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Ultimate crushing strength criteria for GTT NO96 LNG carrier cargo containment system under sloshing load
Ocean Engineering 188 (2019) 106224
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Ultimate crushing strength criteria for GTT NO96 LNG carrier cargo containment system under sloshing load Young IL. Park Department of Naval Architecture and Offshore Engineering, Dong-A University, Republic of Korea
A R T I C L E I N F O
A B S T R A C T
Keywords: LNG carrier cargo containment system (CCS) Finite-element method Failure mode Crushing failure Sloshing
The ultimate crushing strength of a Gaztransport and Technigaz NO96-type liquefied natural gas cargo containment system (NO96 CCS) was investigated through experiments and numerical simulations as one of the possible failure modes. As crushing failure is directly related to material failure along the through-thickness direction of the horizontal members, i.e., the bottom plate of the primary box and the top plate of the second ary box, small-scale experimental crushing test results were applied. Instead of using the test results directly, static failure capacity estimates were obtained via nonlinear finite-element analyses, which accounted for the crushing of the horizontal members and hence also the bending that this introduced in the horizontal members. The crushing-failure capacities were determined by the calculated bending moment in the two horizontal members. The calculated capacity will be used to select a design criterion using the vertical stress at the midheight of the bulkhead panel of the secondary box as an early design objective.
1. Introduction Since the Paris Agreement on climate change, the global liquefied natural gas (LNG) market has grown by 4%–6% annually. According to a report from the International Gas Union (IGU), the global LNG market is growing steadily, with an annual increase of 6.2%. The Asia–Pacific region accounts for 72.4% of the world’s total imports, and the largest importers of LNG have shifted from Japan and Korea to China, India, and the Southeast Asian countries. In particular, the Chinese government aims to increase its share of the global LNG from 6% to 15% by 2030. Moreover, the development of the Floating Storage Regasification Unit (FSRU) is expected to increase the import volumes of small LNGimporting countries and create new importers. The FSRU is considered to be easier and less expensive than LNG land terminals (IGU, 2017). It has lower initial investment cost than land terminals. Additionally, land terminals involve complicated licensing procedures from site selection to construction. With global market trends, technological development, and recovering international oil prices, demand for LNG is increasing, leading to increased orders in shipbuilding, such as LNG carriers and FSRUs (Digital Times, 2018). As LNG is transported in the liquid state at a cryogenic temperature of 163 � C, a cargo hold is installed inside the LNG carrier. The LNG cargo containment system (CCS) is generally classified as an independent-type or membrane-type cargo hold. As shown in Fig. 1, the
membrane-type cargo hold is strong (LNG World Shipping, 2017). The Gaztransport and Technigaz (GTT) NO96 and Mark III are representative membrane-type LNG CCSs (see Fig. 2). The Mark III is a PUF-type cargo hold with horizontal plywood on top of the primary barrier and at the bottom of the secondary barrier, and the R-PUF has it in its interior. There is a triplex between the barriers, and the primary and secondary barriers are fixed with an adhesive. The NO96 is a boxtype cargo hold with a box structure through the plywood, and it uses perlite or glass wool (GW) for the interior. There is invar steel between the barriers, and each barrier is fixed to the hull with a binding device. The NO96 LO3, NO96 LO3þ, Mark III Flex, and Mark V have been modified to reduce the boil-off gas (GTT, 2017). A major advantage of a membrane-type LNG CCS, such as the GTT MARK III and NO96, compared with an independent-type CCS is the flexibility for increasing the size of the cargo tank. However, the risk of sloshing damage to the membranes is an important factor to consider when selecting a membrane-type LNG CCS, particularly for offshore applications, e.g., LNG FPSO and FSRU. To assess the risk of sloshing damage and determine the feasibility of a membrane-type LNG CCS, several factors must be considered. For instance, adequate design capacities and parameters must be considered for each CCS and each design choice applied. Daewoo Shipbuilding & Marine Engineering Co., Ltd. developed a two-row tank arrangement for the NO96 CCS for LNG offshore
E-mail address: [email protected]. https://doi.org/10.1016/j.oceaneng.2019.106224 Received 15 November 2018; Received in revised form 15 July 2019; Accepted 16 July 2019 Available online 8 August 2019 0029-8018/© 2019 Elsevier Ltd. All rights reserved.
