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https://link.springer.com/article/10.1007/s11661-016-3631-3
DOI: 10.1007/s11661-016-3631-3
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DIRETORIA DE TRATAMENTO DA INFORMAÇÃO Cidade Universitária Zeferino Vaz Barão Geraldo
CEP 13083-970 – Campinas SP Fone: (19) 3521-6493
Physical Simulation of a Duplex Stainless Steel
Friction Stir Welding by the Numerical and
Experimental Analysis of Hot Torsion Tests
EDUARDO BERTONI DA FONSECA, TIAGO FELIPE ABREU SANTOS, SERGIO TONINI BUTTON, and ANTONIO JOSE RAMIREZ
Physical simulation of friction stir welding (FSW) by means of hot torsion tests was performed
on UNS S32205 duplex stainless steel. A thermomechanical simulator Gleeble 3800Ò with a
custom-built liquid nitrogen cooling system was employed to reproduce the thermal cycle measured during FSW and carry out the torsion tests. Microstructures were compared by means of light optical microscopy and electron backscatter diffraction. True strain and strain rate were calculated by numerical simulation of the torsion tests. Thermomechanically affected zone
(TMAZ) was reproduced at peak temperature of 1303 K (1030°C), rotational speeds of 52.4
rad s1(500 rpm) and 74.5 rad s1(750 rpm), and 0.5 to 0.75 revolutions, which represent strain
rate between 10 and 16 s1and true strain between 0.5 and 0.8. Strong grain refinement, similar
to the one observed in the stir zone (SZ), was attained at peak temperature of 1403 K (1130°C),
rotational speed of 74.5 rad s1 (750 rpm), and 1.2 revolution, which represent strain rate of
19 s1 and true strain of 1.3. Continuous dynamic recrystallization in ferrite and dynamic
recrystallization in austenite were observed in the TMAZ simulation. At higher temperature, dynamic recovery of austenite was also observed.
DOI: 10.1007/s11661-016-3631-3
Ó The Minerals, Metals & Materials Society and ASM International 2016
I. INTRODUCTION
F
RICTION stir welding (FSW) is a solid-statejoining technique, which has been object of intense research in the last few decades, not only due to the good quality of the resulting joints but also for eliminating solidification-related problems, such as voids, cracks, and segregation.[1] Industrial use of FSW is still focused in aluminum alloys, especially for the aeronautic, marine, and ground transportation fields. However, many researchers have proved the feasibility of applying FSW to joining steel,[2] stainless steel,[3]and titanium alloy,[4]for example.
Despite the intense research, FSW still requires a better understanding on the thermomechanical history to which the material is subjected during welding. Some
researchers were able to measure the thermal history
during welding, providing insightful relationships
between the resulting microstructure and the tempera-tures and cooling rates measured.[4,5] However, true strain and strain rate of the material are not yet fully understood. Finite element method (FEM) and compu-tational fluid dynamics (CFD) were both applied to address this issue, with incongruous results.[6] FEM simulations are limited in true strain and require remeshing after large deformation. On the other hand, CFD simulations consider the material as a high-vis-cosity fluid, which results in high strain rates and few data on true strain. As shown in TableI,[7–16]reported values of strain and strain rate calculated via CFD and FEM oscillate in a wide range.
Physical simulation consists in observing the resulting microstructure of a material after processing and trying to reproduce it by means of controlled tests. Because different thermomechanical paths may lead to similar microstructures, knowledge on the system being studied is instrumental. Forest et al. employed hot compression tests to simulate HSLA-65 steel friction stir welding and reported limitations to the experimental technique.[17] On the other hand, researchers who employed torsion tests for physical simulation of FSW had better results, which is justified by the fact that shearing is the main deformation mechanism taking place during FSW.[18,19] Moreover, torsion tests can avoid friction and achieve high strains at elevated strain rates before fracture. Physical simulations of FSW of HSLA-65 steel,[18,19] 304L and 310 stainless steels,[20] Ti alloy,[21] and EDUARDO BERTONI DA FONSECA, formerly M.Sc. Student
with the School of Mechanical Engineering, University of Campinas, Campinas, SP, Brazil, is now M.Sc. Graduate with the Brazilian Nanotechnology National Laboratory LNNano, CNPEM, Campinas, SP, Brazil. Contact e-mail: [email protected] TIAGO FELIPE ABREU SANTOS, Professor, formerly with the Brazilian Nanotechnology National Laboratory LNNano, CNPEM, and the School of Mechanical Engineering, University of Campinas, is now with the Department of Mechanical Engineering, Federal University of Pernambuco, Recife, PE, Brazil. SERGIO TONINI BUTTON, Professor, is with the School of Mechanical Engineering, University of Campinas. ANTONIO JOSE RAMIREZ, Professor, formerly with the Brazilian Nanotechnology National Laboratory LNNano, CNPEM, is now with the Department of Materials Science and Engineering, The Ohio State University, Columbus, OH.
