Estimation of Constraint Factor on the Relationship Between J ...

Estimation of Constraint Factor on the

Relationship Between J Integral and

CTOD for Offshore Structural Steel

Weldments

Dong-Hyun MoonDepartment of Naval Architecture and Ocean Engineering,

Pusan National University,

Busan 609-735, South Korea

e-mail: [email protected]

Deok-Geun KimDepartment of Naval Architecture and Ocean Engineering,

Pusan National University,

Busan 609-735, South Korea

e-mail: [email protected]

Jeong-Soo LeeTechnology Research Institute,

Total Marine Service Co., Ltd.,

Busan 600-814, South Korea

e-mail: [email protected]

Jae-Myung LeeDepartment of Naval Architecture and Ocean Engineering,

Pusan National University,

Busan 609-735, South Korea

e-mail: [email protected]

Myung-Hyun Kim1

Department of Naval Architecture and Ocean Engineering,

Pusan National University,

Busan 609-735, South Korea

e-mail: [email protected]

Offshore structures are exposed to severe operating conditionsbecause energy resource development has recently extendedtoward deeper seabed and lower temperature regions. Hence,fracture toughness evaluation for very thick and high strengthsteels is one of the most important parameters required for thestructural integrity assessment of offshore structures. Fracturetoughness is known as a property which describes the ability of amaterial containing a crack to resist unstable brittle fracture.Crack tip opening displacement (CTOD) and J integral are themost commonly employed parameters as fracture criteria in elas-tic plastic fracture mechanics (EPFM). There have been extensiveresearch efforts to clarify the relationship between CTOD and Jintegral in elastic plastic regime. Plastic constraint factor (PCF)in the relationship between CTOD and J integral can serve as aparameter to characterize constraint effects in fracture involvingplastic deformation. In this regard, the characteristics of the PCFare of significant importance in EPFM analysis. In this study, weevaluated fracture toughness of American Petroleum Institute(API) 2 W Gr. 50 steel in terms of CTOD in various temperaturesusing single edge notched bend (SENB) specimens. Test speci-mens are fabricated by submerged arc welding (SAW) and flux

cored arc welding (FCAW). In addition, CTOD values are com-pared to absorbed impact energy with respect to the weld metal(WM) and heat affected zone (HAZ). Then, we investigated PCFswith respect to several regions of the weldment at various temper-atures. Experimental values of PCFs were calculated and thencompared against the predicted values according to the AmericanSociety for Testing and Materials (ASTM) standard. CTOD valuesof WM by SAW is found to be about three times higher than thatof FCAW at �10 �C, and CTOD values calculated by the ASTMstandard are approximately 30% lower than the CTOD accordingto British Standard (BS). In addition, the maximum of 40% dis-crepancy is observed in PCFs obtained between the experimentand the predicted values according to the ASTM standard. Thismay lead to too conservative fracture toughness estimation for thewelded joints of API 2 W Gr. 50 steel when using PCF by ASTM.Based on the accurate estimated PCF values obtained from thisstudy, it is believed that rational fracture design of offshore struc-tures is possible. [DOI: 10.1115/1.4031668]

1 Introduction

Offshore structures are exposed to severe operating conditionsbecause energy resource development has recently extended to-ward deeper seabed and lower temperature regions. Therefore,offshore structures for developing oil and gas resources demandhigh performance steels with heavy thickness and high strength.FCAW and SAW are mainly used in offshore industry among var-ious welding processes. These welding processes offer severaladvantages to the application in offshore structures. FCAW can beemployed for all welding position, and there is no limit in thematerial thickness to be welded. In case of SAW, welded jointsare known to possess excellent uniformity, ductility, corrosionresistance, and impact strength.

Welding, however, may introduce material imbalance, residualstress, and various defects. They may lead to structural failure(e.g., brittle fracture) and design life reduction [1]. Thus, offshorestructures with welding defects are more prone to brittle fractureat low temperature region. In order to prevent brittle fracture, it isimportant that WM have sufficient fracture toughness. In particu-lar, adequate fracture toughness at HAZ is one of the most impor-tant properties of welded joints in offshore structures for a safetyassurance. In this regard, API recommended practice requiresCTOD value above 0.38 mm for HAZ at �10 �C [2].

