Weld toe modification using spherical-tip WC tool FSP in fatigue strength improvement of high-strength low-alloy steel joints

Materials and Design 160 (2018) 1019–1028

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Weld toe modification using spherical-tip WC tool FSP in fatigue strength improvement of high-strength low-alloy steel joints Hajime Yamamoto ⁎, Yoshikazu Danno, Kazuhiro Ito, Yoshiki Mikami, Hidetoshi Fujii Joining and Welding Research Institute, Osaka University, 11-1 Mihogaoka, Ibaraki, Osaka 567-0047, Japan

H I G H L I G H T S

G R A P H I C A L

A B S T R A C T

• FSP using a spherical tip tool was applied to the weld toe of HSLA steel joints. • The weld toe geometry and microstructure were successfully modified without defect formation. • Tool wear on the surface induced compressive residual stress. • Fatigue strength was improved by FSP, while was influenced by surface roughness related to tool travel speed.

a r t i c l e

i n f o

Article history: Received 22 June 2018 Received in revised form 10 October 2018 Accepted 26 October 2018 Available online 28 October 2018 Keywords: Friction stir processing Post-weld treatment High-strength low-alloy steel Fatigue strength Weld geometry Residual stress

a b s t r a c t Fatigue strength of fusion-welded joints is lower than that of the base metal due to stress concentration, tensile residual stress, and microstructural degradation at the weld toe. To improve these issues, friction stir processing (FSP) using a spherical tip tool was directly applied to the weld toe of high-strength low-alloy steel joints. The toe geometry was successfully modified without defect formation, resulting in 25% reduction in fatigue notch factor in comparison to the as-welded joints. Significant grain refinement due to FSP increased hardness beneath the toe surface. In addition, compressive residual stress related to tool wear was produced on the FSP surface. Bending fatigue strength was improved by these benefits, with dependence on the tool travel speed. Fatigue crack initiation occurred at the toe surface for all the joints. Although high travel speed FSP produced serrated surface resulting in degradation of FSP improvement in fatigue strength, low travel speed FSP contributed to reduction of surface roughness leading to maximum FSP improvement of 50% in fatigue strength. The results obtained in this study suggest the encouraging prospect of direct application of FSP to weld toes as new post-weld treatment for steel joints. © 2018 Elsevier Ltd. All rights reserved.

1. Introduction Fatigue strength of fusion-welded joints is lower than that of the base metal because of stress concentration, tensile residual stress, and microstructural degradation generated at the weld toe. To benefit fully from advanced materials exhibiting high strength, various post-weld treatments have been proposed, and generally divided into two groups of residual stress modification, such as peening [1–3] and post-weld ⁎ Corresponding author. E-mail address: [email protected] (H. Yamamoto).

https://doi.org/10.1016/j.matdes.2018.10.036 0264-1275/© 2018 Elsevier Ltd. All rights reserved.

heat treatment [4], and weld geometry modification such as grinding [5–7] and remelting [8–10]. The former is more effective than the latter under low fatigue stress levels, whereas it seems that the former does not show any benefit under high stress levels, because of residual stress relaxation during fatigue [11]. On the contrary, reduction of stress concentration at the weld toe can improve fatigue life regardless of the stress level. In addition, remelting methods using arc plasma or laser as a heat source have advantages of not only smoothing the weld surface profile, but also removing internal defects or existing cracks without reducing plate thickness [9]. It has been reported that tungsten inert gas (TIG) dressing causes softening around the remelted zone,

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H. Yamamoto et al. / Materials and Design 160 (2018) 1019–1028

which impedes large improvement of fatigue strength for the highstrength steel welds [10]. Moreover, brittle structure due to martensitic or bainitic transformation can be formed depending on the cooling rate and hardenability in the remelted zone for carbon steels. On the other hand, friction stir processing (FSP) can be used as a new post-weld treatment to improve fatigue strength of the fusion-welded joints. It is known that FSP is an effective surface modification technique to refine and homogenize microstructure of various metallic materials in a solid state, based on the principle of friction stir welding (FSW) using a rotational tool [12]. Mechanical properties such as strength, ductility, and fracture toughness of a stir zone (SZ) produced locally by FSP can become superior to those of the base metal [13–15]. Furthermore, material flow during FSP contributes to elimination of internal defects that can act as crack initiation sites in addition to grain refinement, resulting in the improvement of fatigue properties of cast alloys [16–18]. With the use of these advantages, Costa et al. reported that the application of FSP increased fatigue strength of several aluminum alloy joints fabricated using metal inert gas welding [19–22]. The reason was explained in terms of grain refinement, geometry modification, and elimination of weld defects such as porosity and overlap at the toe. Therefore, FSP can provide grain refinement in addition to the other benefits provided by conventional methods. However, use of the FSP technique has so far been limited to a few light metal alloy welds. For carbon steels used as major structural materials, it might possibly be effective for not only reducing the stress concentration at the toe without microstructural degradation, but also for modifying soft or brittle structure formed in the heat-affected zone (HAZ). Mechanical properties of the SZ formed in the carbon steels can be significantly influenced by phase transformation [23,24] or tool wear [25,26] during FSP, depending on the process conditions such as tool rotational speed and travel speed, and an understanding of these behaviors, which are largely different from those of aluminum alloys, is essential for the application. In the present study, FSP using a spherical tip tool was performed directly around the weld toe of high-strength low-alloy (HSLA) steel joints, and feasibility of fatigue strength improvement was investigated in terms of weld geometry modification, residual stress modification, grain refinement strengthening, and surface roughness produced by FSP. 2. Experimental procedures Chemical compositions of 10-mm-thick HSLA steel plates and MXZ200 filler metals used to obtain butt-welded joints in this study are shown in Table 1. Si and Mn contents in the filler metals are more than those in the steel plates, and act actively as deoxidizers to prevent porosity formation in the weld metal. Three-pass CO2 gas-shielded arc welding was conducted on V-groove with a root gap of 5 mm between the plates under the conditions indicated in Table 2. An FSP tool was made of WC–6%Co, and had a 15-mm-diameter spherical tip with a curvature radius of 10 mm as shown in Fig. 1(a). FSP was performed on positions 2 mm away from both right- and leftside of the weld toe, where the rotational axis was located (Fig. 1(b)). The tool was tilted to the welding direction by 3°, and operated at a travel speed of 100–500 mm/min with a clockwise rotation at a rotational speed of 200–800 rpm in Ar atmosphere. Cross-sectional observation of the as-welded specimens with and without FSP was conducted using optical microscopy (OM) and scanning electron microscopy (SEM). The cross section of the specimens was etched with 2% nital solution after mechanical polishing. Main