Manuscript submitted February 5, 2015. Article published online July 6, 2016
nickel-based alloys[22]were studied, and they produced good results using hot torsion tests.
A special geometry of specimen was employed in these tests in order to reproduce the high cooling rate observed during FSW.[18] Nonetheless, this geometry results in a thermal gradient along the reduced section, and thus, the simple equations for strain and strain rate calculations cannot be applied. Moreover, it was pointed out that some materials undergo a polymorphic
transformation during cooling, which masks the
deformed microstructure. Other materials, such as the nickel-based alloys, do not have a polymorphic trans-formation and their thermomechanically affected zone (TMAZ) was successfully reproduced.[22] Nevertheless, the refined grain size observed in the stir zone (SZ) was not reproduced by the torsion tests.
Duplex stainless steels (DSS) also do not present a full polymorphic transformation during cooling from the temperatures achieved in FSW, which makes them good candidates for physical simulation, which is the objec-tive of this work. Although ferrite and austenite fraction changes with temperature, and so does their deforma-tion resistance, most of the microstructural evoludeforma-tion due to deformation is conserved after cooling.
FSW of DSS was reported by a number of authors who presented results with a strongly refined SZ,[3] balanced microstructure,[23]good mechanical properties, and absence of heat-affected zone (HAZ) and detrimen-tal secondary phases.[23,24]The TMAZ undergoes a less severe plastic deformation and still preserves the base metal (BM) elongated microstructure. FSW creates an asymmetric welded joint, with advancing and retreating sides defined according to the relative velocities on each side of the tool. On the advancing side (AS), where the direction of travel speed and the tangential rotational speed vectors are coincident, there is a sharp interface between the SZ and TMAZ. In opposition, on the retreating side (RS), where the direction of the vectors is opposed, there is a subtle and gradual transition between the SZ and BM, being difficult to identify the TMAZ.
DSS are designed to have similar volume fractions of ferrite and austenite, which provides good mechanical and environmental performances.[25,26] Because of the coexistence of phases, a complex mechanism takes place during hot deformation of DSS, as each phase has a distinct behavior. Due to the differences in stacking fault energy (SFE), slip systems, and diffusivity, strain par-titioning takes place at the early stages of deforma-tion.[27–29] Thus, ferrite accommodates most of the strain in the hot working temperature range. Strain is then transferred to austenite as deformation increases.
It is well known that austenite, which has a lower SFE, undergoes limited dynamic recovery (DRV) and softens mainly due to discontinuous dynamic recrystal-lization (DDRX) when nucleation sites are avail-able.[29–31] Another possible softening mechanism in austenite is subgrain coalescence. Ferrite, however, has a higher SFE, which leads to intensive DRV. At high strains, continuous dynamic recrystallization (CDRX) may become the acting softening mechanism, in which the increase of deformation causes high-angle grain boundaries to be formed due to the progressive rotation of subgrains.[29,32] There remains a certain controversy over the softening mechanisms in ferrite, since some authors suggest the occurrence of DDRX in certain hot deformation conditions.
As each mechanism is responsible for a different microstructural change, one can understand the hot deformation behavior and microstructural changes of the material after torsion tests. The main objective and contribution of this study is the use of torsion tests to obtain microstructures similar to the ones observed in FSW and correlate them to the acting softening mech-anisms. Moreover, strain and strain rate of the FSW process were calculated via numerical analysis.
