Effects of Temperature and Crack Tip Constraint on Cleavage Fracture Toughness in the Weld Thermal
Simulated X80 Pipeline Steel
J Xua, Z L Zhangb, E Østbyc, B Nyhusc, D B Suna a School of materials science and engineering, University of Science and Technology Beijing (USTB)
Beijing, China b Department of Structural Engineering, Norwegian University of Science and Technology (NTNU)
Trondheim, Norway c SINTEF, Materials and Chemistry
Trondheim, Norway
ABSTRACT This paper studies the effects of temperature and crack tip constraint on cleavage fracture toughness of weld thermal simulated X80 pipeline steel at low temperatures. A large number of fracture toughness (as denoted by CTOD) tests together with 3D finite element analyses are performed using single edge notched bending (SENB) and tension (SENT) specimens with different crack depths at -90 and ℃ -30℃. Coarse-grained heat-affected zone (CGHAZ) is considered as the material microstructure in preparation of weld thermal simulated fracture mechanics specimens. The applicability of RKR model as the local fracture criterion in prediction of cleavage fracture toughness is examined. KEY WORDS: Fracture toughness, coarse-grained heat affected zone (CGHAZ), X80 pipeline steel, weld thermal simulation, finite element analysis (FEA). INTRODUCTION Fracture toughness testing of ferritic structural steels reveals a significant effect of crack depth and loading mode (bending vs. tension) on fracture toughness values (Wallin, 1985; Sorem, et al, 1991; Joyce and Link, 1997). These studies show significant elevations in fracture toughness (characterized as KIC or JC) for specimens containing shallow cracks and/or subjected to tensile loading. With increasing loads in a finite body, such as a cracked specimen or a structural component, the crack-tip plastic zone is increasingly influenced by the nearby traction free boundary which relaxes the near-tip stress levels well below the values corresponding to small scale yielding (SSY) conditions. This phenomenon, often termed loss of constraint (Dodds, Shih and Anderson, 1993; Nevalainen and Dodds, 1995; Dodds, Ruggieri and Koppenhoefer, 1997), contributes to the apparent increased toughness of shallow cracked and tension loaded geometries observed in fracture mechanics testing. Once high triaxiality conditions no longer apply, larger crack driving forces in the finite body are necessary to generate a highly stressed region ahead of the crack-tip sufficient to trigger
cleavage. Since the crack tip constraint affects the stress distribution around a crack and invalidates the use of single parameter characterization of the crack-tip stress field. Therefore, the second parameter, for example T-stress (Betegon and Hancock, 1991) and Q-parameter (O’Dowd and Shih, 1991), has been developed to further characterize the crack-tip stress field and quantify constraint levels for various geometries and loading configurations. Macroscopic fracture toughness values (KIC or JC), measured experimentally over the ductile-to-brittle transition (DBT) range of ferritic steels consistently exhibit a large amount of scatter (e.g. see Sorem, et al, 1991; Wallin, 1984). This DBT regime has been widely investigated and discussed by many researchers (Lambert-Perlade, Gourgues and Pineau, 2004; Pineau, 2006 and 2008; Ortner, 2006; Hausild, Berdin and Bompard, 2005). Among all these works addressing temperature effect on cleavage fracture mechanism, it has been found that the cleavage mechanisms were changing with temperature. Although they respectively pointed out that a deeper knowledge of the cleavage mechanisms at higher temperature is needed, the results strongly indicate that a single micromechanism model for cleavage fracture may not apply for the whole transition region. The heat-affected zone (HAZ) of a weldment is generally considered as the weakest part and is crucial in the failure of steel structures. The reason of that is its heterogeneous microstructure owing to the sensitive variation with the thermal cycle experienced during welding process. Therefore, treatment of brittle fracture in weldment and HAZ is challenging. Both lead to large local variations in fracture toughness and local constraint effect due to the difference in material properties. In the 90s’ there was a significant focus on characterizing the local stress fields in weldment and HAZ. SINTEF/NTNU developed the so-called J-Q-M theory (see Zhang et al, 1996) where both constraint effects due to geometry and material mismatch were included in the characterization of the local stress field ahead of the crack-tip. In this study, the effects of temperature and crack tip constraint on cleavage fracture toughness in weld thermal simulated X80 pipeline steel is studied. Weld thermal simulation has been used to produce the microstructural properties of coarse-grained HAZ (CGHAZ) in X80. A large number of CTOD (crack tip opening displacement) tests are carried out at -90 and ℃ -30℃. Single edge notched bending (SENB
