Experimental and numerical study on local mechanical properties and failure analysis of laser welded DP980 steels

Materials Science & Engineering A 680 (2017) 378–387

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Experimental and numerical study on local mechanical properties and failure analysis of laser welded DP980 steels

crossmark

Qiang Jiaa, Wei Guoa, Weidong Lia,＊, Peng Penga,b, Ying Zhua, Guisheng Zouc, Yun Pengd, Zhiling Tiand a

School of Mechanical Engineering and Automation, Beihang University, Beijing 100191, China International Research Institute for Multidisciplinary Science, Beihang University, Beijing 100191, China c Department of Mechanical Engineering, Tsinghua University, Beijing 100084, China d Welding Research Institute, Central Iron and Steel Research Institute, Beijing 100081, China b

A R T I C L E I N F O

A BS T RAC T

Keywords: Laser welding Heat-affected zone Finite element analysis Thermal simulation Mechanical properties Failure location

DP980 steel was welded using Nd:YAG laser with low heat input (4 m/min) and high heat input (1 m/min). The ultimate tensile strength (UTS) of the low heat input welded joint reached 99.7% of base metal, while the joint with high heat input reached 95.6%. The sub-critical heat affected zone (HAZ) was softened with obvious yielding platform, and higher heat input caused lower strength of local HAZs. The low heat input welded joint fractured at base metal, while the joint with high heat input failed at sub-critical HAZ with significantly lower fracture strain. The simulation results revealed that UTS of joint decreased with the increase of HAZ width and HAZ softening degree. Moreover, there were critical values of HAZ width and softening degree to predict the failure location of the joint. With the increase of fusion zone width, the UTS of joint increased gradually closing to that of base metal, and the joint tended to fail at base metal due to biaxial stress intensification effect.

1. Introduction

Extensive studies have been conducted on the mechanical performance of fusion zone (FZ) and HAZ in terms of AHSS welds. Conventional experiment methods to characterize local properties have been reported using mini tensile samples, which are machined directly from FZ and HAZ of the joint, respectively [6,7]. However, laser welding provides much narrower FZ and HAZ compared with conventional arc welding, making such mini tensile samples are difficult to machine because of insufficient FZ and HAZ in the joints. On the other hand, it is inaccurate for assuming a homogeneous property of HAZ using mini tensile sample method, as the HAZ microstructure varies along with the thermal gradient in real situation [8]. The tensile test based on a “rule of mixture” is another method commonly adopted to extract properties of the FZ and HAZ based on the assumption of isostrain, of which the weld metal is parallel to the loading direction [9]. Lee et al. extracted the average mechanical properties of the weld bead and the HAZ of tailed-welded blanks via the subsize samples [10]. Abdullah et al. obtained the weld properties from tailed-welded blank using a similar approach [11]. In their study, four different sized samples were tested, and more accurate results were found using smaller samples owing to the larger proportion of weld in the cross section of sample. The tensile test combined with the “rule of mixture” is an appealing method for obtaining local properties due to its

The environment friendly vehicles demand less fuel consumption as well as lower CO2 emissions. The ever-increasing use of advanced high strength steel (AHSS) is a promising way of lightening car body in the automotive industry. Ferrite-martensite dual-phase (DP) steel is one of the AHSS families widely adopted for body-in-white [1,2]. DP steels consist of martensite islands embedded in ferritic matrix. These martensite islands contribute to the strength of DP steel, while the ductility arises from ferrite [3]. The combination of martensite and ferrite offers higher initial work hardening rate along with considerably uniform elongation compared to conventional steels [4]. Welding is the mostly used joining technique in the automobile industry. Among various welding methods, laser welding is playing a vital role in the joining of AHSS due to its flexibility and high energy density [5]. The microstructures of joint are transformed locally under welding thermal cycles, and then the mechanical properties change correspondingly. The safety of vehicles is closely related to the mechanical behavior of welded joint. Thus, characterizing and understanding of local microstructures and mechanical properties of welded joints is important for modeling and predicting the overall mechanical behavior of DP welded joints [6].

