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Welding effect on tensile strength of grade S690Qsteel butt joint

Chen, Cheng; Chiew, Sing‑Ping; Zhao, Ming‑Shan; Lee, Chi‑King; Fung, Tat Ching

2018

Chen, C., Chiew, S.‑P., Zhao, M.‑S., Lee, C.‑K., & Fung, T. C. (2019). Welding effect on tensilestrength of grade S690Q steel butt joint. Journal of Constructional Steel Research, 153,153‑168. doi:10.1016/j.jcsr.2018.10.009

https://hdl.handle.net/10356/142816

https://doi.org/10.1016/j.jcsr.2018.10.009

© 2018 Elsevier Ltd. All rights reserved. This paper was published in Journal ofConstructional Steel Research and is made available with permission of Elsevier Ltd.

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1

Welding effect on tensile strength of grade S690Q steel butt joint

Cheng Chen a, Sing-Ping Chiew b, *, Ming-Shan Zhao b, Chi-King Lee c, Tat-Ching Fung a

a School of Civil and Environmental Engineering, Nanyang Technological University, Singapore

b Singapore Institute of Technology, Singapore

c School of Engineering and Information Technology, University of New South Wales Canberra,

Australia

ABSTRACT

In this study, welding effects on the tensile strength of high strength steel butt joints were

investigated experimentally and numerically. Two butt joints were fabricated by joining two 8

mm thick reheated, quenched and tempered S690Q high strength steel (HSS) plates together

using Shielded Metal Arc Welding with different welding heat inputs. In the experimental

study, post-welding microstructure and micro-hardness tests performed for the butt joints

confirmed the existence of a soft layer within the heat-affected zone. The size of this soft layer

increased when higher heat input was adopted. Subsequent coupon tensile tests showed that

the tensile strength of the butt joints deteriorated with the increase of the welding heat input.

In numerical study, a finite element (FE) simulation considering the non-uniform material

properties heat-affected zone was implemented to complement the experimental study.

Comparisons between experimental and numerical results were carried out for the temperature

history curves, failure modes and stress-strain curves for the coupons manufactured from the

butt joints. It was found that the proposed simulation method could reflect the influence of

welding process on the tensile strength of the S690Q HSS butt joints tested in this study.

Keywords: High strength steel, Butt joint, Tensile strength, Vickers hardness test,

Microstructure

1. INTRODUCTION

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Structural steel has a pivotal position in civil engineering. It is widely employed in the

constructions of high-rise buildings, bridges, stadiums and offshore structures due to its high

strength-to-weight ratio, easy prefabrication properties and attractive appearance. Currently,

most of the existed steel structures are constructed by using normal strength steel with nominal

yield strength up to 460 MPa. However, with higher requirements in architectural and structural

design, normal strength steel is not always capable of delivering the most sustainable and

economic solution in many applications in which the self-weight of the structure is critical [1].

As a result, high strength steel (HSS) with yield stress higher than 460 MPa recently has

attracted much attention due to its remarkable strength to weight ratio and its high potential to

offer a lower combined material and fabrication costs due to the materials and weight reduction.

Owing to its good weldability, the reheated, quenched and tempered (RQT) low alloy steel is

one of the most favourable types of HSS and it has been used in many structural engineering

projects. Since it is easier to achieve uniform material properties during the manufacturing

process, RQT HSS is still mainly available in the plate form. As a result, HSS plates are often

welded to form built-up sections and welding is unavoidable in the fabrication of all HSS

sections in which severe temperature change is generated around the weld area. Towards this

end, as the RQT hardening process employed during the manufacturing of HSS makes it more

susceptible [2, 3] to high heat input than normal strength steel and its mechanical properties

are often adversely affected by welding [4, 5], there are concerns about the mechanical

performance of welded RQT HSS components.

Many experimental and numerical investigations showed that the welding thermal cycle often

changes the microstructure of RQT HSS. In particular, a fine grain heat affected zone (FGHAZ)

[6] will be created. This FGHAZ is often referred as a “soft layer” due to its inferior mechanical

properties compared with the base material. In addition, higher heat input is generally

accompanied by wider and worsen mechanical properties in the HAZ [7-11]. Meanwhile, it

was shown that the strength and size of the soft layer has influence on the strength of RQT HSS

joints [12]. By assuming a constant strength of the entire HAZ, numerical analysis predicted

that the tensile strength of HSS butt joints would decrease with the increase of the width of

HAZ. Hochhauser and Rauch [13] found that the width of soft layer ranged from 0.33 to 0.6

times of the specimen thickness, and the reduction of tensile strength of butt joint was 3 % to

8 % of its original strength. When conducting numerical simulation of the welding process,

one of the most complicated modelling parameters is the material property of the HAZ. In

general, due to the complex welding thermal cycles and non-uniform cooling rate, the material

properties within the HAZ are highly non-uniform. In many numerical investigations, the non-

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uniform HAZ is simplified as a uniform area [14]. The size of the HAZ is estimated by using

hardness test, and simplified material properties would be then assigned [15,16] during

numerical modelling. The width of HAZ was increased in simulation when a higher welding

heat input was employed during the welding of the butt joints. Obviously, such simplified

simulation method cannot fully describe the material properties variation within the HAZ and

thus may not be able to fully reflect the influence of welding on mechanical behaviour of HSS

butt joint accurately.

