Welded aluminium alloy connections: test results and BS8118

Thin-Walled Structures 36 (2000) 265–287 www.elsevier.com/locate/tws

Welded aluminium alloy connections: test results and BS8118 T.K. Chan a

b

a,*

, R.F.D. Porter Goff

b

School of Civil and Structural Engineering, Nanyang Technological University, Nanyang Avenue, Singapore 639798, Singapore Department of Engineering, Cambridge University, Trumpington Street, Cambridge CB2 1PZ, UK Received 3 August 1999; accepted 4 January 2000

Abstract Heat-treatable aluminium alloys are known to suffer from severe loss of strength in the vicinity of welds. This investigation examines the structural effect of a reduced strength zone adjacent to the welds of cruciform connections and assesses its effects on the load carrying capacity, ductility and mode of failure. The results indicate that the load carrying capacity for short connection lengths are lower than values predicted with current methods of design and that the mode of failure is usually with a distinctive dish-like necking around the welded finger within the RSZ of the joint.  2000 Elsevier Science Ltd. All rights reserved. Keywords: Heat-affected zones; Reduced-strength zone; Welded aluminium joints

1. Introduction Aluminium alloys made of the 5000 (Al-Mg), 6000 (Al-Mg-Si) and 7000 (Al-Zn) series are now commonly joined by welding. However, the effect of welding on the mechanical properties can be drastic depending on the composition of the alloy and its subsequent heat-treatment. Therefore, the use of butt or fillet welds in a direction perpendicular to the load path effectively reduces the load carrying capacity of the member to the minimum strength in the heat-affected zone (HAZ). A more efficient connection is to introduce the load from a parallel narrow bar into the member with the forces being carried through shear, in both the welds and the heat-affected zone.

* Corresponding author. Tel.: +65-790-5283; fax: +65-791-0676. E-mail address: [email protected] (T.K. Chan). 0263-8231/00/$ - see front matter  2000 Elsevier Science Ltd. All rights reserved. PII: S 0 2 6 3 - 8 2 3 1 ( 0 0 ) 0 0 0 0 6 - 9

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Nomenclature d ⌺d Fpred Fult kz xA xB z0 zRSNA zOA a d h sult s0.002 ⑀ult

width of heat path summation of widths of all heat paths available to the weld predicted failure load ultimate failure load softening factor distance of point A from weld centreline distance of point B from weld centreline basic value of heat affected zone width of RSNA zone measured from weld centreline width of OA zone measured from weld centreline effect of interpass temperature effect of 3 or more heat flow paths effect of heat build-up due to free edges or nearby welds ultimate stress proof stress, stress corresponding to 0.2% permanent strain ultimate strain

The length of the bar can be designed to carry loads close to the full capacity of the member. This type of connection is observed in light-weight demountable aluminium bridges. The approach to this investigation is to examine the structural effect of a reducedstrength zone (RSZ) adjacent to the welds of a load diffusing joint where the load is introduced through a narrow bar into a flat plate. The effect of this weakened zone was investigated experimentally with welded cruciform specimens characteristic of the narrow bar to plate connection and with finite element analyses. The emphasis is on assessing the effects of an RSZ on the load carrying capacity, mode of load transfer, ductile behaviour and the mode of failure, with a view to develop guidelines for the design of these bar-to-plate connection in weldable aluminium alloys.

2. Heat-affected and reduced strength zones The heat from the weld usually produces a temperature that exceeds the solubility limit of the alloy (about 350°C for the Al-4.0%Zn-2.0%Mg alloy); the region immediately around the weld thus reverts to its as-quenched condition and suffers some degradation of strength. However, the material in this region is able to age naturally at room temperature but it does not regain the strength achieved by artificial precipitation hardening. Further into the HAZ, the high temperatures tend to severely overage the alloy, thus reducing the strength drastically with no recovery. On the

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outer edge of the HAZ, the temperature reached only just exceeds that of the previous precipitation heat treatment and changes in property or structure are marginal. The extent of the HAZ was initially defined by Kelsey [1] based on a hardness scan across the mid-thickness of a transverse section of butt joint or a bead-on-plate as shown in Fig. 1. Two points, A and B, at distances of xA and xB respectively from the centreline of the weld is defined: A being the point of maximum strength loss and B, at the edge of the HAZ. Tensile test on strips taken across a transverse section of the welded sheet or plate indicate that the position of fracture in the tensile test, i.e. the weakest point, coincided with the position of point A as indicated by the hardness scans. There is little published data on the extent and severity of the heat-affected zone (HAZ) for heat-treatable alloys [2–5] and even fewer for the 7000-series alloy. Robertson [6] has reported results for the extent of HAZ in Al-4.0%Zn-2.0%Mg alloy, for a limited range of weld sizes. A study of the extent and severity of the HAZ of Al-4.0%Zn-2.0%Mg alloy (a heat-treatable alloy in the fully-heat treated stage) [7– 9] indicate that the extents of the HAZ are closely correlated to the ratio of area of weld deposit over the total connected thickness. In evaluating the effect of the HAZ on the strength of a member, a “reducedstrength zone” (RSZ) (smaller than the HAZ) can be assumed, such that the strength