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Correspondence to: Beihang University, XueYuan Road No.37, HaiDian District, Beijing, China. E-mail address: [email protected] (W. Li).

http://dx.doi.org/10.1016/j.msea.2016.10.121 Received 5 August 2016; Received in revised form 29 October 2016; Accepted 31 October 2016 Available online 05 November 2016 0921-5093/ © 2016 Elsevier B.V. All rights reserved.

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(a)

simplicity, but only the average properties of the weld or the HAZ could be acquired. As a result, the property variation of the HAZ was ignored, whereas it is important for numerical simulation. Local constitutive behavior of HAZ could be obtained by scaling stress-strain curve according to the hardness profile, where the variation of the HAZ properties was considered [12–14]. For example, Pavlina and Van Tyne found that the tensile strengths of steels over the range of 450–2350 MPa presented a linear correlation with hardness [15]. However, Rojek et al. pointed out that this method underestimates the yield stress in the range of small plastic deformation and overestimates the yield stress in the range of large plastic deformation [16]. The newly developed digital image correlation (DIC) has been applied in measuring local strain during tensile test [17–19]. The corresponding local stress is based on the iso-stress assumption of the whole sample. Only part of the stress-strain curve can be obtained for the hardened zone, as little plastic deformation would occur there. Thermal simulation is an effective method with precise temperature control that can reproduce a relatively large HAZ samples with homogeneous microstructure. This feature allows for conventional mechanical properties test of HAZ [20]. Goodall et al. have examined the toughness of thermal simulated HAZ of arc-welded X80 line pipe steel by Charpy impact test [21]. Dancette et al. have investigated the local microstructure and constitutive behavior of spot welds of DP steels experimentally with a Gleeble 3500 thermo-mechanical simulator [22]. However, the welding thermal cycles are difficult to measure for the thermal simulation, as the HAZ is too narrow to locate several thermocouples at different positions accurately. In this work, therefore, the thermal cycles experienced during laser welding of DP980 steel were identified using finite element (FE) analysis. The thermal cycles were then reproduced using thermo-mechanical simulator to investigate the local constitutive behaviors of the joint. The overall tensile behavior of the joint was evaluated by experiment and FE analysis. Particular attention was paid to the influence of heat input on the failure location of the joint.

Laser beam

Argon gas 45°

Weld

90°

Welding sample

Clamping system

Fig. 1. (a) Schematic illustration of laser welding; (b) Experimental setup of thermal simulation. Table 2 Welding parameters used in the study. Laser power (kW)

Welding speed (m/ min)

Focus distance (mm)

Head angle (°)

Beam diameter (mm)

Shielding gas, Ar (L/min)

3.5

1 and 4

0

0

0.45

16

weld zone in order to ensure precision. To simplify the model, material was assumed as isotropic and the surface of welding pool was flat, and the ambient temperature was 25 ℃. The cone-shaped body heat source with Gaussian distributed thermal energy density was employed in the simulation [23]. The dimensions of the FZ was measured by experiments and taken as the parameters of the heat source model. The heat flux in the predefined Cartesian coordinate system can be expressed as

2. Experimental procedures 2.1. Materials and laser welding process Hot dip galvanized DP980 steel with 1.2 mm thickness was used in the experiments. The chemical compositions and mechanical properties of the base metal (BM) are shown in Table 1. The milling machined steel sheets (100 mm×200 mm) were welded in butt-joint configuration, as shown in Fig. 1a. Laser welding was conducted using a Nd:YAG laser system (TRUMPF HL4006D), with the welding parameters presented in Table 2. Low heat input joint was welded at the welding speed of 4 m/min, and high heat input was achieved at 1 m/min. Highpurity argon (99.99%) was employed as shielding gas on the top surface of the steel sheets at 16 L/min.

Q(x, y, z) =

2ηPe −r 2 (r − ri)z exp ( 2 ); r 2 = x2 + y 2 ; r0 = ri + e h πr 2e h r0

(1)

where η is thermal efficiency, P is the laser power, re is the half width of FZ at the top surface, ri is the half width of the FZ at the bottom surface, and h is the welding penetration. The boundary conditions including convection and radiation were imposed to all the free surfaces in the model, while the symmetrical plane was thought to be adiabatic. The physical properties and thermo-physical properties of DP980 steel are listed in Table 3 [24,25] and Table 4 [26], respectively.