The main objective of this paper is to investigate the influence of welding on the mechanical

performance of HSS butt joints experimentally and numerically. In the experimental study, two

joints were formed by 8 mm thick S690Q HSS plates by using Shielded Metal Arc Welding

(SMAW) with different welding heat inputs. After the joints were fabricated, coupons for direct

tensile test were carefully cut out from the welded joints. Microstructure and micro-hardness

tests were first employed to identify the extends and the mechanical properties of the soft layer.

Direct tensile tests of the coupons were then conducted to find out the impacts of welding on

the tensile behaviour of the butt joints. During the numerical analysis, a new simulation

procedure for joint strength prediction was proposed and its accuracy was validated by

comparing the modelling results with the tensile tests results. Furthermore, a preliminary

parametric study was conducted to investigate the impacts of different welding parameters on

the tensile strength of S690Q HSS butt joints with different plate thicknesses.

2. EXPERIMENTAL STUDY

2.1 Material and joint fabrication

The S690Q HSS employed in this study has a nominal yield strength of 690 MPa and a tensile

strength between 790 MPa and 930 MPa. These plates were manufactured by a refined

quenching and tempering technology for the formation of tempered martensite. In general, they

exhibit better homogeneity in through-thickness mechanical properties compared with

conventional quenched and tempered steel plates. The steel plates comply with EN 10025-6

grade S690Q specification [17] and is equivalent to the ASTM A514 steel [18].

The weld material LB-80L was used as filler for welding and it satisfies the AWS A5.5

E11018-G [19] filler metal specification. The mechanical properties of the S690Q HSS and

filler weld metal are listed in Table 1. The corresponding chemical composition and carbon

equivalent (CE) are listed in Table 2. It can be seen that the strength of weld metal matches

with that of S690Q HSS. Moreover, the HSS plates used have similar chemical composition

when compared with normal strength mild steel. Therefore, it has excellent weldability,

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especially in the sense of reduced preheating requirements and reduced susceptibility to cold

cracking [20,21].

Table 1. Mechanical properties of S690 HSS and weld metal

Material Elastic modulus

(GPa)

Yield strength

(MPa)

Ultimate tensile strength

(MPa)

Elongation

(%)

S690Q HSS 209 745 838 14.5

LB-80L 210 767 848 22.0

Table 2. Chemical composition of S690 HSS and weld metal

Elements C Si Mn P S Ni Mo CE

S690Q HSS 0.14 0.40 1.35 0.012 0.003 0.01 0.12 0.40

LB-80L 0.05 0.56 1.37 0.009 0.005 2.77 0.76 0.33

In this study, two butt joints were fabricated using 8 mm thick S690Q HSS plates by SMAW.

Two different electrodes with diameter of 3.2 mm and 5.0 mm were utilized to achieve different

welding heat inputs. The two butt joint specimens were respectively named as BJ-3.2 and BJ-

5.0 for 3.2 mm and 5.0 mm electrodes welding. The temperature history near the weld was

recorded by two thermocouples. The plate configuration, thermocouple positions and a welded

joint are shown in Figure 1. The distance between the closest thermocouple and the top edge

of the weld is denoted as d=3 mm for joint BJ-3.2. During welding of BJ-3.2, the measured

peak temperature is 1139.4 ºC which is very close to the measurement limit (1200 ºC) of the

thermocouple used. As it is expected that during the welding of BJ-5.0 higher heat input will

be used, it was worried that the peak temperature at the same position may exceed 1200 ºC

which could damage the thermocouple. Hence, during the welding of BJ-5.0, d was increased

to 5 mm. Eventually, a peak temperature of 851.5 ºC was recorded at the thermocouple for joint

BJ-5.0 with d=5mm. The t8/5 cooling time (time for the joint to cool down from 800 ºC to 500

ºC) was 21.9 s and 37.4 s for joint BJ-3.2 and joint BJ-5.0, respectively. The welding was

carried out by an experienced qualified welder who was asked to keep the welding speed as

constant as possible. However, there were still some small differences in welding speed for

different welding passes. The start and finish time of each welding pass were recorded during

welding and the average welding speed was obtained by dividing the welding length by the

welding time. Subsequent tensile test results revealed that the mechanical behaviours of the

two specimens are almost the same. Hence, it can be deduced that similar welding heat input

and welding speed are achieved at the two positions.

It should be noted that while the welding speeds of the 3.2 mm and 5 mm electrode was similar,

the time of the peak temperature was more than 300 s in Figure 2a but less than 200 s in Figure

5

2b. It was because during the welding of BJ-3.2, one electrode was used up after the first

welding pass and it took time for the welder to install a new electrode before the second pass

was started. While for the welding of BJ-5.0, the second pass was started immediately after the

first pass was completed.

During welding, thermal deformation was limited by three clips, and two passes were used for

both joints. The corresponding welding parameters for each pass are listed in Table 3. Note

that the heat inputs listed in Table 3 are calculated according to EN 1011-1[22]. In addition,

the worker welded the joints by traditional welding method used for mild steel, and the welding

heat inputs were not controlled specially for S690Q HSS.

