Advances in Materials Science and Engineering
Volume 2019 (2019), Article ID 2983506, 13 pages
https://doi.org/10.1155/2019/2983506

Research Article

Analysis on Fracture Toughness of the L360QS/N08825 Bimetallic Composite Pipe Welded Joint

Bin Wang, Hongxia Lan, and Bobo Lei

School of Materials Science and Engineering, Southwest Petroleum University, Chengdu 610500, China

Correspondence should be addressed to Hongxia Lan; lanhongxia@163.com

Received 20 February 2019; Revised 27 August 2019; Accepted 6 September 2019; Published 7 October 2019

Guest Editor: Stanislav Seitl

Copyright © 2019 Bin Wang et al. This is an open access article distributed under the Creative Commons Attribution License, which permits unrestricted use, distribution, and reproduction in any medium, provided the original work is properly cited.

Abstract.

The fracture toughness of the weld and heat-affected zone (HAZ) of the L360QS/N08825 composite pipe welded joint was evaluated by a crack tip opening displacement (CTOD) test. The fracture morphology, microstructure, and inclusion near fracture zones were observed by means of scanning electron microscopy (SEM), transmission electron microscopy (TEM), and energy dispersive spectroscopy (EDS). The grain size and grain orientation of the crack propagation zone in the weld were investigated by electron back-scattered diffraction (EBSD). The results revealed that the average CTOD values of the weld and HAZ samples were relatively high, and a greater dispersion of CTOD values of the HAZ samples is due to the pop-in phenomenon in the P–V curve. The fracture surfaces of the weld and HAZ samples showed the characteristics of ductile fracture to a certain extent, whereas the fracture of the CTOD sample with the pop-in phenomenon exhibited a quasicleavage feature. High-density dislocation and a large number of inclusions were observed in the near fracture zone of the weld and HAZ samples. The stress concentration, caused by hindering the dislocation slip, was the main reason for microcrack formation. The existence of large-size grains and large-scale small-angle grain boundary in the weld implies that the cracks propagate toward the weld.

1. Introduction

Bimetallic composite pipes have been increasingly applied in the chemical industry, for environmental protection, for oil and gas transportation, and in the nuclear industry [1–3] because of their several benefits such as a reduction in material costs and a combination of good strength and excellent corrosion resistance of the two dissimilar materials [4, 5]. Because of great differences in the chemical composition, linear expansion coefficient, and thermal conductivity between the two dissimilar metals, it is easy for the bimetallic composite pipe welded joint to have defects and residual stress [6–8], which results in cracking [9–11]. The crack propagates until fracture occurs, thereby increasing the risk of oil and gas transportation and reducing the service life of the pipeline [12–14]. It is very important to investigate the fracture toughness of the bimetallic composite pipe welded joint to ensure the safety of construction projects.

At present, crack tip opening displacement (CTOD) is thought to be the essential fracture parameter to assess the fracture behavior of steel and welded joints based on nonlinear elastic fracture mechanics [15]. As the performance index and fracture criterion for crack initiation and propagation analysis of steel and welded joints, CTOD can not only effectively evaluate the fracture toughness of steel and welded joints, but it can also provide an experimental basis for the reliability and security assessment of construction [16, 17] through CTOD tests.