2.2. FE simulation of laser welding process Laser welding process was numerically simulated to estimate the thermal cycles experienced at different local zones of the weld. The commercial FE software Abaqus 6.14 was employed for the simulation. Since the welded sheets are symmetrical about the weld center, only half of the weld was modeled to save calculating time. The calculation domain was 100 mm×200 mm×1.2 mm, which was fine meshed in

2.3. Thermal simulation Thermal simulation of HAZ was carried out on a Gleeble1500 system, and the experimental setup is shown in Fig. 1b. Samples were machined into the dimension of 100 mm×30 mm×1.2 mm, and the

Table 1 Chemical compositions (wt%) and tensile properties of the DP980 steel. Material

C

Mn

Si

Al

Cr

Mo

Cu

Fe

YS (MPa)

UTS (MPa)

Elongation (%)

DP980

0.15

1.5

0.31

0.05

0.02

0.05

0.02

Balance

638.5

986.9

15.8

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Table 3 Physical properties of DP980 steel. Properties

Values

Density(kg/m3) Ac1 (℃) Ac3 (℃) Solidus temperature (℃) Liquidus temperature (℃) Boiling temperature (℃) Latent heat of fusion (J/kg) Specific heat of liquid (J/(kg K)) Effective thermal conductivity of liquid (W/(m K))

7840 696 860 1468 1514 2827 2.46 ×105 682 109.2

Table 4 Thermo-physical properties of DP980 steel. Temperature (℃)

25

125

325

525

725

925

1225

Specific heat (J/kg K) Thermal conductivity (W/m K)

449 69.5

491 66.5

572 55.1

675 44.3

961 40.5

607 36.6

651 36.6

length direction was along the rolling direction. The samples were precoated with glass lubricant to prevent oxidation at high temperatures. A closed-loop temperature control was achieved by a K-type thermocouple welded at the middle of the sample. To prevent overshooting of the temperature, the maximum heating rate of 300 ℃/s was adopted. Once the peak temperature reached, different cooling rates of welds with high heat input (1 m/min) and low heat input (4 m/min) were simulated using compressed air and water ejected with compressed air, respectively. Fig. 3. Microstructure, temperature field and hardness distribution of the weld cross section of (a),(b) low heat input (LHI) and (c),(d) high heat input (HHI).

2.4. Microstructure and mechanical property testing Metallographic samples were prepared by standard metallographic procedure followed by etching with 3% nital solution. Microstructures of different welds were observed using optical microscope (OM, OLYMPUS DP72) and scanning electron microscope (SEM, JSM 6010). Vickers microhardness was measured on the etched joint samples at middle depth from the surface of the plate by a microhardness tester (FM-800) at a load of 300 g and 15 s dwell time. At least five hardness points were tested and averaged for each thermal simulated sample. Tensile tests were performed at room temperature using a computerized INSTRON 8801 testing system at a constant velocity of 1.5 mm/min. The geometry and dimensions of the tensile sample are

displayed in Fig. 2a. Transverse tensile samples of the joint were machined according to ASTM: E8M standards, of which the weld was positioned at the center of the tensile samples. And the strain was measured using an extensometer with a gauge length of 25 mm. Constrained by sample size, the thermal simulated samples for tensile test were machined following a nonstandard size. The FZ tensile sample was cut from the weld zone directly. A special customized extensometer with a gauge length of 5 mm was used to measure the strain for these nonstandard tensile samples, as shown in Fig. 2b.

Hydraulic clamp Tensile sample

Customized extensometer

Fig. 2. (a) Geometry and dimensions of the tensile samples (mm) and (b) tensile test with customized extensometer.

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Fig. 4. Local thermal cycles of welding simulation with (a) low heat input (b) high heat input, and thermal histories setting for HAZ thermal simulation with (c) low heat input (d) high heat input.

