous microstructures and microhardness profiles, and lower residual stresses were achieved within the stirred zones.

Keywords:
Alloy 625
Ni-based alloy
Friction stir welding
X-ray diffraction
Residual stress
Microstructure
Full Text
1Introduction

New challenges of the petroleum industry include exploitation of deep-water wells and increasingly harsh environments, such as those encountered in the coasts of Brazil and West Africa [1]. Deep wells generally are very corrosive presenting higher temperatures, pressures, and low pH, thus requiring high-performance materials. In this context, for the subsea technology, rigid clad pipes have been used to transport oil to the production platform. To protect those pipelines against corrosion from salt and gases such as H2S and CO2, the use of a corrosive resistant alloy (CRA) is a solution. This corrosion resistant layer is frequently made by a Nickel-based alloy 625 (UNS N06625) [2], commonly known as Inconel 625®, which is an austenitic non-magnetic alloy primarily strengthened by a solid solution.

Alloy 625 is often used in the soft annealed (Grade I) or solution annealed (Grade II) conditions according to the standards ASTM B446 [3] and ASTM B564 [4]. When higher mechanical resistance is necessary, the soft annealed condition is usually selected. In the soft annealed condition, this alloy presents an austenitic matrix with a dispersion of residual M(C,N) type carbonitrides rich in Nb and Ti, and M6C and M23C6 types carbides both rich in Mo and Cr (at grain boundaries). Although this alloy is considered fully austenitic without precipitates, it is not entirely stable. After long time heat treatments, or depending upon the thermomechanical process route, microstructural constituents such as MC, M6C, M23C6, Laves, γ″ phase (Ni3Nb), δ phase and even some residual amounts of γ′ phase (Ni3(Al,Ti,Cr,Nb)) can be generated [2,5–8].

In the metals industry, welding as a manufacturing process is an inevitable step. Consequently, a solid-state joining process having low heat-input is a promising way for joining Ni-based alloys aiming at sound welds. One of the most popular solid-state methods is friction stir welding (FSW) [9], which has a vast potential to replace traditional fusion welding of Nickel-based alloys, notably if the tool cost and effectiveness are improved.

Different process parameters have been used in FSW of distinct alloys [10–14]. Kwon et al. [10] reported that friction stir processing produced higher grain sizes in AA 1050 when the tool rotational speed was increased from 560 to 1840rpm at constant welding speed of 155mm/min. Sato et al. [11] observed grain sizes of 5.9, 9.2, and 17.8μm in FSW of AA 6063 at tool rotational speeds of 800, 1220, 2450rpm, respectively being all of them with constant welding speed of 360mm/min. Wang et al. [12] also evaluated the effect of rotational speed on the microstructure in FSW joints of AZ31 with various rotational speeds (800, 1000, 1200, 1400 and 1600rpm). Their results indicated that the grain size enhances with the increasing rotational rate. Zhang et al. [13] presented the microstructural characteristics in friction stir welds of Ti–6Al–4V and noted that both the prior grain size and α colony size increase with increasing rotational speed. Gunter et al. [14] analyzed the grain sizes in friction stir processed 304L and found that the weld processed at 80rpm–50mm/min had the smallest average grain size (2.8μm), while the specimen produced at 250rpm–100mm/min had the largest average grain size (11.4μm). Thus, it can be observed that the changes in FSW parameters, especially the tool rotational speed, were mainly investigated for low melting point alloys.

In any manufacturing process, welding included, residual stress is an important concern. FSW introduces significant levels of residual stresses, even being a relatively fast joining process, which usually takes place below the material's melting point [15]. Excessive residual stress levels can lead to significant distortion and premature failure. For low melting point alloys, the effect of FSW on residual stresses has been analyzed several times [16–22]. However, the residual stress states in friction stir welds of high melting point alloys, such as alloy 625, were not profoundly evaluated and, therefore, are not totally understood.

This work aims at understanding the residual stresses in friction stir welds of alloy 625. Therefore, these characteristics were investigated by XRD and top surface macro and microstructural observations for different process conditions of tool rotational speed and welding speed.

2Experimental procedure

The base material used in this study was an alloy 625 with 3.2mm thickness, and its chemical composition is presented in Table 1. This is a typical clad material (internal corrosion protective layer) of API pipelines. The chemical data was given by supplier's certificate according to DIN 17744 09/2002 [23].