A few studies of fracture toughness of welding materials have been done by means of CTOD tests. Wang et al. [18] studied the CTOD fracture toughness of the weld and heat-affected zone (HAZ) of X80 steel at 0°C according to the BS7448 standard and found that the CTOD test could be used to evaluate effectively the fracture toughness of welded joints. Miao et al. [19] conducted a CTOD test on a submerged arc-welded joint of extrathick ocean structural steel S355G10+N. It was shown that the fracture toughness of the S355G10+N welded joint is excellent, and the welding process can be directly used to construct a marine steel structure. In addition, it was suggested that high stress should be avoided in the weld zone owing to the lowest toughness of the weld. Coronado and Cerón [20] studied CTOD for welded joints of three welding procedures and found that the fracture parameter CTOD was correlated with the fracture surface and microstructures. Leng et al. [21] studied the relationship between the fracture toughness and microstructure of S335G10+N welded joints by CTOD, a Charpy impact test, and scanning electron microscopy (SEM), and they found that the change trend of the CTOD value at the weld and fusion line with a notch was similar to that of the impact value. In addition, there is a big difference in the average grain sizes of S3557, S3556, and S3559 near-fracture microstructures. It was found that the larger the average grain size, the smaller the CTOD value. Wang et al. [22] studied the fracture properties and microstructure of A508/316L welded joints. The results showed that the fracture mechanism of A508, 316L base materials, and the HAZ of 316L is typical ductile fracture of the nucleation, growth, and coalescence of micropores. The fracture mode of the HAZ of A508 and the interface area of A508/52Mb composed of martensite is ductile-brittle mixed fracture, and the presence of martensite leads to low crack growth resistance. At the same time, it was found that the orientation of columnar austenite crystal can affect significantly the fracture mechanism and crack growth resistance in the weld zone. Guo et al. [23] also studied the fracture toughness and microstructure characteristics of various regions of 9Cr/CrMoV welded joints. The research results show that the fracture toughness of the CrMoV side is better than that of the 9Cr side because the microstructure of the former is fine austenite and bainite and the latter is coarse martensite. Li et al. [9] studied the crack and fracture properties of Fe₃Al/Cr18 Ni8 welded joints. The test shows that the crack initiation is in the fusion zone of the Fe₃Al side, in which there is much dislocation; most of the cracks propagate along the fusion zone, but a few cracks propagate horizontally to the HAZ and end at the weld. This is because of γ + δ organization in this area. Ju et al. [24] investigated the variations between ductile-to-brittle transition temperature (DBTT) and CTOD within the HAZ of API X65 pipeline steel. Both values varied dramatically with the distance from the fusion zone, and the region near the fusion zone exhibited the lowest CTOD and highest DBTT, which is possibly caused by the increasing portion of coarse grain in the HAZ. Because most mechanical components are subjected to complex loading conditions with varying magnitude and direction, Mokhtarishirazabad et al. [25] successfully evaluated the overload effect of biaxial fatigue cracking in terms of crack growth rate, crack opening displacement, and stress intensity factor, providing a new method to study the effect of applying overload cycle on the behavior of a crack under cyclic biaxial loading. The CTOD test can not only be used as the basis for selecting the toughness of piping and offshore materials, but also to provide the test basis for evaluating the safety and reliability of structures.

A few research efforts have been made to examine the fracture toughness for a bimetallic composite pipe via CTOD tests because the application of bimetallic composite pipes is becoming more and more extensive. To ensure the safety of bimetallic composite pipe welded joints during operation, in the present study, an experimental investigation was carried out using CTOD tests to evaluate the fracture toughness of the weld and the HAZ in an L360QS/N08825 bimetallic composite pipe welded joint according to BS7448 [26–28], ISO12135 (2002) [29], and ISO15653 (2010) [30] fracture toughness test standards at room temperature. The fracture morphology, microstructure, and inclusion near fracture zones were observed by means of SEM, TEM, and EDS, and the differences in grain size and grain orientation of the crack propagation area were studied by the EBSD technology. The primary aim was to study the fracture toughness characteristics of each part of the welded joint to provide a theoretical reference for engineering applications.

2. Materials and Methods

The cladding and base materials were made of N08825 and L360QS, respectively. The chemical compositions of the L360QS/N08825 bimetallic composite pipe formed by centrifugal casting are given in Table 1. The L360QS/N08825 bimetallic composite pipe circumferential weld was welded using tungsten inert gas (TIG) method through a D/T-HW350 automatic welder with a U-shaped groove shown in Figure 1. The filler metal was an ERNiCrMo-3 welding wire, whose chemical composition is given in Table 2. The diameter of the L360QS/N08825 bimetallic composite pipe welded joint was 610 mm, and the wall thickness was (20 + 3) mm.

Table 1: Chemical composition of L360QS and N08825 (mass fraction/%).

Material
C
Si
Mn
P
S
Ni
Cr
Cu
Mo
V
Ti

N08825
0.022
0.14
0.47
0.015
0.007
41.14
22.07
1.47
3.13
—
—

L360QS
0.07
0.30
1.06
0.012
0.001
0.090
0.090
0.060
0.040
0.040
0.002

Figure 1: Bimetallic composite pipe welded joint.

Table 2: Chemical composition of ERNiCrMo-3 (mass fraction/%).

Material
Ni
C
Cr
Mo
Nb
Fe
Ti
Si

ERNiCrMo-3
64.43
0.011
22.2
9.13
3.53
0.19
0.23
0.05

2.1. Sample Preparation

The fracture toughness test was conducted according to the BS7448 standard to obtain the CTOD value of the welded joint. Three-point bend specimens of the SE(B) with fatigue precrack in the weld and the HAZ were prepared, and the form of the SE(B) is shown in Figure 2. Thickness (B), width (W), and length (L) are the dimensions of the specimen, M is the machining notch, m is the length of the precrack generated by fatigue, and N is the notch thickness.

Figure 2: Three-point bending specimen SE(B).

For estimating each parameter, the ISO12135 (2002) and ISO15653 (2010) standards were used to intercept samples of the HAZ and weld, respectively. The size of the weld sample was B × 2B, and the notch direction was NP. The size of the HAZ sample was B × B, and the notch direction was NQ. The schematic diagram of the sample interception is shown in Figure 3, where N is the vertical weld direction, P is the direction of the parallel weld, and Q is the direction of weld thickness. NP implies that the direction of the sample length is a vertical weld and the direction of crack propagation is a parallel weld. NQ represents that the direction of sample length is a vertical weld and the direction of crack propagation is in the direction of the weld thickness. There were three samples to be tested for the weld and for the HAZ: W1, W2, and W3 and H1, H2, and H3, respectively.