II. EXPERIMENTAL PROCEDURE
The chemical composition of the studied UNS S32205
DSS is shown in TableII. The base metal displayed
Table I. Strain and Strain Rate Reported in the Literature
References Material Method
FSW Parameters Strain (–) Strain rate (s1) Obs. Rotational Speed (rpm) Welding Speed (mm min1) 7 AA6061-T6 FEM 290 to 390 120 to 240 20 to 24 — 8 AA2024-T3 FEM — — 133* 1000*
9 AA6061-T6 FEM 390 to 690 120 to 600 20 to 78 — strain values: SZ: >20 TMAZ: 16 to 20 HAZ: <16
10 AA2139-T8 FEM 800 to 2000 50 to 300 12 25 butt joint
11 AA2024-T3 FEM — 120 2,4* — 12 AA6061 FEM 650 100 to 300 80 — 13 AA2024-T351 CFD — — 6** — 14 AISI 1018 CFD 450 25 — 40 15 AA2524 CFD 300 126 10 to 5 10 16 AA5083-H18 CFD 1000 to 1500 100 to 300 200 to 300 2000 to 3000 * Maximum value. ** Mean value.
elongated lamellae of austenite in a ferrite matrix, as it is
observed in Figure1. Torsion specimens were machined
from 19-mm-thick plates according to the design pro-posed by Norton, which is schematically shown in Figure2.[18] This specimen has a through hole that allows for internal flow of a cooling gas in order to achieve the high cooling rates observed in FSW. Specimens were machined along the plates of austenite so that shearing occurred in hot torsion tests perpen-dicularly to the normal of the plates, as it happens in FSW.
Hot torsion tests were performed in a
thermome-chanical simulator Gleeble 3800Ò with a custom-built
liquid nitrogen (LN2) cooling system, as shown in
Figure3. The cooling system was developed in order to reproduce the thermal history measured during FSW.[5] Type K thermocouples were capacitor discharge welded in the middle of the reduced sections for temperature
control, as shown in Figure2. The specimens were
heated at a constant rate of 100 K s1 up to the test temperature, and held for 0.5 seconds before torsion. After deformation, specimens were cooled down to room temperature according to the measured FSW
cooling profile using the LN2 cooling system, which
creates a flow of nitrogen inside the specimen.
Torsion test temperatures were determined based on the work of Santos et al.,[5] who performed FSW on 6 and 10 mm plates of UNS S32205 and measured their thermal cycle using thermocouples, a pyrometer, and an infrared camera. A finite element model was used to validate the temperatures and to predict maximum temperature and cooling rates during FSW. Based on
these results, a peak temperature of 1303 K (1030°C)
was used for the physical simulation of the TMAZ and
1403 K (1130°C) for the SZ. These values correspond
to a position close to a steady-state regime of FSW. The number of revolutions was limited by the geometry of the specimen, which did not withstand more than 1.2 revolution before plastic instability. For the TMAZ simulations, specimens were rotated 0.5 or 0.75 revolu-tion, based on the results from preliminary tests. For the SZ, however, larger strains are necessary, and thus, 1.0 and 1.2 revolutions were applied to the specimen. Rotational speed was kept at 52.4 or 78.5 rad s1 (500 or 750 rpm) based on the work by Sinfield.[19]
All specimens were cut along the torsion axis and prepared using standard metallographic procedures for light optical microscopy. Electrolytic etching was per-formed in a solution of 40 vol pct. HNO3 in distilled
water in two steps: at 1.5 V for 60 seconds and at 1.0 V for 480 seconds. Images were taken close to the outer surface of the specimen, where the strain is maximized, and halfway through the reduced section length, where the temperature was controlled, as shown schematically
in Figure2. In all photomicrographs from torsion tests, the torsion axis is horizontal. Volume fraction of ferrite was determined by digital image analysis using ImageJÒ. Simulated microstructures were compared to actual FSW zones observed in butt-weld joint performed in 6-mm-thick plates using a step spiral tool of 40 pct metallic matrix of WRe (25 pct Re) and polycrystalline cubic boron nitride (PCBN) reinforcement, with shoul-der diameter of 25 mm and pin length of 6 mm.[24,33] Welding was performed in downward force control mode at 20.9 rad s1(200 rpm) of tool rotational speed, 1.67 mm s1(100 mm minute1) of welding speed, and 37 kN of downward force. Samples for metallography were taken from positions analogue to where the
Table II. Chemical Composition (Weight Percentage) of UNS S32205 Duplex Stainless Steel
C Si Mn Cr Ni Mo N P S Fe
0.030 0.32 1.84 22.4 5.3 2.9 0.145 0.034 <0.002 bal.
Fig. 1—Base metal of UNS S32205 duplex stainless steel composed of austenite lamellae (light phase) in ferrite matrix (dark phase) with nearly the same volume fraction.