Fig. 1. Hardness profile across a weld bead on plate specimen indicating positions of points A and B (left), and two-zone reduced strength zone model; OA zone and RSNA zone (right).

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of the welded members can be obtained by assuming suitable mechanical properties within this region. This leads to the definition of a two-zone reduced strength model: zone A defined as the re-solution treated and naturally aged (RSNA) zone extending to a distance of zRSNA=xA from the centre-line of the weld, and zone B, defined as the over-aged (OA) zone of width zOA=0.5(xA+xB), as shown in Fig. 1. A summary of the extents and severity of the HAZ is included in Appendix A.

3. Experimental study Two test series were carried out in this investigation. The first series of specimens was designed to investigate the behaviour of a single finger cruciform with various welding configurations. The second series was to evaluate different finger profiles and plate widths. The cruciform joints were made of the Al-4.0%Zn-2.0%Mg heattreatable aluminium alloy in the fully heat-treated condition. The specimens were automatically welded using the MIG process with continuous recording of current and voltage across the arc. 3.1. Series 1 specimens In the first series, a total of seven specimens were welded up and tested in tension (see Fig. 2). The cruciforms were fabricated from plates of 300×100×4.75 mm slotted into finger plates of 300×100×13 mm effectively providing fingers of 100 mm length. The weld sizes and welding configurations for the specimens are tabulated in Table 1. The plate was slotted into the centre of the finger and held together with clamps during welding. After welding, all the specimens were kept at 20°C for at least 14 days before testing. Testing was carried out on a 1000 kN Avery universal testing machine. As the main concern was the ultimate load carrying capacity of the specimens, no displacement or strain measurements were taken. 3.2. Series 2 specimens In the second series, a total of 11 specimens were made as tabulated in Table 2 and shown in Fig. 3. The cruciforms were fabricated from plates of 400×200×4.75 mm slotted into finger plates of 300×100×13 mm providing fingers of 150 mm length. As in the previous series, the plate was slotted into the finger and clamped together during welding. The welds were laid in a continuous run; welding down one side of the finger, around the tip, and back along the opposite side of the finger for a total weld length of 313 mm (150+13+150 mm). The narrower set of cruciforms were initially made using 200 mm wide plates and with identical welding procedures. The plate was trimmed to 100 mm after welding. The 200 mm plate was used instead of a 100 mm plate to reduce the extent of the HAZ. This means that the HAZ in all series 2 specimens are identical. After welding, the fingers on some specimens were cut to the required finger profile (see Table 2). The finger was cut from the weld toe back up to a point 150 mm from the weld toe; four with a straight cut

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Fig. 2.

269

Series 1 cruciform specimens.

producing a triangular profile and another four with a parabolic profile. The profiles are shown in Fig. 4 and Fig. 5. The load-deflection behaviour was measured for all specimens and every other specimen was strain gauged. Two specimens of the same width were made for each finger profile and strain gauges (linear and rosette gauges) were affixed to one of the two for a record of the strains around the weld toe and in the HAZ along the finger. The displacement of the cruciform joint under load was measured by placing

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Table 1 Dimension, welds and failure load for Series 1 specimens

Specimen

No. of welds

Length of each weld (mm)

Weld End Interpass area weld temp. (mm2)

Fult (kN) Failure mode

savg (MPa)

1 2 3 4 5 6 7

4 4 4 2 4 2 2

80 80 80 80+13+80 80 80+13+80 80+13+80

No No No Yes No Yes Yes

171.0 116.0 156.0 142.0 121.0 205.0 184.5

360 244 328 299 255 216 195

20 20 ⬎20 20 20 20 20

33 33 33 33 33 33 33

Failure through HAZ Crack through crater Failure through HAZ Failure through HAZ Crack through crater Crack at weld toe Crack at weld toe

Table 2 Dimensions, welds and failure loads for Series 2 specimens Specimen Finger profile

Width (mm)

Interpass Fult (kN) temp.