(a)

(b)

Figure 3: Schematic diagram of CTOD sample interception. (a) HAZ; (b) weld.

2.2. Machining of Mechanical Notch

The cutting line of the specimens was marked in accordance with the BS7448 standard. For the HAZ specimens, the line should be drawn on the fusion line to ensure that the crack-tip position is not more than 0.5 mm from the fusion line. The line should coincide with the center line for the weld specimens. In addition, the vertical angle between the plane of the cutting line and the sample cutting surface should be 90° ± 5° when the notch is being processed. An electrical-discharge-machining (EDM) wire-cutting process with molybdenum wire of 0.08 mm diameter was used to machine the notch on the side of the N08825 alloy along the thickness direction of the specimen. The machining depth of the notch was approximately 45% of the specimen thickness. The angle of the notch root was less than 60°, and the radius of the notch root was 0.12 mm.

2.3. Experimental Procedure

The fatigue precracking was carried out on the side of the N08825 alloy using a PLG-200 high-frequency fatigue-testing instrument with a vibration frequency between 115 and 117 Hz at room temperature (20°C). The load ratio was set at 0.1. For the weld and HAZ samples, the loading span was set as 160 and 80 mm, respectively. Meanwhile, maximum fatigue precracking loads of 22.00 and 35.00 kN were applied for the weld and HAZ samples, respectively. The time of fatigue precracking for each sample was approximately 0.4 h. According to the BS7448 standard, crack growth should not be too fast when fabricating the fatigue precracking. In the last 1.3 mm, the fatigue load ratio can be appropriately increased to avoid the growth of the surface crack being larger than that of the internal crack. Loading force needs to be strictly controlled to prevent the plastic deformation of the crack tip during the process of fatigue precracking and ensure that the initial crack length a₀ of all samples should be within the range of 0.10–0.45W (HAZ) or 0.45–0.70W (weld), where W is sample width (mm) [30].

CTOD experiments were performed with a DDL 300 universal testing machine at room temperature; at the same time, the P–V curve was automatically recorded (P is applied load, and V is the crack tip opening displacement). A typical P–V curve is shown in Figure 4. The parameters of the experiment were: speed of application of the load, 0.1 mm/min; span, 160 mm for the weld and 80 mm for the HAZ; and sampling interval for equipment data acquisition, 0.1 s. The initial crack length (a₀) [28] and CTOD value (δ) [26] are calculated according to the following equation:

where is the initial crack length of the i-th test point () (the measure of is shown in Figure 5), is the applied load, is span of SE(B), is Poisson’s ratio, is elastic modulus, is function for (), stands for the plastic component of CTOD, is yield strength, and is blade thickness.

Figure 4: Typical P–V curve.

Figure 5: Measure of (i = 2, 3, 4, …, 8).

After the CTOD test, the validity of the CTOD samples was determined according to the BS7448, ISO12135 (2002), and ISO15653 (2010) standards, and specific steps were performed based on standard procedures from the literature [21]. The analysis showed that the CTOD samples of the welded joints were effective.

The fracture surface of the samples was analyzed by SEM using secondary electrons, while the composition of fracture inclusions was analyzed by EDS. The SEM sample was etched by a solution of CH₃OH + HNO₃ with the proportion of 3 : 1. TEM was used to observe and analyze the microstructure morphology and the distribution of dislocations in the near fracture of the sample, and the acceleration voltage was 200 kV. The piece of TEM observational sample was cut from the middle thickness region of the specimens along the direction perpendicular to the machining notch, as shown in Figure 6. After mechanical thinning to 50 µm, the TEM specimen was exposed to a double electrolytic jet in an electrolytic double spray device, and the electrolyte was 75% CH₃OH + 25% HNO₃ (volume fraction). To understand further the direction of crack propagation from the perspective of crystallography, the crystal structure, grain size, and grain boundary misorientation of the weld zone were observed and analyzed using EBSD technology.

Figure 6: TEM observation of CTOD near-fracture sample.

3. Results and Discussion

3.1. Discussion of CTOD Value

Table 3 summarizes the data concerning the fracture toughness tests. The average CTOD value (δm) of the weld samples is 1.3232 mm, and the minimum value is 1.2022 mm. Meanwhile, the average CTOD value of the HAZ samples is 1.5119 mm, and the minimum value is 0.9787 mm.

Table 3: CTOD values and parameters of specimens of L360QS/N08825 welded joint (20°C).