Table 1.

Chemical composition of alloy 625 base material (% by weight).

  Ni  Cr  Fe  Mo  Nb  Co  Mn  Al  Ti  Si 
Alloy 625  Bal.  21.7  4.7  8.6  3.38  0.03  0.09  0.13  0.18  0.18  0.015 
Standard UNS N06625  >58  20–23  <5  8–10  3.15–4.15  <1  <0.5  <0.4  <0.4  <0.5  <0.1 

In order to remove oxides and contaminants, the specimens were grinded on the top and adjoining surfaces with grid emery paper. This preparation is an essential step to improve the weld surface quality, but this procedure certainly has influenced the residual stresses [24]. Before FSW, the sheets were degreased and cleaned with ethanol. FSW was carried out on a vertical milling machine fitted with servomotors and automated control systems.

A threaded polycrystalline cubic boron nitride (pcBN) tool with W-Re binder, shoulder of 25mm in diameter, and pin length of 3mm was used. The direction of the tool rotational was counterclockwise, and then the left side of the weld bead was called retreating side (RS) while the right side was called advancing side (AS). Finally, to obtain the welded joints, the tool was tilted 1.5° forward from the vertical axis, and an argon atmosphere was used to minimize surface oxidation during FSW.

Welding Ni-based alloys is not as easy as welding traditional stainless steels due to their high strength at elevated temperatures as well as the propensity to work hardening. Therefore, in the practice of establishing the suitable FSW parameters, four samples have been welded and they were divided into two groups named as follow: high rpm (parameter I and II) and low rpm (parameter III and IV). Table 2 presents the main welding parameters. The rotational speed was changed from 1200 to 200rpm, welding speed was in the range between 1 and 1.5mm/s, and axial force was kept constant. In this context, the high rpm conditions presented flash, some top surface signs of oxidation and marks and unacceptable levels of surface roughness. For the low rpm conditions, otherwise, the welded joints displayed homogeneous top surface without much flash and an excellent appearance. Finally, the process parameters range was selected according to our workgroup experience on FSW of stainless steels used in similar engineering applications.

Table 2.

Overview of FSW process parameters and welded depth achieved.

Process parameter  Welding speed, mm/s  Axial force, kN  Rotational speed, rpm  Weld surface quality  Welded depth, mm 
High rpm welds
60  1200  Medium  2.52 
II  60  1000  Medium  2.40 
Low rpm welds
III  60  200  Excellent  1.85 
IV  1.5  60  200  Good  1.83 

Macro and microstructure observations were done after standard metallographic procedure, which included grinding with abrasive papers, polishing, and etching. The specimens were etched with Glyceregia reagent (10ml HCl, 10ml HNO3, 0.5ml of glycerine). The metallographic investigation was performed by OM and SEM. OM analysis was realized in a Zeiss AxioVision A1 microscope. SEM examinations used JEOL JXA8230 scanning electron microprobe.

Mechanical properties were analyzed by microhardness instead of tensile testing due to the lack of penetration observed in the joints. However, this non-complete welded depth was also noted in many other works [25–28] which can occur due to the high work hardening coefficients of high strength materials [29]. Hence, the microhardness profiles were measured on the top surface using Vickers indentations of 500g and dwell time of 10s.

X-ray diffraction measurements were performed by GE-Seifert-Charon-M diffractometer using an X-ray tube of Cr-Kα radiation (Fig. 1). The specimens for residual stresses investigations had dimensions of 300mm×80mm×3.2mm with the weld centerline located in the center of the 300mm dimension. Fig. 2 shows a schematic sketch detailing the measurement positions.

Fig. 1.

GE-Seifert-Charon-M diffractometer with an X-ray tube of Cr-Kα radiation.

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Fig. 2.

Schematic sketch of residual stress investigations in friction stir welds of alloy 625.

(0.12MB).

