applicability of pjp groove welding to beam...

APPLICABILITY OF PJP GROOVE WELDING TO BEAM-COLUMN CONNECIONS UNDER SEISMIC LOADS

Yoshiaki Kurobane and Koji Azuma, Sojo University, Kumamoto, Japan

Yuji Makino, Kumamoto University, Kumamoto, Japan

ABSTRACT

The applicability of partial joint penetration (PJP) groove welding to beam-column connections subjected to seismic loads is the main concern of this study. PJP groove welding can be significantly more economical than complete joint penetration (CJP) groove welding. However, the PJP groove welded joints inevitably contain sharp notches at the roots of welds, which may induce a non-ductile failure, especially when subjected to cyclic loading. Four full-sized beam-to-column connections were tested under cyclic loads. When the unfused regions created by PJP groove welds were reinforced by fillet welds so that the welded joints have sufficient cross-sectional areas, connections showed sufficient strength to withstand large inter-story drift angle demands. Strains sustained at points around internal discontinuities were found to be low because of greater cross-sectional areas of welded joints compared with the cross-sectional area of the beam flanges. Both the test results and the fracture mechanics-based assessment demonstrated that it is unlikely to initiate brittle fracture at these discontinuities. One connection, which failed prematurely due to a lack of penetration in the PJP welds, also showed that a brittle fracture was unlikely because tips of unfused gaps were at a low level of stress triaxiality.

INTRODUCTION The 1994 Northridge and 1955 Kobe Earthquakes revealed that beam-to-column connections in steel moment resisting frames were susceptible to brittle fracture. In Northridge brittle fractures most frequently occurred at the beam bottom flange groove welds with cracks initiating at roots of the welds. Unfused regions between steel backup bars and column flanges created sharp notches at the weld roots. In Kobe cracks frequently initiated at toes of the beam copes. The occurrences of these cracks were found to be reduced by improving profiles of the beam copes. Cracks also frequently emerged from notch roots formed by the steel weld tabs and beam flanges at the terminations of CJP groove welded joints. The improvements of the weld tab regions were found to be made by replacing the steel tabs by flux tabs or removing the steel tabs after welding and then grinding smooth the ends of the welded joints. However, the improvement of the weld tabs only is insufficient to avoid a brittle fracture initiating from the terminations of the CJP welds at the beam ends. Further improvements in welded joints are required by some means. The shop-welded connections proposed here utilize PJP groove welds with reinforcing fillet welds for the welded joints at the beam flange ends. The beam webs are fillet-welded to the column flanges. Backing bars and beam copes are unnecessary for

Connections in Steel Structures V - Amsterdam - June 3-4, 2004 367

these connections. After flux weld tabs are removed, superimposed fillet welds are returned continuously around the edges of the beam flanges. Therefore, by using PJP welding, improvements can be made on 3 representative locations where a brittle fracture was susceptible to occur. The cost of these improvements is minimal. When an appropriate size of reinforcing fillet welds is selected, the strength demand per unit area on the welded joints can be lower than the strength demand on unreinforced CJP welded joints. The past investigations into the applicability of PJP groove welded joints to beam-to-column connections supported these possible advantages over conventional CJP welded connections (1,2). PJP groove welded joints, however, inevitably contain discontinuities that may act as sharp notches. Thus, the design strength of PJP groove welds recommended in the Eurocode 3 (3) and AISC LRFD Specification (4) are generally equivalent to the design strength of fillet welds. Eurocode 3 allows that a tee-butt joints, consisting of a pair of PJP groove welds reinforced by superimposed fillet welds, is calculated as a CJP groove weld, if the total nominal throat thickness is not less than the thickness t of the part forming the stem of the tee joint, although the stringent limitation that the unwelded gap is not more than (t/5) or 3 mm is imposed. The AIJ LSD Recommendations (5), in contrast, specify the design strength of PJP welds equivalent to that of CJP welds. The design strength according to AIJ is based on extensive test results for plate-to-plate welded joints. However, details of the joints and loading conditions are different between the beam-to-column joints and plate-to-plate joints. Especially, the former joints are subjected to cyclic loading. Further experimental investigations are required to establish a reliable design methodology for PJP welded beam-to-column joints. This paper presents test results for two connections, previously reported, and for two additional connections. Based on these test results, the behavior and design of PJP welded beam-to-column connections will be discussed hereafter. SPECIMENS AND TESTING PROCEDURES Full-size beam-to-column connections with PJP groove welded joints were tested to failure under cyclic loads. Four specimens, two with wide-flange section columns, designated as H1 and H2, and two with RHS columns, designated as R3 and R4, were tested. All the specimens were of one-sided connections, reproducing beam-to-exterior column connection assemblies (See Fig.1).