Fig. 5. SEM microstructures of (a) BM, (b) FZ of low heat input (LHI) and (c) FZ of high heat input (HHI) (M: martensite; F: ferrite).

changes of welded joint on the condition of high heat input. The high heat input weld exhibited wider FZ and deeper hardness drop in softened zone compared to low heat input. It should be noted that this local hardness difference between low heat input and high heat input welds might have correlation with the potential failure initiation of the joints [5]. Local thermal cycles of low heat input and high heat input welds based on numerical simulation are shown in Fig. 4a and b. The cooling time from 800 ℃ to 500 ℃, noted as t8/5 time, is important to evaluate the cooling rate during welding. The average t8/5 in HAZ of FE simulated welding was approximately 0.62 s for low heat input, while 4.35 s for high heat input. The peak temperature selected for HAZ thermal simulation were 1100 ℃ (CGHAZ), 900 ℃ (FGHAZ), 800 ℃ (ICHAZ) and 700 ℃ (SCHAZ), respectively. Fig. 4c and d illustrate the thermal histories of each zone during HAZ thermal simulation. The

3. Local characterization 3.1. Local thermal cycles Fig. 3a presents a correspondence between the observed microstructure and the calculated temperature field of weld cross section with low heat input. The peak temperature in the FZ exceeded the boiling temperature of BM resulting in key-hole effect during laser welding. On both sides of the FZ, the inhomogeneous HAZ was divided into four zones as the descending of peak temperature, namely coarsegrained HAZ (CGHAZ), fine-grained HAZ (FGHAZ), inter-critical HAZ (ICHAZ) and sub-critical HAZ (SCHAZ) [8]. The hardness profile of low heat input weld (Fig. 3b) varied due to the local microstructure evolutions, where the FGHAZ exhibited the highest hardness and the SCHAZ was softened. Similarly, Fig. 3c and d illustrate the local 381

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Fig. 6. Microstructures of experimental and simulated (a) CGHAZ, (b) FGHAZ, (c) ICHAZ and (d) SCHAZ (LHI: low heat input; HHI: high heat input; M: martensite; F: ferrite; TM: tempered martensite).

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Fig. 9. The YS and UTS of welded joint and local zones (LHI: low heat input; HHI: high heat input).

Fig. 7. Microhardness of experimental and simulated HAZ.

Table 5 Detailed widths of each zone used in the simulation (mm).

average t8/5 of HAZ during thermal simulation was approximately 0.10 s for low heat input and 4.90 s for high heat input, indicating a rapid cooling rate and slow cooling rate were achieved respectively. 3.2. Microstructures and microhardness Microstructure of BM is presented in Fig. 5a. The martensite volume fraction was estimated to be about 57% by Image-Pro Plus. Both FZs in low heat input and high heat input welds exhibited almost fully coarse martensite, as shown in Fig. 5b and c. The disappearance of ferrite was due to the full austenization of original microstructure and rapid cooling during laser welding. However, the low heat input welds had a higher hardness (403 Hv) at FZ than that of high heat input welds (381 Hv) as shown in Fig. 3b and d. As-welded and thermal simulated microstructures of local HAZs are presented in Fig. 6. The representative microstructure of thermal simulated HAZ exhibited similar characters to that of experimental HAZ. Since the welding temperature was far above Ac3 at CGHAZ, the prior austenite grown fast from fine grain and finally became coarse grain martensite (Fig. 6a). The FGHAZ (Fig. 6b) was fully composed of martensite with much smaller grain size than that of CGHAZ. The microstructure of ICHAZ (Fig. 6c) was similar to that of BM, consisting of ferrite matrix with martensite islands. Fig. 6d depicts tempered martensite and original ferrite of SCHAZ. The pre-existing martensite decomposed and precipitated cementite. It was reported that the extent of martensite tempering depended on heating temperature and time [27]. As demonstrated in Fig. 7, the hardness distribution of thermal simulated and experimental HAZs exhibited very similar trend with small deviations. Each simulated local HAZ with low heat input had

Zone

FZ

CGHAZ

FGHAZ

ICHAZ

SCHAZ

Low heat input High heat input

0.800 1.875

0.125 0.250

0.125 0.250

0.125 0.250

0.530 0.875

Fig. 10. The numerical tensile model.