with a/W=0.5) and tension (SENT with a/W=0.3) specimens are selected in order to characterize the constraint effect on the fracture toughness. 3D FEA are employed to model the crack tip stress fields of tested specimens. The true stress-strain curves are obtained through smooth round bar tensile tests at both temperatures for characterizing the material flow properties in 3D FE models. The numerical analyses are compared with the experimental results. EXPERIMENTAL DETAILS Materials Description The material used in this work is the high strength low-alloy (HSLA) X80 grade steel which is increasingly used in the oil and gas transmission pipeline applications. The nominal outer diameter of the pipe is 510 mm, and the nominal wall thickness is 14.6 mm. Table 1 lists the typical chemical composition of this material. Table 1. Chemical composition of X80 steel (%) Steel C Si Mn P S Others
Weld Thermal Simulation The weld thermal simulation technique is based on the reheating small specimens using thermal cycles. The purpose of this paper is to study the fracture mechanical properties of the weld thermal simulated X80 pipeline steel. The single welding cycle simulation is intended to produce the coarse-grained HAZ (CGHAZ) of the welds, which therefore been used in preparation of weld thermal simulated specimens. The specimens were heated to maximum temperature of 1350 ℃ by resistance-heating in a computer-controlled weld thermal simulator and then cooled for 15 sec (Δt8/5). The synthetic CGHAZ microstructure was thereafter produced in a certain region in the specimen where the fatigue pre-crack is introduced after being machined. The True Stress-Strain Curve The true stress-strain curves for material tested (CGHAZ in X80) in this paper have been obtained through the smooth round bar tensile tests at -90 and ℃ -30 ,℃ as shown in Fig. 1. It can be observed that, yield strength for material studied here increases with decreasing temperature. Moreover, temperature shows no significant effect on the material hardening behavior.
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All testing specimens are directly extracted from the X80 pipeline with specimen length following the pipeline longitudinal direction and crack propagation along the pipe thickness. In order to study the crack tip constraint effect on the fracture toughness, SENB and SENT specimens with different crack depths are selected as the fracture mechanics specimens. The geometrical configurations are schematically drawn in Fig. 2 for both SENB and SENT specimens.
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Fig. 2. Specimen configurations. (a) SENB with a/W=0.5; (b) SENT with a/W=0.3. For all specimens, a thickness of B = 10 mm and width W = 10 mm with crack depth (denoted by a), to width ratio of a/W = 0.5 for SENB and a/W = 0.3 for SENT specimens have been considered. The span of the specimen, S, is chosen to be four times of width, W, for SENB (S/W = 4) and L/W = 6 for SENT. The SENB specimens are machined and tested according to the standard of BS 7448 (BS 7448, Part 1: 1991; Part 2: 1997), while SENT testing has been performed in accordance with the “Recommended practice DNV-RP-F108” (Recommended Practice DNV-RP-F108, 2006). The load-CMOD (crack mouth opening displacement) curves were digitally recorded using double clip gauge during the tests. The CTOD values are determined at the maximum load through measured load-CMOD records. 10 parallel tests have been performed at each temperature and specimen geometry. The testing rate is 0.5mm/min of crosshead displacement. After each test, the fatigue pre-cracking length and ductile crack extension that occurred during the test were measured using an optical microscope. NUMERICAL PROCEDURES 3D finite element models for both SENB and SENT specimens were generated using ABAQUS CAE (ABAQUS, 2008) as shown in Fig. 3 (a) and (b). Only one-quarter of the specimen is modeled due to the symmetry with respect to the mid-plane (z = 0) and the crack free surface plane (x = 0). A conventional mesh configuration having a focused ring of elements surrounding the crack front is used with a small notch (with a notch root radius of r = 2 μm) as can be seen in Fig. 3 (c). Along the thickness direction (z-axis), identical planar mesh is repeated from the symmetry plane (i.e. mid-plane with z=0) to the free surface (z = B/2). In order to catch the change of the stress field near the free surface, thickness of continuous element layer is reduced from the mid-plane toward to the free surface. A gradually coarse element mesh in the length-wise of specimen and circumferential direction around the crack tip is applied in order to minimize the element number. 8-node quadratic brick elements
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(13497 nodes and 11450 elements for SENB, 14014 nodes and 11890 elements for SENT) with full integration (ABAQUS: C3D8) are used for FE simulations. The stress-strain response of the material is employed in the 3D models using the true stress-strain curves obtained from the smooth round bar tensile tests at corresponding temperatures.