WF Column (414x405x18x28)

Stiffener (PL-16)

WF Beam (500x200x10x16)

12

12

12

12

4

12

Beam flange

9

7 5

8 8

7

Beam flange

H1

H2

RHS Column (400×400×12)

Diaphragm (PL-25)

WF Beam (500×200×10×16)

12

10

10

10

4

10

Beam flange

8

8 8

8

Beam flange

R3

R4

8 8

Figure 1. Configuration of specimens.

368 Connections in Steel Structures V - Amsterdam - June 3-4, 2004

(a) Specimen H1 (b) Specimen H2

Figure 2. Macro-etch cross sections of PJP welds. The beams are wide-flange sections with the nominal dimensions of 500x200x10x16 mm and with the steel grade JIS (Japanese Industrial Standard) SN490B (specified minimum yield strength=325 N/mm2) for all the specimens. The beam top and bottom flanges were PJP groove-welded to the column flanges using the single-bevel grooves with the groove angle of 45 degrees. The depth of the grooves are 8 mm (one half of the nominal flange thickness) for Specimens H1 and R3 and 4 mm (one quarter of the nominal flange thickness) for Specimens H2 and R4. Reinforcing fillet welds were added all around the perimeters of the beam sections contacting the column faces. The nominal dimensions of the welds are shown in Fig.1. Incomplete penetration of PJP groove welds was found in Specimens H1 and H2 after testing. Sliced sections of the welded joints are shown in Fig. 2. Portions of the weld metal near the root of the weld are not fused with the base metal in Specimen H1. The lack of penetration from the root of the joint was of about 5 mm for Specimen H1 and less than 1 mm for Specimen H2. These specimens were welded in a flat position. Preliminary trials showed that better penetration of the weld metal was achieved by welding in a horizontal position than in a flat position. For Specimens R3 and R4 the beam ends were welded in a horizontal position after placing the columns horizontally. The penetration of the weld metal was satisfactory for these specimens. The GMAW (Gas Metal Arc Welding) electrodes used for fabricating Specimens H1 and H2 were of the grade JIS YGW11 with the specified minimum yield strength of 390 N/mm2, while those for fabricating Specimens R3 and R4 were of the grade JIS YGW18 with the specified minimum yield strength of 430 N/mm2. Table 1. Material properties.

Specimen Coupon extracted from

Thicness (mm)

Yield strength (N/mm2)

Ultimaate tensile strength (N/mm2)

Elongation (%)

H1, H2 Colum flange 28.6 374 518 27 H1, H2 Colum web 19.4 352 509 22 H1, H2 Beam flange 16.3 323 510 22 H1, H2 Beam web 10.4 361 514 19 R3, R4 Column 12.2 404 474 22 R3, R4 Beam flange 15.9 368 526 25 R3, R4 Beam web 9.7 402 541 20 R3, R4 Weld metal - 486 617 26

Beam top flange Column flange

Beam web

Beam top flange Column flange


Table 2. Charpy impact test results. Material VE0 (J) VTre (degree C) VEshelf (J)

Weld (H1, H2) 119 -8.6 194 Base(R1, R2) 221 -46 239 HAZ (R1, R2) 254 -102 246 Weld(R1, R2) 127 -47 147

VE0 = absorbed energy at 0 degree C VTre = energy transition temperature VEshelf = absorbed energy at upper shelf

u 1

L=2400

v 1 v 2

u 2t dH d

Figure 3. Test set-up.