16 17 12 13 14 15

Rectangular Rectangular Triangular Triangular Parabolic Parabolic

200 200 200 200 200 200

20 20 20 20 20 20

318.0 ⬎305.0 290.0 292.5 292.0 290.0

21

Rectangular

100

20

183.0

17a

Rectangular

100

20

182.5

22

Triangular

100

20

183.0

23

Triangular

100

20

186.0

24

Parabolic

100

20

183.5

25

Parabolic

100

20

187.0

Failure mode

savg (MPa)

Cracking at weld toe Connection failure Cracking through weld toe Cracking through weld toe Crack at edge of plate Crack at edge of plate Dish-like crack away from the weld toe Dish-like crack away from the weld toe Dish-like crack+fracture across plate Dish-like crack+fracture across plate Dish-like crack away from the weld toe Dish-like crack+fracture across plate

335 321 305 308 307 305 385 384 385 392 386 394

dial gauges at ends of the specimen. Two sets of gauges were used to eliminate bending effects. The specimens were kept at 20°C for at least 28 days before testing. The cruciforms were then loaded in tension with strain readings and elongations recorded at every 10 kN increment up to the yield load, after which the readings were taken at regular displacement intervals.

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Fig. 3.

Series 2 cruciform specimens.

271

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Fig. 4.

Series 2 cruciform specimens of 200 mm width.

Fig. 5.

Series 2 cruciform specimens of 100 mm width.

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4. Test results 4.1. Series 1 The aim of this series was to observe the effect of different welding configurations on the ultimate load and the corresponding failure mode. The ultimate loads of the cruciform specimens are tabulated in Table 1, with failure mode illustrated in Fig. 2. Specimens 1, 2, 3 and 5, all of which have four 80 mm lengths of 10 mm fillet welds along the sides of the finger, failed at loads ranging from 121 to 171 kN. Specimens 1 and 3, where the welds were laid away from the end of the finger achieved a higher ultimate load than specimens 2 and 5, where the welds were laid in the opposite direction. Specimens 2 and 5 failed with cracks initiated at the crater of the welds. This indicates that weld craters should be located away from positions of high local stress. Specimen 4, which was welded around the end of the finger, failed at an ultimate load of 142 kN. It failed suddenly across the width of the 4.75 mm plate with a curved fracture surface about 30 mm from the tip of the finger (see Fig. 6). The failure clearly occurred within the HAZ of this specimen. Comparing specimen 4 with specimens 1, 2, 3 and 5 which were not welded around the end of the finger indicates that four single run welds, welded away from the finger is most efficient, followed by two continuous U-shaped runs around the end of the finger. The least efficient welding configuration is with four single weld runs, welded towards the end of the finger.

Fig. 6.

Cruciform specimen 4.

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The wider 200mm specimens 6 and 7 failed at 205 kN and 185 kN, respectively. The specimens failed with a crack at the toe of the transverse weld, at the end of the finger. A comparison of the average stress in the plate, in Table 1, shows that the wider cruciform specimens are less efficient than the narrower specimens. 4.2. Series 2 The aim of this series of tests was to examine the effect of three different finger profiles on the ultimate load of cruciform connections. Two sets of cruciform specimens were tested: set 1 consists of 200 mm wide specimens and set 2 consists of 100 mm wide specimens. The test results of sets 1 and 2 are tabulated in Table 2 and the corresponding failure modes illustrated in Fig. 3. The results for the wider 200 mm specimens show that trimming the rectangular finger to either a triangular or parabolic profile has reduced the ultimate load carrying capacity by about 10%. There is little difference in the ultimate loads for the specimens with triangular and parabolic fingers. Specimen 17 did not fail in the welded joint but failed due to insufficient support at its lower end. The failure load of this specimen is expected to be in excess of 305 kN. As this specimen could not be retested with the full 200 mm width, it was trimmed to 100 mm and used as a specimen 17a in set 2. The results in Table 2 show that the profiling of the finger in the narrower specimens does not affect the ultimate load carrying capacity. The variation in the ultimate loads is less than 2.5%. Fig. 7 and Fig. 8 show the load-displacement curves for set 1 and set 2 cruciforms,

Fig. 7.

Load-displacement curve for Series 2 200 mm wide specimens.

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Fig. 8.

