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Influences of welding processes on fatigue life of cruciform joints of pressure vessel grade steels containing LOP defects
Mechanics of Materials 32 (2000) 265±276
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In¯uences of welding processes on fatigue life of cruciform joints of pressure vessel grade steels containing LOP defects V. Balasubramanian *, B. Guha Mechanical Testing Laboratory, Department of Metallurgical Engineering, Indian Institute of Technology, Madras, Chennai 600 036, India Received 5 October 1998; received in revised form 7 September 1999
Abstract The in¯uences of two welding processes, namely, shielded metal arc welding (SMAW) and ¯ux cored arc welding (FCAW), on fatigue life of cruciform joints, containing lack of penetration (LOP) defects, have been studied. Load carrying cruciform joints were fabricated from high strength, quenched and tempered steels of pressure vessel (ASTM 517 ÔFÕ) grade. Fatigue crack growth experiments were carried out in a mechanical resonance vertical pulsator (SCHENCK 200 kN capacity) with a frequency of 30Hz under constant amplitude loading (R 0). It was found that the fatigue lives of the cruciform joints fabricated by the SMAW process were relatively higher than the FCAW counterpart. Ó 2000 Elsevier Science Ltd. All rights reserved. Keywords: Shielded metal arc welding; Flux cored arc welding; Cruciform joint; Lack of penetration; Fatigue life
List of symbols 2a initial LOP length 2W ®llet width 2B specimen width L leg length plate thickness Tp h ®llet angle Dr stress range R stress ratio (rmin =rmax ) crack initiation life Ni Np crack propagation life total fatigue life Nf initial stress intensity factor (SIF) DKi range
* Corresponding author. Tel.: +91-44-2351-365/3801; fax: +91-44-2350-509. E-mailaddress:[email protected](V.Balasubramanian).
m C
Paris constant (crack growth exponent) intercept in Paris equation
1. Introduction The ®llet welded cruciform joints are most common ones in various structures including oshore and nuclear applications. Although the quality of welding has improved over the past decades, welding discontinuities are still unavoidable. The sizes of internal discontinuities present in the weld has a large eect on the measured fatigue life. The most important variable in determining the fatigue life of a ¯awed weldment would seem to be the nature of internal ¯aws contained within the weld and the manner in which these ¯aws interact within the stress ®eld in and around the weld during its fatigue life (Lawrence and Radziminski,
0167-6636/00/$ - see front matter Ó 2000 Elsevier Science Ltd. All rights reserved. PII: S 0 1 6 7 - 6 6 3 6 ( 9 9 ) 0 0 0 4 9 - 6
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1970). Linking the eects of weld defects and failure analysis of weldments pointing towards that the fatigue alone is considered to account for most of the disruptive failures and often precedes the onset of brittle failure (Tsai, 1986). The fatigue resistance of the weld metal and heat aected zone of various steels are better than or equal to the base metal (Vosikovsky, 1980). However, problems arise when there is an abrupt change in section by excess weld reinforcement, undercut, inclusion of slag or lack of penetration (LOP) or fusion (Blodgett, 1992). Many of the fatigue failures that occur in welded joints involve fatigue cracking from severe imperfections which are actually an inherent part of the joint (Maddox, 1993). There are two types of cracking in a ®llet welded joints: (i) root cracking, and (ii) toe cracking. Fatigue cracks initiate at the ®llet weld toe when the ®llet weld size is large enough and initiate at weld root when weld size is inadequate (Gurney, 1979). The type and severity of weld defects are in¯uenced by the welding process itself. Automatic welding processes are favourable over manual processes for several reasons such as increased productivity, lower costs, etc. However, a twofold overall increment in fatigue lives was observed for manual welds with respect to comparable automatic welds. This was mainly attributed to a higher stress concentration eect as a result of poor weld pro®les in case of automatic welding, e.g., submerged arc welding (Otegui et al., 1991). Further, the mechanised submerged arc welding in average gives the steepest weld to base metal transition regions. But the manual and semi-automatic processes produce welds with a more smooth transition to the base metal (Overbeeke et al., 1991). Moreover, in automatic welding process, the mis-alignment is an important manufacturing defect. The bending stresses induced by the misalignment of welded parts signi®cantly in¯uence the fatigue lifes of weldments by decreasing mainly the fatigue crack initiation lives. This aects mostly the fatigue resistance of external discontinuities such as weld toes but only slightly reduces the fatigue resistance of internal discontinuities such as LOP. Therefore, a weld which should fail at an internal discontinuity may fail at
an external discontinuity because of bending stresses (Jakubcsak and Glinka, 1986). Fatigue life prediction of welded joints is very complex, costly and time consuming. This is due to its complex joint geometry, number of stress concentration points and heterogeneous weld metal property making the joint. So to avoid the costly and complex procedure, traditionally, the fatigue life of the joint for structural applications followed the S±N type of approach covered by BS 5400 and IIW (Hobbacher, 1988). But for critical structural applications where both initiation and propagation behaviour are equally important for the purpose of safety, then the fracture mechanics approach is more appropriate in the place of traditional methods to predict the fatigue life of the component (Allen et al., 1988). Hence, it is customary to predict the fatigue life of welded joint with defect in terms of crack growth parameters such as da/dN against DK obtained by crack growth experiment. Practically, these types of data merely indicate the fatigue crack growth behaviour of the weldments and do not predict the actual fatigue life. From the literature (Smith and Smith, 1983; Branco et al., 1985; Testin et al., 1987), it is evident that the most of the investigations on fatigue life prediction of the ®llet welded joints are based on toe failure. Very few investigators (Usami and Kusumoto, 1978; Frank and Fisher, 1979; Guha, 1994) have studied the fatigue behaviour of ®llet welded joints failing from the root region. Hence, this investigation has been carried out to study the in¯uence of welding processes, namely shielded metal arc welding (SMAW) and ¯ux cored arc welding (FCAW) processes, on fatigue life of cruciform joints, failing from the root LOP region. 2. Experimental Pressure vessel grade steel (ASTM 517 ÔFÕ) of weldable quality in the form of rolled plates of 8 mm thickness have been used as the base material throughout the investigation. These steels are used for welded construction of all kinds such as pressure vessels, penstocks, bridges and structures as well as transport vehicles, hoisting and
V. Balasubramanian, B. Guha / Mechanics of Materials 32 (2000) 265±276
267
various root faces enabled the joints to have dierent LOP lengths after welding. The ®llet leg lengths were varied from 6 to 8 mm and LOP length between 7 and 8 mm. The dimensions of cruciform joints are shown in Fig. 1. All necessary care was taken to avoid joint distortions and the joints were made without applying any clamping devices. The fatigue crack growth experiments were conducted in a mechanical resonance controlled vertical pulsator (200 kN capacity) with a frequency of 30 Hz under constant amplitude loading (R 0). Before loading, the specimen surface near the LOP was polished to enable the crack growth measurement. A travelling microscope was used to monitor the crack length with an accuracy of 0.01 mm. The specimen was loaded at a particular stress level (range) and crack initiation and its subsequent propagation from LOP defect was recorded from time to time until complete failure of the specimen. A similar crack growth experiment was conducted on a number of specimens at various stress levels and experimental data (crack initiation life, crack propagation life and total fatigue life) were recorded. Table 1 shows the chemical composition of base metal and weld metal. The mechanical properties of the base metal and weld metal are given in Table 2. The welding process parameters used to fabricate the joints are presented in Table 3.
