simulation results with anisotropy enhanced plasticity and GTN model for X100 suggest that the material behavior can be reproduced well using plasticity model ...
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Proceedings of the Twenty-second (2012) International Offshore and Polar Engineering Conference Rhodes, Greece, June 17–22, 2012 Copyright © 2012 by the International Society of Offshore and Polar Engineers (ISOPE) ISBN 978-1-880653-94–4 (Set); ISSN 1098-6189 (Set)
Failure Modeling of Pipeline X100 Material in Transition Region A. Nonn
C. Kalwa
Salzgitter Mannesmann Forschung GmbH Duisburg, Germany
Europipe GmbH Mülheim an der Ruhr, Germany
Nonn and Kalwa, 2010; Tanguy, Luu, Perrin, Pineau and Besson, 2008). The experimental database comprised e.g. tests on round notched bar specimens and fracture mechanics tests covering a wide range of stress states. The damage evolution induced by the ductile fracture mechanisms of void nucleation, growth and coalescence has been described by the most prominent Gurson damage model (Gurson, 1977), which has been extended by Tvergaard and Needleman (Tvergaard, 1982; Tvergaard and Needleman, 1984) to account for the damage evolution caused by void interaction and coalescence. In the last two decades the Gurson model has been modified frequently to capture different effects, e.g. strain rate (Mühlich, Brocks and Siegmund, 1998), kinematic hardening (Leblond, Perrin, Devaux, 1995; Mühlich and Brocks, 2003; Mear and Hutchinson, 1985) shear-stress dominated state (Jackiewicz, 2011; Nahshon and Hutchinson, 2008; Nielsen and Tvergaard, 2011), anisotropy (Benzerga and Besson, 2001; Tanguy, Luu, Perrin, Pineau and Besson, 2008; Chen and Dong , 2008), void shape (Gologanu, Leblond and Devaux, 1993), etc.
ABSTRACT This paper focuses on the characterization of the fracture performance of X100 material in transition temperature region using both experimental and numerical methods. The ductile fracture has been analyzed using tests on round notched bar specimens and standard fracture mechanics tests performed at room temperature. In previous publications the damage model Gurson-Tvergaard-Needleman (GTN) has been applied and verified by existing experimental data to describe ductile fracture behavior. The brittle fracture and the fracture in temperature transition region have been studied by means of deep and shallow notched SENB specimens at two different temperatures T=80°C and -40°C. Besides elastic-plastic analyses to quantify constraint levels for different initial crack configurations at the onset of cleavage fracture, the brittle failure has been described using modified Beremin model. The influence of the stable crack growth on the cleavage failure probability in temperature transition region has been captured by coupling the ductile fracture model (GTN) with the modified Beremin model. Finally, examples have been presented for the practical application of the numerical results on the fracture assessment of the flawed high-strength pipelines.
Due to the manufacturing process, the X100 material displays anisotropy of material properties (strength and toughness), which should be taken into account in the plasticity and damage model. The simulation results with anisotropy enhanced plasticity and GTN model for X100 suggest that the material behavior can be reproduced well using plasticity model for plastic flow and isotropic damage model (Tanguy, Luu, Perrin, Pineau and Besson, 2008). The numerical study by Nonn and Kalwa (2010) has shown the importance of considering the volume fraction of secondary voids fN in the GTN model to describe the ductile fracture behavior of high-strength X100 steel. Within this study a reasonable estimation of the ductile fracture has been achieved with the GTN model by using the same damage parameter set for different types of specimens, e.g. round notched bar and fracture mechanics specimens. This result seems to be surprising given the outcome from several studies (Brocks, Sun and Hönig, 1995; Benzerga, Besson and Pineau, 1999; Pardoen and Hutchinson, 2000; Zhang, Thaulow and Odegard, 2000), which show the dependence of the critical porosity on stress triaxiality and strain hardening.
KEY WORDS: ductile-brittle fracture; high-strength pipeline material; damage model
INTRODUCTION The increasing exploration of natural gas resources in remote areas with harsh climate has lead to the development of linepipe steel grades, which should demonstrate sufficient strength and toughness properties also in low temperature regions under various loading conditions. Although the governing failure mechanisms for the steel material have been studied intensively over the last decades, there is still missing comprehension about the influence and contribution of different microstructure entities on the onset of ductile and brittle failure especially for the newly developed high-strength linpipe steel grades, such as X100 material.