XRD experiments were executed with the starting point at the base material (on the RS), passing by the weld centerline and then finishing at base material (on the AS). The diffracted beam slit had an aperture of 2mm and a 20° linear detector GE-Meteor-1D. Residual stress evaluation used the conventional sin2ψ method with Bragg-Brentano Geometry. The slope of a linear regression through the measured data points gives the value of residual stress. The standard deviation of residual stress measurements stems from deviations of X-ray diffraction line positions at 21 different angular positions (ψ-angles) to the calculated regression line. The total measured length was 80mm. As the size of the incident X-ray beam was 2mm, the irradiated areas touched each other. Hence, the residual stress levels were extensively examined in two directions (longitudinal and transverse). During the welding, inhomogeneous plastic strains may be introduced into the workpiece [30]. In this way, the full width at half maximum (FWHM) describes second-order stress related to dislocations density distributions. Starting from the XRD results, the FWHM analyzes were done to compare the process parameters evaluated.

3Results and discussions3.1Top surface macro and microstructural features

Fig. 3 presents the alloy 625 base material microstructure. Fig. 3a shows recrystallized grains where it can be observed that the previous austenitic grains of the soft annealing process are well marked by precipitates. In Fig. 3b, it can be observed M(C,N) type carbonitrides rich in Nb and Ti dispersed in the matrix and a large amount of intergranular carbides. Fig. 3c displays the intergranular carbides precipitation indicating two types, i.e., M6C and M23C6 usually rich in Mo and Cr, respectively. The identification of these carbides was made by EDS, and their occurrence is expected to this alloy [2,5,27,28,31].

Fig. 3.

Alloy 625 base material. (a and b) OM micrographs and (c) SEM micrograph. The alloy was etched with Glyceregia.

(0.88MB).

Top surface macrographs for the four conditions are presented in Fig. 4. All the joints showed lack of penetration due to the pin being shorter than the material thickness and the wear evolution in FSW trials. FSW using high rpm promoted some debris at the top surface and an irregular aspect. Only few parts of these weld beads showed a consistent appearance (Fig. 4a and b). The degradation mechanisms of pcBN tool were analyzed [32] and their findings showed that the tool wear behavior changes due to FSW process parameters. As can be seen in Fig. 4c, the parameter III (low rpm) reached a very good top surface appearance and selected as the best weld. Also, this suitable process condition was suggested in the literature [2,24,32].

Fig. 4.

Top surface macrographs of welded joints. (a) Parameter I, (b) parameter II, (c) parameter III and (d) parameter IV.

(0.27MB).

The material flow and cross-section macrostructures changed as a function of process parameters, and these results were reported in other works [2,32]. These findings revealed that the welded depth was influenced by the welding parameters as well as the tool wear progress during FSW development to this alloy. Specially for Ni-based alloys, FSW attempts to find the optimized welding parameters and sound welds can be very challenging. In the current study, the FSW application led to the recrystallization and finer grains (Figs. 5 and 6). However, it should also be noted that these alloys can keep their mechanical strength even at high temperatures [2,32] and, therefore, they are difficult to deform as well as to become plastic. These difficulties can also be somehow influenced by their low thermal conductivity [33–35].

Fig. 5.

SEM top surface micrographs of the stir zone. (a and b) Parameter I and (c and d) parameter II.

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Fig. 6.

SEM top surface micrographs of the stir zone. (a and b) Parameter III and (c and d) parameter IV.

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Ni-based alloys tend to rapidly work hardening [33,34] which would influence any deformation process applied to them. In this context, Ni-based alloys are generally considered difficult to machine, especially at high cutting speeds [33,34]. Hence, Fig. 5 shows the friction stir welds processed by higher heat input (high rpm – tool rotational speeds of 1200 and 1000rpm, respectively) and their heterogeneous macrostructures. It was verified that these welds presented some grains not fully recrystallized in the joint cross section [2]. Therefore, it was more challenging to weld at high rpm (higher temperatures), and these conditions occasioned severe tool wear [32]. On the other hand, the use of low rpm (200rpm) promoted relatively lower heat inputs and more homogeneous macrostructures [2,32] (Fig. 6). Consequently, as for machining Ni-based alloys, it is recommended to use relatively low process temperatures (low rpm in this case) to avoid excessive tool wear and to achieve a suitable weld condition [2,32].

Finally, Fig. 5 shows the top surface micrographs of the stir zone for the parameters I and II, which corresponds to higher heat inputs (high rpm). It was found a considerable quantity of porosity and carbides. In this sense, the joint fatigue properties could be deteriorated due to porosities. Besides, the carbides were disposed as inter and intragranular manner in the matrix. Moreover, because of porosity and several carbides observed, it is suggested that high and excessive temperatures were obtained. In other words, local melting occurred and, therefore, temperatures of around 1200°C were reached.