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applications to reduce the sloshing design load (Ryu et al., 2016). Various experimental and numerical studies have been performed on the GTT MARK III-type CCS (Kim et al., 2010; Chun et al., 2009). Arswendy and Moan (2015) conducted experimental and numerical studies to determine the buckling and crushing strengths for a T-shaped plywood specimen as a part of the NO96 CCS. Chun et al. (2011), Hwang et al. (2014), Lee and Shin (2014), Lee et al. (2012), Nho et al. (2012), Pillon et al. (2009), Yoo et al. (2011), and Wang et al. (2012) performed structural evaluations under sloshing impact loads to define design sloshing pressure and/or ultimate strength of a MARK III-type LNG CCS depending on several failure modes. Dobashi and Usami (2012) proposed a design dynamic amplification factor for a NO96-type CCS. Lee et al. (2011) and Wang et al. (2009) investigated the general structural behavior of a NO96-type CCS under sloshing impact via simulations. However, few studies have been performed on the NO96 CCS, particularly for early-design purposes. Park and Lee (2018) proposed design dynamic buckling strength capacities for the critical failure mode of a box-type LNG CCS, such as the NO96 CCS. DNVGL (2016) identifies four design failure modes of the NO96 CCS: shear, bending failures at the cover panel of the primary box, buckling failure of the bulkhead panel of the primary and secondary boxes, and crushing failure at the intersection of the primary and secondary box bulkheads. However, it applies a comparative approach to a 138000-m3 an LNG carrier, which is a proven design and target vessel. In this study, the failure modes of the NO96 CCS are classified, and criteria are developed for ultimate crushing failure via numerical sim ulations tuned through experiments based on risk. The results will aid the initial design of LNG cargo holds and improve currently used holds. First, the failure modes of the box-type cargo hold as proposed by classification societies are compared, and several failure modes are added. To confirm the additions and the new list, a load due to sloshing is selected for a static-load test of an independent box-type cargo hold. A nonlinear finite-element (FE) analysis is performed to validate the experimental results. The experiments and FE analysis are also performed for additional failure modes, and design criteria are presented.
In this study, the design capacity of the crushing of plates at the bulkhead intersections was investigated numerically. 3. Experimental tests for crushing-failure analysis of plywood components Arswendy and Moan (2015) performed experimental tests of the crushing and buckling failure modes of plywood, which is the major strengthening part of a box-type LNG CCS. Fig. 5 shows the test spec imen configurations that they used for the crushing test, where (a), (b), and (c) correspond to crushing tests of a plywood plate and plywood bulkhead, of a plywood plate and steel bulkhead, and of a steel plate and plywood bulkhead, respectively. As the major failure mode of crushing for a box-type NO96 CCS should be related to horizontal member failure and vertical bulkhead member failure, the test configuration (a) of Fig. 5 was selected in this study. Fig. 6 shows the relationship between the vertical load and displacement for the (a) configuration. 4. FE simulations Proper interpretation of the test results requires knowledge of the correlation between the loads applied to the NO96 CCS and those applied to the test specimen under which the local stress at the inter section of the plates is similar. Thus, it is necessary to determine whether there is any difference in the stress concentration between the test setup and the NO96 CCS during crushing failure. To investigate the nonlinear behavior related to crushing failure at the intersection of the vertical bulkheads of the primary and the sec ondary boxes, a nonlinear static FE analysis was performed using pre vious experimental results (Arswendy and Moan, 2015) for cases (a) and (b) of Fig. 5, as the main focus of the present study was the crushing failure of the horizontal member. For the FE simulations, the commercial FE code ABAQUS was employed, with the specifications presented in Table 2. The material stiffness depends on the design temperature, plywood grain direction (orthotropic material properties), and bending/mem brane characteristics, as shown in Table 3 (DNVGL, 2016). An elastic modulus of 205800 MPa and a Poisson’s ratio 0.3 were considered for steel part in case (b) of Fig. 5. The indentation of the vertical bulkhead plates into the horizontal plate was localized to small regions directly