Specimens H1 and H2 have continuity plates welded to the column flanges and webs at the positions of beam flanges. Specimens R3 and R4 have internal diaphragms CJP welded to the columns at the positions of beam flanges. Material properties determined by tensile coupon tests and measured plate thicknesses are summarized in Table 1. All-weld-metal tension testing was conducted for the electrode YGW18, the results of which are also included in Table 1. The notch toughnesses of materials used were measured by Charpy impact testing and summarized in Table 2. Charpy specimens were taken from plates welded under the conditions close to those used for fabricating connection specimens. Notch roots were located at the base metal, weld metal and HAZ. The both ends of the column were fixed to a strong floor, while the cyclic shear load was applied at the end of the beam by a double action hydraulic ram. The loading arrangements are shown in Fig. 3. Lateral bracing systems were provided at two positions of the beam. Displacement measurements taken were not only the horizontal displacement u1 at the loading point but also the vertical displacements v1 and v2 at the column face where the continuity plates or diaphragms exist and the horizontal displacement u2 at the column end. The rotation of the beam θf was calculated by the following equation.

θ f =U1 −U2

L−

V1 −V2Hd − td

(1)

where L and (Hd-td) denote the distance from the loading point to the column face and the


distance between the two continuity plates or diaphragms, respectively. The load was applied in the following sequences: firstly a few cycles of loads were applied in the elastic region and then the amplitude of the beam rotation was increased as 2θp, 4 θp, 6 θp,,…with 2 cycles of load application at each amplitude, up to failure. θp is herein defined as the elastic component of the beam rotation when the beam moment at the column face reaches the fully plastic moment Mp. The elastic rotations measured were slightly different from those calculated using a simple beam theory. The elastic beam rotations were obtained from experimental M versus θ relationships. The values of Mp and θp, calculated using the measured material properties, dimensions and rotations are shown below.

Specimens H1 and H2 Mp = 708 kNm θp = 0.00659 radians Specimens R3 Mp = 793 kNm θp = 0.00847 radians Specimens R4 Mp = 793 kNm θp = 0.00882 radians

HYSTERETIC BEHAVIOR AND FAILURE MODES The moment versus rotation hysteresis loops measured during the test are shown in Fig. 4. The moment is the moment at the column face Mf and is non-dimensionalized by the full-plastic moment Mp of the beam. The moment takes a positive value when the bottom flange is in tension. The rotation θf, which was determined by Eq. 1, is also non-dimensionalized by θp of the beam. Several failure events observed during the test are described below. The numbers shown in the hysteresis loops indicate the occasions when these events occurred. Specimen H1 The beam top flange ruptured along the weld toes (1). The load was reversed. A ductile crack was found at the toes of the welds at the edge of the beam bottom flange (2). Local buckles were found in the beam top flange (3). The cracks in the beam bottom flange grew rapidly (4). Complete rupture of the beam bottom flange (5). Specimen H2 The beam top flange buckled locally (1). Buckles of the beam top flange and web became large (2). The beam bottom flange buckled extensively (3). Small cracks were found along the weld toes in the bottom flange. Specimen R3 The beam top flange buckled locally (1). A crack was initiated along the weld toes in the top flange (2). The bottom flange buckled locally. The crack in the top flange extended over 1 mm. (3). Local buckles of the top flange grew (4). Local buckles of the bottom flange grew and cracks in the top flange extended significantly (5). The top flange buckled accompanying lateral buckling (6). The bottom flange buckled accompanying lateral buckling (7). Specimen R4 The beam bottom flange buckled locally (1). Two cracks were found at the weld toes in the beam bottom flange (2). A crack initiated at the weld toe in the bottom flange (3). Local buckles of the top flange and web progressed. The cracks in the top flange extended (4). The top flange buckled locally and laterally. The cracks in the bottom flange grew extensively (5). Extensive lateral buckling of the beam (6). Specimen H1 sustained a premature rupture of the welded joints owing to incomplete fusion of the weld metal (See Fig. 5). For the remaining 3 specimens the ultimate limit state was governed by combined local and lateral buckling of the beams. In these 3


specimens ductile cracks invariably initiated at the weld toes at the edges of the beam flanges and grew stably along the toes of the welds on the beam flange side (See Fig. 6).

-1.5

-1

-0.5

0

0.5

1

1.5

2

-4 -2 0 2 4 6 8 10θf/θp

Mf/M

p

4

5

1

2 3

-2

-1.5

-1

-0.5

0

0.5

1

1.5

2

-12 -8 -4 0 4 8 12θf/θp

Mf/M

p

3

1 24

(a) Specimen H1 (b) Specimen H2

-1.5

-1

-0.5

0

0.5

1

1.5

-10 -8 -6 -4 -2 0 2 4 6 8 10θf/θp

Mf/M

p

1

23

4

5

6

7-1.5

-1

-0.5

0

0.5

1

1.5

-10 -8 -6 -4 -2 0 2 4 6 8 10θf/θp

Mf/M

p

12

3

4

5

6

(c) Specimen R3 (d) Specimen R4

Figure 4. Moment versus rotation hysteretic curves.