Besides the characterization of ductile fracture, the use of the highstrength pipelines in low temperature region requires the quantification of the fracture resistance against the brittle fracture. In the past decades many papers have focused on developing adequate models to describe the cleavage fracture mechanism in metallic materials. The most popular brittle fracture model has been Weibull-based Beremin model (Beremin 1983; Mudry, 1982) which describes the cleavage fracture
The ductile fracture of this material has been investigated recently experimentally and numerically (Ishikawa, Sueyoshi and Igi, 2010;
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probability by incorporating the weakest link theory. The application of the original Beremin model in several numerical studies (Chaouadi, De Meester, van Walle, Fabry and Van de Velde, 1995; Wiesner and Goldthorpe, 1996; Bernauer, Brocks and Schmitt, 1999) has shown that the model parameters are dependent on the geometry, loading rates and temperature. Hence, the modifications of this model have been proposed in order to capture the influence of the local stress state on the failure probability due to the changing constraint, strain rate or temperature conditions. The improvement of the Beremin model prediction has been reached in most cases by introducing the threshold local stress value σth (Bordet, Karstensen, Knowles and Wiesner, 2005; Bakker and Koers, 1991; Gao, Ruggieri and Dodds, 1998; Petti and Dodds, 2005) by accounting for the plastic strains (Tanguy, 2001) or cracked carbide population (Bernauer, Brocks and Schmitt, 1999) related to the void volume nucleation in the Weibull stress σw definition. Further modification of the Beremin model refers to the formulation of the Weibull parameter, scaling factor σu, in dependence of the temperature. Since this parameter characterizes the material resistance against cleavage fracture, it is assumed that σu will increase due to the facilitated dislocation movements with growing temperature (Rossoll, 1998; Petti and Dodds, 2005). However, as indicated in (Bernauer, Brocks and Schmitt, 1999), there are no consistent statements about the temperature σu dependence (Wiesner and Goldthorpe, 1996).
MATERIAL AND MICROSTRUCTURE Mechanical Properties The X100 material for the investigation originates from large diameter pipe section with outer diameter OD=48” and wall thickness wt=18.4mm. The ferritic-bainitic microstructure of the pipe results from the thermo-mechanical controlled rolling and accelerated cooling (TMCP process) during plate production. The final mechanical properties are achieved in subsequent production of the longitudinally welded pipe in UOE process at EUROPIPE GmbH. The yield Rp0.2 and tensile strength Rm including the hardening exponent nH have been determined by means of tensile tests on the round bar specimens at temperatures T=+20°C and -60°C, see Fig. 1. The measured stressstrain curve at T=-60°C displays discontinues yielding, which leads to an evident dip after extending the curve beyond the necking point. Based on the true stress-strain curves at these two temperatures, additional flow curves have been extrapolated also at T=-40°C and -80°C. These curves will be used as input data for the numerical calculations. It should be noted that the material anisotropy has not been characterized and the mechanical properties have been provided only in transverse T direction. 1200
While, in general, the fracture of the steel material has been well understood and described, there is still insufficient knowledge about the fracture performance of the high-strength pipeline steels, such as API grade X100 in the ductile-brittle transition temperature. In order to close this gap, this paper will focus on the characterization of the fracture behavior of X100 in low shelf and transition region by using both experimental and numerical methods. In the first part, the mechanical properties are determined in terms of flow curves at different temperatures, which will be used as an input for numerical analyses. Since the analyses of the ductile fracture has been already conducted in the previous studies (Nonn and Kalwa, 2010) by fracture mechanics tests on the shallow and deep cracked SENB specimens, the same specimen types are used to quantify the fracture resistance of X100 in low shelf and transition region. Subsequently, the experimental results have been evaluated statistically using Master-curve concept with determination of the reference temperature T0. After the calibration of FE models, the numerical simulations have been performed to calculate the evolution of constraint parameters with increasing crack tip loading and to obtain the constraint-corrected toughness values. Furthermore, the cleavage failure probability has been estimated by Beremin model for different temperatures and geometries. By coupling Beremin with GTN damage model, it should be possible to take into account the influence of the stable crack growth on the failure in the transition region. Finally, the last part of the paper contains the engineering critical assessment (ECA) of the flawed pipe.