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Fig. 3. 3D FE models (1/4 model). (a) SENB with a/W=0.5; (b) SENT with a/W=0.3. (c) Mesh arrangement around the notch tip.
Fig. 4. Finite element mesh of the MBL model. A modified boundary layer (MBL) model solution with T=0 is adopted
herein to represent the reference stress field in calculating the Q-parameter for each case. Due to symmetry, only one half of the model has been used in the MBL model. Fig. 4 shows the global FE mesh for MBL model. The same mesh arrangement in front of the crack tip (with a notch root radius of r = 2 μm) as been used in 3D models is applied in the MBL model. RESULTS AND DISCUSSION This section presents the results of load-CMOD curves obtained both experimentally and numerically, CTOD-values at both temperatures, CTOD-CMOD curves, as well as the applicability of RKR model as local fracture criterion in prediction of cleavage fracture toughness. Load-CMOD Curves The results of load-CMOD curves obtained from both experimental measurement and 3D FE calculation for both SENB and SENT specimens at -90 an℃ d -30℃ are displayed in Fig. 5. Quite good accordance between experiments and numerical simulations has been observed for the load-CMOD curves at both temperatures. The material becomes quite brittle at lower temperature, -90℃, for all specimens. The fracture surface is rather flat and no significant crack growth has been observed from optical microscope observations. For the cases of -30 , only small amounts of crack ℃ extension (mostly less than 0.2 mm) have been observed for both SENB and SENT specimens. No crack growth has been considered in the 3D models for the sake of simplicity. One thing should be noted is that the average crack depth for both SENB and SENT specimens at each temperature is used in the 3D simulation. In addition, it can be seen that the transferability of true εσ − from round thermal simulated tensile bar to the fracture mechanics specimens is quite good.
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Fig. 5. Comparison of measured and calculated load-CMOD curves for all specimens at -90 and ℃ -30℃. CTOD-Temperature The results of fracture toughness (CTOD-value) for all specimens are presented in Fig. 6. It can be seen that, the fracture toughness tends to be scattered at both temperatures. Also, the scatter increases with increasing temperature for both SENB and SENT specimens, which can be explained as a result of the cleavage fracture initiation is preceded by ductile crack extension as have been observed from tests.
Fig. 6. CTODs for CGHAZ in X80 at -90 and ℃ -30℃. On the other hand, the average fracture toughness values at both test temperatures are higher for SENT specimens with shorter crack of a/W=0.3 compared to the SENB specimens with a/W=0.5 as plotted by the solid lines in Fig. 6. Moreover, this difference becomes larger for higher temperature.
The detailed effects of temperature and specimen geometry (as quantitatively characterized by crack tip constraint—Q-parameter) on fracture toughness will be studied in the following. As has been known that, the J-Q methodology (O’Dowd and Shih, 1991) gives a direct measurement of the crack-tip stress field of interest that is related to a reference field, and can therefore describe the evolution of constraint ahead of the crack tip throughout the loading to large-scale yielding, where J sets the deformation level and Q is a stress triaxiality parameter. In this study, the Q-parameter has been used to quantify the crack tip constraint level for each specimen at both temperatures. The Q-parameter (O’Dowd and Shih, 1991) was originally defined as
0
0Re )(
σσσ θθθθ =−
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Q , at 2)//( 0 =σJx , θ = 0. (1)
where θθσ is the opening stress component of interest, 0Re )( =T
fθθσ is
the reference stress component obtained by MBL model solution with T = 0, 0σ is the yield stress, x is the distance from the crack tip along the crack plane (θ = 0). Because of the use of CTOD as the crack driving force in this study, the following definition of Q has been used:
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where specimenθθσ is the opening stress component of specimen at
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Fig. 7. Opening stress distributions ahead of the crack tip at -30℃. (a) SENT; (b) SENB. Only the distribution of the crack tip opening stress (σ22 at 0=θ ) has been studied in this paper. The results of crack tip opening stress
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distribution at different CTODs are presented in Fig. 7 for both SENT and SENB specimens at -30 . The dis℃ tance from the crack tip up to ten times of the CTOD has been plotted. It can be seen that, the opening stress distribution in front of crack tip are nearly parallel to the reference field for SENT specimen, Fig. 7 (a). Whereas, global bending makes the slope of the opening stress distribution gradually deviates from the reference field for SENB specimen, but still quite similar as can be seen in Fig. 7 (b). Similar observations have also been found for specimens at -90 whil℃ e the results are not included herein for the sake of simplicity. Fig. 8 shows the calculated CTOD-Q relationship for both SENB and SENT specimens at -90 and ℃ -30℃. As expected, the Q-parameter decreases with the increase of CTODs. Also, the Q-parameter for SENB specimen is considerably larger than that of SENT at same CTODs for both temperatures, which means the crack tip constraint of SENB specimen is higher than that of SENT as has been previously observed. Additionally, nearly constant Q-parameters are obtained for both temperatures considered at same CTODs. This indicates a weak dependence of temperature on the constraint ahead of crack tip can be expected.