Table 3. Welding parameters of SMAW

Specimen

Electrode

diameter

(mm)

Voltage

(V)

Current

(A)

Welding

speed

pass 1

(mm/s)

Welding

speed

pass 2

(mm/s)

Heat input

pass 1

(kJ/mm)

Heat input

pass 2

(kJ/mm)

BJ-3.2 3.2 35.0 80 1.55 1.60 1.45 1.40

BJ-5.0 5.0 35.0 180 1.50 1.94 3.36 2.60

After the welding was completed, two coupons were cut out from each butt joint for direct

tensile test. Addition weldment covering on the joint welds was first removed to achieve a

uniform cross section of the coupons. In addition, as illustrated in Figure 3, a small block was

cut out from the welding zone for micro-hardness test and microstructure and scanning electron

microscope (SEM) observations. The small block was divided into two parts along the centre

line of the weld and then casted with epoxy into metallurgical specimens. Four metallugrical

specimens were prepared and polished. The surface is finally etched with 2 % nital solution in

order to observe the microstructure clearly.

2.2 Metallurgical investigation

2.2.1 Microstructure study

S690Q HSS mainly consists of martensite which is tempered after hardening treatment at

temperatures below the transformation point A1 (about 730 ºC for steel with carbon content of

0.15 %). The microstructure of S690Q HSS is shown in Figure 4. Theoretically, if martensite

is exposed to temperature between A1 and A3 and then cooled subsequently (similar to the

situation happens during welding), it would be transformed into ferrite and austenite and finally

decomposes into other microstructures with lower hardness such as ferrite, pearlite or cementite.

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Metallurgical examinations were designed to expose such microstructural changes in the HAZ

of the butt joints by light optical microscopy and SEM.

Depending of microstructural appearance, the HAZ of the welded butt joints can be divided to

three sub-HAZs, namely the coarse-grain HAZ (CGHAZ), the fine-grain HAZ (FGHAZ) and

the tempering zone (Figure 5a) [23]. The CGHAZ is formed in the region where a temperature

ranging from A3 (about 830 ºC for steel with carbon content 0.15 %) up to the melting point

(about 1500 ºC) during welding. While, FGHAZ is generally formed at area where the

temperature is higher than the transformation temperature A1 but lower than the temperature

A3. The tempering zone is further away from the weld material, and there is nearly no phase

transformation that has taken place during welding in this area. Figure 5 shows the typical

microstructure distributions of the weld material (Figures 5b and 5c), the HAZ with both the

CGHAZ (Figures 5d and 5e) and the FGHAZ (Figure 5f and 5g). Figure 5d and 5e show that

the microstructure of the CGHAZ is granular bainite. While in the FGHAZ, the microstructure

changes from martensite to ferrite and cementite which has lower hardness compared with

martensite. The microstructure of tempering zone obtained by SEM test is shown in Figure 6,

and it can be seen that some martensite decomposes to ferrite.

2.2.2 Micro-hardness test

Vickers hardness measurement using 500 g force was carried out by using a micro-hardness

tester according to ISO 6507-1 [24]. As shown in Figure 7, three lines of indentations were

made at different thickness of the plate for the cut-out block. Note that the 4mm indentation

line is at the mid-thickness of the block. The interval between two successive indentations is

set as 0.5 mm. As the width of cut out block is 30 mm which is longer than the welding width,

the hardness tests were conducted on the weld material, the CGHAZ, the FGHAZ and the base

material successively.

Figure 8 shows the hardness distributions for joints BJ-3.2 and BJ-5.0. It can be seen that the

hardness results are highly symmetric about the weld centre line. For both joints, two soft layers

with lower hardness are located in the HAZs. In the FGHAZ, the minimum hardness value

(177 HV0.5) is 36.4 % lower than that of the base material, and the minimum hardness value

for both joints are similar. This suggested that the minimum hardness value is almost not

changed with the increase of the welding heat input.

Based on the hardness test, the estimated HAZ size of BJ-3.2 and BJ-5.0 were drawn according

to hardness test, as shown in Figure 9. The plate bevel of the butt joints is confidently

confirmed as one boundary between weld and HAZ, the boundary between HAZ and base

material was established by a three-point-curve. The three points are located in three

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indentation layers separately and corresponding to a hardness value of 280 HV0.5. For joint

BJ-3.2, the average width of the HAZ is found to be 4.35 mm. When the welding heat input

was increased, the average width of HAZ in BJ-5.0 was increased to 7.75 mm.

2.3 Coupon tensile tests

The tensile behaviour of the S690Q HSS butt joints was investigated by direct tensile tests

using the 5900 series universal testing instruments (Figure 10) according to EN 10002-1 [25].

The gauge length of specimen is 50 mm, and the loading rate was set as 0.5mm/min until the

fracture has taken place. The extensometer was employed to capture the deformation of the

specimens. Total four specimens are manufactured by using the S690Q HSS welded butt joints

and were tested in tension.

Eventually, it was found that all the tested butt joints coupons fractured within the HAZ. The

stress-strain curves obtained are shown in Figure 11. The mechanical properties of the coupons

including elastic modulus, the yield strength (taken as the 0.2 % strain offset strengths), the

tensile strength and the fracture strain, are summarized in Table 4.