Table 2 CO2 gas-shielded arc welding parameters. Pass

Welding current, I/A

Arc voltage, V/V

Welding speed, v/mm/min

Sweepback angle, θ (degree)

1st 2nd 3rd

170 280 280

22 32 29

130 200 300

– – 5

elements of the base metal, filler metal and FSP tool were analyzed by electron probe microanalyzer (EPMA) equipped with SEM. Vickers hardness tests were carried out between 0.03 and 2.5 mm lower than the toe surfaces on the cross sections at room temperature (RT) with an applied load of 0.98 N and loading time of 30 s. Transverse residual stress was measured along the direction perpendicular to the weld bead on the surface around the weld toes of fourpoint bending fatigue specimens explained later, using a twodimensional X-ray diffraction (XRD) detector (μ-X360n, Pulstec Industrial Co., Ltd.) employing Cr-Kα radiation with a 1-mm-diameter collimator and incident angle of 34.3°, which was estimated from dspacing of the {211}α-Fe planes based on the cosα technique. In general, fatigue cracks initiate at the weld toe and propagate along the fusion line. The crack propagation is influenced by the transverse residual stress, which was identical with a direction of tensile stress on the weld surface arising from the applied stress in the four-point bending fatigue tests. The X-ray incident angles at the excess weld metal were corrected according to its shape image. The four-point bending fatigue tests were carried out at RT with a sinusoidal waveform of 20 Hz as a function of maximum applied stress with a stress ratio of 0.1. The specimens were prepared by cutting at 10 mm intervals from the welded plates using electrical discharge machine. Outer and inner spans of 10-mm-diameter pins were 90 and 50 mm, respectively. The weld bead was located at the center of the inner pin span, between which the base metals on both right- and left-sides of the weld bead were included. The cross section and fracture surface of the specimens after the fatigue test were observed by OM and SEM to characterize the fracture location and crack initiation site. Surface roughness around the weld toe of a spare specimen (including the fracture location of tested specimens) was measured at a speed of 18 mm/min using SURFCOM 1400G (Tokyo Seimitsu Co., Ltd.). 3. Results 3.1. As-welded joint Fig. 2 shows OM images of the as-welded specimens. The base metal consisted of ferrite-pearlite structure (Fig. 2(c) and (d)). The HAZ was formed around the weld metal with width of about 1–3 mm. The undercut was observed at the toe (Fig. 2(e)), resulting in measured toe radius and angle of about 0.27 mm and 27°, respectively. Microstructure beneath the toe exhibited bainitic ferrite grains as shown in Fig. 2(f). 3.2. Optimization of FSP application to weld toe FSP was conducted on the weld surfaces, and the influences of positioning on the advancing side (AS) and retreating side (RS) were investigated. Fig. 3 shows an OM image and EPMA-Si map (large Si-content difference between the base metal and filler metal) obtained in a typical joint with FSP (obtained with the travel speed of 100 mm/min at the

Table 1 Chemical compositions of the base metal and filler metal (unit mass%). Materials

C

Si

Mn

P

S

Cu

Ni

Cr

Mo

V

Al

Fe

Base metal Filler metal

0.14 0.03

0.23 0.55

1.08 1.55

0.014 0.011

0.006 0.007

0.01 0.02

0.01 0.01

0.02 0.02

– b0.01

0.002 b0.01

0.028 –

Bal. Bal.

H. Yamamoto et al. / Materials and Design 160 (2018) 1019–1028

(a)

15

(b)

1021

(a) FSP tool

FSP tool

(WC-6%Co)

(WC-6%Co)

SR10

2

AS

RS

AS

RS

Flash

2 10 60䢛

70

5

70

(b)

Initial geometry

Fig. 1. Schematic illustrations of (a) the tool geometry and (b) setup for direct application of FSP to the toe surface.

constant rotational speed of 800 rpm). A color change from blue to red indicates an increase in Si content. FSP formed the approximately 1mm-thick SZ around the toe without obvious defects, regardless of whether the positioning was on the AS or RS. On the other hand, there was difference in flash on the base metal side, depending on whether it was on the AS or RS. The EPMA results indicated that the weld metal having high Si content flowed into the base metal side at the RS of the SZ in the left-side of the excess weld metal, while most of the weld metal chipped by the tool became the enormous flash on the base metal side at the RS of the SZ in the right-side of the excess weld metal (Fig. 3(b)). It seems that the decrement of the cross-sectional SZ area in comparison to the initial geometry depicted as a white broken line went to the flash. Thus, the positioning of the AS and RS should be base metal and excess weld metal sides, respectively, to prevent the flash formation. The difference can be attributed to material flow direction and topological asymmetry of the SZ, which is different from a topological symmetry SZ in conventional FSP/FSW on flat plates [27–29]. Fig. 4 shows a process window for FSP in relation to the tool travel speed and rotational speed (v and ω, respectively). The available travel speed could increase with increasing the rotational speed, while the tool was broken into small fragments during FSP with combination of the high-travel and low-rotational speeds, as shown in an OM image of Fig. 4. In the successful cases without tool breakage (○), revolutionary pitches of v/ω, suggesting heat input under the FSP conditions, exhibited minimum and maximum values of 0.125 and 0.625 mm/rev, respectively, obtained from the travel speed of 100 and 500 mm/min at the constant rotational speed of 800 rpm (hereafter, FSP100/800 and FSP500/800). Those two sets of FSP conditions were selected for investigating the benefits obtained, in terms of modification of toe geometry,