Fig. 2—Torsion specimen for physical simulation, which has a through hole for LN2 cooling. A thermocouple was welded to the
temperature profiles were measured. Similar microstruc-tures were identified by light optical microscopy, and the samples were prepared for electron backscatter diffrac-tion (EBSD) in a FEI Quanta 650 scanning electron microscope. Noise reduction was performed on the EBSD maps, and ferrite volume fraction, grain size, and recrystallized phase fraction were calculated and com-pared to data available in the literature.[34] Recrystal-lized phase fraction was determined according to the following steps: (1) grain reconstruction with misorien-tation angle of 5 deg; (2) measurement of internal average misorientation angle within each grain; and (3) classification according to a minimum angle of 1 deg to define a subgrain. Grains were classified as deformed if the average misorientation angle exceeded 1 deg; grains with internal misorientation angle below 1 deg but with subgrain misorientation above 1 deg were considered substructured. All the remaining grains were considered recrystallized.
Since the torsion specimens used were hollow and the temperature was not uniform throughout the reduced section length, it is not possible to use typical equations for calculating strain and strain rate. Thus, numerical simulations of the torsion tests were carried out to determine the actual true strain and strain rate when a microstructure similar to FSW was observed in a sample from the torsion tests. In order to do so, it was necessary to measure the thermal profile developed in the reduced section before torsion.
This was done by submitting torsion specimens to heating and cooling cycles similar to the ones in the torsion tests. Four thermocouples were welded to each specimen in order to acquire the thermal history at different points of the reduced section. The positions of
the thermocouples are shown in Figure 4. Then, the
experimental thermal history was used to determine the temperature distribution along the entire specimen just before torsion, using finite element method (FEM)
software COMSOLÒ v4.2. This distribution obtained
by thermal simulation was employed in the numerical
simulations of the torsion tests, which included the deformation due to torsion. These deformation
simula-tions were performed on FEM software ForgeÒ 2008
and provided information on the true strain and strain rate of the torsion specimens after deformation. Based on the position of each simulated microstructure, their true strain and strain rate were estimated from the simulations.
The numerical simulations were carried out with properties of the DSS steel EN 1.4462, similar in chemical composition and mechanical properties to the S32205 steel, modeled with the Hansel-Spittel equa-tion[35] and constants shown in Eq. [1] to calculate the flow stress (r) as function of work temperature (T), strain (e), and strain rate (_e):
r¼ Aem1TTm9em2em4eð1þ eÞm5Teem7_em3_em8T; ½1 where A= 18023.234; m1=0.00495; m2= 0.02293; m3= 0.13722; m4=0.03375, and m5 to m9= 0 III. RESULTS A. Torsion Tests
The thermal cycle set-up in the thermomechanical simulator was based on FSW experimental measure-ments. The cooling profile measured during welding was fitted by an exponential decay function, which was then discretized in segments and programmed into the
simulator. The LN2 cooling system was activated by
the simulator soon after torsion. Thus, it controlled the temperature by heating up the specimen according to the programmed temperature profile. As a result, the FSW joints thermal cycle and the simulated tempera-ture profile presented a very good agreement.
More-over, the LN2 cooling system provided a good
reproducibility of thermal cycles, as it can be noticed in Figure 5.
Fig. 3—Thermomechanical simulator with custom-built liquid nitro-gen cooling system.
Fig. 4—Position of the thermocouples for measuring the thermal profile in the reduced section before torsion. Four thermocouples were welded near the center of the reduced section. Dimensions in millimeters.
The FSW thermal history shown in Figure5 was acquired by a thermocouple located at the joint line and 2 mm below the upper surface of the plates. However, since any thermocouple located in the SZ is shredded by the tool, the actual SZ temperature had to be estimated based on the FEM simulations by Santos et al.[5] Thus, torsion tests for SZ simulation were
carried out at 1403 K (1130°C), mean temperature
between experimental data and FEM simulation. Sim-ilarly, for the TMAZ simulation, a peak temperature of
1303 K (1030°C) in the hot torsion tests was
employed.
B. Light Optical Microscopy
Advancing and retreating sides of DSS FSWed joints have distinct characteristics due to the relative velocities on each side of the tool, which affects material flow and the final morphology of the cross section. While on the AS a clear interface can be observed between the SZ and the TMAZ, on the RS there is a subtle and gradual transition from the SZ to the BM. Light optical micrographs of both sides of the TMAZ and the center
of the SZ can be observed in Figure6. Elongated
austenite lamellae can be observed on both sides of the TMAZ, but on the RS there are more grain boundaries inside the austenite lamellae, which are more numerous closer to the SZ.