earthmoving equipments which are utilised in dierent climatic conditions (Balasubramanian and Guha, 1998). The rolled plates were cut into the required sizes and pro®les by oxy-fuel cutting and grinding. The initial joint con®guration is obtained by securing the long plates (200 100 mm2 ) and stem plate (200 75 mm2 ) in a cruciform position by tack welding. Subsequently, the ®llets were made between the long plate and stem plate laying weld metal using two dierent welding processes with matching weld metal consumable. SMAW and FCAW processes were chosen to fabricate the joints and multi-pass welding procedures were employed. All the four ®llets forming the joint were made identical leaving an unfused gap between the pair of ®llets. This gap, i.e., LOP was controlled by providing proper root faces, obtained by a prior machining process, known as bevelling (Guha, 1994). The
3. Results and discussions Although the fatigue crack growth experiments were conducted on a large number of specimens having dierent dimensions (a=W 0:3 to 0.4, L=Tp 0:6 to 1.0), only one set of results has been considered here for developing the equations as well as for comparing the two processes in a better
Fig. 1. Dimensions of cruciform specimen. Table 1 Chemical composition (wt%) Material
C
Si
Mn
P
S
Cr
Mo
Ni
Cu
Co
V
Base metal Weld metal (i) SMAW (ii) FCAW
0.19
0.72
0.95
0.01
0.002
0.8
0.35
0.07
0.03
0.004
0.002
0.08 0.08
0.5 0.4
1.6 1.5
0.01 0.02
0.03 0.01
0.35 0.56
0.27 0.44
1.73 2.25
0.1 0.2
0.01 0.02
0.001 0.002
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V. Balasubramanian, B. Guha / Mechanics of Materials 32 (2000) 265±276
Table 2 Mechanical properties Material
Yield strength (MPa)
Tensile strength (MPa)
VickerÕs hardness (v30 kg)
Impact value (J)
Percentage of elongation (%)
Base metal Weld metal (i) SMAW (ii) FCAW
690
790
210
110
19
740 720
845 830
320 280
175 150
24.5 26
Table 3 Welding process parameters Process
Electrode speci®cation (AWS)
Electrode diameter (mm)
Voltage (V)
Current (A)
Welding speed (mm/s)
Heat input (kJ/mm)
(i) SMAW (ii) FCAW
E11018-M E100T5 K4
3.15 1.6
24 30
120 240
2.5 4.0
1.1 1.8
way. The dimensions of the test specimens, taken into account in the present investigation, are as follows: 2a 7:5, 2W 21, L 6:5, Tp 8 mm and h 45 . Fatigue crack growth experiments were conducted at four dierent stress ranges (Dr), i.e., 120, 160, 200 and 240 MPa, and the variations in crack length (2a) with the corresponding number of cycles (N) are plotted as shown in Fig. 2, for SMAW and FCAW joints. From the ®gures, it is evident that the crack growth is comparatively slower in the joints fabricated by the SMAW process than in the FCAW process. 3.1. In¯uence of welding process on crack propagation life (Np ) The fracture mechanics analysis is based on Paris power law (Paris and Erdogan, 1963), da=dN C DKm ;
After re-arranging the equation
da =dN C=2 f m a Drm W m=2ÿ1 ;
where f a is the normalised SIF range which is DK=Dr W 1=2 , and a is the normalised crack length : a=W. The expression for SIF range (DK), at the apex of a root (LOP) defect of a load carrying cruciform joint developed by Frank and Fisher (1979) is DK
m
d 2a=2W C DK Drm ; dN Drm 2W
2
where Dr is the nominal stress range, 2a the lack of penetration defect size, and 2W is the ®llet width as shown in Fig. 1.
Dr 1=2 A1 A2 a pa sec pa =2 ; 1 2 L=Tp 4
where L/Tp is the weld size as de®ned in Fig. 1. A1 and A2 are functions of weld size (L/Tp ): A1 0:528 3:287 L=Tp ÿ 4:361 L=Tp 3
2
4
3:696 L=Tp ÿ 1:874 L=Tp 0:415 L=Tp
1
where da=dN is the crack growth rate, DK the stress intensity factor (SIF) range, and C and m are constants. Eq. (1) can be normalised and adopted for cruciform joints, as in the form given below
3
5
and A2 0:218 2:7717 L=Tp ÿ 10:171 L=Tp
2
13:122 L=Tp 3 ÿ 7:775 L=Tp 4 5
1:785 L=Tp : By normalising and re-arranging, the above equation is simpli®ed (Guha, 1995) as follows: DK Dr f aW 1=2 ;
5
V. Balasubramanian, B. Guha / Mechanics of Materials 32 (2000) 265±276
269
Fig. 3. Eect of L/Tp ratio.
cycles than the FCAW joints, at a particular SIF range value. The term of DK in Eq. (5) for a cruciform joint can be inserted in Eq. (2) to get da Cf m a Drm W m=2 : dN 2W
6
After re-arranging the terms for integration we have Z Z a f da =f m a C=2 Drm W m=2ÿ1 dN ; 7 ai
Fig. 2. Crack growth curves.