Table 4. Mechanical properties of butt joint coupons

Coupon Elastic modulus

(GPa)

Yield strength

(MPa)

Ultimate tensile strength

(MPa)

Elongation at

fracture (%)

S690Q HSS 209 745 838 14.5

BJ-3.2-1 206 672 809 6.2

BJ-3.2-2 209 667 802 6.6

BJ-5.0-1 207 569 750 6.5

BJ-5.0-2 203 562 745 6.5

From the coupon test results, it can be seen that both the yield strength and tensile strength of

the S690Q HSS butt joints tested in this study was decreased as the welding heat input

increased. Table 4 also indicates that the mean values and ranges of yield strength and ultimate

tensile strength for BJ-3.2 are 669 ± 3 MPa and 806 ± 4 MPa, respectively. For BJ-5.0, the

mean values and ranges of yield strength and ultimate tensile strength are 566 ± 4 MPa and

748 ± 3 MPa, respectively. The yield strength of BJ-3.2 decreased by 10.2 % when compared

with that of the base material, and the corresponding tensile strength decreased by 3.8 %. For

the joint BJ-5.0 in which higher heat input was used during the welding, the yield strength and

tensile strength was reduced by 24.1 % and 10.8 %, respectively. Furthermore, the elongation

of the butt joints was reduced by 55.5 % after welding for both joints BJ-3.2 and BJ-5.0. One

reason for the serious deterioration of elongation is the high concentration of the plastic strain

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caused by high heat input and few passes adopted during the welding of the joint [11]. In

addition, due to the existence of the “soft layer”, the deformation of S690Q HSS butt joint was

localized within the HAZ instead of uniformly distributed over the whole gauge length. As a

result, while the strain within the HAZ of the HSS butt joint is larger than the strain within the

HAZ of the specimen with uniform material property, a smaller final deformation overall the

gauge length was recorded.

Based on the EN ISO 18265 [26] guideline for the conversion of hardness-to-tensile strength

value for quenching and tempering steels, a hardness of 280 HV0.5 is corresponding to a tensile

strength of 876 ± 36 MPa. This is within 4 % error when compared with the tensile strength

(838 MPa) obtained from the coupon test. This means that the hardness to strength conversion

value indicated by EN ISO 18265 is reasonable and reliable for S690Q HSS used in this study.

3. NUMERICAL STUDY

Finite element (FE) simulation has been frequently utilized for welding modelling and residual

stress study of welded joints [27-32] as it requires much less resource to conduct with compared

with experimental study. Many FE models used in previous studies, material properties of steel

were assumed to be deteriorated during the heating stage but then recover back to original

values after the welding parts are cooled down to room temperature. While such assumption is

applicable to the welding of normal strength steel joint, it is not applicable for the case of

S690Q HSS joint due to the existence of the soft zone in the HAZ. For those numerical models

in which the effect of soft HAZ is considered, the whole HAZ was often treated as a single

material with homogenous material properties [14]. While the width of HAZ was increased to

reflect the effects of higher welding heat input on the joints, such simplified simulation method

may not truly describe the shape and material properties of the HAZ, and therefore cannot fully

reflect the influence of welding on mechanical behaviour of HSS butt joint.

In view of the deficiencies of previous simulation approaches, in this study a new modelling

procedure is proposed. The new approach adopted the sequentially coupled thermal stress

analysis method in ABAQUS. This sequential coupled thermal stress analysis consists of

alternative rounds of thermal analysis and stress analysis. Thermal cycles of butt joints during

welding are determined by the thermal analysis model with the help of a user-defined

subroutine DFLUX where double-ellipsoid heat source is defined. The obtained temperature

history of butt joint is then imported into the stress analysis model where two other user-defined

subroutines UEXPAN and USDFLD [33] are employed to define the deteriorated material

9

properties within the HAZ based on the imported temperature history. The tensile behaviour of

S690Q HSS butt joints is then finally analysed after the joints cool down to room temperature.

3.1 The FE models used

In the numerical simulation, two 3D solid finite element models, namely the Num-BJ-3.2 and

the Num-BJ-5.0, were created according to the dimensions of the butt joints BJ-3.2 and BJ-5.0,

respectively. Noted that during the welding modelling, welding materials was added according

to the two passes employed in according to the actual welding details, as presented in Figure

12. In addition, the mesh for the coupon tensile test was embedded within the plate and welding

model to facilitate subsequent data transfer for the coupon tension test modelling.

In the thermal analysis model, double-ellipsoid heat source, described as in Equations (1) and

(2), is defined with the help of the user defined subroutine DFLUX [33].

For the front part of heat source

𝑞(𝑥, 𝑦, 𝑧, 𝑡) =6√3𝑄𝑓1

𝑎𝑏𝑐1𝜋√𝜋e

−3x2

a2 −3y2

b2 −3z2

c12

(1)

For the rear part of heat source

𝑞(𝑥, 𝑦, 𝑧, 𝑡) =6√3𝑄𝑓2

𝑎𝑏𝑐2𝜋√𝜋e

−3x2

a2 −3y2

b2 −3z2

c22

(2)

𝑄 = 𝜂𝑈𝐼 (3)

In Equations (1) and (2), f1 and f2 are fractions of deposited heat, and represent the heat

apportionments of heat flux in the front and rear quadrants, respectively. Since more energy is

applied in the front part of heat source, f1 and f2 are respectively defined as 1.33 and 0.67, so

that f1+f2=2. a is the width of semi-axes of heat source and is equal to 4.4 mm and 8.1 mm for

the first and second welding pass, respectively. b is the depth semi-axes of heat source and is

equal to 5.0 mm and 3.0 mm for the first and second welding pass, respectively. c1 and c2 are

the front and rear semi-axes of heat source, respectively. They are assumed to be the same for

the two welding passes and equal to 4.4 mm and 8.8 mm, respectively. The welding speed is

defined based on the measurement in tests. The welding heat input Q is determined by Equation

(3). The value of voltage U and current I are determined according to test values listed in Table

3. Thermal efficient factor 𝜂 is assigned with the value of 0.8 according to EN BS. 1011-1[22].

There are 23996 elements in thermal analysis model. The mesh for the welding materials and

its adjacent areas are refined to capture the steep temperature gradient during welding, and

transition elements are employed to link the coarse and fine mesh zones. The eight-node linear

heat transfer brick (DC3D8) is employed as the FE type for modelling. Regarding thermal

10

boundary conditions, convection and radiation surfaces were assigned to model with film

coefficient of 100 W/(m2˚C) and an emissivity coefficient 0.8 was adopted. The birth and death

element approach [31, 32] was used to simulate the generation of welds, and the time

interruption between two welding passes were also considered and based on the recorded

temperature history in actual welding (Figure 2). The temperature history of thermal analysis

model was then imported into stress analysis as a predefined field.