Figure 5. Specimen H1 after failure.


Figure 6. Ductile tear observed in R4.

ROTATION CAPACITY Although several criteria have been proposed for the assessment of plastic deformation capacities of beam-column assemblies, cumulative plastic deformation factors for the connections discussed herein are evaluated and shown in Table 3. The values of cumulative plastic deformation factors evaluated from the hysteresis loops are comparable to those for connections with improved details (6) and meet ductility requirements for the seismic design. Specimen H1 failed prematurely and is the exception to the above statement. The definition of the cumulative plastic deformation factor can be found in other literature (for example, reference (7)). Table 3. Summary of test results and predictions.

Test results Prediction

Specimen Cumulative plastic deformation factor

Maximum moment at column face (kNm)

Maximum moment at column face (kNm)

Test/Prediction

Top flange -784 -784 1.00 H1

Bottom flange 20.5

1016 781 1.30 Top flange -975 -988 0.99

H2 Bottom flange

64.1 1009 988 1.02

Top flange -1016 -959 1.06 R3

Bottom flange 69.0

1000 959 1.04 Top flange -1014 -959 1.06

R4 Bottom flange

61.9 1046 959 1.09


EVALUATION OF ULTIMATE STRENGTH OF CONNECTIONS Two failure modes were identified in this series of tests: 1. combined local and lateral buckling of beams; and 2. tensile failure of welded joints including ductile crack growth. The ultimate strengths are evaluated focusing only on the latter failure mode. The ultimate strengths of the welded joints at the beam flange ends are calculated mainly using the recommendations by AIJ. The resistance factor is excluded. The AIJ recommendations show no formula for calculating the strength of a welded joint with combined PJP and superimposed fillet welds. Although Eurocode 3 shows a design rule that this type of joint can be calculated as a fillet weld with deep penetration, this rule gives rather conservative predictions as compared with test results.

a1

a2

a3

Beam flange

Beam web

Column flange

Figure 7. Effective throats of PJP and fillet welds. The ultimate strength of the welded joint is calculated as the sum of the ultimate strengths of one PJP weld and two fillet welds, each being calculated as an independent weld (See Fig. 7). Namely, the ultimate tensile strength of the welded joint Pu is calculated by

Pu = a1l1Fw ,u +1.4

3a2l2Fw ,u +

1.43

a3l3Fw ,u (2)

where the symbols a and l denote the effective throat thickness and length of the weld, respectively. The subscripts 1, 2 and 3 distinguish among the PJP weld, superimposed fillet weld and fillet weld on the opposite side of the groove. The factor 1.4 in Eq. 2 is to increase the design strength of the fillet weld when the axis of the weld is perpendicular to the direction of loading. Fw,u represents the ultimate tensile strength of the weld metal. The ultimate strength of the welded joint is governed also by the tensile strength of the beam flange given by Pu = BfltflFfl,u (3) where Bfl, tfl and Ffl,u denote the width, thickness and tensile strength of the beam flange, respectively. Pu takes the smaller value of those given by Eqs. 2 and 3.