where A1 A2 a pa sec pa =21=2 : f a 1 2 L=Tp Fig. 3 shows the normalised SIF range f a vs crack length a , for dierent values of weld size, L/Tp . The calculated values of SIF range of both the processes, for the growing crack, have been plotted in Fig. 4. From the ®gures, it is observed that the SMAW joints endure a larger number of
where ai is the initial defect size after initiation cycles (Ni ) and af is the ®nal defect size at failure, or Ip C=2 Drm W m=2ÿ1 Np ;
8
where Ip is the value of integration for propagation of cycles leading to failure, and Np is the number of cycles to crack propagation. Re-arranging the equation Np 2Ip =C Drm W m=2ÿ1 :
9
The integration was performed on normalised SIF range, f a between initial, ai , to ®nal crack length af with an increment of 0.01. The integration was done for various crack growth exponent
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V. Balasubramanian, B. Guha / Mechanics of Materials 32 (2000) 265±276
Fig. 5. Relationship between Ip and a .
crack length increment against the associated number of cycles to propagation. For all calculations, the ASTM E-647 guidelines were followed. The relationship between the SIF range (DK) and the corresponding crack growth rate, d(2a)/dN on a log±log scale in terms of the best ®t line (BFL), is shown in Fig. 6, for both the processes. The data
Fig. 4. SIF range values for the growing crack.
values (m) and it was also found that the value of integration, Ip , is a function of crack length, a , for all values of m as shown in Fig. 5 in a double log plot. 3.2. Evaluation of m and C The crack growth rate, da=dN , for propagation stage, was calculated considering the slope at the steady-state growth regime at dierent intervals of
Fig. 6. Crack growth rate curves.
V. Balasubramanian, B. Guha / Mechanics of Materials 32 (2000) 265±276
points mostly correspond to the second stage of the sigmoidal relationship of Paris equation (1). The exponent m which is the slope of the line on the log±log plot has been found to be 3.5 and 4.2 for SMAW and FCAW joints, respectively, and the corresponding value of constant C which is the intercept of the line on log±log plot has been found to be 7:6 10ÿ10 and 2:3 10ÿ10 . At a ®xed value of DK, the FCAW joints are showing higher crack growth rate than the SMAW joints. 3.3. In¯uence of welding process on crack initiation life (Ni ) The crack initiation life Ni was evaluated experimentally using the crack ``initiation criteria''. The initiation criterion was based on the assumption that the number of cycles required to grow 1 mm length of crack in excess of its original LOP length at the earlier crack growth stage under particular stress range (Guha, 1995). Similar criteria were adopted by other investigators (Jack and Price, 1971; Testin et al., 1987). Fig. 7 shows the relationship between initial SIF range (DKi ) and crack initiation life (Ni ). It is evident from the ®gure that the initiation lives of both the joints are
271
more or less similar because the crack initiation region (a distance of 1 mm from the LOP tip) is composed of weld metal that is deposited by the SMAW process, commonly known as root run or root pass. The Paris type equation for early crack growth is d 2a=dN Ci DKi m ;
10
where Ci is a constant. For the 1 mm crack growth, the required number of cycles are calculated as 1 mm=Ni Ci DKi Ni 1=Ci DKi Ni C1 DKi
m
ÿm
m
or
or
;
11
where C1 1=Ci . In the above equation, by substituting the experimental values of Ni , the calculated values of Ki and the evaluated values of m, the value of C1 is obtained for both joints and they are presented in the form of initiation life equations: For SMAW process : Ni 1:5 109 DKi
ÿ3:5
; 12
For FCAW process : Ni 6:0 109 DKi
ÿ4:2
: 13
3.4. In¯uence of welding process on total fatigue life (Nf ) The crack growth process is normally comprised of two major stages: (i) crack initiation, and (ii) crack propagation and hence the total fatigue life is considered to be the sum of the cycles spent in these two periods (Lassen, 1990). Therefore, the total fatigue life (Nf ) is evaluated by the following equation: Nf Ni Np :
Fig. 7. Relationship between crack initiation and initial SIF range.
14
Fig. 8 depicts the relationship existing between initial SIF range (DKi ) and the fatigue life values obtained from the experiments. It is evident from the ®gure that though there is not much dierence
272
V. Balasubramanian, B. Guha / Mechanics of Materials 32 (2000) 265±276
sponding initial SIF range (DKi ) values. The fatigue life (Nf ) of the cruciform joint can be predicted by the following equation for the endurance line shown in Fig. 9: Nf 4:2 109 DKi
ÿ3:5
for SMAW joints
15
ÿ4:2
for FCAW joints:
16
and Nf 1:2 109 DKi
in crack initiation life (Ni ) of both the joints, the total fatigue life (Nf ) is showing marked variations which implies that the propagation life (Np ) of both the joints are not same. With the knowledge of crack growth parameters (m and C) and incorporating the crack initiation cycles (Ni ) from Eqs. (12) and (13), the total fatigue lives (Nf ) of both the joints were predicted using Eq. (14) and they are presented as shown in Fig. 9, with corre-
The accuracy of the above developed equations are tested by comparing the predicted fatigue life data with the experimental data and the scatter diagram is shown in Fig. 10. The S±N behaviours of the two weldments are plotted in Fig. 11 (at each stress level, ®ve specimens were tested and the average values are plotted) for the comparison purpose. The variations in the crack growth behaviour and fatigue life of the joints may be due to the dierence in the heat input (welding process parameters) involved during the fabrication. Evans (1982) reported that the following changes in chemical composition, mechanical properties and microstructures were observed as a result of increasing the heat input: (i) the weld metal manganese and silicon content is reduced, (ii) the yield strength and tensile strength decreased, (iii) hardness of as deposited weld metal decreased, (iv) the average width of the columnar grains
Fig. 9. Predicted fatigue life values.
Fig. 10. Scatter diagram.
Fig. 8. Relationship between initial SIF range and fatigue life.