Among the various parameters for the evaluation of fracturetoughness, J integral and CTOD are the most widely used andwell-known fracture toughness parameters. They have been devel-oped separately by ASTM and BS to the context of elastic plasticfracture analysis. In addition, there have been extensive researchefforts to clarify the relationship between CTOD and J integral inelastic plastic regime [2–14]. In general, the relationship betweenthe J integral and CTOD has the following form:

J ¼ mrYSd (1)

where J is J integral value, rYS is the yield stress of the material,m is the PCF, d is the CTOD value [6,7]. A constant factor m inthe relationship between CTOD and J integral given in Eq. (1) isknown to be constraint dependent. In other words, a constant mcan serve as a parameter to characterize the constraint effect at acrack tip. Therefore, the study of constant m in the relationshipbetween J and d is important in EPFM analysis [8,9].

In previous research, Wang [10] and Kirk [11] initially pro-posed the formulation for the calculation of m, respectively. Theconversion of the CTOD from J integral in the ASTM standardknown to provide a significantly different value from the conven-tional geometric CTOD calculated according to the BS standard.This discrepancy between ASTM and BS standards is mainly dueto the constant m [12]. Rafael [13] presented that a constant factorm can be defined in terms of crack size and strength mismatchlevel induced by welding. Huang [14] proposed an empirical m

1Corresponding author.Contributed by the Ocean, Offshore, and Arctic Engineering Division of ASME

for publication in the JOURNAL OF OFFSHORE MECHANICS AND ARCTIC ENGINEERING.Manuscript received June 17, 2013; final manuscript received September 3, 2015;published online October 15, 2015. Assoc. Editor: Lance Manuel.

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factor equation as function of geometry and loading levels. How-ever, the dependence of the constant m on temperature, weldment,and HAZ are not clearly considered in previous researches. Inorder to obtain precise CTOD values, the PCF m should be care-fully examined with respect to the region of welded joints andtemperatures.

In this study, we evaluated fracture toughness of API 2 W Gr.50 steel in terms of CTOD at various temperatures using SENBspecimens which were fabricated by SAW and FCAW, respec-tively. In addition, CTOD values are compared to absorbed impactenergy with respect to the WM and HAZ. Then, we investigatedPCFs with respect to several regions of the welded joints and vari-ous temperatures. Finally, experimental values of PCF were calcu-lated and then compared against the predicted values according tothe ASTM standard [15].

2 Experiment

2.1 Material and Specimens. Test specimens for the evalua-tion of fracture toughness were fabricated from welded panels ofAPI 2 W Gr.50 steel. The welded panels with double bevel grooveare fabricated by FCAW and SAW, respectively. Mechanicalproperties of parent material and welding parameters are summar-ized in Tables 1 and 2, respectively.

In order to obtain the mechanical properties of welded joints byFCAW and SAW, tensile test was carried out according to theInternational Association of Classification Societies (IACS)requirements [16]. The tensile test specimens were extracted fromthe face and the rear portion along the thickness direction in thewelded as shown in Fig. 1.

In order to compare fracture toughness of three differentregions of welded joints of API 2 W Gr.50 steel, a machined notchwas fabricated at WM, coarse grain heat affected zone (CGHAZ)and subcritical heat affected zone (SCHAZ) at correspondinglocations. The locations of machined notch are illustrated inFig. 2. The mechanical properties of welded joints by FCAW andSAW are summarized in Table 3.

Moreover, local compressions are applied prior to fatigue pre-cracking so as to relieve residual stress of the fracture toughnessspecimen. Application of local compression to the ligament belowthe machine notch is often sufficient to reduce the welding resid-ual stresses to low and uniform levels and result in the growth ofan acceptably straight fatigue precrack. The single-edge notchedbend (SENB) specimens were fabricated according to BS 7448standard [17] as shown in Fig. 3, and the dimensions are illus-trated Table 4.