Light optical micrographs of torsion specimens
deformed at 1303 K (1030°C) are shown in Figure7,
which corresponds to the TMAZ simulation tempera-ture. Thus, by comparing the microstructures from
Figures6 and 7, it is possible to point out that
microstructures similar to the TMAZ-AS were obtained
at three different sets of parameters: 52.4 rad s1
(500 rpm) and 0.75 rev., 78.5 rad s1 (750 rpm) and
0.5 rev., and 78.5 rad s1 (750 rpm) and 0.75
rev. Correspondingly, similar microstructures to the
TMAZ-RS were obtained at 78.5 rad s1(750 rpm) and
0.75 rev. Simulated microstructures were first pointed out in terms of morphology and ferrite volume fraction using conventional metallography. Afterward, EBSD tests were carried out in order to properly quantify the comparison in terms of ferrite volume fraction, grain size, and recrystallized phase fraction.
Fig. 5—Thermal cycle measured during FSW and torsion tests.
Fig. 6—(a) Cross section of friction stir-welded UNS S32205, and microstructures at (b) TMAZ-RS, (c) center of SZ, and (d) TMAZ-AS.
On the other hand, the SZ displays strong grain refinement, especially in the austenite, as shown in Figure 6. On the AS, it has been reported average grain sizes ranging from 5 lm to below 1 lm.[3,23,33,36]On the
torsion samples deformed at 1403 K (1130 °C), no
similar microstructure was observed in terms of mor-phology. However, similar ferrite volume fraction and grain size were observed via EBSD, as will be discussed in the next section. The simulated microstructure had large austenite lamellae with few visible grain bound-aries, as it can be seen in Figure8, while even the largest lamellae in the SZ revealed several grain boundaries.
During FSW, the material from the SZ deforms at the same time as temperature rises and decreases. Besides that, there is a mechanical mixing that cannot be simulated using torsion tests, where only shearing takes place. However, such an experiment with thermal cycle associated with deformation could not be performed at the thermomechanical simulator. Moreover, it would become even more complicated to estimate true strain and strain rate from these experiments. Nevertheless, the
TMAZ in FSW is generated due to pure shearing at nearly a constant temperature. Thus, the thermome-chanical cycle of the proposed torsion tests related well to the formation of the TMAZ. In the SZ, on the other hand, there is a tridimensional material flow, which cannot be simulated by torsion tests and which is responsible for the overall morphology of the observed microstructure in FSW.
C. Electron Backscatter Diffraction
The regions of interest selected by light optical microscopy were analyzed using EBSD in order to obtain data on ferrite volume fraction, grain size, and recrystallized fraction. These data were then compared to the values reported by Santos.[34] As stated before, three samples were found to be similar in microstructure
morphology to the TMAZ-AS and one to the
TMAZ-RS, which was corroborated by the ferrite
volume fraction analysis, shown in Table III. Due to
the maximum temperature achieved in FSW and the
Fig. 7—Micrographs of torsion specimens deformed at 1303 K (1030°C): (a) 52.4 rad s1(500 rpm) and 0.75 rev., (b) 78.5 rad s1 (750 rpm)
and 0.5 rev., (c) and (d) 78.5 rad s1(750 rpm) and 0.75 rev. Micrographs (a), (b), and (c) correspond to the physical simulation of TMAZ-AS and (d) of TMAZ-RS.
torsion tests, ferrite volume fraction of the FSW-TMAZ and the torsion-simulated TMAZ were around 50 vol pct, which validates the torsion test temperature as similar to the FSW-TMAZ.
However, the grain size measurements shown in
Table III reveal a stronger grain refinement in the
torsion samples than that in the FSW-TMAZ. More-over, the recrystallized ferrite fraction is similar or larger in the simulated microstructures than that in the
FSW-TMAZ, as it can be observed in TableIII. This
means that deformation parameters promoted more CDRX in ferrite of the torsion specimens than in the TMAZ. This is corroborated by the substructured ferrite
fraction measured by EBSD, which was above
20 vol pct. The smallest value of recrystallized ferrite fraction was measured in the torsion test at 52.4 rad s1 (500 rpm). Rotational speed is related to strain rate; thus, tests at 78.5 rad s1 (750 rpm) promoted defor-mation build-up in ferrite, which favors CDRX.[23]
In austenite, small values of recrystallized fraction were observed, which is consistent with the deformation temperature and is expected from the observation of the
actual TMAZ. At 1303 K (1030°C), the occurrence of
DRV in austenite is limited and dislocations tangle.[26,29] As a result, deformed austenite fraction was above 70 vol pct, indicating load transfer from ferrite to austenite. Nevertheless, the recrystallized austenite frac-tion in the torsion tests is larger than that in the TMAZ, which means that less severe parameters may be better to simulate this FSW zone.