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273
The microstructures of the weldments at two dierent locations, i.e., (i) reheated region of the earlier passes due to subsequent passes, and (ii) non-reheated region of the weld metal, are compared in Figs. 12 and 13. The weld metal microstructure of SMAW process contains a high proportion of ®ne grain acicular ferrite which displays optimum strength and toughness properties. This is attributed to small grain size in which each lath is separated by high angle grain boundaries (Farrar et al., 1974). But in the weld metal microstructure of FCAW process, the formation of large proportion of pro-eutectoid ferrite (grain boundary ferrite), ferrite sideplates or
Fig. 11. S±N curves.
increased, (v) the amount of pro-eutectoid ferrite in weld metal increased at the expense of acicular ferrite, (vi) the lath size of the acicular ferrite increased, and (vii) the grain size of the equiaxed ®ne grained regions increased. A small amount of manganese and silicon is lost through higher oxidation due to high heat input involved in the semi-automatic process. Further, due to high heat input, the cooling rate is comparatively slower in the semi-automatic process and this leads to the softening of the microstructure, as exhibited by the decrease of hardness values (Dixon and Hakansson, 1995). Moreover, the variations in the crack growth behaviour and fatigue life are also due to the multi-passes involved during the fabrication of the joints. In multi-pass deposits, the weld metal consists partly of a tertiary transformation microstructure, i.e., in multi-pass technique, each successive bead tempers previous ones, consequently the secondary microstructures, such as pro-eutectoid (grain boundary) ferrite, side plate ferrite and acicular ferrite, will partly be heated into the region. This extra transformation results in the formation of tertiary microstructure. This structure normally has higher toughness than the non-transformed weld metal, the secondary microstructure (Dolby, 1979; Hoekstra et al., 1986).
Fig. 12. Microstructure of SMAW weld metals: (a) reheated region, and (b) non-reheated region.
274
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Fig. 13. Microstructure of FCAW weld metals: (a) reheated region, and (b) non-reheated region.
martensite has been found to be detrimental to toughness because these structures provide preferential early crack propagation paths, which oer a low resistance during the weld metal cleavage fracture (Levine and Hill, 1977; Tweed and Knott, 1983). Further, the eect of manganese is to promote acicular ferrite at the expense of pro-eutectoid ferrite. Nickel also in¯uences weld metal microstructure through changing the columnar grain size (Zhang and Farrar, 1997). FCAW process weld metals with higher nickel content
Fig. 14. SEM fractographs of SMAW joints: (a) crack initiation, (b) crack propagation, and (c) ®nal fracture.
V. Balasubramanian, B. Guha / Mechanics of Materials 32 (2000) 265±276
275
exhibited poor toughness due to the formation of martensite and the presence of coarsened columnar grains. The fractured surface appearance corresponding to crack initiation, crack propagation and ®nal failure regions as observed by scanning electron microscopy (SEM) are shown in Figs. 14 and 15. The crack initiation regions of both the weld metals are shown in Figs. 14(a) and 15(a), which do not have any distinctive features. Interestingly, the crack propagation region of both the weld metals, compared in Figs. 14(b) and 15(b), shows signi®cant variation in the fracture surface. The fractograph of FCAW joint shows a high density of quasi-cleavage fracture with a larger number of preferential crack initiation sites in the crack propagation region. However, the SMAW joints exhibit a feature of a relatively high energy fracture which shows higher ductility through the presence of more dimples and striations in the crack propagation region compared to FCAW joints. However, the ®nal fracture region, shown in Figs. 14(c) and 15(c), of both the weldments represents the brittle fracture and it is evident from the river pattern of fracture surface.
4. Conclusions
Fig. 15. SEM fractographs of FCAW joints: (a) crack initiation, (b) crack propagation, and (c) ®nal fracture.
1. The fatigue life of SMAW and FCAW cruciform joints of pressure vessel grade steels (ASTM 517 ÔFÕ), containing LOP defects, can be successfully predicted by using the fracture mechanics equations developed in this paper, within a reasonable accuracy. 2. The experimental evidence shows that the microstructural features of deposited weld metal play an important role in fatigue crack growth behaviour and in¯uence the fatigue life of the joints. The joints fabricated by the SMAW process, which contains a high proportion of acicular ferrite in the weld metal microstructre, displayed better fatigue performance than the FCAW joints.
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Acknowledgements The authors are indebted to M/s. Bharat Heavy Electricals Limited (BHEL), Ranipet, Tamil Nadu, India for the material supplied to carryout the investigations. References Allen, R.J., Booth, G.S., Jutla, T., 1988. A review of fatigue crack growth characterisation by linear elastic fracture mechanics Part II. Fat. Frac. Engng. Mater. Struct. 11, 71± 108. Balasubramanian, V., Guha, B., 1998. In¯uence of ¯ux cored arc welded cruciform joint dimensions on fatigue life of ASTM 517 `F' grade steels. Int. J. Pres. Ves. Piping 10, 765± 772. Blodgett, O.W., 1992. Designing weldments for fatigue loading. Weld. J. 7, 39±44. Branco, C.M., Ferreira, J.A.M., Radon, J.C., 1985. Fatigue of ®llet welded joints. Theo. Appl. Frac. Mech. 3, 13±22. Dixon, B., Hakansson, K., 1995. Eect of welding parameters on weld zone toughness and hardness in 690 MPa steel. Weld. Res. Suppl. 4, 122s±132s. Dolby, R.E., 1979. HAZ toughness of structural and pressure vessel steels ± Improvement and prediction. Weld. Res. Suppl. 8, 225s±238s. Evans, G.M., 1982. The eect of heat input on the microstructure and properties of C±Mn all weld metal deposits. Weld. Res. Suppl. 4, 125s±132s. Farrar, R.A., Tuliani, S.S., Norman, S.R., 1974. Relationship between fracture toughness and microstructure of mild steel submerged arc weld metal. Weld Metal Fabrication 42, 68± 73. Frank, K.H., Fisher, J.W., 1979. Fatigue strength of ®llet welded cruciform joints. J. Struct. Div. 105, 1727±1739. Guha, B., 1994. Fatigue life estimate of C±Mn steel welded cruciform joints. Theo. Appl. Frac. Mech. 21, 121±129. Guha, B., 1995. A new fracture mechanics method to predict the fatigue life of MIG welded cruciform joints. Engng. Frac. Mech. 52, 215±229. Gurney, T.R., 1979. An analysis of some fatigue crack propagation data for steels subjected to pulsating tension loading. Weld. Res. Int. 9, 45±52.