T=+20°C T=-40°C T=-60°C T=-80°C
True stress [MPa]
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Steel
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X100
Temp. Rp0.2 Rm Rp0.2/Rm nH [°C] [MPa] [MPa] [%] [---] 20
756
821
92
0.069
-60
784
833
94
0.080
600 0.0
0.2
0.4 0.6 0.8 True plastic strain [-] Fig. 1 Flow curves and mechanical properties of X100
1.0
The microstructure of X100 steel consists of ferrite grains and bainite bands. The islands of martensite-austenite constitutes (MA) and small carbides (Fe3C) have been observed by scanning electron microscope (SEM) on nital etched surfaces, see Fig. 2. As expected from the chemical composition, inclusions of the type calcium sulphide (CaS), titanium nitride (TiN) and aluminum oxides (Al2O3) have been identified using EDX analysis. All these detected second phase particles can be evaluated as potential sites for the ductile and cleavage fracture initiation depending on the temperature level.
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30μm
1:500
1μm
Master Curve Approach
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According to DNV-OS-F101-2010, the characteristic (equivalent) CTOD value for the application of the ECA corresponds to the lowest value from at least 3 test results. However, this methodology for the determination of characteristic values might not be justified with respect to the distinct scattering of the fracture mechanics test data especially in the transition region. One possibility to consider the scatter of the test data in the lower transition region is given by Master Curve approach, which enables statistical evaluation of test data and definition of failure probability levels. The major advantage of this approach lies in the calculation of reference temperature T0, based on which the distribution of the characteristic values can be calculated over transition region. At least 6 valid KJc values obtained at one test temperature or the equivalent thereof are required to derive T0, see ASTM 1921.
martensite iron carbides retained austenite ferrite
Fig. 2 Microstructure of X100 from LOM (left) and SEM (right) analyses
EXPERIMENTAL DATA SENB Fracture Mechanics Tests
The Master-Curves obtained from the evaluation of the fracture mechanics test results for a/W=0.5 and 0.2 at T=-80°C are presented in Fig. 3 and Fig. 4, respectively. The resulting T0 temperature for deepcracked bend specimens is 7°C lower than the testing temperature. For the shallow-cracked configuration, the value of T0 shifts toward lower values by -40°C manifesting significant influence of the constraint reduction on the fracture toughness.
The cleavage fracture of X100 has been analyzed by means of fracture mechanics tests on 20 SENB specimens at T=-80°C and 16 SENB specimens at T=-40°C. In order to quantify the influence of the local stress state on the fracture behavior, one half of the test specimens has been machined with deep (a/W=0.5) and the other half with shallow (a/W=0.2) notch. The dimensions of all SENB specimens are given by square cross section 16x16[mm²].
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The fracture mechanics specimens are sampled in T-S direction, thus no appearance of separations is expected on the fracture surfaces. All deep notched specimens have failed in a brittle manner with CTOD values varying between 0.014 and 0.076mm. By using SEM analysis no ductile fracture characterized by dimples/honeycomb structure has been observed on the fracture surface. Nevertheless, the stretch zone width can be measured associated with the crack tip blunting and amounts up to Δa=0.09mm.
Exp. T0 (MML)
0.5
KJmat [MPa*m ]
400
With low constraint geometry (a/W=0.2), the range between minimum and maximum CTOD value (0.015-0.291mm) increases distinctly when compared to a/W=0.5 due to the failure occurrence in the transition region. While critical toughness values have been determined with 7 specimens failing in brittle manner, the unstable toughness values result from the remaining 3 specimens showing ductile initiation and ductile crack growth with Δa=0.13-0.25mm prior to onset of cleavage fracture.
95% 50%
300
5%
200
100 X100, T0 = -86.7°C, SENB16x16, a/W=0.5
0 -200
-150
-100 -50 Temperature T [°C]
0
50
Fig. 3 Master-Curve for a/W=0.5, T=-80°C 500
The shift of the test temperature from T=-80°C to -40°C leads to the shift from lower to upper transition region and increasing scatter in the experimental CTOD values for a/W=0.5. The resulting CTOD values lie in the range between minimum value of 0.038mm (Δa=0.4mm) and by factor 10 higher maximum value of 0.38 (Δa=0.36mm).
95% 50% Exp. T0 (MML)
0.5
KJmat [MPa*m ]
400
Although upper shelf behavior has been expected by using SENB a/W=0.2 instead of a/W=0.5 at T=-40°C, all toughness values have been determined in the upper transition region. The specimens display significant amount of stable crack growth prior to onset of instability leading to high CTOD values in the range between 0.37mm (Δa=0.41mm) and 0.96mm (Δa=1.07mm). As evident, the minimum CTOD value for a/W=0.2 corresponds approx. to the maximum CTOD value for a/W=0.5 so that there is almost no overlapping in the values for both configurations like in case of T=-80°C.