Samples
Crack location
Mechanical properties
Thickness B (mm)
Width W (mm)
Crack length a₀ (mm)
a₀/W
CTOD value δm (mm)
Average value −δm (mm)

W1
Weld
σs = 370 MPa
19.82
39.88
21.55
0.54
1.2739
1.3232

W2
E = 2.1 × 10⁵ MPa
19.98
39.90
20.59
0.52
1.4935

W3
µ = 0.3
19.96
39.98
21.02
0.53
1.2022

H1
HAZ
σs = 370 MPa
19.94
19.91
4.21
0.21
2.4457
1.5119

H2
E = 2.1 × 10⁵ MPa
19.96
19.94
3.36
0.17
0.9787

H3
µ = 0.3
19.92
19.85
3.77
0.19
1.1113

As Table 3 shows, the δm value of the HAZ is highly discrete. The maximum δm value of 2.4457 is more than twice the minimum δm value of 0.9787, which is thought to be caused by the heterogeneity structure and performance in the HAZ. The lowest CTOD value of H2 may be caused by the pop-in phenomenon in the P–V curve of sample H2 shown in Figure 7. The pop-in phenomenon shows that the load suddenly decreases and the displacement increases in the P–V curve, and then the load and displacement continue to increase until cracking. Meanwhile, when the pop-in phenomenon occurs, a relatively weak crackling sound can be heard during the loading process, and a brittle fracture zone can also be observed on the fracture surface of the sample. In fact, this pop-in phenomenon was derived from the small brittle crack initiated at a local brittle zone. This rapidly propagating brittle crack was subsequently arrested when it propagated into the higher-toughness region that surrounded the local brittle zone [31]. The macroscopic fracture of sample H2 shown in Figure 8 reveals the signs of crack tip instability in the local brittle zone, and the crack propagation path is extremely unstable. Cleavage steps can also be observed in the crack growth zone, which conforms to the pop-in phenomenon characteristics. In the loading process of the CTOD sample, the reduction in resistance to crack propagation of sample H2 resulted from unstable crack propagation, which resulted in a lower δm value. The allowable value of CTOD was set at 0.30 mm by referring to the design value of similar steel structures. The CTOD values of all zones of the above automatic TIG welded joints exceeded the allowable value (0.30 mm), so this process can be used for welding L360QS/N08825 composite pipe.

Figure 7: Pop-in phenomenon in P–V of sample H2.

Figure 8: Macroscopic fracture morphology of sample H2.

3.2. SEM Fractography Analysis of CTOD Sample

SEM images of the fracture surfaces of the weld and HAZ samples are shown in Figures 9 and 10, respectively.

(a)

(b)

(c)

(d)

(e)

(f)

(g)

(h)

(i)

Figure 9: CTOD microscopic fracture photo of weld of L360QS/N08825 bimetallic composite pipe welded joint: (a), (b), (c) W1; (d), (e), (f) W2; (g), (h), and (i) W3.

(a)

(b)

(c)

(d)

(e)

(f)

(g)

(h)

(i)

Figure 10: CTOD microfracture photo of the HAZ of L360QS/N08825 bimetallic composite pipe welded joint: (a), (b), (c) H1; (d), (e), (f) H2; (g), (h), and (i) H3.

Figures 9(a), 9(d), and 9(g) show that there are many shallow dimples and some flat brittle appearance on the fracture surfaces of the crack initiation area of the weld samples, which is believed to be a result of separating the weak grain boundaries with low fracture energy [32]. Figures 9(b), 9(e), and 9(h) are the fracture images of the crack propagation region of the weld sample. In Figures 9(b) and 9(h), the fracture is ductile fracture with large and spherical dimples distributed alternately, and small dimples have side-by-side connections. Figure 9(e) shows the fracture surface that is characterized by parabolic dimples, and inclusions are distributed at the bottom of the dimples. As Figures 9(c), 9(f), and 9(i) show, the fracture surfaces of the final fracture zone consist of equiaxial dimples distributed uniformly, which are typical of microvoid coalescence fracture [33].