Hobbacher, A., 1988. Recommendations for assessment of weld imperfections in respect to fatigue. IIW Doc. IIM-XIII1266-88/XV-659-88. Hoekstra, S., Munning Schmidt-van der Burg, M.A., Ouden, G.D., 1986. Microstructure and notch toughness of ferritic weld metal. Metal Construction 18, 771±775. Jack, A.R., Price, A.T., 1971. The use of crack initiation and growth data in calculation of fatigue lives of specimens containing defects. Metal Construction 3, 416±419. Jakubcsak, H., Glinka, G., 1986. Fatigue analysis of manufacturing defects in weldments. Int. J. Fatigue 2, 51±57. Lassen, T., 1990. The eect of welding process on fatigue crack growth. Weld. Res. Suppl. 2, 75s±81s. Lawrence, F.V., Radziminski, J.B., 1970. Fatigue crack initiation and propagation in high yield strength steel weld metal. Weld. Res. Suppl. 10, 445s±452s. Levine, E., Hill, D.C, 1977. Toughness in HSLA steel weldments. Metal Construction 9, 346±353. Maddox, S.J., 1993. Recent advances in the fatigue assessment of weld imperfections. Weld. J. 7, 42±52. Otegui, J.L., Burns, D.L., Kerr, H.W., Mohaupt, U.H., 1991. Growth coalescence of fatigue cracks at weld toes in steel. Int. J. Pres. Ves. Piping 48, 129±165. Overbeeke, J.L., Alting, T.J., Jonkers, P.A.M., 1991. In¯uence of weld parameters on the fatigue strength under C.A. and V.A. loading of welded joints in HSLA steels. IIW Doc. XII-1404-91. Paris, P.C., Erdogan, F., 1963. A critical analysis of crack propagation laws. J. Basic Engng. 85, 528±538. Smith, I.F.C., Smith, R.A., 1983. Fatigue crack growth in a ®llet welded joint. Engng. Frac. Mech. 18, 861±869. Testin, R.A., Yung, J.Y., Lawrence, F.V., Rice, R.C., 1987. Predicting the fatigue resistance of steel weldments. Weld. Res. Suppl. 4, 93s±98s. Tsai, C.L., 1986. Fitness for service design of welded joints. J. Struct. Engng. 112, 1761±1780. Tweed, J.H., Knott, J.H., 1983. Eect of reheating on microstructure and toughness of C±Mn weld metal. Met. Sci. 17, 45±54. Usami, S., Kusumoto, S., 1978. Fatigue strength at roots of cruciform Tee and Lap joints. Trans. Jpn. Weld. Soc. 9, 1± 10. Vosikovsky, O., 1980. Fatigue crack growth in HY130 steel weldments in air and water. Weld. Res. Suppl. 9, 255s±258s. Zhang, Z., Farrar, R.A., 1997. In¯uence of Mn and Ni on the microstructure and toughness of C±Mn±Ni weld metals. Weld. Res. Suppl. 5, 183s±196s.
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In¯uences of welding processes on fatigue life of cruciform joints of pressure vessel grade steels containing LOP defects V. Balasubramanian *, B. Guha Mechanical Testing Laboratory, Department of Metallurgical Engineering, Indian Institute of Technology, Madras, Chennai 600 036, India Received 5 October 1998; received in revised form 7 September 1999
Abstract The in¯uences of two welding processes, namely, shielded metal arc welding (SMAW) and ¯ux cored arc welding (FCAW), on fatigue life of cruciform joints, containing lack of penetration (LOP) defects, have been studied. Load carrying cruciform joints were fabricated from high strength, quenched and tempered steels of pressure vessel (ASTM 517 ÔFÕ) grade. Fatigue crack growth experiments were carried out in a mechanical resonance vertical pulsator (SCHENCK 200 kN capacity) with a frequency of 30Hz under constant amplitude loading (R 0). It was found that the fatigue lives of the cruciform joints fabricated by the SMAW process were relatively higher than the FCAW counterpart. Ó 2000 Elsevier Science Ltd. All rights reserved. Keywords: Shielded metal arc welding; Flux cored arc welding; Cruciform joint; Lack of penetration; Fatigue life
List of symbols 2a initial LOP length 2W ®llet width 2B specimen width L leg length plate thickness Tp h ®llet angle Dr stress range R stress ratio (rmin =rmax ) crack initiation life Ni Np crack propagation life total fatigue life Nf initial stress intensity factor (SIF) DKi range
* Corresponding author. Tel.: +91-44-2351-365/3801; fax: +91-44-2350-509. E-mailaddress:[email protected](V.Balasubramanian).
m C
Paris constant (crack growth exponent) intercept in Paris equation
1. Introduction The ®llet welded cruciform joints are most common ones in various structures including oshore and nuclear applications. Although the quality of welding has improved over the past decades, welding discontinuities are still unavoidable. The sizes of internal discontinuities present in the weld has a large eect on the measured fatigue life. The most important variable in determining the fatigue life of a ¯awed weldment would seem to be the nature of internal ¯aws contained within the weld and the manner in which these ¯aws interact within the stress ®eld in and around the weld during its fatigue life (Lawrence and Radziminski,
0167-6636/00/$ - see front matter Ó 2000 Elsevier Science Ltd. All rights reserved. PII: S 0 1 6 7 - 6 6 3 6 ( 9 9 ) 0 0 0 4 9 - 6
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1970). Linking the eects of weld defects and failure analysis of weldments pointing towards that the fatigue alone is considered to account for most of the disruptive failures and often precedes the onset of brittle failure (Tsai, 1986). The fatigue resistance of the weld metal and heat aected zone of various steels are better than or equal to the base metal (Vosikovsky, 1980). However, problems arise when there is an abrupt change in section by excess weld reinforcement, undercut, inclusion of slag or lack of penetration (LOP) or fusion (Blodgett, 1992). Many of the fatigue failures that occur in welded joints involve fatigue cracking from severe imperfections which are actually an inherent part of the joint (Maddox, 1993). There are two types of cracking in a ®llet welded joints: (i) root cracking, and (ii) toe cracking. Fatigue cracks initiate at the ®llet weld toe when the ®llet weld size is large enough and initiate at weld root when weld size is inadequate (Gurney, 1979). The type and severity of weld defects are in¯uenced by the welding process itself. Automatic welding processes are favourable over manual processes for several reasons such as increased productivity, lower costs, etc. However, a twofold overall increment in fatigue lives was observed for manual welds with respect to comparable automatic welds. This was mainly attributed to a higher stress concentration eect as a result of poor weld pro®les in case of automatic welding, e.g., submerged arc welding (Otegui et al., 1991). Further, the mechanised submerged arc welding in average gives the steepest weld to base metal transition regions. But the manual and semi-automatic processes produce welds with a more smooth transition to the base metal (Overbeeke et al., 1991). Moreover, in automatic welding process, the mis-alignment is an important manufacturing defect. The bending stresses induced by the misalignment of welded parts signi®cantly in¯uence the fatigue lifes of weldments by decreasing mainly the fatigue crack initiation lives. This aects mostly the fatigue resistance of external discontinuities such as weld toes but only slightly reduces the fatigue resistance of internal discontinuities such as LOP. Therefore, a weld which should fail at an internal discontinuity may fail at