(a) As-welded

Fig. 3. Cross-sectional (a) OM image and (b) EPMA-Si map of a typical joint with FSP (position of the AS and RS is different between left- and right-sides of the excess weld metal). (For interpretation of the references to color in this figure, the reader is referred to the web version of this article.)

microstructure, and residual stress (the RS was located on the excess weld metal). 3.3. Weld toe geometry Fig. 5 shows OM images of the joints with FSP100/800 and FSP500/ 800. The SZ was formed to about the tool diameter and 1–2 mm depth from the SZ surface, a little bit lower than the initial base metal plate surface. The SZ depth was increased with increasing FSP heat input, while seemed to be slightly uneven in the two sides of weld bead and in other cross-sectional images we observed (not shown here). It can be associated with the heat generation and plastic flow due to unstable contact between the spherical tip of tool and the weld toe. It is noted that the difference had little effect on fatigue strength obtained in this study. Similarly, the amount of flash is associated with the FSP heat input, and increases with increasing it. Thus, that of joints with FSP100/800 is larger than that with FSP500/800 (Fig. 5). The joint as shown in Fig. 3 was produced by FSP100/800, and its amount of flash seems to be larger than that of the joint with FSP500/800, but a little smaller than that of the joint with FSP100/800 as shown in Fig. 5. It would be acceptable range in variation of the same FSP condition. The toe radius significantly increased to 3.53 and 3.60 mm in the joints with FSP100/800 and FSP500/800, respectively, in comparison

(c)

(e)

(d)

(f)

b

(b)

Fig. 2. (a) Cross-sectional OM image of the as-welded joint. (b)–(f) Enlarged images.

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H. Yamamoto et al. / Materials and Design 160 (2018) 1019–1028 Table 3 Stress concentration factors at the toe surface.

䕿 Successful

㽢Tool broken

㽢

Travel speed,

/ mm/min

500

䕿 /

㽢

300

䕿

0

200

As-welded w/FSP100/800 w/FSP500/800

䕿

䕿

500

kt ¼ 1 þ 0:165ð tanθÞ0:167 ðt=r Þ1=2

ð1Þ

k f ¼ 1 þ ðkt −1Þ=ð1 þ a=r Þ

ð2Þ

where θ is the toe angle, r is the toe radius, and t is the plate thickness (10 mm). Peterson's material parameter a for steels is obtained from Eq. (3) [30]: a ¼ 1:087  105 Su −2

ð3Þ

(a) w/ FSP100/800

(b)

RS

0.28 3.53 3.60

1.89 1.26 1.26

1.62 1.25 1.25

3.4. Microstructural evaluation

with 0.27 mm of the as-welded joint. On the other hand, the toe angle slightly increased to 36 and 33°, in comparison with 27° in the aswelded joint. To evaluate the influence of FSP on the toe geometry, the elastic stress concentration factor of kt and fatigue notch factor of kf for butt-welded joint were calculated based on Eqs. (1) and (2), respectively [30]:

AS

27 36 33

Fatigue notch factor, kf

Elastic stress concentration factor, kt

800

/ rpm

Fig. 4. Process window for FSP in relation to tool travel and rotational speeds (v and ω, respectively).

b

Toe radius, r/mm

where Su is ultimate bending strength of the notch-root material. The value in this study was obtained as 958 MPa from the bending test for the base metal. The values of kt and kf calculated in all the joints with and without FSP are summarized in Table 3, together with θ and r. Consequently, kf for both joints with FSP exhibited similar values, and decreased by 25% due to the significant increase in the toe radius.

= 0.125 mm/rev

䕿 Rotational speed,

Toe angle, θ (degree)

= 0.625 mm/rev

䕿 /

100

Joints under tests

Cross-sectional OM and SEM images beneath the toe of joints with FSP100/800 and FSP500/800 in high magnification are shown in Fig. 6. A feature of surface after FSP in the SZ was different; a smooth surface was observed in the joint with FSP100/800 (Fig. 6(b) and (c)), in marked contrast to serrated surface with some notches indicated by arrows in the joint with FSP500/800 (Fig. 6(f) and (g)), the size of which corresponds to the tool machining locus. On the other hand, FSP provided significant grain refinement in the SZ. The grain size in the SZ produced by FSP500/800 was smaller than that by FSP100/800 due to the lower heat input (Fig. 6(d) and (h)). Microstructural modification due to FSP on steel welds has already been demonstrated to be effective for fatigue strength improvement by the authors [31,32]. Thermomechanically affected zone (TMAZ) and HAZ in the regions divided by black broken lines in Fig. 6(a) and (e) were formed by FSP100/800 and FSP500/800, respectively, below the SZ with depth of about 500 μm around the toe. The TMAZ and HAZ can be recognized in the joint with FSP500/800, but their boundary in the joint with FSP100/800 was not clear, since the heat input was more associated with plastic flow. The morphologies of the initial weld metal remained slightly in the TMAZ and HAZ, while the grain sizes decreased with getting close to the SZ. It is noted that lath-like structure was observed beneath the

(c) w/ FSP500/800

d

RS AS

AS

RS

RS AS

(d)

Fig. 5. Cross-sectional OM images of joint with (a) FSP100/800 and (c) FSP500/800. (b), (d) Enlarged images of areas surrounded by a rectangular frame in (a) and (c), respectively.