275

Load-displacement curve for Series 2 100 mm wide specimens.

respectively. It is clear that ductility of the joint is highest for specimens with the parabolic finger followed by specimens with the triangular finger for both sets of cruciforms. The rectangular finger is least ductile. The initial stiffness of the cruciforms is similar for all specimens, with yielding occurring at lower loads for the specimens with triangular and parabolic fingers. There were three failure modes observed from the specimens in set 1 (see Fig. 3). The specimen with a rectangular finger (specimen 16) had a dish-like crack away from the weld toe. Specimens with triangular fingers failed through cracking at the weld toe. Specimens with parabolic fingers failed with a crack at the ends of the weld, initiated at the edge of the plate. All the specimens in set 2 failed with a dishlike crack away from the weld toe. The position of the crack, being between 23 to 25 mm from the root of the fillet weld, is believed to be the position of point A on the HAZ model. The dish-like necking was observed at loads of about 90% of the ultimate load in set 2 specimens. Some specimens broke across the plate section as the tensile testing machine was load controlled and the load could not be stopped before total fracture occurs. The crack in specimen 21 of set 2 is shown in Fig. 9. The average stress in the wider cruciforms is 314 MPa, which is only 75% of the ultimate stress of the plate. In comparison, the average stress in the narrower cruciforms is 388 MPa, corresponding to 92% of the strength of the plate. This indicates that the specimens with the 100 mm wide plate are more efficient in carrying load.

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Fig. 9.

Failure of specimen 21 with dish-like crack away from toe of weld.

5. Discussion of test results Tests on single finger cruciforms indicate that the most efficient welding configuration is four single run fillet welds beginning at the end of the finger and terminating at the edge of the plate. Welding around the end of the finger in the cruciform reduces the ultimate load due to a larger HAZ (RSZ) because of extended exposure to weld heat. It was also found that weld defects e.g., crater crack, can substantially reduce the ultimate load of the cruciform joints. The results of tests on cruciforms with different finger profiles indicate that trimming the rectangular finger to either a triangular or parabolic finger to reduce stress concentrations at the end of the fingers, does not increase the load carrying capacity. In the 100 mm wide cruciforms, the failure load was not influenced by finger profile but in the 200 mm wide cruciforms, the failure load dropped by about 10% when the fingers were trimmed.

6. Comparison of experimental results with BS8118 The strength of the experimental cruciforms was compared with the strengths predicted by the procedure set out in Section 6 of BS8118: 1991 Part 1 [10] and from results of the present work. To calculate the plastic strength of the joint, it is neces-

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sary to define two sections in the HAZ of the cruciform. Fig. 10 shows the alternative failure sections 1-1, tension failure in the plate, and 2-2, shear failure in the HAZ along the finger. It is assumed that failure in the welds or fingers is prevented by adequate weld and parent material cross-section, respectively. The strength of the connection is calculated as the minimum of the plastic strength across either section 1-1 or 2-2. The ultimate strength predictions using the procedure described in BS8118 is not very different from the estimates obtained by using the results of the present investigation although there are discrepancies in the actual values for “extent” and “severity” as seen in the Appendix. The predicted and experimental strengths and the “reserve factors” (i.e. the ratio of the experimental strength over the predicted strength) are tabulated in Table 3. Although both methods produced safe values for the narrower cruciform (without any weld defects), they could not account for the low failure loads in the wider cruciforms. From Table 3, it can generally be seen that the 100 mm wide cruciforms failed at higher loads than the calculated values but the 200 mm cruciforms failed at loads lower than predicted. The strongest 200 mm cruciform managed to sustain a load of approximately 90% of the predicted failure load (of the present investigation) and this value dropped to only 53% when the length of connection was short. Some discrepancies in the experimental failure loads in series 1 may be explained

Fig. 10. Location of failure planes 1-1 and 2-2 in the cruciform connection.

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Table 3 Comparison of strengths predicted by BS8118 and results from the present study with experimental results Specimen no.

Series 1 1 2 3 4 5 6 7 Series 2 16 17 12 13 14 15 21 17a 22 23 24 25

Plate width (mm)

Fult (kN)

Fpred (kN)