an external discontinuity because of bending stresses (Jakubcsak and Glinka, 1986). Fatigue life prediction of welded joints is very complex, costly and time consuming. This is due to its complex joint geometry, number of stress concentration points and heterogeneous weld metal property making the joint. So to avoid the costly and complex procedure, traditionally, the fatigue life of the joint for structural applications followed the S±N type of approach covered by BS 5400 and IIW (Hobbacher, 1988). But for critical structural applications where both initiation and propagation behaviour are equally important for the purpose of safety, then the fracture mechanics approach is more appropriate in the place of traditional methods to predict the fatigue life of the component (Allen et al., 1988). Hence, it is customary to predict the fatigue life of welded joint with defect in terms of crack growth parameters such as da/dN against DK obtained by crack growth experiment. Practically, these types of data merely indicate the fatigue crack growth behaviour of the weldments and do not predict the actual fatigue life. From the literature (Smith and Smith, 1983; Branco et al., 1985; Testin et al., 1987), it is evident that the most of the investigations on fatigue life prediction of the ®llet welded joints are based on toe failure. Very few investigators (Usami and Kusumoto, 1978; Frank and Fisher, 1979; Guha, 1994) have studied the fatigue behaviour of ®llet welded joints failing from the root region. Hence, this investigation has been carried out to study the in¯uence of welding processes, namely shielded metal arc welding (SMAW) and ¯ux cored arc welding (FCAW) processes, on fatigue life of cruciform joints, failing from the root LOP region. 2. Experimental Pressure vessel grade steel (ASTM 517 ÔFÕ) of weldable quality in the form of rolled plates of 8 mm thickness have been used as the base material throughout the investigation. These steels are used for welded construction of all kinds such as pressure vessels, penstocks, bridges and structures as well as transport vehicles, hoisting and
V. Balasubramanian, B. Guha / Mechanics of Materials 32 (2000) 265±276
267
various root faces enabled the joints to have dierent LOP lengths after welding. The ®llet leg lengths were varied from 6 to 8 mm and LOP length between 7 and 8 mm. The dimensions of cruciform joints are shown in Fig. 1. All necessary care was taken to avoid joint distortions and the joints were made without applying any clamping devices. The fatigue crack growth experiments were conducted in a mechanical resonance controlled vertical pulsator (200 kN capacity) with a frequency of 30 Hz under constant amplitude loading (R 0). Before loading, the specimen surface near the LOP was polished to enable the crack growth measurement. A travelling microscope was used to monitor the crack length with an accuracy of 0.01 mm. The specimen was loaded at a particular stress level (range) and crack initiation and its subsequent propagation from LOP defect was recorded from time to time until complete failure of the specimen. A similar crack growth experiment was conducted on a number of specimens at various stress levels and experimental data (crack initiation life, crack propagation life and total fatigue life) were recorded. Table 1 shows the chemical composition of base metal and weld metal. The mechanical properties of the base metal and weld metal are given in Table 2. The welding process parameters used to fabricate the joints are presented in Table 3.
earthmoving equipments which are utilised in dierent climatic conditions (Balasubramanian and Guha, 1998). The rolled plates were cut into the required sizes and pro®les by oxy-fuel cutting and grinding. The initial joint con®guration is obtained by securing the long plates (200 100 mm2 ) and stem plate (200 75 mm2 ) in a cruciform position by tack welding. Subsequently, the ®llets were made between the long plate and stem plate laying weld metal using two dierent welding processes with matching weld metal consumable. SMAW and FCAW processes were chosen to fabricate the joints and multi-pass welding procedures were employed. All the four ®llets forming the joint were made identical leaving an unfused gap between the pair of ®llets. This gap, i.e., LOP was controlled by providing proper root faces, obtained by a prior machining process, known as bevelling (Guha, 1994). The
3. Results and discussions Although the fatigue crack growth experiments were conducted on a large number of specimens having dierent dimensions (a=W 0:3 to 0.4, L=Tp 0:6 to 1.0), only one set of results has been considered here for developing the equations as well as for comparing the two processes in a better
Fig. 1. Dimensions of cruciform specimen. Table 1 Chemical composition (wt%) Material
C
Si
Mn
P
S
Cr
Mo
Ni
Cu
Co
V
Base metal Weld metal (i) SMAW (ii) FCAW
0.19
0.72
0.95
0.01
0.002
0.8
0.35
0.07
0.03
0.004
0.002
0.08 0.08
0.5 0.4
1.6 1.5
0.01 0.02
0.03 0.01
0.35 0.56
0.27 0.44
1.73 2.25
0.1 0.2
0.01 0.02
0.001 0.002
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V. Balasubramanian, B. Guha / Mechanics of Materials 32 (2000) 265±276
Table 2 Mechanical properties Material
Yield strength (MPa)
Tensile strength (MPa)
VickerÕs hardness (v30 kg)
Impact value (J)
Percentage of elongation (%)
Base metal Weld metal (i) SMAW (ii) FCAW
690
790
210
110
19
740 720
845 830
320 280
175 150
24.5 26
Table 3 Welding process parameters Process
Electrode speci®cation (AWS)
Electrode diameter (mm)
Voltage (V)
Current (A)
Welding speed (mm/s)
Heat input (kJ/mm)
(i) SMAW (ii) FCAW
E11018-M E100T5 K4
3.15 1.6
24 30
120 240
2.5 4.0
1.1 1.8
way. The dimensions of the test specimens, taken into account in the present investigation, are as follows: 2a 7:5, 2W 21, L 6:5, Tp 8 mm and h 45 . Fatigue crack growth experiments were conducted at four dierent stress ranges (Dr), i.e., 120, 160, 200 and 240 MPa, and the variations in crack length (2a) with the corresponding number of cycles (N) are plotted as shown in Fig. 2, for SMAW and FCAW joints. From the ®gures, it is evident that the crack growth is comparatively slower in the joints fabricated by the SMAW process than in the FCAW process. 3.1. In¯uence of welding process on crack propagation life (Np ) The fracture mechanics analysis is based on Paris power law (Paris and Erdogan, 1963), da=dN C DKm ;
After re-arranging the equation
da =dN C=2 f m a Drm W m=2ÿ1 ;
where f a is the normalised SIF range which is DK=Dr W 1=2 , and a is the normalised crack length : a=W. The expression for SIF range (DK), at the apex of a root (LOP) defect of a load carrying cruciform joint developed by Frank and Fisher (1979) is DK
m
d 2a=2W C DK Drm ; dN Drm 2W
2
where Dr is the nominal stress range, 2a the lack of penetration defect size, and 2W is the ®llet width as shown in Fig. 1.