H. Yamamoto et al. / Materials and Design 160 (2018) 1019–1028

1023

(a) w/ FSP100/800

(b)

(c)

(d)

(e) w/ FSP500/800

(f)

(g)

(h)

Fig. 6. OM images beneath the toe surface of joint with (a) FSP100/800 and (e) FSP500/800. (b)–(d) and (f)–(h) Enlarged SEM images of areas pointed by alphabet characters in (a) and (e), respectively, and (c) and (g) for areas surrounded by a rectangular frame in (b) and (f).

toe surface of the joint with FSP100/800 found at both sides of the center part with refined structure as shown in Fig. 6(b) and at lower left of Fig. 6(c). The areas surrounded by broken lines exhibited contrast different from the refined structure, and about 50 μm in depth (Fig. 6(a)–(c)). To characterize the areas observed at the toe of joint with FSP100/ 800, EPMA elemental analyses were conducted focusing on elements of which the joints and tools consisted. Fig. 7 shows EPMA-W maps of the joints with FSP100/800 and FSP500/800. A color change from blue to red indicates an increase in the W content. The W-rich areas were distributed more extensively and deeply beneath the SZ surface in the joint with FSP100/800 than that with FSP500/800. A W-rich area was also located at the toe of joint with FSP100/800 (Fig. 7(c)). This suggests that tool wear during FSP100/800 with higher heat input occurred more significantly than that during FSP500/800. The W-rich areas correspond to the area with the lath-like structure. This suggests that tool material of WC was decomposed and W alloying widened a martensitic region in continuous cooling transformation diagram during FSP, resulting in enhanced solution hardening due to C and W. Fig. 8(a)–(c) show cross-sectional OM images of an as-welded joint and those with FSP, together with broken lines where Vickers hardness was measured, and the measurement results are shown in Fig. 8(d). In the as-welded joint, the highest hardness of about 200 HV was observed at the weld metal, and the average hardness of 167 HV at the base metal. The grain refinement in the SZ and TMAZ produced by FSP increased the

(a) w/ FSP100/800

hardness. Increase in hardness of the SZ with finer grains produced by FSP500/800 was higher than that by FSP100/800. On the other hand, it is noted that the point 0.03 mm lower than the surface of joint with FSP100/800 exhibited significantly high hardness of about 487 HV, and corresponds with the W-rich areas distributed in the SZ suggesting martensitic transformation with lath-like structure. 3.5. Transverse residual stress The transverse residual stress on the surface was measured by XRD at areas depicted by circles on OM images as shown in Fig. 9(a)–(c), and summaries of the measurements are shown in Fig. 9(d). The positive and negative values indicate tensile and compressive residual stress, respectively. The specimen shape was the same as that prepared for fatigue tests. High tensile residual stresses of about 200 MPa were measured around the center of weld metal in the as-welded specimen, while low tensile residual stress of about 27 MPa was measured at the toe. It is noted that FSP induced high compressive residual stress in almost all the SZ area and in a half region of the SZ area away from the toe in the joints with FSP100/800 (about −500 MPa) and FSP500/800 (about −300 MPa), respectively. The areas correspond with the Wrich areas as shown in the EPMA-W maps (Fig. 7). The W contents and its distribution width seem to depend on depth and width of the W-rich areas, respectively. In contrast, tensile residual stresses similar

(d) w/ FSP500/800

W-rich area

W-rich area

(b)

(c)

(e)

(f)

Fig. 7. EPMA-W maps of joints with (a) FSP100/800 and (d) FSP500/800. (b)–(c) and (e)–(f) Enlarged images of areas indicated by alphabet characters in (a) and (d), respectively. (For interpretation of the references to color in this figure, the reader is referred to the web version of this article.)

1024

H. Yamamoto et al. / Materials and Design 160 (2018) 1019–1028

(a) As-welded

(b) w/ FSP100/800

(c) w/ FSP500/800

transformation, which would be a unique phenomenon that has not been reported in any literature on residual stress distributions induced by FSW [26,33–35]. 3.6. Fatigue properties

(d) 600

SZ

Hardness (HV0.1)

500

Weld metal

TMAZ䡚HAZ

w/ FSP100/800 w/ FSP500/800 As-welded

W-rich area

400 SZ

Weld metal

TMAZ HAZ

300

200

100

Weld metal

HAZ

Base metal

0 0

0.5

1.0

1.5

2.0

2.5

Depth from toe surface, / mm Fig. 8. OM images near weld toe of joints (a) as-welded, (b) with FSP100/800, and (c) with FSP500/800, and (d) Vickers hardness variation with depth from the toe surface measured along a broken line depicted in (a)–(c).

to those obtained in the as-welded joint were observed in the other SZ regions without the W-rich areas. Consequently, the toe in joint with FSP100/800 exhibited a compressive residual stress of 214 MPa, in contrast to the tensile residual stress of 52 MPa in that with FSP500/800. Such W segregation due to tool wear on the SZ surface may be a reason for the induction of compressive residual stress as well as martensitic