Fult/Fpred

BS8118

Present

BS8118

Present

100 100 100 100 100 200 200

171 116 156 142 121 205 185

176 176 137 137 176 288 288

183 183 165 165 183 352 352

0.97 0.66 1.14 1.04 0.69 0.71 0.64

0.93 0.63 0.95 0.86 0.66 0.58 0.53

200 200 200 200 200 200 100 100 100 100 100 100

318 ⬎305 290 293 292 290 183 183 183 186 184 187

352 352 352 352 352 352 154 154 154 154 154 154

363 363 363 363 363 363 165 165 165 165 165 165

0.90 ⬎0.87 0.82 0.83 0.83 0.82 1.19 1.19 1.19 1.21 1.19 1.21

0.88 ⬎0.84 0.80 0.81 0.80 0.80 1.11 1.11 1.11 1.13 1.11 1.13

after detailed examination of the failure surfaces. It has previously been determined that the strength of a full penetration butt weld has a lower strength than the minimum strength in the HAZ. An examination of cruciform 3 indicates that the root of the fillet weld has penetrated beyond the mid-thickness of the 4.75 mm plate on each weld pass effectively making a butt welded connection at the front of the finger. Cruciforms 2 and 5 have previously been discounted as the failure was initiated by a crack in the crater of the weld. The slightly lower test results obtained from cruciform 1 compared to set 2 of series 2 is due to testing only after 14 days of natural ageing. A much lower test result from cruciform 4 compared to series 2 is due to a lack of heat flow paths in cruciform 4 as it is only 100 mm wide during welding whereas the cruciforms in series 2 were all 200 mm wide during welding and later cut to the required 100 mm width. This has clearly increased the ultimate strength of the joint from 142 kN for cruciform 4 to 184 kN for cruciforms 17a and 21-25. The lower failure load for all the 200 mm wide cruciforms may be attributed to a stress and therefore a strain concentration at the front of the welded finger. This consequently induces necking; clearly seen as a dish-like thinning in this region prior to failure. A comparison of specimen 6 and 7 with specimens 16 and 17 tends to indicate that this strain concentration effect is more obvious for short connection lengths.

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7. Finite element study An elastic finite element analysis was initially carried out to study the effects of finger geometry on the stress distribution in the connection. The results of this analysis indicate that a rectangular finger profile generates the largest stress concentration at the tip of the finger. A lower stress concentration is obtained by cutting to triangular profile while a parabolic profile generates only a very small stress increase. An elastic-plastic model of the RSZ was developed and validated against experimental behaviour of a weld bead-on-plate specimen [7]. The validated model was then extended to predict the behaviour of cruciform finger-to-plate connections. The finite element analyses were carried out using MARC software with FEMGEN pre-processing package to generate the meshes. The analysis of cruciform involves a detailed 3-dimensional elastic-plastic analysis of the finger-to-plate connection with a view to clarifying the mechanism of failure. A three-dimensional mesh has to be adopted to enable the three-dimensional nature of the fillet weld and the variation of stress across the height of the finger to be included in the analysis. The three-dimensional mesh generated for use with the finite element program is shown in Fig. 11. The elements chosen are 20 noded threedimensional blocks with quadratic displacement functions. This rapidly converging element was used throughout the model with two elements through the thickness of the finger and plate. However due to symmetry, only one quarter of the connection needs to be modelled. In the welds, the elements were reduced to wedge and tetrahedron shapes simply by repeating node numbers. 7.1. Prediction of ultimate load The prediction of the ultimate load for a bar in tension is simply the product of the ultimate stress and the cross-sectional area. However the prediction of ultimate load for a complex connection with local necking is much more difficult as the finite

Fig. 11. Three-dimensional mesh of the cruciform connection.

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element program is not able to determine the onset of local instability (i.e. necking). A possible simplifying assumption for the prediction of ultimate load is that of limited extension, with failure assumed to occur once the elongation is beyond a certain limiting value. It was found that the ultimate load which correspond closely to the experimental failure load has about 5 to 10 integration points which have an equivalent strain to strain limit ratio equal to or greater than 1.0 (each layer in the 20-noded brick has 9 integration points). It was observed that near this ultimate load, the number of integration points having a strain ratio equal to or greater than 1.0 increases rapidly indicating that the material in this element can not carry any additional load. The location of the failure was determined by observing the positions where the strains achieve the limiting values. 7.2. Series 1 The ultimate loads predicted by the above method are presented in Table 4. In the first series, a total of eight specimens were welded with various weld configurations and the joints tested in tension. All the predictions by the FE model were high for series 1. The failure loads predicted by the FE model were within 20% of Table 4 Comparison of finite element model and experimental ultimate loadsa Specimen

Series 1 1 (100) 2 (100) 3 (100) 4 (100) 5 (100) 6 (200) 7 (200) Series 2 16 (200-R) 17 (200-R) 12 (200-T) 13 (200-T) 14 (200-P) 15 (200-P) 21 (100-R) 17a (100-R) 22 (100-T) 23 (100-T) 24 (100-P) 25 (100-P) a †

Finite element results Fult (kN) Failure zone

Experimental results Fult (kN) Failure zone

Fult FEM/Fult Exp

185.2 185.2 185.8 170.0 189.1 268.4 263.0

OA OA OA OA OA OA OA

171.0 116.0 156.0 142.0 121.0 205.0 184.0

HAZ, end weld toe HAZ, end HAZ, end crater weld toe weld toe

1.08 1.60† 1.19 1.20 1.56† 1.31† 1.43†

326.6 326.6 324.8 324.8 317.0 317.0 174.5 174.5 174.7 174.7 176.0 176.0

OA zone, end OA zone, end OA zone, end OA zone, end RSNA zone, edge RSNA zone, edge OA zone, end OA zone, end OA zone, end OA zone, end OA zone, end OA zone, end