The ultimate flexural strength of the connection at the column face Mf is calculated by Mf = Pu(Hb − t f )+mWweb,net Fweb ,y (4) in which Wweb,net signifies the plastic section modulus of the net area of the beam web considering reduction of the cross section due to the cope holes (no cope hole exists in the present examples) and Fweb,y represents yield strength of the beam web, while the symbol m represents the dimensionless moment capacity of the welded web joints. The second term of the right-hand side of Eq. 4 represents the ultimate flexural load carried by the welded web joint considering flexibility of the column flange. Further details of Eq. 4 are referred to elsewhere (7,8). The ultimate flexural capacities calculated by Eq. 4 are summarized in Table 3. When evaluating ultimate capacities of connections, the measured dimensions, including the sizes of the welds, and material properties were used. All-weld-metal tension testing of the electrode YGW11 was not conducted. However, extensive investigations on material properties of weld metal have already been conducted (9). Especially, many data exist for welded joints with the base metal of SN490B and the filler metal of YGW11. From these data, the data for welded joints having the same heat-input as that for the specimens H1 and H2 were selected. The material properties for the electrode YGW11 that were inferred from the existing data and used for the calculation are: Fw,y=450 N/mm2 and Fw,u=550 N/mm2. The ultimate capacity was governed by the strength of the weld metal only in Specimen H1. The ultimate capacities of the Specimens H2, R3 and R4 were determined by the strength of the base metal (Eq. 3). In Table 3, the rows “top flange” and “bottom flange” distinguish the cases when the top is in tension and when the bottom flange is in tension. As for Specimen H1 the top flange ruptured at the same load as the load predicted. When the load was reversed, the connection supported a much higher load than the predicted load. The reason for this increased resistance is not known. The neutral axis may have moved after failure of the top flange. Specimen H2 showed about the same ultimate loads as the predicted capacities. This specimen sustained extensive buckling but cracks found were small. On the other hand, Specimens R3 and R4 sustained not only extensive buckling but also significant crack growth. The flanges of these two specimens seemed to have nearly reached their tensile capacities. The experimental ultimate loads are a little higher than the predicted ultimate loads. This underprediction is understandable because Eq. 4 is ignoring the effects of biaxial stress state at the beam flange ends and cyclic hardening of materials on the flexural strength of connections. ASSESSMENT OF EFFECTS OF DISCONTINUITIES ON BRITTLE FRACTURE The possibility of a brittle fracture initiating from various discontinuities contained in Specimen H1 and H2 was assessed using the CTOD design curve approach recommended by JWES (10). The results of this assessment have already been reported by Azuma and co-authors. (2). The assessment showed that the required fracture toughness values for preventing a brittle fracture from the roots of the PJP welds of these specimens are much lower than the fracture toughness of the material used. Rather, the possibility of the brittle fracture is higher at the weld toes at the edges of the beam flanges than at the weld roots, although the required notch toughness is still significantly lower than the toughness of the material even at the beam flange edges. In conclusion, the previous investigation indicated that a brittle fracture is unlikely to occur in these specimens. Although the similar assessment on Specimens R3 and R4 has not


been completed yet, the same conclusions as those on Specimens H1 and H2 could be drawn judging from the present test results and the material properties shown in Table 2. However, the previous investigation used the nominal dimensions of the welds for modeling the specimens. A significant lack of penetration was observed in the PJP welds of Specimen H1 as was mentioned earlier. A reanalysis was conducted using measured dimensions of the welds of Specimen H1. The software used is the ABAQUS (2003) general purpose FE package.

Figure 8. Contour plot of von Mises’s equivalent stress in cross section on

the plane of symmetry.

Figure 9. Contour plot of equivalent plastic strain in cross section on plane of symmetry.


As seen in the von Mises’s equivalent stress distributions on the plane of symmetry shown in Fig. 8, stresses concentrate at the two notch roots locating at the tip of the unfused region. A large area of the base metal near the interface with the weld metal is fully yielded at an early stage of loading. As the load is increased further, a slip band in which the von Mises’s equivalent strains reach the order of 20 % is observed in the base metal (See Fig.9). Note that the weld metal is overmatching by about 30 % in yield strength. This band extends diagonally from the unfused region to the surface of the beam flange. Figure 10 shows the J integral versus stress triaxiality Ts curve calculated at one of the notch roots. The stress triaxiality Ts is defined as

Ts =σ h

σ eq (5)

in which σh denotes the hydrostatic stress and σeq denotes the Mises’s equivalent stress. The triaxiality represents the level of plastic constraint at crack tips. As seen in this figure, the stress triaxialiy first reaches a maximum value of about 2.0 under small scale yielding conditions and then decreases as yielding progresses, showing that the constraint is at a very low level under fully yielded conditions. The J integral reached about 250 N/mm at the stage when the connection ruptured during the test.

0

0.5

1

1.5

2

2.5

0 50 100 150 200 250

J (N/mm)

T s

Figure 10. Stress triaxiality Ts versus J integral curve.