Dr 1=2 A1 A2 a pa sec pa =2 ; 1 2 L=Tp 4
where L/Tp is the weld size as de®ned in Fig. 1. A1 and A2 are functions of weld size (L/Tp ): A1 0:528 3:287 L=Tp ÿ 4:361 L=Tp 3
2
4
3:696 L=Tp ÿ 1:874 L=Tp 0:415 L=Tp
1
where da=dN is the crack growth rate, DK the stress intensity factor (SIF) range, and C and m are constants. Eq. (1) can be normalised and adopted for cruciform joints, as in the form given below
3
5
and A2 0:218 2:7717 L=Tp ÿ 10:171 L=Tp
2
13:122 L=Tp 3 ÿ 7:775 L=Tp 4 5
1:785 L=Tp : By normalising and re-arranging, the above equation is simpli®ed (Guha, 1995) as follows: DK Dr f aW 1=2 ;
5
V. Balasubramanian, B. Guha / Mechanics of Materials 32 (2000) 265±276
269
Fig. 3. Eect of L/Tp ratio.
cycles than the FCAW joints, at a particular SIF range value. The term of DK in Eq. (5) for a cruciform joint can be inserted in Eq. (2) to get da Cf m a Drm W m=2 : dN 2W
6
After re-arranging the terms for integration we have Z Z a f da =f m a C=2 Drm W m=2ÿ1 dN ; 7 ai
Fig. 2. Crack growth curves.
where A1 A2 a pa sec pa =21=2 : f a 1 2 L=Tp Fig. 3 shows the normalised SIF range f a vs crack length a , for dierent values of weld size, L/Tp . The calculated values of SIF range of both the processes, for the growing crack, have been plotted in Fig. 4. From the ®gures, it is observed that the SMAW joints endure a larger number of
where ai is the initial defect size after initiation cycles (Ni ) and af is the ®nal defect size at failure, or Ip C=2 Drm W m=2ÿ1 Np ;
8
where Ip is the value of integration for propagation of cycles leading to failure, and Np is the number of cycles to crack propagation. Re-arranging the equation Np 2Ip =C Drm W m=2ÿ1 :
9
The integration was performed on normalised SIF range, f a between initial, ai , to ®nal crack length af with an increment of 0.01. The integration was done for various crack growth exponent
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V. Balasubramanian, B. Guha / Mechanics of Materials 32 (2000) 265±276
Fig. 5. Relationship between Ip and a .
crack length increment against the associated number of cycles to propagation. For all calculations, the ASTM E-647 guidelines were followed. The relationship between the SIF range (DK) and the corresponding crack growth rate, d(2a)/dN on a log±log scale in terms of the best ®t line (BFL), is shown in Fig. 6, for both the processes. The data
Fig. 4. SIF range values for the growing crack.
values (m) and it was also found that the value of integration, Ip , is a function of crack length, a , for all values of m as shown in Fig. 5 in a double log plot. 3.2. Evaluation of m and C The crack growth rate, da=dN , for propagation stage, was calculated considering the slope at the steady-state growth regime at dierent intervals of
Fig. 6. Crack growth rate curves.
V. Balasubramanian, B. Guha / Mechanics of Materials 32 (2000) 265±276
points mostly correspond to the second stage of the sigmoidal relationship of Paris equation (1). The exponent m which is the slope of the line on the log±log plot has been found to be 3.5 and 4.2 for SMAW and FCAW joints, respectively, and the corresponding value of constant C which is the intercept of the line on log±log plot has been found to be 7:6 10ÿ10 and 2:3 10ÿ10 . At a ®xed value of DK, the FCAW joints are showing higher crack growth rate than the SMAW joints. 3.3. In¯uence of welding process on crack initiation life (Ni ) The crack initiation life Ni was evaluated experimentally using the crack ``initiation criteria''. The initiation criterion was based on the assumption that the number of cycles required to grow 1 mm length of crack in excess of its original LOP length at the earlier crack growth stage under particular stress range (Guha, 1995). Similar criteria were adopted by other investigators (Jack and Price, 1971; Testin et al., 1987). Fig. 7 shows the relationship between initial SIF range (DKi ) and crack initiation life (Ni ). It is evident from the ®gure that the initiation lives of both the joints are
271
more or less similar because the crack initiation region (a distance of 1 mm from the LOP tip) is composed of weld metal that is deposited by the SMAW process, commonly known as root run or root pass. The Paris type equation for early crack growth is d 2a=dN Ci DKi m ;
10
where Ci is a constant. For the 1 mm crack growth, the required number of cycles are calculated as 1 mm=Ni Ci DKi Ni 1=Ci DKi Ni C1 DKi
m
ÿm
m
or
or
;
11
where C1 1=Ci . In the above equation, by substituting the experimental values of Ni , the calculated values of Ki and the evaluated values of m, the value of C1 is obtained for both joints and they are presented in the form of initiation life equations: For SMAW process : Ni 1:5 109 DKi
ÿ3:5
; 12
For FCAW process : Ni 6:0 109 DKi
ÿ4:2
: 13
3.4. In¯uence of welding process on total fatigue life (Nf ) The crack growth process is normally comprised of two major stages: (i) crack initiation, and (ii) crack propagation and hence the total fatigue life is considered to be the sum of the cycles spent in these two periods (Lassen, 1990). Therefore, the total fatigue life (Nf ) is evaluated by the following equation: Nf Ni Np :
Fig. 7. Relationship between crack initiation and initial SIF range.