(a) As-welded

Toe

Bending fatigue performance of these joints with FSP was investigated, and relationship between applied nominal stress amplitudes and number of cycles to failure (S-N diagram) is shown in Fig. 10, together with those of the as-welded joints and base metal specimens. Fatigue life was significantly improved by FSP in comparison with that of the as-welded joints at all stress amplitudes. The number of cycles to failure of 2 × 106 cycles was set as fatigue limit, at which the stress amplitudes became 225 and 200 MPa (arrows mean “without failure”) in the joints with FSP100/800 and FSP500/800, respectively. These improvement rates were estimated to be about 50 and 33% in comparison to those of the as-welded joint. On the other hand, fatigue life in the high stress amplitude in the joints with FSP100/800 was similar to that in the base metal specimens. Fatigue life could not be measured for the base metal specimens in the highest stress amplitude same as joints with FSP, which is similar to the yield stress of base metal. The joints with FSP had the excess weld metal, which was acted as a reinforcement member, and their weld toes were hardened. Thus, strength of the welded joints with FSP was higher than that of the base metal specimens, and the tendency of S-N curves in higher stress amplitude near their yield stress was different, and the comparison has no point. Consequently, it can be concluded that FSP is an effective technique to improve the fatigue strength of butt-welded steel joints, similar to aluminum alloy joints [19–22]. To understand the difference in stress amplitudes between the joints with FSP100/800 and FSP500/800, failure positions and stress concentration effects were investigated. Fig. 11 shows cross-sectional OM images of three kinds of typical joints after fatigue failure, for which all the joints failed at their toe, probably caused by local stress

(d) 1000

As-welded w/ FSP500/800 w/ FSP100/800

/ MPa

800

(b) w/ FSP100/800

-214 MPa ( = -7 mm)

(c) w/ FSP500/800

Weld metal

400

Transverse residual stress,

+27 MPa ( = -10 mm)

HAZ

Base metal

600

200

0 -200 -400

-600

AS

RS

SZ (FSP500/800)

-800

AS

RS

SZ (FSP100/800) -1000

-25

+52 MPa ( = -8 mm)

-20

-15

-10

Distance from weld center,

-5

0

/ mm

Fig. 9. OM images of joint surface (a) as-welded, (b) with FSP100/800, and (c) with FSP500/800, and (d) transverse residual stress variation with distance from the weld center measured on circles depicted in (a)–(c).

H. Yamamoto et al. / Materials and Design 160 (2018) 1019–1028

500

500

= 0.1

Stress ratio:

/ MPa

/ MPa

Stress ratio:

= 0.1

400

Local stress amplitude,

,

,

400

Nominal stress amplitude,

1025

300

200

50% UP

Base metal w/ FSP100/800 w/ FSP500/800 As-welded 100 10000 104

100000 105

1000000 106

Number of cycles to failure,

300 13% UP

200 Base metal w/ FSP100/800 w/ FSP500/800 As-welded

10000000 107

(Cycles)

100

104 Fig. 10. S-N diagram obtained in four-point bending fatigue tests for joints as-weld, with FSP100/800, and with FSP500/800, together with that of the base metal.

concentration. The local stress amplitude of σa, loc at the toe can be calculated using the following Eq. (4): σ a;loc ¼ σ a;nom ∙k f

ð4Þ

where the nominal stress amplitude of σa, nom can be obtained as in Fig. 10 and the coefficient of kf is shown in Table 3. Fig. 12 shows a modified S-N diagram for all the joints, plotting σa, loc instead of σa, nom. It is noted that the geometric factor related to local stress concentration at the toe was eliminated in the modified S-N diagram. Consequently, σa, loc at fatigue limit of the joints with FSP100/800 was still about 13% higher than that of the as-welded joints, and similar to that of the base metal, while the figure suggests that the weld geometry modification was the major factor in the fatigue strength improvement. This indicates that grain refinement strengthening and compressive residual stress attributed to the W-rich areas beneath the toe contributed to about 13% improvement in the fatigue strength (FSP100/800). In contrast, there might be external factors involved in the degradation of the improvement due to FSP500/800 in addition to less W-rich areas, especially near the toe, although the grain refinement strengthening was larger than that due to FSP100/800. To understand the external factors, cross-sectional OM images and EPMA-W maps around the fatigue failure positions at toe were obtained as shown in Fig. 13. The W-rich areas were observed at the toe surface of joint with FSP100/800, but not observed at that with FSP500/800, and the areas prevented crack initiation. On the other hand, the toe surface of the joint with FSP100/800 was not fully covered with W-rich areas as shown in Figs. 6(b), (c), and 7(c). The crack tends to initiate at the matrix or/and boundaries between the matrix and W-rich areas because of their large hardness difference, in the joint with FSP100/800 (Fig. 13 (b)). Although no microcrack was observed at either side of failure position in the joint with FSP100/800, some microcracks at the left side of failure position were observed at the toe surface without W-rich areas in the joint with FSP500/800 (Fig. 13(c) and (d)). This indicates

(a) As-welded

(b) w/ FSP100/800

105

106

Number of cycles to failure,

107

(Cycles)

Fig. 12. Modified S-N diagram for all joints, plotting local stress amplitude of σa, loc instead of nominal stress amplitude of σa, nom.

that W-rich areas at the toe surface can be resistance to fatigue crack initiation. 3.7. Fatigue fracture surface Fig. 14 shows SEM images of the fatigue fracture surface obtained at stress amplitude of 300 MPa for all the joints. The fracture surfaces consisted of the SZ and TMAZ, in order from the toe surface, and the SZ depth was about 500 μm. The fracture surface in the SZ produced by FSP (Fig. 14(d) and (f)) exhibited finer wrinkling than that of the as-welded joint (Fig. 14(b)), which is attributed to smaller striation spacing corresponding to crack propagation distance per one cyclic load. This suggests that the grain refinement due to FSP is effective to enhance resistance to the crack propagation. In the joint with FSP100/ 800, a delaminated zone underneath the topmost W-rich area was observed, perhaps because large hardness difference between the topmost W-rich area and the underneath induced sudden crack propagation underneath (Fig. 14(d)). The toe surface was fully covered with more Wrich area, which indicates positive utilization of the tool wear; higher fatigue strength can be obtained due to the prevention of the crack initiation. On the other hand, it can be seen that there are a lot of ratchet marks, which indicate the boundary between two adjacent failure planes, in the joints with FSP100/800 and FSP500/800 (Fig. 14(c), (e), and (f)). The number of ratchet marks depends on the number of crack initiation sites. The number per fracture surface width in both joints is shown in Fig. 15. Consequently, the number of ratchet marks in the joint with FSP500/800 as shown in Fig. 14(e) was more than two times of that with FSP100/800, and this trend was similar at the different stress amplitude of 250 MPa. This suggests that the fatigue crack can initiate easily from the surface of joint with FSP500/800, as some microcracks near the failure position as seen in Fig. 13(c).