2.2 Fracture Toughness Test Procedure. According to BS7448 standard [17], a series of fracture toughness test was carriedout by three-point bending to determine J integral and CTOD,simultaneously. Fatigue precracking was performed at room tem-perature under constant stress ratio (R¼ 0.1) before three-pointbending test. For each specimen, the fatigue precrack was gener-ated in the maximum length of 5 mm, and the final a0=W ratio(including notch) was set to below 0.55. Here, a0 and W are cracklength and specimen width, respectively. The three-point bendingtest was conducted from room temperature to �80 �C in order toobtain CTOD and J integral values at various temperatures. A clipgauge was attached at the machined notch of each specimen tomeasure crack mouth opening displacement (CMOD). The testmachine used in this study is a hydraulic structural test machine

(IST-8800, INSTRON). The three-point test setup is shown inFig. 4.

CTOD value and J integral are calculated according to thefollowing equations according to BS 7448 standard [17]

d ¼ FS

BW1:5� f

a0

W

� �� �2 1� t2ð Þ2rYSE

þ 0:4 W � a0ð ÞVp

0:4W þ 0:6a0 þ z(2)

J ¼ FS

BW1:5� f

a0

W

� �� �2 1� t2ð ÞE

þ 2Up

B W � a0ð Þ (3)

fa0

W

� �

¼3

a0

W

� �0:5

1:99� a0

W

� �1� a0

W

� �� �2:15� 3:93a0

Wþ 2:7a0

2

W2

� �

2 1þ 2a0

W

� �1� a0

W

� �1:5

(4)

where a0 and W are crack length and specimen width, respec-tively. Here, B is the specimen thickness. In addition, t is Pois-son’s ratio, Vp is the plastic component of CMOD and Up is theplastic component of area under plot of force (F) versus CMOD.

On the other hand, CTOD is converted from J integral usingEq. (5) according to ASTM E1820 standard, and J integral isobtained using gpl and Apl in the following equations [15]:

d ¼ J

mrY(5)

J ¼ Jel þ Jpl ¼K2 1� t2ð Þ

Eþ

gplApl

BN W � a0ð Þ (6)

gpl ¼ 3667� 2:199a0

W

� �þ 0:437

a0

W

� �2

(7)

where rY is average of the yield and tensile strength K is stressintensity factor (SIF), gpl is plastic eta factor for J integral calcula-tion by ASTM E1820 and Apl is plastic component of the areaunder plot of force (F) versus CMOD. Here, a0 and W are cracklength and specimen width, respectively. In addition, BN is thespecimen thickness.

A constant factor m in Eq. (5), provided in ASTM E1820 [15],is calculated by the following equation:

Table 1 Mechanical properties of parent material

Material Mechanical properties

API2 W Gr.50

Tensile strength(MPa)

Yield strength(MPa)

Elongation(%)

517 404 35

Table 2 Parameters of welding condition

Weldingprocess

Heat input(kJ/cm)

Current(A)

Voltage(V)

Speed(cpm)

SAW 45 747 29.7 29.8FCAW 15 292 30.6 36.4

Fig. 1 Schematic diagram of tensile test specimen

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m ¼ A0 � A1

rYS

rTS

� �þ A2

rYS

rTS

� �2

� A3

rYS

rTS

� �3

A0 ¼ 3:18� 0:22 a0=Wð ÞA1 ¼ 4:32� 2:23 a0=Wð ÞA2 ¼ 4:44� 0:29 a0=Wð ÞA3 ¼ 2:05� 1:06 a0=Wð Þ

(8)

where m is the PCF, rYS and rTS are yield and tensile strengths ofthe specimen, respectively. In addition, a0 and W are crack lengthand specimen width, respectively.

ASTM and BS standards provide different methodology for cal-culating CTOD. As described earlier, CTOD is converted from Jintegral in the ASTM standard. Therefore, the converted CTODfrom J integral by the ASTM standard may provide a differentvalue from the conventional geometric CTOD based on the BSstandard [17].