14
Fig. 8 depicts the relationship existing between initial SIF range (DKi ) and the fatigue life values obtained from the experiments. It is evident from the ®gure that though there is not much dierence
272
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sponding initial SIF range (DKi ) values. The fatigue life (Nf ) of the cruciform joint can be predicted by the following equation for the endurance line shown in Fig. 9: Nf 4:2 109 DKi
ÿ3:5
for SMAW joints
15
ÿ4:2
for FCAW joints:
16
and Nf 1:2 109 DKi
in crack initiation life (Ni ) of both the joints, the total fatigue life (Nf ) is showing marked variations which implies that the propagation life (Np ) of both the joints are not same. With the knowledge of crack growth parameters (m and C) and incorporating the crack initiation cycles (Ni ) from Eqs. (12) and (13), the total fatigue lives (Nf ) of both the joints were predicted using Eq. (14) and they are presented as shown in Fig. 9, with corre-
The accuracy of the above developed equations are tested by comparing the predicted fatigue life data with the experimental data and the scatter diagram is shown in Fig. 10. The S±N behaviours of the two weldments are plotted in Fig. 11 (at each stress level, ®ve specimens were tested and the average values are plotted) for the comparison purpose. The variations in the crack growth behaviour and fatigue life of the joints may be due to the dierence in the heat input (welding process parameters) involved during the fabrication. Evans (1982) reported that the following changes in chemical composition, mechanical properties and microstructures were observed as a result of increasing the heat input: (i) the weld metal manganese and silicon content is reduced, (ii) the yield strength and tensile strength decreased, (iii) hardness of as deposited weld metal decreased, (iv) the average width of the columnar grains
Fig. 9. Predicted fatigue life values.
Fig. 10. Scatter diagram.
Fig. 8. Relationship between initial SIF range and fatigue life.
V. Balasubramanian, B. Guha / Mechanics of Materials 32 (2000) 265±276
273
The microstructures of the weldments at two dierent locations, i.e., (i) reheated region of the earlier passes due to subsequent passes, and (ii) non-reheated region of the weld metal, are compared in Figs. 12 and 13. The weld metal microstructure of SMAW process contains a high proportion of ®ne grain acicular ferrite which displays optimum strength and toughness properties. This is attributed to small grain size in which each lath is separated by high angle grain boundaries (Farrar et al., 1974). But in the weld metal microstructure of FCAW process, the formation of large proportion of pro-eutectoid ferrite (grain boundary ferrite), ferrite sideplates or
Fig. 11. S±N curves.
increased, (v) the amount of pro-eutectoid ferrite in weld metal increased at the expense of acicular ferrite, (vi) the lath size of the acicular ferrite increased, and (vii) the grain size of the equiaxed ®ne grained regions increased. A small amount of manganese and silicon is lost through higher oxidation due to high heat input involved in the semi-automatic process. Further, due to high heat input, the cooling rate is comparatively slower in the semi-automatic process and this leads to the softening of the microstructure, as exhibited by the decrease of hardness values (Dixon and Hakansson, 1995). Moreover, the variations in the crack growth behaviour and fatigue life are also due to the multi-passes involved during the fabrication of the joints. In multi-pass deposits, the weld metal consists partly of a tertiary transformation microstructure, i.e., in multi-pass technique, each successive bead tempers previous ones, consequently the secondary microstructures, such as pro-eutectoid (grain boundary) ferrite, side plate ferrite and acicular ferrite, will partly be heated into the region. This extra transformation results in the formation of tertiary microstructure. This structure normally has higher toughness than the non-transformed weld metal, the secondary microstructure (Dolby, 1979; Hoekstra et al., 1986).
Fig. 12. Microstructure of SMAW weld metals: (a) reheated region, and (b) non-reheated region.
274
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Fig. 13. Microstructure of FCAW weld metals: (a) reheated region, and (b) non-reheated region.
martensite has been found to be detrimental to toughness because these structures provide preferential early crack propagation paths, which oer a low resistance during the weld metal cleavage fracture (Levine and Hill, 1977; Tweed and Knott, 1983). Further, the eect of manganese is to promote acicular ferrite at the expense of pro-eutectoid ferrite. Nickel also in¯uences weld metal microstructure through changing the columnar grain size (Zhang and Farrar, 1997). FCAW process weld metals with higher nickel content
Fig. 14. SEM fractographs of SMAW joints: (a) crack initiation, (b) crack propagation, and (c) ®nal fracture.
V. Balasubramanian, B. Guha / Mechanics of Materials 32 (2000) 265±276
275
exhibited poor toughness due to the formation of martensite and the presence of coarsened columnar grains. The fractured surface appearance corresponding to crack initiation, crack propagation and ®nal failure regions as observed by scanning electron microscopy (SEM) are shown in Figs. 14 and 15. The crack initiation regions of both the weld metals are shown in Figs. 14(a) and 15(a), which do not have any distinctive features. Interestingly, the crack propagation region of both the weld metals, compared in Figs. 14(b) and 15(b), shows signi®cant variation in the fracture surface. The fractograph of FCAW joint shows a high density of quasi-cleavage fracture with a larger number of preferential crack initiation sites in the crack propagation region. However, the SMAW joints exhibit a feature of a relatively high energy fracture which shows higher ductility through the presence of more dimples and striations in the crack propagation region compared to FCAW joints. However, the ®nal fracture region, shown in Figs. 14(c) and 15(c), of both the weldments represents the brittle fracture and it is evident from the river pattern of fracture surface.
4. Conclusions
Fig. 15. SEM fractographs of FCAW joints: (a) crack initiation, (b) crack propagation, and (c) ®nal fracture.
1. The fatigue life of SMAW and FCAW cruciform joints of pressure vessel grade steels (ASTM 517 ÔFÕ), containing LOP defects, can be successfully predicted by using the fracture mechanics equations developed in this paper, within a reasonable accuracy. 2. The experimental evidence shows that the microstructural features of deposited weld metal play an important role in fatigue crack growth behaviour and in¯uence the fatigue life of the joints. The joints fabricated by the SMAW process, which contains a high proportion of acicular ferrite in the weld metal microstructre, displayed better fatigue performance than the FCAW joints.
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