(c) w/ FSP500/800

Fig. 11. OM images of typical joints (a) as-welded, (b) with FSP100/800, and (c) with FSP500/800 after bending fatigue fracture.

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(a) w/ FSP100/800

(c) w/ FSP500/800

W-rich area

(b)

(d)

Fig. 13. (a), (c) OM images and (b), (d) EPMA-W maps around the fatigue crack initiation site in joints with FSP100/800 and FSP500/800.

3.8. Surface roughness The σa, loc of the joints with FSP500/800 were lower than those of aswelded joints in the modified S-N diagram (Fig. 12); the degradation of improvement due to FSP500/800 is attributed to not only loss of the Wrich area at the toe surface, but other factors as well. To understand the difference in resistance to the fatigue crack initiation at the toe between the two sets of FSP conditions, surface roughness was measured, and the results are shown in Fig. 16. Surface roughness varied periodically under both sets of FSP conditions. The amplitude and frequency of the surface roughness became much smaller and higher, respectively, in the joint with FSP100/800 than that with FSP500/800. The wavelength approximately corresponds to the weld pitch. The maximum height roughness, which indicates the largest amplitude, was about 3.62 and 12.43 μm in the joints with FSP100/800 and FSP500/800, respectively. The higher maximum height roughness in the joint with FSP500/800 was related to the serrated surface as seen in Fig. 6(g) and more ratchet marks as

(a) As-welded

(c) w/ FSP100/800 SZ

(b)

(d)

indicated in Fig. 15, resulting in more crack initiation sites at the surface. The decrease in the tool travel speed contributed to reduction in the surface roughness, indicating that materials removed ahead of the tool were accumulated behind the tool, and a lower travel speed meant sufficient time was available for materials that accumulated behind the tool to have leveled off on the surface. Thus, optimization focusing on the tool travel rotational speeds should lead to further improvement in fatigue strength of the steel joints with FSP in terms of the surface roughness as well as tool wear and grain refinement around the toe. 4. Discussion As mentioned in the introduction, grain refinement strengthening and weld geometry modification have been obtained by FSP for the fatigue strength improvement of several aluminum alloy joints. The weld geometry modification due to FSP using the spherical tip tool was the major factor in the fatigue strength improvement of the welded

(e) w/ FSP500/800 SZ

(f)

Fig. 14. SEM images of the fatigue fracture surface obtained at stress amplitude of 300 MPa for joints (a) as-welded, (c) with FSP100/800, and (e) with FSP500/800. (b), (d), and (f) Enlarged images of areas surrounded by rectangular frames in (a), (c), and (e), respectively.

Number o f ratchet marks per f racture surf ace width,

/ mm-1

H. Yamamoto et al. / Materials and Design 160 (2018) 1019–1028

4 w/ FSP100/800 w/ FSP500/800

3

2

1

0 250 Nominal stress amplitude,

300 / MPa

Fig. 15. Number of ratchet marks per fracture surface width at stress amplitudes of 250 and 300 MPa for joints with FSP100/800 and FSP500/800.

HSLA steel joints with FSP, while the benefit due to the tool wear was also obtained in this study, in addition to grain refinement that is a specific factor of FSP. The hard W-rich area at the processed surface could be one of resistances to fatigue crack initiation as seen in the joint with FSP100/800, due to high hardness induced by martensitic transformation as well as grain refinement at the processed surface. Even though a fatigue crack initiated at the surface, high compressive residual stress would prevent the crack propagation. The two main factors leading to the fatigue strength improvement are associated with martensitic transformation observed only in the W-rich area at the processed surface. It is well known for W alloying in steel materials to increase hardenability caused by martensitic transformation, and in addition martensitic 12.5

(a) w/ FSP100/800

10.0

/ μm

5.0

Height,

7.5

0

3.62 μm

-2.5 -5.0

0.125 mm/rev

-10.0 -12.5 0 12.5

0.5

1.0

1.5

2.0

2.5

3.0

3.5

(b) w/ FSP500/800

10.0

4.0

4.5

5.0

FSP was applied directly to the weld toe of the HSLA steel joints using a spherical tip tool. The weld toe geometry and microstructure were successfully modified without defect formation. Compressive residual stress was also produced, and its magnitude and distribution width depended on FSP condition. Bending fatigue strength and life were improved by FSP, but varied with the tool travel speed (FSP100/ 800 and FSP500/800). The important findings are as follows:

= 12.43 μm

1) FSP resulted in 25% reduction in the fatigue notch factor due to a significant increase in the toe radius in comparison to that of the aswelded joints. The effect did not vary with tool travel speed.