3 Results and Discussion

3.1 Preliminary Test Results. Prior to facture toughnesstests, Charpy impact test and hardness test were carried outaccording to IACS requirements [16]. The geometry and dimen-sion of specimen for Charpy impact test are shown in Fig. 5.

The hardness values shown in Fig. 6 represent the average val-ues from a number of indentations. As illustrated in Fig. 6, thehardness values obtained from the center region are higher thanthose of the face and the rear regions in the welded joints. In addi-tion, the hardness values measured from FCAW are higher thanthose from SAW. The reason for this appears to be related to therelatively rapid cooling rate of FCAW due to lower heat inputthan that of SAW [18].

Fig. 2 Machined notch of fracture toughness test specimens

Table 3 Mechanical properties of welded joints

Welding process Tensile strength (MPa) Yield strength (MPa)

SAW 584 506FCAW 637 583

Fig. 3 Schematic diagram of facture toughness test specimen

Table 4 Dimension of SENB specimens

Notchlocations

Thickness(mm)

Width(mm)

Totallength (mm)

Notch length(mm)

WM CGHAZ SCHAZ 75 75 365 33

Fig. 4 Three-point bending test setup Fig. 5 Schematic diagram of Charpy impact test specimen

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Average values of absorbed energy with respect to each weldprocess and to various temperatures are shown in Fig. 7. Theabsorbed energy of the specimen by SAW is higher than that byFCAW. Based on the result, we assumed that this is due to thefraction of acicular ferrite. Typically, the fraction of acicular fer-rite may affect fracture toughness, and the fracture toughness isknown to increase with the increase of the fraction of acicular fer-rite [18]. To investigate the assumption, the fraction of acicularferrite is investigated by scanning electron microscope as sum-marized in Table 5. As expected, FCAW has lower acicular ferritethan SAW, and this is found to be consistent with the assumption.

3.2 Fracture Toughness Test Results. CTOD and J integralvalues with respect to each welding process and various tempera-tures are shown in Figs. 8 and 9. Both CTOD and J integral valuesof SAW at WM are higher than those of FCAW. In case of SAW,higher CTOD values are observed in the order of WM, SCHAZand CGHAZ as illustrated in Fig. 9. As shown in Figs. 8 and 9,the CTOD values at �10 �C are satisfied against the requirementof API RP 2Z [2] while the difference of CTOD values betweenWM and HAZs is big. Typical API RP 2 Z requirement in WMand HAZ is above 0.38 mm. Moreover, the CTOD and J integralvalues at WM exhibit a similar qualitative tendency compared to

the absorbed energy values obtained from Charpy impact test atWM. The ductile to brittle transition temperature (DBTT) withrespect to the welding processes, FCAW and SAW, is also illus-trated in Figs. 8 and 9, respectively. It is observed that for eachspecimen, the fracture toughness dramatically decreases belowDBTT. The DBTT of WM by FCAW and SAW is approximately�20 �C and �70 �C, respectively. The DBTT of CGHAZ andSCHAZ by FCAW is �65 �C and �25 �C, and that of SAW is�40 �C and �60 �C, respectively.

Figure 10 compares the CTOD values calculated based on theequations provided by ASTM [15] and BS standards [17], respec-tively. As shown in Fig. 10, CTOD values calculated by theASTM standard appear approximately 30% lower than those by

Fig. 6 Hardness profile with respect to welding process

Fig. 7 The result of Charpy impact test with respect to various temperatures

Table 5 Fraction of acicular ferrite depending on weldingprocesses

Weldingprocess

Acicularferrite (%)

Grain boundary ferriteand side plate ferrite (%)

SAW 97.25 2.75FCAW 95.29 4.71

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the BS standard. In other words, the ASTM standard tends to givesmaller CTOD values in comparison to the CTOD values calcu-lated by the BS standard. The result is consistent with the previousresearch by Tagawa [12]. In this regard, it is reasonable to assume

that the inconsistency between ASTM and BS standards is mainlydue to the discrepancy in the constant factor m values. In Sec. 3.3,a more detailed discussion of PCFs obtained from this studyfollows.