Top surface micrographs of parameters III and IV are displayed in Fig. 6. The welded joints made with parameters III and IV (lower heat inputs) showed a top surface microstructure with only few M(C, N) carbonitrides, and it has to be mentioned that these were not found at grain boundaries in an observation by SEM. Therefore, it can be considered that relatively low FSW process temperatures were achieved. In general, as the tool rotational speed decreased (from 1200 to 200rpm), the grain refinement increased due to greater deformation and effective recrystallization. When the welding speed was altered, from 1 to 1.5mm/s, the most refined grains were observed. Hence, smallest grains can be achieved by increasing welding speed, which promotes higher cooling rate [25].

3.2Microhardness profiles

Fig. 7 presents the top surface microhardness profiles as a function of welding parameters. It can be observed that the weld bead was smaller for the parameters I and II (Fig. 4a and b), while the parameters III and IV displayed a large weld bead (Fig. 4c and d).

Fig. 7.

Top surface microhardness profiles for the welds processed at high rpm and low rpm.

(0.24MB).

The base material and stir zone microhardness levels ranged between 268–310HV and 335–385HV, respectively. However, the microhardness profiles were notably influenced by the process parameters chosen. In general, the grain size reduction was the main responsible for enhanced microhardness, a fact which has been reported by other authors [25–28]. In this sense, lower tool rotational speeds would promote higher hardness and tensile strengths [36]. Therefore, improved mechanical properties were also reported in friction-stir-welded Ni-based alloys [25–28].

As can be noted in Fig. 7, while the high rpm group of welds presented a microhardness scatter along the weld region, the low rpm group of welds achieved a more homogeneous profile in this zone. Also, parameter II reached the highest microhardness value (385HV). Still, from the microhardness profiles and microstructural features, it was verified that the parameters I and II led to several carbides locally formed and grain refinement in the stirred zone. Moreover, the hardness difference between the RS and AS for these welds was higher. Nevertheless, considering the parameters III and IV, it can be indicated that the best weld condition (parameter III) is mainly related to grain refinement and not through carbides precipitation. Hence, it is suggested that the friction based processes and their applications to alloy 625 result in superior joint properties with smaller grain size and relatively homogeneous weld microstructures.

3.3X-ray diffraction analyzes of residual stresses

Longitudinal residual stress results can be seen in Fig. 8, where it can be observed that the base material has relatively homogeneous residual stresses of around −650MPa. The longitudinal tensile residual stress peak increased for higher tool rotational speeds. In addition, as previously reported in [24], the tensile residual stresses peak was found in the longitudinal direction. When considering the stirred zone, longitudinal residual stresses are tensile for the parameters I and II (high rpm) with relatively higher values of 150 and 200MPa, respectively. For the set of parameters III and IV (low rpm), at the same region, the residual stress presented a near zero residual stress state. These residual stresses distributions kept this behavior for distances from −10 to 10mm.

Fig. 8.

Surface longitudinal residual stresses distributions as a function of FSW process parameters.

(0.22MB).

The comparison between parameters I and II and parameters III and IV shows differences up to 200MPa at the weld centerline. Changing the welding speed from 1 to 1.5mm/s not significantly modified the residual stresses. Hence, it is evident that the welding parameters set, especially an increase in the tool rotational speed [15], affects the residual stress magnitude near the weld centerline. Finally, the process asymmetry nature can be recognized for the parameters (I, II, III and IV) by dissimilarities between residual stress states on the RS and AS.

Transverse residual stresses are shown in Fig. 9, where the base material displays residual stresses of around −590MPa. Moreover, as can be verified, the residual stress peak is almost the same for the parameters I and II (around of 200MPa). It was also observed that an increase in the tool rotational speed led to higher residual stresses. Previous works have shown this tendency [15]. At the weld centerline, the results showed that decreasing the tool rotational speed from 1200 to 200rpm consequently changed the residual stress distributions from −100 to −200MPa. Thus, near the weld centerline, the transverse residual stress distributions are all compressive.

Fig. 9.

Surface transverse residual stresses distributions as a function of FSW process parameters.

(0.21MB).