As it has been discussed in the previous section, torsion tests were not able to reproduce the microstruc-ture morphology of the FSW-SZ, although strong grain refinement was observed. Torsion tests performed at
1403 K (1130 °C) produced balanced microstructures,
with a ferrite volume fraction close to 50 vol pct, which is slightly below the expected values for the SZ. However, average grain sizes of 1 lm were observed in the torsion sample, similar to the SZ, as shown in Table III.
Fig. 8—Micrograph of torsion specimens deformed at 1403 K (1130°C), 78.5 rad s1(750 rpm), and 1.2 rev.
Table III. Ferrite Volume Fraction, Grain Sizes, and Recrystallized Fractions of FSW Zones and Torsion Samples Condition Temperature [K (° C)] Rotational Speed [rad s 1 (rpm)] Number of Revolutions
Ferrite Volume Fraction (Percent)
Grain Size* (l m) Recrystal- lized Frac-tion (Per-cent) Numerical Simulation ac a c True Strain (–) Strain Rate (s 1 ) FSW TMAZ-AS 52.0 4.5 ± 0.2 3.8 ± 0.2 31 1 Hot torsion 1303 (1030) 52.4 (500) 0.75 49.8 2.6 ± 0.2 1.7 ± 0.1 27 13 0.7 10 1303 (1030) 78.5 (750) 0.5 47.6 2.9 ± 0.5 1.8 ± 0.2 52 8 0.5 16 1303 (1030) 78.5 (750) 0.75 46.8 2.6 ± 0.3 1.6 ± 0.1 65 11 0.7 16 FSW TMAZ-RS 54.8 4.5 ± 0.2 2.9 ± 0.1 37 8 Hot torsion 1303 (1030) 78.5 (750) 0.75 50.5 2.5 ± 0.1 1.0 ± 0.1 75 17 0.8 14 FSW SZ-AS 55.2 1.1 ± 0.1 0.9 ± 0.1 50 35 FSW SZ-RS 62.5 1.6 ± 0.1 1.0 ± 0.1 56 29 Hot torsion 1403 (1130) 78.5 (750) 1.2 50.2 1.1 ± 0.1 1.3 ± 0.1 4 2 3 1.3 19 * Confide nce inte rval of 95 pc t for grain size measu rement s.
Recrystallized austenite fraction of the torsion sample was just below the actual values reported for the SZ. However, an elevated fraction of substructured austenite
was observed, indicating the occurrence of DRV. On the other hand, ferrite was mainly deformed, and thus, the recrystallized fraction was almost negligible. Due to the elevated strain, ferrite became entrapped between austenite lamellae in a bamboo-type microstructure, in which the distance between the lamellae became com-parable to the ferrite subgrain size, as it can be seen in
Figure9. The material flow during FSW fragments the
austenite lamellae and enhances the mobility of ferrite subgrains, which may lead to a more extensive CDRX and smaller ferrite deformed fraction. On the other hand, during torsion tests shearing mostly elongates the austenite lamellae instead of fragmenting them. In consequence, CDRX in ferrite may be impaired due to the low mobility of the subgrains, in a bamboo-type microstructure.
D. Numerical Simulation
Figures10(a) and (b) show the results of numerical
simulation of one of the torsion specimens. As it was done in SectionIII–B, the longitudinal section of the specimens was analyzed. Three profiles (P1, P2, and P3) were traced across the longitudinal section length [as shown schematically in Figures10(a) and (b)] in order to obtain true strain and strain rate values, shown in
Figures10(c) and (d). Based on the position of each
observed microstructure (Figures7and8) relative to the
Fig. 9—Ferrite inverse pole figure of torsion sample deformed at 1403 K (1130°C): 78.5 rad s1(750 rpm) and 1.2 rev.