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Fig. 8. CTOD vs. Q for both SENB and SENT specimens at -90 and℃ -30℃. (a) SENB; (b) SENT. CTOD-CMOD Curves In this subsection, the CTOD vs. CMOD relationship gained from both experiments and numerical simulations are plotted in Fig. 9 for both SENB and SENT specimens. It can be seen that, the experimental results for SENB specimens at both temperatures considered in this study can be quite well predicted by 3D FE analyses. As for the SENT specimens, there is less good agreement between experimental measurements and 3D simulations, especially for the higher temperature, -30℃. This is a remaining issue and more efforts are needed in further work.
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Fig. 9. CTOD vs. CMOD from experimental tests and 3D FE analyses for both SENB and SENT specimens at -90 and ℃ -30℃. Applicability of RKR model Once the crack-tip stress fields are determined from the numerical FEA, a fracture criterion is needed to characterize the fracture event. A simple model relating fracture stress σf to fracture toughness through a micro-structurally determined characteristic distance rc has been proposed by Ritchie, Knott and Rice (RKR) (Ritchie, Knott and Rice, 1973), who
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postulated that cleavage failure occurs when the stress ahead of the crack tip exceeds σf over a characteristic distance. The fracture stress σf has been studied extensively (Curry, 1980; Chen, et al, 1990; Wang, et al, 2001; Yang, et al, 2003; Wu and Knott, 2004) and is considered to be nearly independent of test temperature and taken as a parameter of fracture criterion connecting the macroscopic fracture toughness in steel microstructure. The critical distance rc in the RKR model is considered as the size of the fracture process zone and is argued to be of the order of two to three grains dimension (Ritchie, Knott and Rice, 1973; Ritchie, Server and Wulleart, 1979). The RKR model can therefore connect the macroscopic fracture toughness with the local material microstructure through the critical local fracture stress σf. The aim of the present study is to examine the applicability of RKR model as local fracture criterion in prediction of the cleavage fracture toughness of CGHAZ in X80. From FEA, the normal tensile stress σ22 ahead of the notch tip can be obtained, as displayed in Fig. 10 for both SENB and SENT specimens at -90 and ℃ -30℃ for δ = 0.02 mm. It can be seen that, the opening stress distribution in front of the notch tip increases with decreasing temperature for both SENB and SENT specimens, which in return indicates a decrease of fracture toughness with decreasing temperature can be expected. Similar trends have also been observed for various CTOD-values.