318.0 ⬎305.0 290.0 292.5 292.0 290.0 183.0 182.5 183.0 186.0 183.5 187.0

HAZ, end – Weld, edge Weld, edge Weld, edge Weld, edge HAZ, end HAZ, end HAZ, end HAZ, end HAZ, end HAZ, end

1.03 1.07 1.12 1.11 1.09 1.09 0.95 0.96 0.95 0.94 0.96 0.94

zone, zone, zone, zone, zone, zone, zone,

end end end end end end end

Failure due to weld defects or initiated at weld toe.

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the experimental values for specimens 1, 3, and 4, the prediction for specimen 1 being within 8%. Failure in the case of specimens 2 and 5 was caused by defects in the welds. Crater cracks and lack of fusion at the start of the welds precipitated failure at loads substantially lower than those predicted by the model. Specimens 6 and 7 also failed at loads lower than the predicted failure loads with a crack at the weld toe as opposed to the OA zone as suggested by the model. In this respect the FE model cannot predict failures induced by weld or other defects. The results emphasise the need for good welding practice and proper inspection in the fabrication process and where appropriate weld dressing should be carried out. 7.3. Series 2 The aim of this series of tests was to examine the effect of three different finger profiles on the ultimate load of cruciform connections. The ratio of the predicted load to the experimental failure load for all 200 mm wide specimens was between 1.03 and 1.12. There was a slight decrease in the predicted load carrying capacity when the fingers were cut to a triangular profile and dropped a further 3% when trimmed to a parabolic profile. The ratio of the predicted load to the experimental failure load for 100 mm wide specimens was between 0.94 and 0.96. There was very little difference in the predicted ultimate loads for all three finger profiles. The predicted failure zone for the 200 mm wide specimens with rectangular and triangular fingers was in the OA zone of the plate at the end of the finger and in the RSNA zone at the edge of the plate for the specimens with a parabolic finger. Although failure in the specimen with a rectangular finger (Specimen 16) was not with a dishlike crack in the OA zone, there was considerable thinning in the HAZ indicating that the HAZ region was nearing failure. The FE model also indicated that the location of failure of the specimen with a parabolic finger was at the edge of the plate as opposed to the end of the finger. The FE model predicted that the specimen with a triangular finger should have failed in the HAZ at the end of the finger but failure occurred at the edge of the plate for both specimens 12 and 13. A closer examination of the FE results indicate that the region at the edge of the plate is also nearing failure. The FE model predicted accurately that the location of failure was at the HAZ at the end of the finger for all 100 mm wide specimens. 7.4. Load deflection behaviour The load-deflection curves were measured on specimens of Series 2 only. The experimental and FE predictions of the load-displacement behaviour of the 200 mm and 100 mm wide specimens of Series 2 are plotted in Fig. 7 and Fig. 8, respectively. It can be seen that the initial stiffness of both the experimental and the predicted load-displacement curves are similar for both the 200 mm and 100 mm wide specimens. The point of failure of the predicted load-deflection curve is at the end of respective solid lines for the various specimens. The dashed lines extending beyond

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the point of failure is only to indicate the behaviour of the specimen if the strain limiting criterion is not applied. All the predicted load-deflection curves for the 200 mm wide specimens exhibit a higher yield point and failure load than the experimental curves. This indicates that the specimens were less stiff in the plastic load region as compared to the predictions. The FE model also successfully modeled the increasing ductility of the connection when the fingers were cut to a triangular and parabolic profile. The FE model, however, predicted failure loads lower than the experimental failure loads for the 100 mm wide specimens and also a lower ductility than observed. The predicted load-deflection behaviour was reasonably good for the specimens with rectangular and triangular fingers but was too stiff in the plastic load region. It is also important to point out that the load-deflection curves for the 100 mm wide specimens are within a few percent of its ultimate plastic load of about 183 kN but the curves for the wider 200 mm specimens are still on the strain hardening portion. This indicates that the strain limiting criterion is affecting the load carrying capacity of the wider 200 mm specimens more than the 100 mm specimens. 7.5. Strain concentrations—-series 2 Contour plots of the strain intensity have been produced for a few specimens i.e. specimen 16 (200-R) and specimen 21 (100-R). The plots of the strain intensities in the plate are presented as a ratio of the strain over the ultimate strain in Fig. 12. The corresponding positions of the fillet weld and RSNA and OA zones are as indicated. Although the strain intensities for the welds are not presented, they were also examined within the FE procedure for possible strain limitations. 7.5.1. 200 mm wide specimens Examination of the contours of Fig. 12 for specimen 16 (200-R) for the last load increment show that the strain is concentrated within the OA zone at the end of the finger with only a low strain intensity in the HAZ along the weld. There is no significant straining in the parent metal away from the HAZ. For specimen 12 (200Triangular finger), the high strain intensities are still prevailing at the OA zone at the end of the finger but has now appeared at the edge of the plate, albeit at a slightly lower intensity compared to the first location. For specimen 14 (200-Parabolic finger), the location of the greatest strain intensity has moved to the edge of the plate and contributed to a reduction of the strain intensity at the OA zone at the end of the finger. 7.5.2. 100 mm wide specimens Fig. 12 for specimen 21 (100-R) shows a high strain concentration in the OA zone at the end of the finger. As the load increased, additional straining was confined to the OA zone in this location. There was no indication of any excessive straining in the HAZ along the weld or in any other locations. There was no significant difference in the strain intensity plots for specimen 22 (100-T) and specimen 24 (100-P) from the specimen with a rectangular finger.