slightly higher hardness than that of the corresponding experimental HAZ. For high heat input, each simulated local HAZ had slightly lower hardness than that of the corresponding experimental HAZ. These deviations resulted from the slightly different cooling rates between laser welding and thermal simulation. The hardness variations of local HAZs mainly depend on the volume fraction and size of martensite. Both experimental and simulated CGHAZs with low heat input displayed higher hardness than that of high heat input because of different martensite sizes caused by different cooling rates. Similarly, this is also applicable to the simulated FGHAZ shown in Fig. 6b. The FGHAZ exhibited higher hardness than CGHAZ owning to the effect of reinforcement by fine grains (Fig. 7). The SCHAZ, known as softened zone, exhibited lower hardness due to tempered martensite [28]. This hardness valley was influenced by heat input, e.g., the cementite precipitated more suffi-

Fig. 8. Representative engineering stress - strain curves of (a) low heat input and (b) high heat input.

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Fig. 14. The simplified tensile model of welded joint.

ciently under higher heat input and leaded to lower hardness.

3.3. Tensile behavior of local zones and overall joint The tensile behavior of FZ and local HAZs is illustrated in Fig. 8. The high heat input welded joint presented a significantly lower fracture strain (6.2%) compared with low heat input (12.8%). This decreased ductility was resulted from the premature fracture at SCHAZ as Farabi et al. reported [29]. It should be noted that SCHAZ exhibited an obvious yielding platform, whereas the other zones presented continuous deformation behavior. The true stress-strain curves were transferred from the engineering stress-strain curves through Eq. (2) and Eq. (3),

Fig. 11. Experimental and simulated stress-strain curves of (a) low heat input and (b) high heat input welded joints.

S = (1 + ε)*σ

(2)

e=ln(1 + ε)

(3)

where S is true stress; σ is engineering stress; e is true stain; ε is engineering strain. The tensile strength of welded joints and local zones is shown in Fig. 9. The ultimate tensile strength (UTS) of the low heat input welded joint reached 99.7% of BM, and exhibited slightly higher yield strength (YS) than that of BM due to hardened FZ. In terms of the high heat input welded joint, the UTS reached 95.6% of BM with obviously increased YS. This was resulted from the softened SCHAZ. In accordance with hardness profile, the FZ, CGHAZ and FGHAZ all presented higher YS and UTS than that of BM because of a higher volume fraction of martensite at those zones [28]. Moreover, the fine grained martensite contributed to the highest YS and UTS of FGHAZ across joint, while ICHAZ had the lowest YS across joint due to the presence of soft ferrite. The SCHAZ was softened with the lowest UTS across joint, which might become the weakest zone during tensile loading. It is worth noting that each zone of low heat input welded joint exhibited higher YS and UTS than that of the corresponding zone of high heat input welded joint.

Fig. 12. Failed tensile samples and simulated strain distribution of (a) low heat input and (b) high heat input welded joints.

Fig. 13. The σe, σ1, σ2, σ3 vs. strain variation of SCHAZ and BM: (a) low heat input and (b) high heat input.

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Fig. 15. Effect of HAZ width and softening degree on the (a) UTS and (b) failure location of welded joint.

Fig. 16. Effect of FZ width on (a) UTS and failure location of joint and (b) effective stress (β=0.15, 0.4 mm-HAZ).

simulation revealed that the necking locations were in agreement with experimental observation as shown in Fig. 12. In the Mises yield criterion, the yielding occurs when

4. Tensile behavior of the welded joint 4.1. FE modeling To investigate the effect of heat input on tensile behavior of DP980 steel joint, both low and high heat input welded joints were simulated, respectively. To simplify the FE analysis of tensile behavior of butt joints, following basic assumptions were applied:

σe = σy

(4)

σe={(1/2)[(σ1 − σ2 )2 + (σ2 − σ3)2 + (σ3 − σ1)2 ]}1/2

(5)

where σy is yield strength; σe is effective stress; σ1, σ2, σ3 are principal stresses in longitudinal, width and thickness directions, respectively. For the low heat input welded joint, strain localized in BM firstly due to its lowest YS across joint. It should be noted that SCHAZ had the lowest UTS, but the joint failed at BM. During tensile test shown in Fig. 13a, the σe was almost equal to σ1 for BM because the BM was under uniaxial stress state. On the other hand, the SCHAZ was under biaxial stress state due to the constraint of hardened FZ. Thus, for the SCHAZ, the σe was lower than σ1 due to the presence of σ2 with the increasing strain. Thereby, the σe of BM reached the UTS firstly and the joint failed at BM due to the biaxial stress intensification effect [30]. For high heat input welded joint, the SCHAZ was wider with lower UTS than that of low heat input case. Further, the biaxial stress intensification effect was insufficient to counteract the HAZ softening effect. Therefore, the σe of SCHAZ reached its UTS firstly whereas the σe of BM below its UTS, as shown in Fig. 13b, and finally fracture occurred at SCHAZ.