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Fig. 10. Opening stress distribution ahead of notch tip at -90 and ℃ -30℃. (a) SENB; (b) SENT. After determining the fracture stress, σf, and the characteristic distance, rc, the fracture toughness of the tested materials as a function of temperature could be predicted from the opening stress distribution in front of crack tip. In general, the characteristic distance has to be determined empirically since no consistent relationship exists between this distance and microstructural parameters, e.g. grain size (Curry and Knott, 1978). The results of CTOD vs. temperature from RKR model together with experimental measurements for both SENB and SENT specimens were presented in Fig. 11, where Fig. 11 (a) for fixing distance with various
fracture stresses whereas Fig. 11 (b) for fixing fracture stress with various distances. In these two pictures, the symbols are measured CTOD-values from tests at both temperatures while the solid and broken lines are the calculated fracture toughness values according to the RKR model. As can be observed from Fig. 11, the RKR model was successful in evaluating the fracture toughness at lower temperature (-90℃) with certain fracture stress at a certain distance (Fig. 11 a) or with certain distance at a certain fracture stress (Fig. 11 b). Similar trends have also been obtained for SENT specimen by varying fracture stresses and fixing distance as in Fig. 11 (a), so does SENB specimen with fixing fracture stress while varying distances as in Fig. 11 (b). However, for the cases at higher temperature, -30 ,℃ where fracture toughness and its scatter sharply increase with increasing temperature, this simple model seems failed to depict this uprising behavior in toughness values.
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Fig. 11. CTOD vs. temperature from RKR model and experimental results at -90 and ℃ -30℃. (a) SENB, rc =0.25mm; (b) SENT, σf =1850MPa.
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Fig. 12. CTOD vs. temperature from RKR model at -90 and ℃ -30℃, σf = 1850MPa, rc = 0.3mm. By assuming the same cleavage mechanism, so as for RKR model can
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operate at both temperatures, the same critical local fracture stress σf and critical distance rc (e.g. σf =1850 MPa and rc = 0.3 mm) in the RKR model has been employed for both SENB and SENT specimens, therefore allowing a connection and comparison to be established between these two specimen configurations. Fig. 12 shows the predicted CTOD vs. temperature according to RKR model for both SENB and SENT specimens together with the average values of CTODs obtained experimentally at each temperature. It can be observed that, the predicted facture toughness is higher for SENT specimens than that of SENB as observed from experimental results (see Fig. 6). Additionally, RKR model shows a much better prediction of fracture toughness for SENB than SENT specimens at both temperatures. Discussion In order to picture the scatter in fracture toughness values with temperature, a probabilistic treatment, rather than a deterministic one, e.g. Weibull distribution (Beremin, 1983) and Master Curve (ASTM E1921-95, 2004), has been investigated by many researchers. From the previous investigations (e.g. Heerens et al, 1993; Landes, 1993; Anderson et al, 1994), it can be known that, in spite of the fact that, e.g. three-parameter Weibull distribution describes in a good way with the sets of experimental data, it has no theoretical basis. Moreover, the experimental and analytical results demonstrated that the Weibull parameters are strongly influenced by temperature and specimen geometry. Further work is required to incorporate suitable plastic constraint (Ruggieri and Dodds, 1996; Ruggieri, Gao and Dodds, 2000; Gao, Ruggieri and Dodds, 1998; Gao and Dodds, 2000) and/or temperature dependent parameter(s) into the equation for the calculation of the Weibull stress to make it truly transferable between different specimen geometries. On the other hand, the treatment of prediction of fracture behavior in HAZ and weldments is a challenging issue. A more fundamental understanding of the evolution of the stress/strain fields in front of a crack that grows by a ductile mechanism is needed. Anyhow, in order to establish a better understanding between microstructure and the micromechanical parameters, metallographic and fractographic studies, such as the scanning electron microscopy (SEM) investigations on the fracture surface and cleavage initiation sites, the electron backscattering diffraction (EBSD) technique in determining the crystallographic orientations of initiation cleavage facets at different temperatures, should be utilized to give more informative and helpful insight into the physical mechanism according to the microstructural features. CONCLUSIONS It has been found that the transferability of the true stress-strain curves from round thermal simulated tensile bar to the fracture mechanics specimens is fairly well. Temperature shows no significant effect on the material hardening behavior, which indicates the crack tip constraint is less dependent on the temperature as also observed from 3D FE analyses. The predicted load-CMOD curves from 3D models are in good accordance with experimental results. As for the local fracture parameter, as depicted with CTOD-CMOD relationship, the experimental results for SENB specimens can be quite well simulated by 3D simulations. For SENT specimens, there is less good agreement between experiments and simulations especially for cases at higher temperature, but still quite similar. This is a remaining issue and much work is needed to further modify the model (e.g. the influences of local inhomogeneity of microstructure in HAZ and ductile crack propagation) and thereby improving the validity and accuracy of simulations with respect to the experiments. By combining the RKR model as the local fracture criterion, the
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