FATIGUE ASSESSMENT OF BILGE KNUCKLEJOINT OF VLCC ACCORDING TO JTP/JBP RULES

Transcription

1 FATIGUE ASSESSMENT OF BILGE KNUCKLEJOINT OF VLCC ACCORDING TO JTP/JBP RULES 1. DESCRIPTION OF TEST AND EXPERIMENTAL RESULTS The model was a bilge knuckle section for a double hull VLCC in approximately 1/3 scale, Ship Research Panel 245 (2001). The model was about 6m in length, 5m in width, 3.6m in height and 20tons in steel weight. To make stress distributions of the model similar to those of the actual ship, a three-floor space in the longitudinal direction was modelled. The model was fixed to a rigid wall at the double hull side, with the ship s bottom being upside and the inner bottom being downside. The load was applied by three syncronised hydraulic jacks on the centreline of the double bottom, see Fig. 1. The model was built from mild steel in accordance with NK rules. In order to initiate cracks only at the bilge knuckle section of the centre floor, the weld leg length at the bilge knuckle section at the other floors was increased, and the weld toes were ground smooth. Fig. 1. Bilge knuckle model. The fatigue loading was constructed from a block loading simulating variable amplitude loading conditions during sea-going service. The typical mean stress alternation for a VLCC is a tensile mean stress in full load condition and a compressive mean stress in ballast condition. The applied load cycles were therefore composed of five blocks with a tensile mean stress and five similar block with a compressive mean stress. The load history is shown in Table 1 and Fig. 2. One loading set consisted of 10 blocks, and the number of cycles in each block was 10,000, i.e., the number of cycles of each set was Fig. 2. Block program loading. 3-1

2 Table 1. Details of the block loading program. Block no P/2 (kn) Pm (kn) Measured hot spot stress amplitude during the test is shown in Fig. 3. Surface fatigue cracks only initiated at the weld toe of bilge knuckle part of centre floor. After 2x10 5 cycles the hot spot stress at this location started to decrease, indicating crack initiation. Applying the 5 % drop criterion, fatigue crack initiation had taken place at 4.3x10 5 cycles. Fig. 3. Hot Spot Stress Amplitude. The test ran for 1.23x10 6 cycles. At this stage the fatigue crack surface was investigated. Contours of the crack at various stages are shown in Fig. 4. At 38 or 48x10 4 cycles of fluctuating load, many surface fatigue cracks had been initiated at the weld toe. Subsequently they were combined into a shallow and broad surface crack, which propagated from around the centre of the transverse floor in the depth direction. At end of test the crack was nearly through thickness. In the following, we assume that the crack initiation life, Nc, is 4.3 x10 5, and the through thickness life, Nf, is 1.23 x10 6. Fig. 4. Crack growth behaviour. 3-2

3 2. EVALUATION OF HOT-SPOT STRESS The VLCC bilge knuckle model shown in Fig. 5 was analysed using design procedures of the 2nd draft of JTP / JBP rules. In the FE calculation of Hot-Spot Stress (H.S.S.), plating members are modelled by NASTRAN 4-node thin shell elements. The size of the FE mesh close to the intersection was t by t. Fig. 5. Shell FE model used in HSS calculations. The shape of the weld bead is shown in Fig. 6. It is shown that the frank angle is about 45 degree. Because the plates intersect at an angle of 45 degree, this joint can be modelled as shown in Fig. 7, and x wt defined in JTP rule becomes zero in this case. Therefore, stress at a distance of 0.5t from theintersection line, where t is the net thickness of the inner bottom plate, was used as H.S.S. used in JTP procedure. In this case, H.S.S. corresponding to a jack load of 1kN is MPa. Fig. 6. Shape of the weld bead. Fig. 7. Simplified model of the weld joint. 3-3

4 For JBP procedure, H.S.S. assessment using 1.5t-0.5t linear extrapolation to the intersection line was applied. A weld notch factor K f =1.30 and a correction factor for material f material =0.909 were applied. In this case, H.S.S. corresponding to a jack load of 1kN is MPa, and a correction factor for knuckle shape 0.8 was applied. The stress in the vicinity of the intersection was also calculated by the solid FE model shown in Fig. 8. This solid model was embedded in the shell FE model as shown in this figure. Adaptive P-method and a shape function suitable for the treatment of stress singularity were applied in this calculation. The surface stress was evaluated by calculating the strain components at the corner points of solid elements. Fig. 8. Embedded solid FE model. The comparisons of the surface (top and bottom) and membrane stresses calculated by the shell and solid models are shown in Fig. 9. It is shown that the shell FE calculation gives a fairly good agreement with the solid FE calculation. Fig. 9. Comparison of the stress calculated by shell and solid FE models. 3. FATIGUE STRENGTH ASSESSMENT For JTP procedure, Class D SN data of Table 2 was used. The effect of mean stress was considered by assuming a stress range equal to the tensile component plus 60% of the compressive component. For JBP procedure, UK-Den basic B curve shown in Table 3 was used. The procedure for mean stress correction, in 3-4

5 which shake-down of residual stress is taken into account, was applied. The results from analyses are shown in Table 4 and Fig. 10. References [1] Ship Research Panel 245. (2001). Study on Ship Structural Life of the Double Hull Tanker. Shipping Research Association of Japan (in Japanese). Table 2. JTP Basic SN Data, In-Air. Class K 1 m Standard Deviation K 2 S q log 10 log e log 10 log e N/mm 2 D 3.988E E Table 3. JBP SN Data (UK-Den basic B curve). m C m C -2.Stdev x x10 21 mean x x10 21 ( N = C σ m ; σ (10 7 C) l/m C 0 σ m0 ; Table 4. Calculated Fatigue Damage for JTP/JBP. Damage factor N for D=1.0 Initiation Through thickness N D=1 /N c N D=1 /N f N c =4.3x10 5 N f =1.23x10 6 JTP mean 1.1x JTP -2.Stdev 4.2x JBP mean 1.1x JBP -2.Stdev 4.3x Fig. 10. Comparison of experimental and calculated fatigue lives. 3-5