(1) The butt-welded joint consisted of FZ, HAZ and BM, and the HAZ could be further divided into CGHAZ, FGHAZ, ICHAZ and SCHAZ. (2) The mechanical properties of each zone were uniform. Detailed widths of each zone of the joint were measured and listed in Table 5. The numerical model of tensile test is shown in Fig. 10. The calculation was performed using Abaqus 6.14 software and the model was meshed with an 8-node hexahedral element (SOLID45) with a denser mesh in these parallel zones than others. The tensile behavior was achieved by pulling one end with a displacement of 6.0 mm, while another end was fixed. 4.2. FE results and validation

4.3. Factors affecting failure location

Fig. 11 displays experimental and FE simulated stress-strain curves of welded joints. Experimental and FE simulation results all matched well in both such low heat input and high heat input welded joints. The errors of stress-strain curves were mainly from the homogeneous microstructure assumption of four local HAZs and slightly different cooling rates during thermal simulation. The strain distribution in FE

The presences of FZ and HAZ in welded joint affected the failure location during tensile test greatly. The hardened FZ could constrain plastic deformation of softened SCHAZ. This constraint effect might change the failure location from SCHAZ to BM when heat input is 385

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Acknowledgements

suitable. To evaluate factors affecting failure locations, HAZ softening degree, width of softened HAZ and width of FZ were considered using numerical simulation with a simplified model. This simplified tensile model consisted of hardened FZ with softened HAZ and BM on both sides, as shown in Fig. 14. Here, the HAZ softening degree was defined as

β=1−

p q

This work was supported by the International Science and Technology Cooperation Program of China (No. 2013DFR50590, 2015DFA51460), the National Natural Science Foundation of China (Nos. 50705050, 51605019 and 51675030). References

(6)

where β reflects the reduction of HAZ strength, and UTS of SCHAZ and BM were defined as p and q, respectively. In this study, when a FZ width is 1.2 mm, the influences of HAZ width (0.1–0.7 mm) and HAZ softening degree (0.05–0.40) on UTS and failure location of joint are shown in Fig. 15 a and b. It can be seen that UTS of the joint decreased with the increase of HAZ width, and the joint failed at HAZ when the HAZ was wide enough. Similarly, increasing of HAZ softening degree caused the reduction of the UTS of joint, where the joint was more likely to fail at HAZ rather than BM. Fig. 16a displays the influences of FZ width on constraint effect under the condition of β=0.15, 0.4 mm-HAZ. With the increase of FZ width, the UTS of joint increased gradually closing to that of BM. The failure location shifted from HAZ to BM when such FZ was wide enough. Fig. 16b presents the variations of effective stress σe across the joint in longitudinal direction (β=0.15, 0.4 mm-HAZ), which reflects the constraint effect of different FZ widths. It can be found that peak values of σe existed at the interface between the FZ and HAZ, whereas the σe was significantly reduced at HAZ. When principal stress σ1=700 MPa, the lowest σe at HAZ exhibited nearly no difference between 0.8 mm-FZ and 1.6 mm-FZ. When this principal stress σ1 increased to 860 MPa, the lowest σe of 1.6 mm-FZ at HAZ was obviously lower than that of 0.8 mm-FZ, indicating a much obvious constraint effect of 1.6 mm-FZ. As a result, the HAZ with 0.8 mm-FZ was more likely to reach its UTS and fail at HAZ, while the joint with 1.6 mm-FZ would fracture at BM due to stronger biaxial stress intensification effect.