For the direct tensile test stress analysis model, as the FE mesh for the coupon was embedded

within the mesh employed for welding simulation, the nodal temperature history obtained

during thermal analysis model can be directly assigned to the nodes of the coupon mesh. After

the temperature history data was imported, the tensile test process was simulated as standard

non-linear material and large deformation quasi-static analysis to predict the mechanical

performance of butt joint coupon. The C3D8 element which is an 8-node linear brick was

employed for coupon test modelling, and the physical boundary conditions of numerical

models are determined according to the experiment set up, as shown in Figure 13.

3.2 Material properties

During the sequential coupled thermal stress analysis of the welding process, both the thermal

properties and mechanical properties for the S690Q HSS were defined as functions of the

temperature. The thermal properties of S690Q HSS consist of density, thermal conductivity,

specific heat and latent heat. The density is assumed to be 7850 kg/m3. The latent heat of HSS

is defined as 300 kJ/kg with liquidus temperature 1480 ˚C and solidus temperature 1430 ˚C.

The specific heat ca and conductivity a are determined according to the equations given in

EC3: part 1-2 [34].

The specific heat of steel ca (J/(kg˚C)) is determined by the following equations from EC3 [34]:

For 20 ℃ ≤ 𝑇 < 600 ℃;

𝑐𝑎 = 425 + 7.73 × 10−1𝑇 − 1.69 × 10−3𝑇2 + 2.22 × 10−6𝑇3 (4)

For 600 ℃ ≤ 𝑇 < 735 ℃;

𝑐𝑎 = 666 +13002

738−𝑇 (5)

For 735 ℃ ≤ 𝑇 < 900 ℃;

𝑐𝑎 = 545 +17820

𝑇−731 (6)

For 900 ℃ ≤ 𝑇 ≤ 1200 ℃;

𝑐𝑎 = 650 (7)

The thermal conductivity of steel a (W/(m˚C)) is defined from the following equations [34]:

11

For 20 ℃ ≤ 𝑇 < 800 ℃;

𝜆𝑎 = 54 − 3.3310−2𝑇 (8)

For 800 ℃ ≤ 𝑇 ≤ 1200 ℃;

𝜆𝑎 = 27.3 (9)

As one of the mechanical properties, the expansion coefficient T (mechanical property) is also

derived from EC3: part 1-2 [34] and defined by the user-defined subroutine UEXPAN during

stress analysis as follows:

For 20°C ≤ T < 750°C;

𝜀T = 1.2 × 10−5T + 0.4 × 10−8T2 − 2.416 × 10−4 (10)

For 750°C ≤ T ≤ 860°C;

𝜀T = 1.2 × 10−2 (11)

For 750°C < T ≤ 1200°C;

𝜀T = 2 × 10−5T − 6.2 × 10−3 (12)

For the most important mechanical property, the stress-strain curves at elevated temperatures

and after cooling down from different elevated temperatures are defined based on experimental

results [3], as shown in Figure 14a. During the heating stage of welding, the stress-strain

curves at elevated temperatures are determined for every node depending on its temperature.

After the welding is completed and butt joint is cooled down, the nodal material properties are

adjusted by another user-defined subroutine USDFLD. The nodal material properties are

determined by referring to the post heating stress-strain curves (Figure 14b) obtained from

Zhao et al. [4] for different peak temperature attained during welding.

It should be noted that in the FE model for the direct tensile test analysis, the post-heating

stress-strain curves shown in Figure 14b, cannot be directly used to retrieve material properties

of the HAZ and further interpolations are still necessary. The reason is that as additional

weldment material is added during the welding and these addition materials will eventually

achieve an ultimate tensile strength and hardness compatible with the S690Q HSS.

From Figure 8, the minimum hardness in the HAZ is 177 HV0.5 and it is corresponding a

tensile strength of 566 MPa [26]. As the tensile strength of S690Q HSS is 552 MPa after

exposure to 800 °C (Figure 14b), this implies that the minimum hardness location will

experienced a peak temperature of approximately equal to (but slightly higher than) 800 °C

Hence, it is defined in the subroutine USDFLD that the material properties of S690Q HSS at

12

the minimum hardness location will follow the post-heating stress-strain curves corresponding

to the 800 °C curve shown in Figure 14b.

For the region between the minimum hardness location and the outer boundary of the HAZ

where the maximum temperature reached during welding is only 400 °C and the hardness of

the HSS (280 HV0.5) is not affected by the heat input during welding, the material properties

of the S690Q HSS will be calculated by linearly interpolating the post-heating stress-strain

curves between 400 °C and 800 °C based on the hardness value of the materials obtained from

the hardness test.