It can also be observed in Fig. 9 that the parameters I and II (high rpm) have more significant differences in the residual stress distributions between RS and AS. Therefore, these residual stress profiles are more asymmetric. Hence, it may be related to their temperatures locally reached as well as specific material flows. These results are in a good agreement with the microhardness profile, where the high rpm welds showed a large variation in the microhardness along the weld. On the other hand, for the parameters III and IV (low rpm), the residual stresses between RS and AS sides are only slightly asymmetric, which indicates a more uniform material flow. This smaller difference between RS and AS was also verified in the microhardness distributions. As observed in the longitudinal residual stresses, the variation of welding speed from 1 to 1.5mm/s has not promoted higher modifications in the residual stress levels.

The FWHM analyzes of XRD peaks shows a different micro residual stress state due to the stress relaxation occasioned in FSW. Therefore, the FSW thermal and deformation gradients affected the FWHM distribution. The base material showed a FWHM of 2.75° which narrowed significantly for the stirred zone. Consequently, the recrystallization and finer grains led to an FWHM of approximately 0.8°, as can be observed in Fig. 10. However, even having different material flow and microstructures, all the joints possessed a similar tendency and peak breadth value in the weld zone, as noted elsewhere [22].

Fig. 10.

Full width at half maximum (FWHM) distributions of XRD peaks as a function of FSW process parameters.

(0.19MB).

Furthermore, the comparison between the base material and the set of parameters I and II shows some changes in FWHM after ±20mm distance from the weld centerline (on the RS). Also, the assessment between the base material and the welds produced by the parameters III and IV displays no considerable differences after ±20mm distance. Hence, based on these findings, it is noted that FSW can modify the initial FHWM base material magnitude.

Alloy 625 base material has a relatively homogeneous residual stress distribution as well as FWHM magnitude even presenting several carbides on its microstructure. Thus, these residual stress findings are mainly related to its previous manufacturing process (rolling and soft annealing). Since FSW is an asymmetric thermomechanical process, it was expected that both RS and AS would have different residual stress states. Therefore, the residual stress distributions changed as a function of the tool rotational speed and the welding speed. Moreover, the clamping system, frequently used in FSW, may exert an effect in the residual stresses [16]. Besides, the abrasive cleaning process employed before welding may also have influenced the residual stresses [24,37]. In addition, many characteristics of FSW such as severe plastic deformation (SPD), recrystallization and grain refinement play an import role in the residual stresses of friction-stir-welded alloy 625. Finally, it was interestingly seen that the weld corresponding to the parameter III (low rpm), which achieved the best top surface appearance, presented the lowest residual stress in the stirred zone.

4Conclusions

The present study was focused on characterizing the influence of process parameters on residual stress distributions of friction-stir-welded alloy 625. Therefore, the findings of the current investigation can be overviewed as follows:

  • The base material microstructure was composed by M(C,N) carbonitrides, M6C and M23C6 carbides. The welds made at high rpm (1200 and 1000rpm) presented refined grains, high quantity of carbides as well as porous formed in the weld zone which indicates that excessive temperatures were reached. On the other hand, generally, the welds made at low rpm (200rpm) showed the smallest grains, more homogenous microstructures and microhardness profiles, and only few M(C, N) carbonitrides in the stirred zone (observed by SEM). Hence, it can be assumed that the low rpm conditions achieved relatively low process temperatures enough to promote effective recrystallization and sound welds.

  • The tool rotational speed considerably affects the residual stresses in the weld zone. For high rpm welds, higher differences between the residual stress states of RS and AS were observed. On the other hand, for low rpm welds, the residual stress of RS and AS only slightly changed. The weld which presented an excellent top surface quality (parameter III, 200rpm) achieved the most compressive residual stress (−240MPa) in the stirred zone. Overall, in general, decreasing the tool rotational speed promoted lower residual stress in the stirred zone and adequate joints. Otherwise, increasing the tool rotational speed led to more tensile residual stresses at the weld centerline and unacceptable joints. Lastly, the change in welding speed, from 1mm/s to 1.5mm/s, does not cause higher variations in the residual stresses.

Conflicts of interest

The authors declare no conflicts of interest.

Acknowledgments

The authors of this work would like to thank CNPq (National Council for Scientific and Technological Development) through the Science Without Borders program and CAPES (Brazil) for their support (grant number 152437/2014-7).