[1] W. Xu, D. Westerbaan, S.S. Nayak, D.L. Chen, F. Goodwin, Y. Zhou, Tensile and fatigue properties of fiber laser welded high strength low alloy and DP980 dualphase steel joints, Mater. Des. 43 (2013) 373–383. [2] D. Parkes, D. Westerbaan, S.S. Nayak, Y. Zhou, F. Goodwin, S. Bhole, D.L. Chen, Tensile properties of fiber laser welded joints of high strength low alloy and dualphase steels at warm and low temperatures, Mater. Des. 56 (2014) 193–199. [3] A.P. Pierman, O. Bouaziz, T. Pardoen, P.J. Jacques, L. Brassart, The influence of microstructure and composition on the plastic behaviour of dual-phase steels, Acta Mater. 73 (2014) 298–311. [4] N. Farabi, D.L. Chen, Y. Zhou, Microstructure and mechanical properties of laser welded dissimilar DP600/DP980 dual-phase steel joints, J. Alloy. Compd. 509 (2011) 982–989. [5] J.H. Lee, S.H. Park, H.S. Kwon, G.S. Kim, C.S. Lee, Laser, tungsten inert gas, and metal active gas welding of DP780 steel: comparison of hardness, tensile properties and fatigue resistance, Mater. Des. 64 (2014) 559–565. [6] J. Ma, F. Kong, W. Liu, B. Carlson, R. Kovacevic, Study on the strength and failure modes of laser welded galvanized DP980 steel lap joints, J. Mater. Process. Technol. 214 (2014) 1696–1709. [7] A.F. Mark, J.A. Francis, H. Dai, M. Turski, P.R. Hurrell, S.K. Bate, J.R. Kornmeier, P.J. Withers, On the evolution of local material properties and residual stress in a three-pass SA508 steel weld, Acta Mater. 60 (2012) 3268–3278. [8] Q. Jia, W. Guo, P. Peng, M. Li, Y. Zhu, G. Zou, Microstructure- and strain ratedependent tensile behavior of fiber laser-welded DP980 steel joint, J. Mater. Eng. Perform. 25 (2016) 668–676. [9] K. Abdullah, P.M. Wild, J.J. Jeswiet, A. Ghasempoor, Tensile testing for weld deformation properties in similar gage tailor welded blanks using the rule of mixtures, J. Mater. Process. Technol. 112 (2001) 91–97. [10] W. Lee, K. Chung, D. Kim, J. Kim, C. Kim, K. Okamoto, R.H. Wagoner, K. Chung, Experimental and numerical study on formability of friction stir welded TWB sheets based on hemispherical dome stretch tests, Int. J. Plast. 25 (2009) 1626–1654. [11] K. Abdullah, P.M. Wild, J.J. Jeswiet, A. Ghasempoor, Tensile testing for weld deformation properties in similar gage tailor welded blanks using the rule of mixtures, J. Mater. Process. Technol. 112 (2001) 91–97. [12] X. Kong, Q. Yang, B. Li, G. Rothwell, R. English, X.J. Ren, Numerical study of strengths of spot-welded joints of steel, Mater. Des. 29 (2008) 1554–1561. [13] Y.P. Yang, S.S. Babu, F. Orth, W. Peterson, Integrated computational model to predict mechanical behaviour of spot weld, Sci. Technol. Weld. Join. 13 (2008) 232–239. [14] H. Lee, J. Choi, Overload analysis and fatigue life prediction of spot-welded specimens using an effective J-integral, Mech. Mater. 37 (2005) 19–32. [15] E.J. Pavlina, C.J. Van Tyne, Correlation of yield strength and tensile strength with hardness for steels, J. Mater. Eng. Perform. 17 (2008) 888–893. [16] J. Rojek, M. Hyrcza-Michalska, A. Bokota, W. Piekarska, Determination of mechanical properties of the weld zone in tailor-welded blanks, Arch. Civ. Mech. Eng. 12 (2012) 156–162. [17] G. Li, F. Xu, G. Sun, Q. Li, Identification of mechanical properties of the weld line by combining 3D digital image correlation with inverse modeling procedure, Int. J. Adv. Manuf. Technol. 74 (2014) 893–905. [18] C. Leitão, I. Galvão, R.M. Leal, D.M. Rodrigues, Determination of local constitutive properties of aluminium friction stir welds using digital image correlation, Mater. Des. 33 (2012) 69–74. [19] H. Lee, C. Kim, J. Song, An evaluation of global and local tensile properties of friction-stir welded DP980 dual-phase steel joints using a digital image correlation method, Materials 8 (2015) 8424–8436. [20] C.Wei, J.Zhang, S.Yang, W.Tao, F.Wu, W.Xia, Experiment-based regional characterization of HAZ mechanical properties for laser welding, The International Journal of Advanced Manufacturing Technology, 2015. [21] G.R. Goodall, J. Gianetto, J. Bowker, M. Brochu, Thermal simulation of HAZ regions in modern high strength steel, Can. Metall. Quart. 51 (2012) 58–66. [22] S. Dancette, V. Massardier-Jourdan, D. Fabrègue, J. Merlin, T. Dupuy, M. Bouzekri, HAZ microstructures and local mechanical properties of high strength steels resistance spot welds, ISIJ Int. 51 (2011) 99–107. [23] L. Chuan, Z. Jianxun, N. Jing, Numerical and experimental analysis of residual stresses in full-penetration laser beam welding of Ti6Al4V alloy, Rare Met. Mater. Eng. 38 (2009) 1317–1320. [24] X. Li, L. Wang, L. Yang, J. Wang, K. Li, Modeling of temperature field and pool formation during linear laser welding of DP1000 steel, J. Mater. Process. Technol. 214 (2014) 1844–1851. [25] M.Matsushita, Y.Kitani, R.Ikeda, K.Oi, H.Fujii, Development of friction stir welding of high strength steel sheet, Science and Technology of Welding and Joining, 2013. [26] F.K.E.F.Shanglu Yang. Dynamic simulation of temperature field in hybrid laserGTAW welding of galvanized steels in a gap-free lap joint configuration., the ASME