Fig. 10—Numerical simulation of torsion test for 1403 K (1130°C), 78.5 rad s1(750 rpm), and 1.2 rev. Longitudinal section of the torsion specimen showing (a) true strain and (b) strain rate. P1, P2, and P3 are schematic profiles, which were used to obtain graphs of (c) true strain and (d) strain rate.
specimen longitudinal section, the values for true strain and strain rate were obtained from the numerical simulations, and are presented in TableIII.
Maximum strain rate of 30 s1was observed near the jaws of the thermomechanical simulator. In spite of the different torsion parameters used in each test, a good agreement of true strain and strain rate values was found in the analyzed FSW zones, i.e., specimens deformed under different conditions generated similar microstructures because at some point, they achieved similar values of true strain and strain rate. Thus, TMAZ was simulated with true strain between 0.5 and 0.8 and strain rate between 10 and 16 s1. On the other hand, the SZ was simulated with true strain of 1.3 and strain rate of 19 s1.
IV. DISCUSSION
It is known that similar microstructures can be obtained via different processing routes. However, it was taken care to reproduce conditions as similar as possible to the ones in FSW during the determination of the parameters for the torsion tests. Based on the thermal history acquired during FSW and a number of studies available in the literature, the temperature profile for different zones is the most certain parameter. On the other hand, deformation during FSW is not yet fully understood.
It is reasonable to state that the SZ undergoes more severe plastic deformation than the TMAZ, based on the microstructure and the fact that a tridimensional material flow develops during welding, which could not be simulated in the torsion tests. Physical simulation is limited by the geometry of the specimen and the capabilities of the thermomechanical simulator. Never-theless, it was still possible to attain a microstructure with some characteristics of the SZ, such as the strong grain refinement. Moreover, it was possible to point out the reason for the difference in the microstructure: larger strains would promote the fragmentation of the lamel-lae[26]but were limited by the specimen geometry.
On the other hand, the torsion-simulated TMAZ and the FSW-TMAZ had similar morphology and other comparable characteristics. The TMAZ is more satis-factorily reproduced because in this zone the material is deformed at once and nearly at a constant temperature, alike the torsion tests.
In the SZ, however, the material is deformed during heating, a condition that cannot be created in the thermomechanical simulator. Moreover, the treatment of the data in this case would be much more intricate. Nevertheless, torsion tests reproduce the main deforma-tion mode observed in the SZ of FSW, which is shearing.[36]
Previous work on Ni-based alloys was not successful in correctly reproducing the microstructures observed in friction stir welding or processing.[22] Authors pointed out that, in spite of no on-cooling transformation, the grain size after torsion tests did not reproduce the zones observed in FSW and that more trial and error exper-imentation could lead to microstructure replication.
Another work on Ti-6Al-4V obtained reasonable match-ing between FSW and hot torsion tests above and below b-transus temperature.[37] However, true strain was not estimated for each microstructure. Instead, stress–strain curves were used to propose a constitutive model for computational models of FSW.[38]
The present work succeeds in proposing a method for determining the true strain and strain rate based on the torsion tests and the simulated microstructures. These data are instrumental in the development of numerical simulation of FSW and for a better understanding of this welding process.
V. CONCLUSIONS
The following conclusions can be drawn from this work:
1. Numerical simulation of the torsion tests provided
reasonable values for true strain and strain rate in spite of the specimen geometry.
2. Both sides of the TMAZ were reproduced by
hot torsion tests at 1303 K (1030°C), 52.4 to
78.5 rad s1 (500 to 750 rpm), 0.5 to 0.75 rev.
Microstructures were similar in terms of
morphol-ogy and ferrite volume fraction. Grain size
and recrystallized fraction measurements revealed a stronger grain refinement in the simulated microstructures than that in the actual TMAZ. Numerical simulation of the torsion tests for TMAZ simulation resulted in true strain between
0.5 and 0.8 and strain rate between 10 and 16 s1.
3. Torsion tests performed at 1403 K (1130°C),
78.5 rad s1 (750 rpm), 1.2 rev. resulted in strong
grain refinement, as observed in the SZ with grain sizes of the order of 1 lm. However, the morphol-ogy of the simulated zone did not match the actual SZ, since a pure shear condition is attained in the torsion tests, while in FSW there is a tridimensional material flow. Numerical simulation of the torsion tests for SZ simulation resulted in true strain of 1.3
and strain rate of 19 s1.
ACKNOWLEDGMENTS
The authors would like to acknowledge FAPESP and CNPq for scholarships and Aperam for material donation. Research supported by Petrobras and ANP.
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