References
[1]
Rolls-Royce Holdings: annual report. 65 Buckingham Gate, London.
(2012),
[2]
G.V.B. Lemos, S. Hanke, J.F. Dos Santos, L. Bergmann, A. Reguly, T.R. Strohaecker.
Progress in friction stir welding of Ni alloys.
Sci Technol Weld Join, 22 (2017), pp. 643-657
[3]
ASME Standard SB 446. Specification for nickel–chromium–molybdenum–columbium alloy (UNS N06625) rod and bar. New York.
American Society of Mechanical Engineers, (2008),
[4]
ASME Standard SB 564. EN-SB-564 specification for nickel alloy forgings. New York.
American Society of Mechanical Engineers, (2008),
[5]
A.B. Farina.
Efeito do teor de ferro e do tratamento térmico na microestrutura e propriedades da liga UNS N06625.
Polytechnic School of the University of São Paulo, Metallurgical and Materials Engineering Department, (2014),
[6]
H.L. Eiselstein, J. Gadbut.
Matrix-stiffened alloy. US 3,160,500 patent application. United States. December.
(1964),
[7]
H.L. Eiselstein, D.J. Tillack.
The invention and definition of alloy 625.
Superalloys 718, 625 and various derivatives,
[8]
S. Floreen, G.E. Fuchs, W.J. Yang.
The metallurgy of alloy 625.
Superalloys 718, 625, 706 and various derivatives,
[9]
W.M. Thomas, E.D. Nicholas, J.C. Needham, M. Murch, P. Temple-Smith, C.J. Dawes.
Friction-stir butt welding, GB patent no. 9125978.8, international patent application no. PCT/GB92/02203.
(1991),
[10]
Y.J. Kwon, I. Shigematsu, N. Saito.
Mechanical properties of fine-grained aluminum alloy produced by friction stir process.
Scr Mater, 49 (2003), pp. 785-789
[11]
Y.S. Sato, M. Urata, H. Kokawa.
Parameters controlling microstructure and hardness during friction-stir welding of precipitation-hardenable aluminum alloy 6063.
Metall Mater Trans A, 33 (2002), pp. 625-635
[12]
W. Wang, D. Deng, Z. Mao, Y. Tong, Y. Ran.
Influence of tool rotation rates on temperature profiles and mechanical properties of friction stir welded AZ31 magnesium alloy.
Int J Adv Manuf Technol, 88 (2017), pp. 2191-2200
[13]
Y. Zhang, Y.S. Sato, H. Kokawa, S.H.C. Park, S. Hirano.
Microstructural characteristics and mechanical properties of Ti–6Al–4V friction stir welds.
Mater Sci Eng A, 485 (2008), pp. 448-455
[14]
C. Gunter, M.P. Miles, F.C. Liu, T.W. Nelson.
Solid state crack repair by friction stir processing in 304L stainless steel.
J Mater Sci Technol, 34 (2018), pp. 140-147
[15]
N. Kumar, R.S. Mishra, J.A. Baumann.
Residual stresses in friction stir welding.
Butterworth-Heinemann, (2013), pp. 50-51 http://dx.doi.org/10.1016/C2013-0-09884-2
[16]
V. Richter-Trummer, E. Suzano, M. Beltrão, A. Roos, J.F. Dos Santos, P.M.S.T. De Castro.
Influence of the FSW clamping force on the final distortion and residual stress field.
Mater Sci Eng A, 538 (2012), pp. 81-88
[17]
A. Steuwer, S.J. Barnes, J. Altenkirch, R. Johnson, P.J. Withers.
Friction stir welding of HSLA-65 steel: Part II. The influence of weld speed and tool material on the residual stress distribution and tool wear.
Metall Mater Trans A, 43 (2011), pp. 2356-2365
[18]
T. Terasaki, T. Akiyama.
Mechanical behaviour of joints in FSW: residual stress, inherent strain and heat input generated by friction stir welding.
Weld World, 47 (2003), pp. 24-31
[19]
M. Peel, A. Steuwer, M. Preuss, P.J. Withers.
Microstructure, mechanical properties and residual stresses as a function of welding speed in aluminium AA5083 friction stir welds.
Acta Mater, 51 (2003), pp. 4791-4801
[20]