7.5

/ μm

transformation expansion leaves compressive residual stress. Thus, the tool wear during FSP seems to be followed by its decomposition into W and C, probably due to high pressure and heat produced by FSP. Simultaneously, W elements were dissolved into the topmost of SZ at the toe of HSLA steel joints. Based on this discussion, effects of FSP condition on the tool wear, WC decomposition, and W alloying in the SZ related to high hardness and compressive residual stress should be analyzed in the future work. There is evidence that two obvious peaks of W 4f5/2 (~33 eV) and W 4f7/2 (~31 eV) were observed in a portion (around the W 4f) of a X-ray photoelectron spectroscopy profile of the W-rich area of SZ surface beneath a oxide scale in a stainless steel with FSP using a WC tool, suggesting metallic W in the processed surface (not yet submitted). However, maximum hardness at the topmost area of SZ center in the joint with FSP100/800 was much higher than that resulting from martensitic area (not shown here), suggesting that hard WC due to the tool wear remained on the surface in addition to W alloying leading to the martensitic transformation. Thus, it is the future work to analyze how much those factors contribute to fatigue strength improvement. On the other hand, some microcracks occurred at the toe without the W-rich area in the joint with FSP500/800, where is beside the fatigue failure position similar to that with the W-rich area in the joint with FSP100/800 that was the SZ edge. This suggests that the W-rich area at a weld toe plays an important role in preventing the fatigue crack initiation and propagation. However, large surface roughness produced by FSP500/800 resulted in more crack initiation sites at the surface, leading to degradation of fatigue strength improvement. Thus, the toe without the W-rich area cannot always necessarily degrade preventing crack initiation. The effect of surface roughness produced by FSP on fatigue properties is also a specific factor of FSP, which found in this study. The larger surface roughness becomes, the lower fatigue strength would be obtained. The relationship between FSP condition and surface roughness has not been clear yet. The new two factors of the tool wear and surface roughness are needed further analyses, such as identifying mechanisms of WC decomposition and W alloying in the SZ, and clarifying relationship between FSP surface roughness and fatigue property degradation. Among them, the tool wear is generally considered to be avoided in FSP due to the damage of the tool and the increased cost, but if the toe surface is fully covered with more W-rich area, which indicates positive utilization of the tool wear as a new line of thought, higher fatigue strength would be obtained in the future studies. 5. Conclusions

2.5

-7.5

Height,

1027

5.0 2.5 0 -2.5 -5.0 -7.5 -10.0 -12.5

0.625 mm/rev 0

0.5

1.0

1.5

2.0

2.5

Distance,

3.0

3.5

4.0

4.5

5.0

/ mm

Fig. 16. Surface roughness around the toe surface of joints with (a) FSP100/800 and (b) FSP500/800.

2) FSP increased hardness due to significant grain refinement and formation of lath-like structure beneath the toe surface of joints with FSP100/800. The increase in the SZ with finer microstructure produced by FSP500/800 was higher than that by FSP100/800. On the other hand, the W-rich areas distributed beneath the SZ surface due to the tool wear induced martensitic transformation in the joint with FSP100/800, and its hardness was significantly high (about 487 HV). 3) FSP induced compressive residual stress on the SZ surface, corresponding to the positions of W-rich areas. The W contents and

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width of the W-rich areas depended on the FSP conditions, and the toe surface of joint with FSP100/800 exhibited a compressive residual stress of 214 MPa, in contrast to the tensile residual stress of 52 MPa at that with FSP500/800. 4) Bending fatigue life of the joints with FSP was longer than that of the as-welded joints at all the applied stress amplitudes. The stress amplitudes without failure at 2 × 106 cycles were about 225 and 200 MPa in the joints with FSP100/800 and FSP500/800, and they were estimated to be about 50 and 33% higher, respectively, in comparison to that of the as-welded joint. 5) Fatigue crack initiation occurred at the toe surface for all the joints. On the other hand, the W-rich area at the toe surface in the joint with FSP100/800 contributed to the prevention of crack initiation, while the higher maximum height roughness would more easily induce crack initiation in the joint with FSP500/800 than in the aswelded joint. CRediT authorship contribution statement Hajime Yamamoto: Microstructure and fracture analyses, Surface roughness measurements, Corresponding author. Yoshikazu Danno: Specimen preparations, Bending fatigue tests, Microstructure and fracture analyses, Vickers hardness tests. Kazuhiro Ito: Funding, Discussion. Yoshiki Mikami: Funding, Discussion. Hidetoshi Fujii: Funding, Discussion. Acknowledgements One of the authors, Mr. Yamamoto, gratefully acknowledges the financial support from Program for Leading Graduate Schools: Interactive Materials Science Cadet Program in Osaka University. The authors would like to thank Mr. Yoshihisa Mishima at X-ray Residual Stress Measurement Center for the residual stress measurement. This paper is partially based on results obtained from a future pioneering project commissioned by the New Energy and Industrial Technology Development Organization (NEDO). References [1] W. Ting, W. Dongpo, H. Lixing, Z. Yufeng, Discussion on fatigue design of welded joints enhanced by ultrasonic peening treatment (UPT), Int. J. Fatigue 31 (2009) 644–650. [2] Y. Sakino, Y. Sano, R. Sumiya, Y.-C. Kim, Major factor causing improvement in fatigue strength of butt welded steel joints after laser peening without coating, Sci. Technol. Weld. Join. 17 (2012) 402–407. [3] J. Berg, N. Stranghöner, Fatigue behaviour of high frequency hammer peened ultra high strength steels, Int. J. Fatigue 82 ( (2016) 35–48. [4] H.T. Kang, Y.-L. Lee, X.J. Sun, Effects of residual stress and heat treatment on fatigue strength of weldments, Mater. Sci. Eng. A 497 (2008) 37–43. [5] N. Ye, T. Moan, Improving fatigue life for aluminium cruciform joints by weld toe grinding, Fatigue Fract. Eng. Mater. Struct. 31 (2007) 152–163. [6] R. Baptista, V. Infante, C.M. Branco, Study of the fatigue behavior in welded joints of stainless steels treated by weld toe grinding and subjected to salt water corrosion, Int. J. Fatigue 30 (2008) 453–462. [7] Z. Fu, B. Ji, X. Kong, X. Chen, Grinding treatment effect on rib-to-roof weld fatigue performance of steel bridge decks, J. Constr. Steel Res. 129 (2017) 163–170. [8] T. Dahle, Design fatigue strength of TIG-dressed welded joints in high-strength steels subjected to spectrum loading, Int. J. Fatigue 20 (1998) 677–681. [9] A.L. Ramalho, J.A.M. Ferreira, C.A.G.M. Branco, Fatigue behaviour of T welded joints rehabilitated by tungsten inert gas and plasma dressing, Mater. Des. 32 (2011) 4705–4713.