The numerical analysis demonstrated that the roots of the PJP welds are nearly under a plane stress state. The loss of constraint leads to enhanced resistance to the cleavage fracture. Although the notch toughness of the base metal of Specimen H1 is not measured, the J integral of 250 N/mm is not large enough to induce cleavage fracture (See reference (11)). Therefore, it is unlikely to have a brittle fracture starting from the roots of the PJP welds. If a ductile tensile failure occurs after extensive yielding of the joint, the ultimate strength of the joint can be predicted based on a simple plastic analysis, like the analysis presented in the previous section. CONCLUSIONS Testing of 4 full-scale beam-to-column connections was conducted under cyclic loads. The connections have PJP welded joints with reinforcing fillet welds between the beam flanges and column flanges and two sided fillet welds between the beam webs and


column flanges. Three connections of these four showed sufficient strength and plastic deformation capacity. Plastic hinge formed at the beam ends accompanying local and lateral buckling of the beams. Ductile cracks initiated at the weld toes at the edges of the beam flanges and grew stably. The previous investigation indicates that it is unlikely that brittle fractures start from the roots of the PJP welds or cracks at the edges of the beam flanges. One connection sustained a ductile tensile failure prematurely. This premature failure was caused by the lack of penetration in the PJP welded joints. It was found important to select an appropriate joint detail, welding position and other welding conditions for achieving sufficient penetration in PJP welded joints. The evaluation of the ultimate strength of the connections based on a simple plastic analysis was found to be accurate and applicable to the connection design. A non-linear FE analysis showed that tips of unfused gaps existing in PJP welded tee-butt joints are not subjected to high plastic constraint. The stresses were not high enough to induce a brittle fracture starting from the tips. ACKNOWLEDGEMENTS The proposals made here are based on joint experimental programs between Sojo and Kumamoto Universities. The authors wish to thank H. Shinde, M.Sc. student and the other 4th year students for their hard work in laboratories. This work was partly supported by the Japanese Society for the Promotion of Science Grant-in-Aid for Scientific Research under the number 13650645. REFERENCES (1) Azuma, K., Kurobane, Y. and Makino, Y., (2000). Cyclic testing of beam-to-column

connections with weld defects and assessment of safety of numerically modeled connections from brittle fracture. Engineering Structures, Vol. 22, No. 12, pp.1596-1608.

(2) Azuma, K., Kurobane, Y., Dale, K. and Makino, Y., (2003). Full-scale testing of beam-to-column connections with partial joint penetration groove welded joints. Tubular Structures X, M.A. Jurrieta, A. Alonso and J.A. Chica eds., Balkema, Lisse, The Netherlands, pp. 419-427.

(3) CEN, (1992). Eurocode 3: Design of steel structures, ENV 1993-1-1: Part 1.1 General rules and rules for buildings. Comité Européen de Normalisation, Brussels, Belgium.

(4) AISC, (2000). Load and resistance factor design specification for structural steel buildings. American Institute of Steel Construction, Chicago, Ill., USA.

(5) AIJ, (1998). Recommendation for limit state design of steel structures. Architectural Institute of Japan, Tokyo, Japan. (in Japanese)

(6) Kurobane, Y., Azuma, K. and Makino, Y., (2003). Fully restrained beam-to-RHS column connections with improved details. Tubular Structures X, M.A. Jurrieta, A. Alonso and J.A. Chica eds., Balkema, Lisse, The Netherlands, pp. 439-446.

(7) Kurobane, Y., Packer, J.A., Wardenier, J. and Yeomans, N.F., (2004). Design guide for structural hollow section column connections. CIDECT/Verlag TÜV Rheinland,


Köln, Germany. in press (8) AIJ, (2001). Recommendations for the design of structural steel connections.

Architectural Institute of Japan, Tokyo, Japan. (in Japanese) (9) Mukai, A., Nakano, T., Okamoto, H. and Morita, K.,( 2000). Investigation on MAG

welding wires for building. Journal of Constructional Steel, Japanese Society of Steel Construction, Vol. 7, No. 26, pp. 13-25. (in Japanese)

(10) JWES, (2000). Method for assessment of brittle fracture in steel weldments subjected to cyclic and dynamic large straining, WES-TR2808, The Japan Welding Engineering Society, Tokyo, Japan. (in Japanese)

(11) Iwashita, T., Kurobane, Y., Azuma, K. and Makino, Y., (2003). Prediction of brittle fracture initiating at ends of CJP welded joints with defects: study into applicability of failure assessment diagram approach. Engineering Structures, Vol. 25, Issue 14, pp. 1815-1826.


applicability of pjp groove welding to beam...

Documents