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Fig. 12. Strain intensity contours at last load increment; (a) material allocation, (b) specimen 16, and (c) specimen 21.

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8. Discussion Finite element analysis of these cruciform joints has shown that the strains are highly concentrated in the weaker OA zone at the end of the fingers. This localised strain concentration does not spread into other regions to distribute the loads as more plastic deformation occurs but instead strains the OA zone to its ultimate strain and consequently fails in this OA zone. At failure, the uniform longitudinal stress in the 100 mm plate is higher than the stress in the 200 mm plate. From the FE analysis, the stress in the 100 mm plate is between 367 and 371 MPa whereas the stress in the 200 mm plate is between 305 and 344 MPa. This result indicates that the 100 mm specimen is more efficient in carrying the load from a finger to a plate. This is because the HAZ in the 100 mm plate, being 75% of the total width, has achieved its limiting strain and the plate is carrying maximum load over a greater proportion of the plate. The HAZ in the 200 mm plate is only 38% the total width of the plate and thus is carrying maximum load only across a small proportion of the plate, the remaining width is at a lower stress level. With the narrower 100 mm wide specimens, the width of the HAZ is 75% of the total width of the plate. It is therefore not surprising that trimming the end of the finger could not reduce the proportion of load being transferred to the finger through the HAZ at the end of the finger. The plate is not sufficiently wide to transfer load through shear along the sides of the finger. The 200 mm wide specimens on the other hand have a HAZ width only 38% of the total plate width. It is therefore possible to reduce the proportion of load being transferred through the HAZ at the end of the finger and increase the proportion transferred through shear along the sides of the finger by trimming the rectangular finger. The trimming of the rectangular finger has reduced the strain concentration in the HAZ at the end of the finger but instead increased the strain concentration at the edge of the plate and precipitated failure there. This clearly shows that there are two locations where strain concentrations may concentrate and has highlighted the fact that reducing the stress concentration at one end may increase the stress concentration at the other end with equally damaging results. A possible method of improving the load carrying capacity is to increase the length of connection and this idea will be developed in future studies.

9. Conclusions The main points that can be derived from the experimental investigation and finite element analyses of a welded connection with a reduced strength zone are as follows: The results indicate that trimming the rectangular finger to either a triangular or parabolic finger does not increase the load carrying capacity. The most efficient welding configuration is four single run fillet welds beginning at the end of the finger and terminating at the edge of the plate. The welds should be laid with full interpass cooling. Welding around the end of the finger in the cruciform reduces the ultimate load due to a larger HAZ because of extended welding time. Plastic strength calculations across alternative failure planes may not be adequate

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for determining the ultimate load of welded cruciform connection in aluminium alloy, have been found to over-estimate the strengths of some connections. The failure load for short connection lengths tends to be much lower than the predicted strength compared to a longer connection length for the same plate width. The mode of failure of the welded cruciform is usually with a distinctive dishlike necking around the welded finger within the RSZ of the joint. The analyses have shown that the strain intensities were high in the OA zone in front of the finger for all specimens and ultimately caused the weak zone to fail. An FE model was proposed to predict the failure load to within 12% of the actual loads by adopting a strain-limiting criterion for the material in the OA zone. Other possible modes of failure associated with welding defects or improper welding procedures are cracks initiated at weld craters which can significantly reduce the load carrying capacity and excessive penetration of the weld over thin plates which in effect make a lower strength butt weld rather than a fillet connection. Acknowledgements We wish to express our gratitude to the Defence Research Agency (Military Division), U.K. for providing funding to support this work and to extend our thanks to Mr D. Webber and Mr T. Arnitt for their helpful advice and discussions. Appendix A The extent and severity of strength loss of the Al-4.0%Zn-2.0%Mg alloy were investigated in this project and were reported earlier [7–9], but are tabulated in Tables 5 and 6 here to illustrate its application to the welded cruciform connections. These Table 5 Estimation of the extent of strength loss Method