For the region between the minimum hardness location and the inner boundary of the HAZ, it

is assumed that 1350 °C is the maximum temperature reached for the CGHAZ [10-11]. As the

maximum temperature reached in this region should be between 800 °C and 1350 °C and an

increase in material hardness is observed from the hardness test (Figure 8), this indicated that

the mechanical properties of materials in this region must be improved from the minimum

hardness location the inner boundary of HAZ. As a result, in order to reflect such an

improvement mechanical property in this region, the material property was assumed to increase

linearly with the increase of hardness so that at the 1350 °C location, the mechanical properties

of the materials are the same as the base metal HSS materials properties.

Finally, for the region between the weld centre and the CGHAZ where the weldment material

(LB-80L) is deposited during welding, since the material properties of the weldment after

cooling are similar to the HSS (they have very similar hardness value), the base metal HSS

materials properties are assigned to materials in this region. The mechanical properties of HAZ

are indicated in term of hardness as shown in Figure 15.

3.3 Verification of numerical model

The proposed simulation method is verified by comparing the model predictions with the

experimental results in terms of the temperature history records, the failure mode of the coupon

tension test and the stress-strain curves of butt joints during direct tensile test. Figure 16 shows

the typical temperature distribution of the butt joint during welding, and the temperature history

is extracted and compared with the test data in Figure 17. The numerical results are generally

in good agreement with the experimental data especially in the peak temperature and cooling

rate. However, it should be noted that the difference between measured and simulated

temperature increase slighting with the distance from the welding. The reason is that the

thermal material properties and thermal boundary conditions used in the modelling were based

on the suggestions by EC3. While such suggestions are reasonable and realistic, there are

13

inevitability some differences between them and the actual conditions and values of the HSS

used. Therefore, it is inevitable that the temperature field predicted by the model is not exactly

the same as the actual measurement. Obviously, such modelling error will accumulate and

increase as the distance from the welding position (i.e. heat source) or welding time increase

and results in high errors as shown in Fig. 18.

The failure models of S690Q HSS butt joints were also obtained by FE analysis and are shown

in Figure 18. Same as test results, the FE model predicted that the coupon was failed by

necking and fracture at one of the HAZ. Besides the failure mode, the stress-strain curves

obtained from FE modelling shown reasonably good consistency with the corresponding test

results, as shown in Figure 19. The yield stress fy and ultimate tensile strength fu of tested

coupons and numerical models are compared and listed in Table 5. It is found that the proposed

simulation method can predict the tensile strength of S690Q HSS butt joint precisely. The

errors for the ultimate tensile strength are only 3.94 % and 2.50 % for BJ-3.2 and BJ-5.0,

respectively. Regarding the yield strength, the difference between finite element model and test

results is slight greater. The error of yield strength are 5.04 % and 16.5 % for joint BJ-3.2 and

joint BJ-5.0, respectively. Hence, it can be concluded that the proposed simulation method can

be predict the influence of welding on the tensile performance of S690 HSS butt joint

reasonably well.

Table 5. Comparison of mechanical properties for butt joints BJ-3.2 Num-BJ-3.2 Error (%) BJ-5.0 Num-BJ-5.0 Error (%)

fy (MPa) 669 703 5 565 660 16

fu (MPa) 806 774 4 745 726 3

3.4 Parametric study

Based on the validated simulation model, a small scale parametric study is carried out to

investigate the effect of welding on butt joints with different plate thickness. FE models for

four additional butt joints with 12 mm and 16 mm plate thickness are built. For each model,

two different weld heat inputs of 1.425 kJ/mm and 2.98 kJ/mm per welding pass are employed.

Note that 1.425 kJ/mm is the average heat input recorded of joint BJ-3.2. 2.98 kJ/mm is the

average heat input of joint BJ-5.0 with the average heat input defined as

𝐻𝑒𝑎𝑡 𝑖𝑛𝑝𝑢𝑡 𝑜𝑓 𝑝𝑎𝑠𝑠 1+ℎ𝑒𝑎𝑡 𝑖𝑛𝑝𝑢𝑡 𝑜𝑓 𝑝𝑎𝑠𝑠 2

2. The welding pass configurations for 12 mm and 16 mm

butt joints are presented in Figure 20. There are four welding passes for 12 mm joints, and

14

seven passes for 16 mm joints. The material property definition is the same as the validated

model.

For all the four numerical models created, it is found that the failure mode is the same as the 8

mm butt joints tested. That is, the failure positions are all located at the soft HAZ layer. The

characteristic strength of modelled butt joints, including yield strength and tensile strength, are

summarized in Table 6. In addition, for each model, the yield strength reduction factor

Ry(𝑌𝑖𝑒𝑙𝑑 𝑠𝑡𝑟𝑒𝑛𝑔𝑡ℎ 𝑜𝑓 𝑏𝑢𝑡𝑡 𝑗𝑜𝑖𝑛𝑡

𝑌𝑖𝑒𝑙𝑑 𝑠𝑡𝑟𝑒𝑛𝑔𝑡ℎ 𝑜𝑓 𝑆690𝑄 𝑏𝑎𝑠𝑒 𝑚𝑒𝑡𝑎𝑙) and the tensile strength reduction factor

Ru(𝑇𝑒𝑛𝑠𝑖𝑙𝑒 𝑠𝑡𝑟𝑒𝑛𝑔𝑡ℎ 𝑜𝑓 𝑏𝑢𝑡𝑡 𝑗𝑜𝑖𝑛𝑡

𝑇𝑒𝑛𝑠𝑖𝑙𝑒 𝑠𝑡𝑟𝑒𝑛𝑔𝑡ℎ 𝑜𝑓 𝑅𝑄𝑇 𝑆690) are calculated and their values are summarized in Figure 21.