Observations of the fracture surfaces of the crack initiation area of the HAZ samples are shown in Figures 10(a), 10(d), and 10(g). Figure 10(a) shows that the shape of the dimples is parabolic, and the size is nonuniform. The number of dimples in Figure 10(g) is less than that in Figure 10(a), which indicates that the toughness of sample H3 is weaker than that of sample H1 in accordance with the CTOD value (δm) listed in Table 3. The fracture morphology of the crack initiation area of H2 in Figure 10(d) is a quasicleavage fracture, and macroscopic cracking occurs between some cleavage surfaces. The occurrence of macroscopic cracks may be caused by cracking along grain boundaries or brittle phases. The SEM fracture analysis of the crack propagation region of sample H1 revealed a ductile fracture with large and deep dimples in a parabolic shape, as shown in Figure 10(b). The extensive and deep dimples in the fracture morphology of H1 indicate that a great deal of energy has been consumed during the fracture process [34], contributing to a high CTOD value (δm). Some shallow dimples can be seen in the crack propagation region of sample H2 (Figure 10(e)), which is characteristic of local ductile fracture. Figure 10(h) shows that the fracture surfaces of H3 had a ductile appearance characterized by equiaxial dimples that are typical of microvoid coalescence fracture. Figures 10(c), 10(f), and 10(i) show high-resolution fractographs of the final fracture zone of the HAZ samples. Figure 10(c) shows that many parabolic dimples were distributed on the fracture surface of H1, and the fracture had a ductile fracture morphology. The fracture surfaces of the final fracture zone of H2 consisted of brittle fracture with quasicleavage characteristics, including a fluvial pattern, as shown in Figure 10(f). This fracture was accompanied by less plastic deformation and energy consumption, and it corresponds to lower crack growth resistance. In the final fracture zone of H3, some small dimples and brittle facets can be seen, as shown in Figure 10(i), and this is characteristic of mixed ductile–brittle fracture.

As shown in Figure 10, there is a clear difference in the fracture surface characteristics between sample H2 with the pop-in phenomenon and samples H1 and H3 without the pop-in phenomenon. The fracture morphology of the crack initiation zone and the final fracture zone of samples without the pop-in phenomenon is the ductile dimple. However, the fracture morphology of the same area of the samples with the pop-in phenomenon is quasicleavage. This further confirms that samples containing the pop-in phenomenon have lower CTOD values compared with samples without the pop-in phenomenon in line with data shown in Table 3.

The morphology and composition of inclusions in the fracture surface of the CTOD sample of the HAZ are shown in Figure 11. The spherical inclusions are located at the bottom of the dimple and are identified as oxides, including Al₂O₃ and CaO, by EDS analysis. The inclusion is a favorable position for the formation of dimples, and the stress concentration around the inclusion generates a microstress field, which leads to the formation and propagation of cracks until ultimate fracture.

(a)

(b)

Figure 11: SEM of morphology and composition of inclusions in the fracture zone of the HAZ.

3.3. TEM Analysis of CTOD Sample near Fracture Zone

The microstructure of the near fracture zone of CTOD fracture samples of the L360QS/N08825 composite pipe welded joint is shown in Figure 12. Figures 12(a) and 12(b) show that the microstructure of the near fracture zone of the HAZ samples is mainly dendritic austenite with a clear dendritic boundary, and the branch width is between 181.61 and 620.69 nm. The appearance of the dendritic austenite microstructure of the near fracture zone of the HAZ samples is because the location of the CTOD fracture crack propagation zone is in the weld. At the same time, there is a high dislocation density inside the austenite grain and along the grain boundary, and dislocation entanglement was also observed in the grain. As reported in the literature [35], dislocation slips forward under the action of external forces. When the dislocation moves to the grain boundary, it is obstructed and causes local stress concentration. The stress concentration at the grain boundary becomes the crack source that eventually develops into a microcrack with the increase of the deformation degree, which eventually leads to fracture after crack growth, propagation, and unstable propagation. In addition, a large number of precipitated phases with a relatively uniform size of approximately 15 nm on the dislocation line were also observed inside the dendritic austenite grain.

(a)

(b)

(c)

(d)

Figure 12: TEM images of near fracture zones of CTOD samples: (a, b) HAZ and (c, d) weld.

Figures 12(c) and 12(d) show that the microstructure morphology of the near fracture zone of the weld samples is similar to that of the HAZ samples. The major microstructure is dendritic austenite with a clear dendritic boundary, and the branch width is in the range of 228.99 to 1660.00 nm. Furthermore, high-density dislocation and precipitated phases with different sizes ranging from 5.88 to 17.65 nm were also observed in the austenite grain and grain boundary.

The above analysis shows that the high-density dislocation was distributed in the grain and grain boundary, providing conditions for the initiation of microcracks that go through the growth and propagation process until fracturing eventually. The relevant literature [36] indicates that the precipitations on the dislocation line could hinder dislocation motion effectively and delay the initiation and propagation of cracks to some extent, which is beneficial to the toughness of materials. The appearance of a mass of precipitations in the HAZ and weld might be one of the reasons why the average δm value of the CTOD samples was higher than the standard value (0.30 mm).

In general, the inclusions in the matrix of the metal material welded joint play an important role in the fracture process as the nucleus center of microvoid formation and aggregation. Figure 13 shows the distribution of inclusions and EDS analysis of the near fracture zone of the HAZ samples. Figure 13 shows that there are approximately 40 spherical inclusions with different sizes between 3 and 8 µm in diameter in the area of 400 × 400 µm. These inclusions were distributed in the grain and the grain boundary, and they were composed of sulfide and oxide, mainly including TiO₂ and CaS.