Extent HAZ/RSZ

Chan [7]

Aw Aw Extent of HAZ defined by xA=9.6 , xB=14.2 ⌺d ⌺d Two-zone RSZ model: Zone A defined as the re-solution treated and naturally aged (RSNA) zone extending to a distance of zRSNA=xA from the centreline of the weld, and Zone B, defined as the over-aged (OA) zone of width zOA=0.5(xA+xB), as shown in Fig. 2. RSZ increased by a factor of about 1.80 for double pass welds Extent of HAZ, z=dahz0, where d—effect of 3 or more heat flow paths, a—effect of interpass temperature, h—effect of heat build-up due to free edges or nearby welds, and z0—basic value

BS8118: 1991 [10]

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Table 6 Estimation of the severity of strength loss Method

Severity of strength loss

Chan [7]

The proof stress is more severely affected by the weld heat than the ultimate stress. Softening factors for the ultimate and proof stresses are as follows: for single pass welds, sult=0.845 of sult(parentmetal) and s0.002=0.751 of s0.002(parentmetal) for double pass welds, sult=0.781 of sult(parentmetal) and s0.002=0.632 of s0.002(parentmetal) Severity is defined by a softening factor, kz which is applied to the limiting stress for local capacity. kz=0.8 for single pass welds with cooling, and kz=0.6 for continuous welding around the finger

BS8118: 1991 [10]

predictions are presented in parallel to the recommendations contained in BS8118:1991 Structural use of aluminium [10]. Calculations for the extent of the HAZ of the cruciform specimens using the measured weld area of 33 mm2, for three heat flow paths with a multiplication factor of 1.8 for double welds with a very short time interval, produced values of xA=25mm, xB=37mm leading to zRSNA=25mm and zOA=31mm. A summary of the material properties and zone boundaries for series 2 specimens, where the welds were laid in a continuous run, is shown in Table 7.

References [1] Kelsey RA. Effects of heat input on welds in aluminium alloy 7039. Welding Journal Research 1971;50(Supplement 12):507–14. [2] Bartle PM. The metallurgical factors affecting the properties of welds in Al-Zn-Mg alloys. In: Proceedings of the Conference on Weldable Al-Zn-Mg alloys, The Welding Institute, 1969:5–13. [3] Bartle PM, Young JG. Weldability of the new Al-Zn-Mg alloys. In: Proceedings of the Second Commonwealth Welding Conference, The Welding Institute, 1965:246–53.

Table 7 Material properties and zone boundaries for Series 2 specimens Description

Location

s0.002 (MPa)

sult (MPa)

⑀ult

Weld metal RSNA OA—double Parent metal

in the weld 0ⱕxⱕ25 mm 25ⱕxⱕ31 mm 31 mmⱕx

124.5 250.0 228.0 361.0

286.0 405.0 325.4 416.6

0.205 0.200 0.116 0.089

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[4] Cullimore MSG. Static design of welded joints to BS 8118: structural use of aluminium. In: Proceedings of the Fifth International Conference on Aluminium Weldments, 1992. [5] Hill HN, Clark JW, Brungraber RJ. Design of welded aluminium structures. Journal of the Structural Division, Proceedings of the American Society of Civil Engineers 1960;86:101–24. [6] Robertson I. Strength loss in welded aluminium structures, Ph.D thesis. Cambridge University, Department of Engineering, 1985. [7] Chan TK. Stress concentrations in weld heat affected zones in aluminium-zinc-magnesium alloy. Ph.D thesis. Cambridge University, Department of Engineering, 1992. [8] Chan TK, Porter Goff RFD. Extent and severity of heat-affected zones in multiple pass welds. In: Vincent L, editor. Proceedings of MAT-TEC 91, 1991:305–10. [9] Chan TK, Porter Goff RFD. Experimental investigation into the structural behaviour of a welded Al-Zn-Mg alloy joint. In: Proceedings of the Third International Conference on Aluminium Alloys: Their Physical and Mechanical Properties, 1992:69–74. [10] British Standards Institution. BS 8118: Structural use of aluminium, Part 1: Code of practice for design, 1991.