3.3 PCF. As discussed, we investigated PCFs with respect toseveral different regions of welded joints at various temperatures.Experimental values of PCF calculated from Eq. (1) are comparedagainst the predicted values according to the ASTM E1820 [15]standard as shown in Eq. (8). Figs. 11–13 show the PCFs calcu-lated at WM and HAZs based on Eqs. (1) and (8), respectively. As

Fig. 8 Comparison of fracture toughness with respect to vari-ous temperatures (FCAW)

Fig. 9 Comparison of fracture toughness with respect to vari-ous temperatures (SAW)

Fig. 10 Comparison of CTOD with respect to ASTM and BSstandards

Fig. 11 Variation of the PCFs for various temperatures (WM)

Fig. 12 Variation of the PCFs for various temperatures(CGHAZ)

Fig. 13 Variation of the PCFs for various temperatures(SCHAZ)

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illustrated in Figs. 11–13, the predicted PCFs by the ASTM stand-ard are higher than the experimental PCFs obtained from Eq. (1).It can be seen that the experimental m values differ from the pre-diction made by the ASTM standard, and the discrepancy betweenPCFs obtained from experiment and those by the ASTM standardis approximately 40%. As shown in Figs. 11–13, the variation ofthe predicted PCFs according to the ASTM standard is almostnegligible regardless of the location of welded joints. Hence, thePCFs calculated by the ASTM standard do not effectively reflectthe effect of various temperatures and welded joints.

The average value of PCFs investigated in this work is summar-ized in Fig. 14. As shown in Fig. 14, the experimental PCFs ofSAW are slightly lower than those of FCAW except at CGHAZ.For practical purpose, the PCF values for welded joints of API2 W Gr. 50 steel are suggested to be in the range between 1.1 and1.2 from this study.

4 Conclusions

In this study, fracture toughness of welded joints with API 2 WGr. 50 steel was evaluated with respect to various temperaturesusing SENB specimens. In addition, PCFs are investigated withrespect to several different regions of welded joints such as WM,CGHAZ, and SCHAZ at various temperatures. Experimental val-ues of PCF were calculated by the general relationship between Jintegral and CTOD, and then compared against the predicted PCFvalues according to the ASTM standard. The major findings fromthis study are summarized in the following:

� The fracture toughness of SAW is higher than that ofFCAW. This appears to be related to the fraction of acicularferrite, and the corresponding fractions of acicular ferritefor FCAW and SAW are 95% and 97%, respectively. Incase of SAW, higher CTOD values are observed in the orderof WM, SCHAZ, and CGHAZ.

� CTOD values calculated by the ASTM standard are approx-imately 30% lower than those by the BS standard. The dis-crepancy is due to the lack of sufficient information of thePCFs for welded joints of API 2 W Gr.50 at varioustemperatures.

� The discrepancy between PCFs obtained from the experi-ment and those calculated by the ASTM standard is approx-imately 40%. The variation of the predicted PCFs accordingto the ASTM standard is almost negligible regardless of thelocation of welded joints. In this study, the practical PCFvalues are experimentally obtained and examined againstthe values by ASTM.

� The practical PCF value for welded joints of API 2 W Gr.50 steel is found to be in the range between 1.1 and 1.2.With the PCF values with the improved accuracy found inthis study, it is believed that more rational fracture designfor welded joints of API 2 W Gr. 50 steel is possible.

Acknowledgment

This work was supported by the National Research Foundationof Korea (NRF) grant funded by the Korea Government (MSIP)through GCRC-SOP (No. 2011-0030013). Moreover, this workwas supported by the National Research Foundation of Korea(NRF) grant funded by the Korea Government (MEST) (No.2011-0016804) and the Human Resource Training Program forRegional Innovation and Creativity through the Ministry ofEducation and National Research Foundation of Korea (NRF-2014H1C1A1066889).

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Fig. 14 Variation of PCFs for WM and HAZs

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