5. Conclusions (1) The numerical simulation of laser welding showed that the t8/5 at HAZ of low heat input and high heat input are approximately 0.62 s and 4.35 s, respectively. The microstructure and hardness of thermal simulated HAZ exhibited similar characters to that of experimental HAZ. (2) The UTS of the low heat input and high heat input welded joint reached 99.7% and 95.6% of BM, respectively. The fine grained martensite of FGHAZ contributed to the highest YS and UTS across joint, and the ICHAZ had the lowest YS. The SCHAZ was softened with the lowest UTS across joint. Each zone of low heat input exhibited higher YS and UTS than that of the corresponding zone of high heat input. (3) Numerical simulations for the tensile properties of the joints agreed well with experiments. The high heat input welded joint fractured at SCHAZ with significantly lower fracture strain. The low heat input welded joint failed at BM due to biaxial stress intensification effect. (4) The UTS of joint decreased with the increase of HAZ width, and the joint failed at SCHAZ when HAZ was wide enough. The increase of HAZ strength enhanced the UTS of joint, and such joint was more likely to fail at BM as the HAZ softening degree decreased. (5) With the increase of FZ width, the UTS of joint increased gradually closing to that of BM, and the joint was more likely to fail at BM. The joint with 1.6 mm-FZ exhibited a much obvious constraint effect than that of 0.8 mm-FZ when the tensile stress reached 860 MPa.

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[29] N. Farabi, D.L. Chen, J. Li, Y. Zhou, S.J. Dong, Microstructure and mechanical properties of laser welded DP600 steel joints, Mater. Sci. Eng.: A 527 (2010) 1215–1222. [30] Q. Jia, W. Guo, W. Li, Y. Zhu, P. Peng, G. Zou, Microstructure and tensile behavior of fiber laser-welded blanks of DP600 and DP980 steels, J. Mater. Process. Technol. 236 (2016) 73–83.

2009 International Manufacturing Science and Engineering Conference, West Lafayette, Indiana, USA, 2009. [27] T. Waterschoot, K. Verbeken, B.C. De Cooman, Tempering kinetics of the martensitic phase in DP steel, ISIJ Int. 46 (2006) 138–146. [28] M. Xia, E. Biro, Z. Tian, Y.N. Zhou, Effects of heat input and martensite on HAZ softening in laser welding of dual phase steels, ISIJ Int. 48 (2008) 809–814.

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