[10] T. Skriko, M. Ghafouri, T. Björk, Fatigue strength of TIG-dressed ultra-high-strength steel fillet weld joints at high stress ratio, Int. J. Fatigue 94 (2017) 110–120. [11] H.C. Yıldırım, Recent results on fatigue strength improvement of high-strength steel welded joints, Int. J. Fatigue 101 (2017) 408–420. [12] R.S. Mishra, Z.Y. Ma, Friction stir welding and processing, Mater. Sci. Eng. R 50 (2005) 1–78. [13] D. Yadav, R. Bauri, Effect of friction stir processing on microstructure and mechanical properties of aluminium, Mater. Sci. Eng. A 539 (2012) 85–92. [14] D.M. Sekban, S.M. Aktarer, P. Xue, Z.Y. Ma, G. Purcek, Impact toughness of friction stir processed low carbon steel used in shipbuilding, Mater. Sci. Eng. A 672 (2016) 40–48. [15] P. Xue, B.B. Wang, F.F. Chen, W.G. Wang, B.L. Xiao, Z.Y. Ma, Microstructure and mechanical properties of friction stir processed cu with an ideal ultrafine-grained structure, Mater. Charact. 121 (2016) 187–194. [16] D.R. Ni, D. Wang, A.H. Feng, G. Yao, Z.Y. Ma, Enhancing the high-cycle fatigue strength of Mg–9Al–1Zn casting by friction stir processing, Scr. Mater. 61 (2009) 568–571. [17] S. Jana, R.S. Mishra, J.B. Baumann, G. Grant, Effect of friction stir processing on fatigue behavior of an investment cast Al–7Si–0.6 Mg alloy, Acta Mater. 58 (2010) 989–1003. [18] R. Kapoor, K. Kandasamy, R.S. Mishra, J.A. Baumann, G. Grant, Effect of friction stir processing on the tensile and fatigue behavior of a cast A206 alloy, Mater. Sci. Eng. A 561 (2013) 159–166. [19] J. da Silva, J.M. Costa, A. Loureiro, J.M. Ferreira, Fatigue behaviour of AA6082-T6 MIG welded butt joints improved by friction stir processing, Mater. Des. 51 (2013) 315–322. [20] L.P. Borrego, J.D. Costa, J.S. Jesus, A.R. Loureiro, J.M. Ferreira, Fatigue life improvement by friction stir processing of 5083 aluminium alloy MIG butt welds, Theor. Appl. Fract. Mech. 70 (2014) 68–74. [21] J.D.M. Costa, J.S. Jesus, A. Loureiro, J.A.M. Ferreira, L.P. Borrego, Fatigue life improvement of mig welded aluminium T-joints by friction stir processing, Int. J. Fatigue 61 (2014) 244–254. [22] J.S. Jesus, J.M. Costa, A. Loureiro, J.M. Ferreira, Fatigue strength improvement of GMAW T-welds in AA5083 by friction-stir processing, Int. J. Fatigue 61 (2017) 124–134. [23] H. Fujii, L. Cui, N. Tsuji, M. Maeda, K. Nakata, K. Nogi, Friction stir welding of carbon steels, Mater. Sci. Eng. A 429 (2006) 50–57. [24] L. Cui, H. Fujii, N. Tsuji, K. Nakata, K. Nogi, R. Ikeda, M. Matsushita, Transformation in stir zone of friction stir welded carbon steels with different carbon contents, ISIJ Int. 47 (2007) 299–306. [25] S.H.C. Park, Y.S. Sato, H. Kokawa, K. Okamoto, S. Hirano, M. Inagaki, Boride formation induced by pcBN tool wear in friction-stir-welded stainless steels, Metall. Mater. Trans. A 40 (2009) 625–636. [26] 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 (2012) 2356–2365. [27] A.P. Reynolds, Flow visualization and simulation in FSW, Scr. Mater. 58 (2008) 338–342. [28] Z. Zhang, H.W. Zhang, Numerical studies on controlling of process parameters in friction stir welding, J. Mater. Process. Technol. 209 (2009) 241–270. [29] X.H. Zeng, P. Xue, D. Wang, D.R. Ni, B.L. Xiao, Z.Y. Ma, Effect of processing parameters on plastic flow and defect formation in friction-stir-welded aluminum alloy, Metall. Mater. Trans. A 49 (2018) 2674–2683. [30] J.-Y. Yung, F.V. Lawren, Analytical and graphical aids for the fatigue design of weldments, Fatigue Fract. Eng. Mater. Struct. 8 (1985) 223–241. [31] K. Ito, T. Okuda, R. Ueji, H. Fujii, C. Shiga, Increase of bending fatigue resistance for tungsten inert gas welded SS400 steel plates using friction stir processing, Mater. Des. 61 (2014) 275–280. [32] H. Yamamoto, K. Ito, Effects of microstructural modification using friction stir processing on fatigue strength of butt-welded joints for high-strength steels, Mater. Sci. Appl. 9 (2018) 625–636. [33] H. Polezhayeva, A.I. Toumpis, A.M. Galloway, L. Molter, B. Ahmad, M.E. Fitzpatrick, Fatigue performance of friction stir welded marine grade steel, Int. J. Fatigue 81 (2015) 162–170. [34] 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) 313–321. [35] J.W. Sowards, T. Gnäupel-Herold, J.D. McColskey, V.F. Pereira, A.J. Ramirez, Characterization of mechanical properties, fatigue-crack propagation, and residual stresses in a microalloyed pipeline-steel friction-stir weld, Mater. Des. 88 (2015) 632–642.