Table 6. Characteristic strength of butt joints predicted by numerical model

Thickness

(mm) 8 12 16

Heat input

(kJ/mm) 1.425 2.98 1.425 2.98 1.425 2.98

fy (MPa) 703 660 721 692 723 704

fu (MPa) 774 726 800 772 804 779

Note: For the base metal S690Q HSS fy=745 MPa and fu=838 MPa

From the reduction factor curves shown in Figure 21, the modelling results suggested that for

a given welding heat input, the influence of welding on the strength of butt joint will be reduced

with the increase of butt joint thickness. When the plate thickness was increased from 8 mm to

16 mm, the Ry and Ru were respectively increased from 0.914 to 0.940 and from 0.910 to 0.946

for a heat input of 1.425 kJ/mm. For the heat input of 2.98 kJ/mm, the Ry and Ru were increased

from 0.858 to 0.916 and from 0.855 to 0.917, respectively. These results could be caused by

the higher heat conduction effect for a thicker plate. For a given welding heat input per pass,

when the plate thickness is increased, as the weld pass size relative to the plate thickness is

reduced, a higher portion of heat lost will be contributed by conduction. As a result, the cooling

rate near the HAZ is also increased and the area affected by high temperature is smaller

compared with a thinner plate. Eventually, the size of the soft FGHAZ becomes smaller and

results in better performance.

4. CONCLUSIONS

This paper presents a study on the influence of welding on mechanical performance of HSS

butt joints by experimental test and numerically analysis. In the experimental study, two joints

are fabricated by welding together two 8 mm thick reheated, quenched and tempered S690Q

high strength steel (HSS) plate by shielded metal arc welding (SMAW) with different welding

heat inputs per welding pass. Microstructure and micro-hardness tests for butt joints confirmed

15

that the welding heat input generated a weaker and soft layer of fine grain heat affected zone

(FGHAZ) due to microstructure changing from tempered martensite to ferrite and cementite.

The existence of the soft layer directly impairs tensile strength of the butt joints, and the

deterioration becomes more serious with increase of welding heat input as it will increase the

size of the FGHAZ. Subsequent tensile test of the coupon showed that this weaker HAZ as the

failure location of the joint under tension. However, the exact fracture location within the HAZ

was not precisely determined in this study and therefore a more detailed examination of the

exact facture location within the HAZ could be a good topic for future research. For the 8 mm

S690Q HSS butt joints tested, it is found the lowest hardness of soft layer FGHAZ was not

sensitive to the increase of welding heat input per pass. In the numerical analysis part, a new

simulation is proposed to predict the direct tensile strength of the welding coupon obtained

from the experimental study. The accuracy of the proposed model is then validated by

comparing with the test results. Furthermore, by using the validated model, a small scale

parametric study is conducted to investigate the effect of welding on the tensile strength of

S690Q HSS butt joints with different plate thickness. The parametric study eventually showed

that if the heat input per pass is kept constant, the joint strength reduction in direct tensile test

is reduced as the butt joint thickness is increased.

16

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18

Figure 1. Plate configuration and thermocouple positions

(a) Temperature history of BJ-3.2

(b) Temperature history of BJ-5.0

Figure 2. Temperature history of butt joints during welding

0

200

400

600

800

1000

1200

1400

0 100 200 300 400 500 600 700 800 900 1000 1100 1200

Tem

per

ature

(°C

)

Time (s)

Position 1

Position 2

0

200

400

600

800

1000

0 100 200 300 400 500 600 700 800 900 1000 1100 1200

Tem

per

ature

(°C

)

Time (s)

Position 1Position 2

19

Figure 3 Dimension of coupons and cut out block for hardness, micro-hardness and tensile tests

30

20

20

20

20

91

29

11

0

20

20

20

20

11

0

61.1 8.9 60 8.9 61.1

200

R12.5

R12.5

R12.5

R12

.5

12

.5 ±

0.2

60.061.1 61.1

200.0

8

20

.0

R12.5 R

12.5

8.98.9

8.9 61.18.961.1

8

200.0

3 28.528.5

60°

Unit: mm

20

(a) light optical microscopy image

(b) SEM image

Figure 4. Microstructure of S690Q HSS

21

22

(a) light optical microscopy image of HAZ

(b) light optical microscopy image of weld material

23

(c) SEM image of weld material

(d) light optical microscopy image of CGHAZ

24

(e) SEM image of CGHAZ

(f) light optical microscopy image of FGHAZ

25

(g) SEM image of FGHAZ

Figure 5. Microstructures of weld material, CGHAZ and FGHAZ

Figure 6. SEM image showing the microstructure of tempering zone

26

Figure 7. Indentation points distribution

(a) hardness variation along 1 mm indentation line for BJ-3.2

(b) hardness variation along 4 mm indentation line for BJ-3.2

150

180

210

240

270

300

330

360

-15 -12 -9 -6 -3 0 3 6 9 12 15

Hard

nes

s (H

V0

.5)

Distance (mm)