adjacent to the edges of the bulkhead plate. An accurate FE representation of this localized damage required an extremely fine element mesh. Instead, an elastoplastic ma terial model for the thickness behavior of the horizontal plates was calibrated to provide an equivalent overall inelastic stiffness charac teristic of the bulkhead intersections for the mesh density used in the model. The crushable-foam model available in ABAQUS was applied to allow for volumetric plastic compressibility of the material. This was necessary to obtain reasonable inelastic deformation modes along the thickness direction of the horizontal plates. The incompressible plastic behavior of the standard metal plasticity model resulted in the unphysical locking of the elements of the horizontal plates located above and beneath the vertical bulkhead plates; consequently, the inelastic deformation was forced to occur in the adjacent element. Because the crushable-foam model is only applicable to isotropic materials in ABAQUS applications, four-node shell elements with orthotropic material properties were embedded into the solid element region, as shown in Fig. 7. The parameters of this stiffness-equivalent elastoplastic crushablefoam model were determined via FE simulations of the experimental test (Arswendy and Moan, 2015). The through-thickness elastic modulus, material yield stress, and plastic tangent modulus were tuned such that the simulated load–displacement response fitted the test re sults, as shown in Table 4. A comparison of the test and simulation
2. NO96 CCS The GTT NO96 is a box-type membrane LNG CCS containing plywood panels for strength, GW or perlite for insulation, and other components, e.g., a securing device and a membrane. Various NO96 CCSs have been developed to increase the thickness of the internal bulkheads or increase the number of internal members to bear greater load, as shown in Fig. 3 and Table 1. In this study, the crushing strength was evaluated for CCS#2, #3, and #4. The DNVGL (2016) and Lloyd’s Register (2009) identify the following failure modes for the NO96 CCS as shown in Fig. 4. 1) Bending/shear (including combined shear and bending) failure of the cover plate of the primary box 2) Buckling failure of the vertical bulkheads (internal and external bulkheads) 3) Crushing failure A. Bulkhead intersections B. Resin rope 4) Failure related to the box-securing system A. Bulkheads of the corner border box—cleats at the 90� transverse corner zone B. Bulkheads of the corner border box—cleats at the 135� longitu dinal corner zone 5) Failure connected to the membrane tongue attachment 2
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somewhat Fig. 11 shows the load vs. deflection curves under out-plane pressure for the two aforementioned FE models. The outputs are identical. As the FE model with the 3D solid elements and two-dimensional (2D) shell embedded was implemented for the horizontal member of the CCS—1) bottom plate of the primary box and 2) top plate of the secondary box—the out-plane behavior was more critical than the inplane behavior. Therefore, it was found to be acceptable to apply these concepts to the FE model. Elastoplastic FE analyses of a flat NO96 CCS were performed. The Table 1 General specifications of the NO96 CCS for each case study. Fig. 1. Global trends of LNG carrier CCSs.
results is shown in Fig. 6. Fig. 8 shows a comparison of the test results of Arswendy and Moan (2015) with the numerical simulations. In this study, the nonlinear material properties for the through-thickness direction, which is for crushing failure on the horizontal member, were calibrated according to case (a) (plywood–plywood), and the same material properties were considered for case (b). The comparison between the experimental test of case (b) and the numerical simulation reveals somewhat conservative behavior in the simulation; however, this can be acceptable for design purposes. Because the crushable-foam model is only applicable to isotropic materials in ABAQUS applications, four-node shell elements with orthotropic material properties were embedded into the solid element region, as shown in Fig. 7. Several FE analyses were performed to confirm whether the struc tural behavior of the embedded shell in the solid model was practically reasonable. Two FE models were employed: orthotropic shell elements (left figure of Fig. 9) and orthotropic shell elements with threedimensional (3D) solid elements (right figure of Fig. 9). The elastic buckling under in-plane compressive loads was evalu ated, as shown in Fig. 10. The buckling strength in the right frame of Fig. 10 is 12% higher than that in the left frame, which could be
Case No.