(a)

(b)

Figure 13: SEM images of inclusions and compositions in the near fracture zone of the HAZ.

The distribution of inclusions and EDS analysis of the near fracture zone of the weld samples are shown in Figure 14. Many irregularly shaped inclusions with nonuniform size were distributed unevenly on the weld substrate. The EDS results reveal that the composition of inclusions is mainly SiO₂ oxide. The occurrence of stress concentration happens easily and is common around the inclusions under the action of external force, which increases the opportunity for microcracks to form in the material. Meanwhile, high-density distributed inclusions make the connection of microcracks easier, which eventually leads to the fracture of components after the process of growth and propagation of cracks [35].

(a)

(b)

Figure 14: SEM images of inclusions and compositions in the near fracture zone of weld.

3.4. EBSD Analysis of Weld Zone

Metallography observation on the surface perpendicular to the fracture surface showed that fatigue precracks of HAZ and weld samples propagated into the weld zone during the CTOD test. To reveal the relationship between the direction of crack propagation and microstructure and the grain size and grain boundary misorientation from the perspective of crystallography, EBSD technology was used to analyze the grain size and grain boundary misorientation in the weld zone.

The strength and toughness of materials are closely related to the sizes of grains. Generally, the strength and toughness of materials are improved as the size of the metal grain decreases [37, 38]. Therefore, it is necessary to study the grain size of the crack propagation zone in the weld. Figure 15(a) shows the grain size distribution in the weld zone, and the color of each grain is selected randomly. As the histogram of the weld grain size distribution in Figure 15(b) shows, there are 1172 grains, and the average grain size is 95.2 µm. Based on the above data analysis from Figure 15, the average grain size in the weld zone is relatively large, which leads to a decrease in toughness and promotes crack propagation in this area.

(a)

(b)

Figure 15: Grain size of weld: (a) grain size and (b) histogram.

To reveal the crystallographic parameters of the weld microstructure further, the grain orientation distribution of the weld (the same color represents the same grain orientation) and the grain boundary misorientation distribution diagram of the weld are shown in Figures 16 and 17, respectively. Mathematical statistics show that there is an approximately 81% small-angle grain boundary (2°–15°) in the weld zone, and the remaining 19% is a high-angle grain boundary distributed within 49°–62° randomly. This result indicates that the grain boundary misorientation of the weld zone is small, and the appearance of high-angle grain boundaries (>15°) results from the accumulation of small-angle grain boundaries [39]. According to the relevant literature [40], the existence of small-angle grain boundaries leads to the decline of material toughness, which is conducive to the propagation of cracks. Therefore, a large number of small-angle grain boundaries in the weld zone reduce the resistance to crack propagation, which explains why cracks propagated into the weld zone during the CTOD test.

(a)

(b)

Figure 16: Grain orientation distribution of weld.

Figure 17: Grain boundary misorientation distribution histogram of weld.

4. Conclusions

In the present work, the fracture toughness characteristics of bimetallic composite pipe welded joints were studied. The following results were obtained:

(1)The CTOD values of the HAZ and weld samples were high; nevertheless, the greater dispersion of CTOD values of the HAZ samples resulted from the pop-in phenomenon in the P–V curve of sample H2.(2)The fracture surfaces of the L360QS/N08825 composite pipe welded joint were characterized by ductile fracture to a certain extent, while the fracture surface of the CTOD sample in the HAZ region containing the pop-in phenomenon had a quasicleavage morphology.(3)The microstructure of the near fracture zone of the HAZ and weld samples was mainly dendritic austenite, and there were high-density dislocation and dislocation entanglement in the grains and grain boundary. A large number of spherical or irregular inclusions were observed, which is the main reason for the appearance of microcracks. Meanwhile, the high-density inclusions distributed on the matrix made the connection of microcracks easier, which eventually leads to the fracture of components after the process of growth and propagation of cracks.(4)The average grain size of the weld structure was 95.2 µm, and small-angle grain boundaries accounted for approximately 81%. A larger grain size and a larger proportion of small-angle grain boundary can reduce the resistance to crack propagation and promote crack propagation.

Data Availability

The data used to support the findings of this study are included within the article.

Conflicts of Interest

There are no conflicts of interest regarding the publication of this paper.

References

F. Jiang, K. Zhao, and J. Sun, “Evaluation of interfacial crack growth in bimaterial metallic joints loaded by symmetric three-point bending,” International Journal of Pressure Vessels and Piping, vol. 80, no. 2, pp. 129–137, 2003.

R. Kacar and M. Acarer, “An investigation on the explosive cladding of 316L stainless steel-din-P355GH steel,” Journal of Materials Processing Technology, vol. 152, no. 1, pp. 91–96, 2004.