BJ-3.2-1mm

WeldHAZ HAZBase material Base material

150

180

210

240

270

300

330

360

-15 -12 -9 -6 -3 0 3 6 9 12 15

Ha

rdn

ess

(HV

0.5

)

Distance (mm)

BJ-3.2-4mmWeldHAZ HAZBase material Base material

27

(c) Hardness variation along 7 mm indentation line for BJ-3.2

(d) Hardness variation along 1 mm indentation line for BJ-5.0

150

180

210

240

270

300

330

360

-15 -12 -9 -6 -3 0 3 6 9 12 15

Ha

rdn

ess

(HV

0.5

)

Distance (mm)

BJ-3.2-7mmWeldHAZ HAZBase material Base material

150

180

210

240

270

300

330

360

-15 -12 -9 -6 -3 0 3 6 9 12 15

Hard

nes

s (H

V0.5

)

Distance (mm)

BJ-5.0-1mmWeldHAZ HAZ

Base

material

Base

material

28

(e) Hardness variation along 4 mm indentation line for BJ-5.0

(f) Hardness variation along 7 mm indentation line for BJ-5.0

Figure 8. Measured hardness values of two butt joints

150

180

210

240

270

300

330

360

-15 -12 -9 -6 -3 0 3 6 9 12 15

Ha

rdn

ess

(HV

0.5

)

Distance (mm)

BJ-5.0-4mmWeldHAZ HAZ

Base

material

Base

material

150

180

210

240

270

300

330

360

-15 -12 -9 -6 -3 0 3 6 9 12 15

Hard

nes

s (H

V0.5

)

Distance (mm)

BJ-5.0-7mmWeldHAZ HAZ

Base

material

Base

material

3 5.6

3.1

5.6

12.23.1

HAZ WM BM

12.26.2 6.2

39.3 9.3

BJ-3.2 BJ-5.0

8 8

Unit:mm

29

Figure 9. HAZ boundaries for Specimens BJ-3.2 and BJ-5.0

Figure 10. The 5900 series universal testing instruments

Figure 11. Stress-strain curves of butt joint coupons

0

100

200

300

400

500

600

700

800

900

1000

0 2 4 6 8 10 12 14 16

Str

ess

(MP

a)

Strain (%)

BJ-3.2-1

BJ-3.2-2

BJ-5.0-1

BJ-5.0-2

S690Q HSS

30

Figure 12. Numerical model of S690Q HSS butt joint

31

Figure 13. 3D finite element model for coupon stress analysis model

32

(a) Stress-strain curves of RQT S690 HSS at elevated temperatures

(b) Post-heating stress-strain curves of S690Q HSS after cooling down from different elevated

temperatures

0

100

200

300

400

500

600

700

800

900

1000

0 2 4 6 8 10 12 14 16

s (

MP

a)

(%)

25°C

100°C

200°C

300°C

400°C

450°C

500°C

600°C

700°C

800°C

0

100

200

300

400

500

600

700

800

900

1000

0 3 6 9 12 15 18 21 24 27 30

s(M

Pa)

(%)

25°C 400°C

600°C 800°C

900°C 1000°C

33

Figure. 14. Stress-stain curves defined in stress analysis model

Figure 15. Hardness diagram of numerical model

Figure 16. Temperature distribution of Num-BJ-3.2

34

(a) Comparison of temperature history for BJ-3.2 for thermocouple located at 3mm and 8mm from the

welding

(b) Comparison of temperature history for BJ-5.0 for thermocouple located at 5mm and 10mm from

the welding

Figure 17. Comparison of temperature history

0

200

400

600

800

1000

1200

1400

0 200 400 600 800 1000 1200 1400

Tem

per

atu

re (

°C)

Time (s)

BJ-3.2-3mm

BJ-3.2-8mm

Num-BJ-3.2-3mm

Num-BJ-3.2-8mm

0

200

400

600

800

1000

1200

0 200 400 600 800 1000 1200

Tem

per

ature

(°C

)

Time (s)

BJ-5.0-5mm

BJ-5.0-10mm

Num-BJ-5.0-5mm

Num-BJ-5.0-10mm

35

(a) Failure mode of BJ-3.2

(b) Failure mode of BJ-5.0

Figure 18. Comparison of failure mode

36

(a) Comparison of stress-strain curves for BJ-3.2

(b) Comparison of stress-strain curves for BJ-5.0

Figure 19. Comparison of stress-strain curves

0

100

200

300

400

500

600

700

800

900

1000

0 1 2 3 4 5 6 7

Str

ess

(MP

a)

Strain (%)

BJ-3.2-1

BJ-3.2-2

Num-BJ-3.2

0

100

200

300

400

500

600

700

800

900

1000

0 1 2 3 4 5 6 7

Str

ess

(MP

a)

Strain (%)

BJ-5.0-1

BJ-5.0-2

Num-BJ-5.0

37

(a) Welding passes for 12mm butt joints

(b) Welding passes for 16mm butt joints

Figure 20. Welding pass configuration for 12mm and 16mm butt joints

38

(a) Welding heat input 1.425kJ/mm

(b) Welding heat input 2.98kJ/mm

Figure 21. Reduction factor of yield strength and ultimate strength

0.9

0.92

0.94

0.96

0.98

1

6 8 10 12 14 16 18

Red

uct

ion

fac

tor

Thickness (mm)

Ry

Ru

0.8

0.84

0.88

0.92

0.96

1

6 8 10 12 14 16 18

Red

uct

ion f

acto

r

Thickness (mm)

Ry

Ru