Primary box Thickness of cover plate
Thickness of BHD
Secondary box Thickness of BHD
No. of combs
CCS#1 CCS#2 CCS#3 CCS#4
12 mm 12 mm � 2 12 mm � 2 12 mm � 2
9 12 12 15
9 12 12 15
0 0 1 2
Fig. 4. Failure modes considered for the NO96 CCS (DNVGL, 2016).
Fig. 2. Overview of GTT Mark III and NO96 LNG CCSs (source: http://www.gtt.fr/).
Fig. 3. Insulation box of the NO96 CCS for various reinforcement designs (source: www.gtt.fr). 3
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Fig. 5. Crushing-test configurations (Arswendy and Moan, 2015).
Fig. 6. Relationship between the vertical load and displacement for the test configuration shown in Fig. 5 (Arswendy and Moan, 2015).
Table 3 Stiffness properties of the integrated orthotropic material for plywood panels (DNVGL, 2016).
Table 2 Specifications of the FE analysis. Item
Notes
Software (solver)
ABAQUS/Standard (V.6.14)
For static-buckling calculation
Element type
4-node shell
Reduced integration for strength member, i.e., plywood panel Spring connection, i.e., stapling Dependent on operational temperature, i.e., 163 to 20 � C
Material type Boundary conditions Surface contact
2-node beam elements Homogeneous orthotropic type 4 edges—simply supported Kinematic hard contact option
Parameter
Temperature 20 � C
Em,1 [MPa] Em,2 [MPa] Em,3 [MPa] Gm,12 [MPa] Gm,13 [MPa] Gm,23 [MPa]
ν12 ν 13 ν 23
Eb,1 9 mm [MPa] Eb,2 9 mm [MPa] Eb,1 12 mm [MPa] Eb,2 12 mm [MPa] Density [ton/mm3] σ F [MPa]
4
9450 8000 820 790 325 260 0.1 0.1 0.1 10950 6550 10450 7000 6.8e-10 43.0
163 � C 13200 11200 1800 2900 700 550 0.1 0.1 0.1 15350 9150 14650 9800 6.8e-10 65.0
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Fig. 7. Schematic of the crushable foam for the FE model.
the horizontal members, i.e., the bottom plate of the primary box and the top plate of the secondary box. The shell elements were constrained to the solid element mesh such that the composite mesh represented the total stiffness of the plates. The boundary conditions of the FE model were imposed at the resin support locations as simple support. FE analyses for five cases with different loads were conducted (refer to Fig. 13). The main objective of this study was to develop strength criteria for the crushing behavior of the NO96 CCS. Therefore, the structural response was investigated for various loading conditions. The vertical stresses in the bulkheads of the boxes were considered to be representative of the forces transferred from the primary box to the secondary box through the bulkhead intersections. Therefore, the stress was chosen as a candidate for the acceptance criterion of the crushing failure. Fig. 14 and Table 5 show the stress check points, which could have been representative strength criteria for the five cases of the load. Figs. 15–29 show the stress behavior under the applied pressure at the five check points for CCS#2, CCS#3, and CCS#4. The stress
Table 4 Material properties for the calibrated crushable-foam model. Material properties for crushable foams Elastic Crushable foam Crushable-foam hardening
Elastic modulus [MPa] Poisson’s ratio Compressive yield stress ratio Hydrostatic yield stress ratio Yield stress [MPa] Uniaxial plastic strain
80 0.1 0.9 0.9 10–28 0.0–0.93
analyses considered the thickness inelastic behavior of the primary-box bottom plate and the secondary-box top plate using the properties of the calibrated material (refer to Table 4). The quarter-size FE models of the insulation boxes used in the ana lyses are shown in Fig. 12, for CCS#2, CCS#3, and CCS#4 with different numbers of combs at the secondary box. The model was built using eight 3D node brick elements combined with four 2D node shell elements for
Fig. 8. Comparison of the test results with the numerical simulation results for the simplified model (left: plywood–plywood, case (a); right: steel–plywood, case (b) (Arswendy and Moan, 2015)).