X. Qian, Y. Wang, J. Y. Richard Liew, and M.-H. Zhang, “A load-indentation formulation for cement composite filled pipe-in-pipe structures,” Engineering Structures, vol. 92, pp. 84–100, 2015.

L.-J. Zhang, Q. Pei, J.-X. Zhang, Z.-Y. Bi, and P.-C. Li, “Study on the microstructure and mechanical properties of explosive welded 2205/X65 bimetallic sheet,” Materials & Design, vol. 64, no. 9, pp. 462–476, 2014.

S. Bagherzadeh, B. Mollaei-Dariani, and K. Malekzadeh, “Theoretical study on hydro-mechanical deep drawing process of bimetallic sheets and experimental observations,” Journal of Materials Processing Technology, vol. 212, no. 9, pp. 1840–1849, 2012.

A. Joseph, S. K. Rai, T. Jayakumar, and N. Murugan, “Evaluation of residual stresses in dissimilar weld joints,” International Journal of Pressure Vessels and Piping, vol. 82, no. 9, pp. 700–705, 2005.

C. Jang, J. Lee, J. Sung Kim, and T. Eun Jin, “Mechanical property variation within Inconel 82/182 dissimilar metal weld between low alloy steel and 316 stainless steel,” International Journal of Pressure Vessels and Piping, vol. 85, no. 9, pp. 635–646, 2008.

A. Mortezaie and M. Shamanian, “An assessment of microstructure, mechanical properties and corrosion resistance of dissimilar welds between Inconel 718 and 310S austenitic stainless steel,” International Journal of Pressure Vessels and Piping, vol. 116, no. 1, pp. 37–46, 2014.

Y. Li, H. Ma, and J. Wang, “A study of crack and fracture on the welding joint of Fe₃Al and Cr18–Ni8 stainless steel,” Materials Science and Engineering: A, vol. 528, no. 13-14, pp. 4343–4347, 2011.

R. Strubbia, S. Hereñú, A. Giertler, I. Alvarez-Armas, and U. Krupp, “Experimental characterization of short crack nucleation and growth during cycling in lean duplex stainless steels,” International Journal of Fatigue, vol. 65, pp. 58–63, 2014.

H. Naffakh, M. Shamanian, and F. Ashrafizadeh, “Dissimilar welding of AISI 310 austenitic stainless steel to nickel-based alloy Inconel 657,” Journal of Materials Processing Technology, vol. 209, no. 7, pp. 3628–3639, 2009.

H. T. Wang, G. Z. Wang, F. Z. Xuan, and S. T. Tu, “An experimental investigation of local fracture resistance and crack growth paths in a dissimilar metal welded joint,” Materials & Design, vol. 44, pp. 179–189, 2013.

S. Kumar, P. K. Singh, K. N. Karn, and V. Bhasin, “Experimental investigation of local tensile and fracture resistance behaviour of dissimilar metal weld joint: SA508 Gr.3 Cl.1 and SA312 Type 304LN,” Fatigue & Fracture of Engineering Materials & Structures, vol. 40, no. 2, pp. 190–206, 2016.

X. Deng, F. Lu, H. Cui, X. Tang, and Z. Li, “Microstructure correlation and fatigue crack growth behavior in dissimilar 9Cr/CrMoV welded joint,” Materials Science and Engineering: A, vol. 651, pp. 1018–1030, 2016.

J. W. Hutchinson, “Fundamentals of the phenomenological theory of nonlinear fracture mechanics,” Journal of Applied Mechanics, vol. 50, no. 4b, pp. 1042–1051, 1983.

Z. T. Fang, B. Sun, and C. R. Li, “Experimental study on CTOD fracture toughness of welded joints of low temperature steel,” Advanced Materials Research, vol. 328-330, pp. 1272–1276, 2011.

A. H. Elsayed, M. M. Megahed, A. A. Sadek, and K. M. Abouelela, “Fracture toughness characterization of austempered ductile iron produced using both conventional and two-step austempering processes,” Materials & Design, vol. 30, no. 6, pp. 1866–1877, 2009.

P. Wang, M. J. Hu, and E. Dang, “Study on CTOD fracture toughness of welded joint of X80 marine drilling riser,” Advanced Materials Research, vol. 228-229, pp. 1163–1168, 2011.

Z. M. Miao, T. Miao, F. X. Qiu, S. W. Leng, and L. N. Niu, “CTOD testing and evaluation for weld joint for heavy thickness offshore structure steel S355G10+N,” Applied Mechanics and Materials, vol. 117–119, pp. 1597–1601, 2011.

J. J. Coronado and C. Cerón, “Fracture mechanisms of CTOD samples of submerged and flux cored arc welding,” Theoretical and Applied Fracture Mechanics, vol. 53, no. 2, pp. 145–151, 2010.