A.P. Reynolds, W. Tang, T. Gnaupel-Herold, H. Prask.
Structure, properties, and residual stress of 304L stainless steel friction stir welds.
Scr Mater, 48 (2003), pp. 1289-1294
[21]
P. Staron, M. Koçak, S. Williams, A. Wescott.
Residual stress in friction stir-welded Al sheets.
Phys B Condens Matter, 350 (2004), pp. E491-E493
[22]
L.N. Brewer, M.S. Bennett, B.W. Baker, E.A. Payzant, L.M. Sochalski-Kolbus.
Characterization of residual stress as a function of friction stir welding parameters in oxide dispersion strengthened (ODS) steel MA956.
Mater Sci Eng A, 647 (2015), pp. 313-321
[23]
DIN 17744:2002-09, Wrought nickel alloys with molybdenum and chromium – chemical composition.
Deutsches Institut für Normung, Beuth Verlag GmbH, (2002),
[24]
G.V.B. Lemos, R.M. Nunes, P. Doll, L. Bergmann, T.R. Strohaecker, J.F. Dos Santos.
Avaliação das Tensões Residuais em Juntas Soldadas de Inconel 625 Obtidas Através da Soldagem por Fricção e Mistura Mecânica.
Soldagem Insp, 22 (2017), pp. 35-45
[25]
K.H. Song, H. Fujii, K. Nakata.
Evaluation of grain refinement and mechanical property on friction stir welded Inconel 600.
Mater Trans, 50 (2009), pp. 832-836
[26]
K.H. Song, H. Fujii, K. Nakata.
Effect of welding speed on microstructural and mechanical properties of friction stir welded Inconel 600.
Mater Des, 30 (2009), pp. 3972-3978
[27]
K.H. Song, K. Nakata.
Effect of precipitation on post-heat-treated Inconel 625 alloy after friction stir welding.
Mater Des, 31 (2010), pp. 2942-2947
[28]
K.H. Song, K. Nakata.
Mechanical properties of friction-stir-welded Inconel 625 alloy.
Mater Trans, 50 (2009), pp. 2498-2501
[29]
K. Gangwar, M. Ramulu.
Friction stir welding of titanium alloys: a review.
Mater Des, 141 (2018), pp. 230-255
[30]
R.C. Wimpory, P.S. May, N.P. O’Dowd, G.A. Webster, D.J. Smith, E. Kingston.
Measurement of residual stresses in T-plate weldments.
J Strain Anal Eng Des, 38 (2003), pp. 349
[31]
R. Lackner, G. Mori, R. Egger, F. Winter, M. Albu, W. Grogger.
Sensitization of as rolled and stable annealed alloy 625.
Berg Huettenmaenn Monatsh, 159 (2014), pp. 12
[32]
S. Hanke, G.V.B. Lemos, L. Bergmann, D. Martinazzi, J.F. Dos Santos, T.R. Strohaecker.
Degradation mechanisms of pcBN tool material during friction stir welding of Ni-base alloy 625.
Wear, 376–377 (2017), pp. 403-408
[33]
V. Bushlya, O. Gutnichenko, J. Zhou.
Effects of cutting speed when turning age hardened Inconel 718 with PCBN tools of binderless and low-CBN grades.
Mach Sci Technol, 17 (2013), pp. 497-523
[34]
D. Loureiro, A.E. Diniz, A.B. Farina, S. Delijaicov.
The influence of turning parameters on surface integrity of nickel alloy 625.
Proc Inst Mech Eng Part B J Eng Manuf, (2016), pp. 1-11
[35]
M. Zielinska, M.Z. Yavorska, M. Porêba, J. Sieniawski.
Thermal properties of cast nickel-based superalloys.
Archiv Mater Sci Eng, 44 (2010), pp. 35-38
[36]
A.S. Golezani, R.V. R Barenji, A. Heidarzadeh, H. Pouraliakbar.
Elucidating of tool rotational speed in friction stir welding of 7020-T6 aluminum alloy.
Int J Adv Manuf Technol, 81 (2015), pp. 1155-1164
[37]
G.V.B. Lemos, P.H.C.P. Cunha, R.M. Nunes, L. Bergmann, J.F. Dos Santos, T. Clarke.
Residual stress and microstructural features of friction-stir-welded GL E36 shipbuilding steel.
Mater Sci Technol, 34 (2017), pp. 95-103
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Journal of Materials Research and Technology

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