5%
300
200
100 X100, T0 = -127.5°C, SENB16x16, a/W=0.2
0 -200
-150
-100 -50 Temperature T [°C]
0
50
Fig. 4 Master-Curve for a/W=0.2, T=-80°C According to following equation (Wallin, 2001), which has been included in the FITNET procedure, the reference temperature T0cal for SENB a/W=0.2 can be estimated to -113°C due to the constraint effect:
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T0cal ≈ T0 deep +
Tstress °C with 10 MPa
On the other hand, the determined T0 values for a/W=0.2 with test data at T=-80°C (T0=-127.5°C) and T=-40°C (T0=-125.7°C) deviate only slightly by 1.8°C. Although the standard ASTM 1921 requires 6 valid test results at one temperature for calculation of T0, it should be noted that more significant results can be achieved when considering the complete test data from both tests at T=-80°C and -40°C. In case of a/W=0.5, the difference in T0 values between test data at T=-80°C (T0=-94°1C) and T=-40°C (T0=-86°7C) is with 7°C higher than for a/W=0.2. Nevertheless, this difference lies in the range of the expected deviation.
(1)
T0deep-temperature obtained from high constrained specimens Tstress-constraint parameter The difference between the calculated and Master-Curve evaluated T0 values is ΔT=14.5°C. Hence, by applying the Eq. 1 with T0deep of SENB a/W=0.5, the calculated reference temperature T0cal for SENB a/W=0.2 is by 14.5°C higher than experimentally determined T0 and can be used for the conservative estimation of the fracture toughness values while still including constraint benefits.
By using the Eq. 1 the reference temperature T0cal for SENB a/W=0.2 is calculated to -119°C. While in case of T=-80°C, the T0 value has been overestimated by 14°C, the Eq. 1 provides a better approximation to Master-Curve temperature with only 8°C. Hence, regarding the toughness values for the ECA according to FITNET procedure, the reference temperature T0cal=-119°C determined at T=-40°C is suitable to obtain more accurate but still conservative results.
The Master Curve approach has also been employed to determine the reference temperature T0 for SENB a/W=0.5 and 0.2 tested at T=-40°C, see Fig. 5 and Fig. 6. The T0 values amount to -94.1°C and -125.7°C for a/W=0.5 and 0.2, respectively. Hence, the constraint reduction leads to by 8°C lower shift in T0 values (ΔT0=32°C) when compared to the shift obtained from tests at T=-80°C (ΔT0=40°C).
NUMERICAL ANALYSES
500 95%
Evaluation of the local stress state 50%
5%
In order to illustrate the advantage of lower constraint geometry for providing more realistic stress state and thus toughness values, elasticplastic FE calculations have been performed for both deep and shallow notch geometries. As a deliverable of these calculations, the local stress state has been quantified in terms of constraint parameter Q (constraint analyses). The latter parameter can be used within the FITNET procedure for the constraint corrected safety assessment of flawed linepipe.
Exp. T0 (MML)
0.5
KJmat [MPa*m ]
400
300
200
100
50
The required 3D FE models have been created using Abaqus/Standard and meshed with 8-nodes solid elements (C3D8). Fig. 7 shows the typical FE mesh with especially fine meshed area around the crack tip and minimum element size of 0.005mm. By considering the symmetric conditions in length and thickness directions, the analysis has been conducted only on a fourth of the specimen.
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Fig. 7 FE model with typical mesh (left) and plastic zones (red coloured areas) (right) of SENB specimen with a/W=0.2
X100, T0 = -94.1°C, SENB16x16, a/W=0.5
0 -200
-150
-100 -50 Temperature T [°C]
0
Fig. 5 Master-Curve for a/W=0.5, T=-40°C 500 95%
50% 5%
Exp. T0 (MML)
0.5
KJmat [MPa*m ]
400
300
200
100 X100, T0 = -125.7°C, SENB16x16, a/W=0.2
0 -200
-150
-100 -50 Temperature T [°C]
0
Fig. 6 Master-Curve for a/W=0.2, T=-40°C
The load-deformation behavior can be well reproduced by numerical calculations, see Fig. 8 for SENB a/W=0.5 at T=-80°C. A good correspondence between global experimental and numerical results has to be provided in order to proceed with subsequent analyses of the local stress/strain field.
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(Q=-0.14) in the standard SENB specimen can be maintained for the crack tip loading J