S. W. Leng, Z. M. Miao, F. X. Qiu, L. N. Niu, and T. Miao, “Analysis of the relationship between CTOD toughness and micromechanism of marine steel weld joints,” Applied Mechanics and Materials, vol. 117-119, pp. 1867–1873, 2011.

H. T. Wang, G. Z. Wang, F. Z. Xuan, and S. T. Tu, “Fracture mechanism of a dissimilar metal welded joint in nuclear power plant,” Engineering Failure Analysis, vol. 28, pp. 134–148, 2013.

Q. Guo, F. Lu, X. Liu, R. Yang, H. Cui, and Y. Gao, “Correlation of microstructure and fracture toughness of advanced 9Cr/CrMoV dissimilarly welded joint,” Materials Science and Engineering: A, vol. 638, pp. 240–250, 2015.

J.-B. Ju, W.-s. Kim, and J.-i. Jang, “Variations in DBTT and CTOD within weld heat-affected zone of API X65 pipeline steel,” Materials Science and Engineering: A, vol. 546, pp. 258–262, 2012.

M. Mokhtarishirazabad, P. Lopez-Crespo, B. Moreno, A. Lopez-Moreno, and M. Zanganeh, “Optical and analytical investigation of overloads in biaxial fatigue cracks,” International Journal of Fatigue, vol. 100, pp. 583–590, 2017.

BS7448, Fracture Mechanics Toughness Test Part 1. Method for Determination of KIC, Critical CTOD and Critical J Values of Metallic Materials, British Standards Institution, London, UK, 1991.

BS7448, Fracture Mechanics Toughness Test Part 2. Method for Determination of KIC, Critical CTOD and Critical J Values of Welds in Metallic Materials, British Standards Institution, London, UK, 1997.

BS7448, Fracture Mechanics Toughness Test Part 4. Method for Determination of Fracture Resistance Curves and Initiation Values for Stable Crack Extension in Metallic Materials, British Standards Institution, London, UK, 1997.

ISO 12135-2002 (E), Metallic Materials-Unified Method of Test for the Determination of Quasistatic Fracture Toughness, International Standardization Organization, Geneva, Switzerland, 2002.

ISO 12135-2010 (E), Metallic Materials-Method of Test for the Determination of Quasistatic Fracture Toughness of Welds, International Standards Organization, Geneva, Switzerland, 2010.

Y. Yamashita and S. Kanna, “Assessment of pop-in significance in heterogeneous weld heat-affected zone using finite element analyses,” Procedia Materials Science, vol. 3, pp. 991–996, 2014.

J. Yang and L. Wang, “Fracture mechanism of cracks in the weakest location of dissimilar metal welded joint under the interaction effect of in-plane and out-of-plane constraints,” Engineering Fracture Mechanics, vol. 192, pp. 12–23, 2018.

R. Sunder, W. J. Porter, and N. E. Ashbaugh, “Fatigue voids and their significance,” Fatigue Fracture of Engineering Materials and Structures, vol. 25, no. 11, pp. 1015–1024, 2002.

J. Yang, G. Z. Wang, F. Z. Xuan, S. T. Tu, and C. J. Liu, “Out-of-plane constraint effect on local fracture resistance of a dissimilar metal welded joint,” Materials & Design, vol. 55, no. 1, pp. 542–550, 2014.

B. I. Zongyue, J. Yang, N. Jing, and J. Zhang, “Fracture toughness of welded joints of X100 high-strength pipeline steel,” Acta Metallurgica Sinica, vol. 49, no. 5, pp. 576–582, 2013.

M. Wang, Z. Zhou, H. Sun, H. Hu, and S. Li, “Microstructural observation and tensile properties of ODS-304 austenitic steel,” Materials Science and Engineering: A, vol. 559, no. 3, pp. 287–292, 2013.

S. Nafisi, M. A. Arafin, L. Collins, and J. Szpunar, “Texture and mechanical properties of API ×100 steel manufactured under various thermomechanical cycles,” Materials Science and Engineering: A, vol. 531, no. 5, pp. 2–11, 2012.

X. Zhang, Z. Jiang, S. Li, and J. Fan, “Effect of effective grain size and grain boundary of large misorientation on upper shelf energy in pipeline steels,” Journal of Wuhan University of Technology-Mater. Sci. Ed., vol. 31, no. 3, pp. 606–610, 2016.

B. Wang, B. B. Lei, J. X. Zhu, Q. Feng, L. Wang, and D. Wu, “EBSD study on microstructure and texture of friction stir welded AA5052-O and AA6061-T6 dissimilar joint,” Materials & Design, vol. 87, pp. 593–599, 2015.

C. Zhang, Q. Wang, J. Ren et al., “Effect of martensitic morphology on mechanical properties of an as-quenched and tempered 25CrMo48V steel,” Materials Science and Engineering: A, vol. 534, pp. 339–346, 2012.