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Fig. 9. 2 FE models: left, orthotropic shell model; right, 3D solid element þ 2D shell model.
Fig. 10. Elastic buckling strength for the in-plane compressive load comparison (orthotropic shell model – left: 2.65 MPa, shell–solid combination).
Fig. 11. Out-plane pressure behavior for an out-plane uniform pressure.
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Fig. 12. 1/4-sized FE model of the NO96 CCS (left: CCS#2; middle: CCS#3; right: CCS#4); arrow: grain direction.
responses show that the redistribution of forces away from the bulkhead intersections and toward the adjacent parts of the bulkhead edges occurred in the structure owing to the bending of the horizontal bottom plate of the primary box and the cover plate of the secondary box. Owing to the crushing deformation at the intersections of the pri mary and secondary bulkheads, loads were redistributed from the bulkhead intersection to the remainder of their edges from the bending capacity of the bottom plate of the primary box and the cover plate of the secondary box. Both plates (the bottom plate of the primary box and the cover plate of the secondary box) were bent along their two in-plane directions—the strong and weak directions—and the capacity and effi ciency of the plates to redistribute loads was dependent on them maintaining their bending strength. To define the ultimate bending
capacities of the horizontal members (the bottom plate of the primary box and the cover plate of the secondary box), four additional stress responses were observed, as shown in Table 6 and Fig. 30. The strong/weak directions of the bending-stress contours of the bottom plate of the primary box and the cover plate of the secondary box for a full patch load are shown in Figs. 31–34. The criteria for the ultimate bending strength in Table 7, depending on the plywood grain direction, i.e., the strong or weak direction, were applied. The critical design pressures for each loading case on CCS#2, CCS#3, and CCS#4 were determined by Table 7: 1100 and 760 Nmm/mm for the strong and weak directions of the bending capacities for the bottom plate of the primary box, respectively, and 1850 and 1380 Nmm/mm for
Fig. 13. Cases of the FE analysis, with different patch loads. 7
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the strong and weak directions of the bending capacities for the top plate of the secondary box, respectively, as shown in Figs. 35–46. From Figs. 35–46, the following conclusions are drawn. - The structure had a limited capacity to transfer the applied force. The most critical patch-load configurations occurred in the case of a uniform load. - The stress responses of the bulkheads of the primary box, the bulk heads of the secondary box, and the hotspots near the horizontal plates exhibited similar behavior. 5. Acceptance criteria for crushing failure Crushing strength should be measured in terms of the nominal stress in the vertical bulkhead of the NO96 CCS, as the basic designer conducts FE calculations using 2D shell-element modeling. Crushing deformation between the primary and secondary boxes leads to the failure of the horizontal plates, i.e., the bottom plate of the primary box and the cover plate of the secondary box. Loads applied on the cover plate of the primary box are redistributed from the intersection of the bulkheads to the remainder of their edges by the horizontal plates, i.e., the bottom plate of the primary box and the cover plate of the secondary box. Both the horizontal plates were bent along their two inplane directions, and the ability and efficiency of the plates to redis tribute the loads were dependent on the bending and shear integrity of the horizontal plates. Plate failure effect analysis was performed, as summarized in Table 8 and Table 9, to assess the conditions of the critical failure of the inter section of the bulkheads. Bending was a more critical failure mechanism than shear for the plates. Therefore, the total capacity of the bulkhead intersections was evaluated via a bending-capacity assessment. The failure effect analysis in Tables 8 and 9 indicates that #7 and #8 were critical effects. They represent the worst case of the acceptable consequence. The system was induced to significantly limit the redis tribution of forces to the edge of the bulkhead. Therefore, the acceptance criterion for the bulkhead intersections was determined by #7 and #8. Each capacity of the bending mode for each of the two plates was estimated according to the pressure–bending stress curves obtained via
Fig. 14. Stress-output check points in the primary and secondary boxes.
Table 5 Stress-response check points on the vertical bulkhead member. ID
Description
Location
A B C D E
Nominal stress at mid-height location of vertical bulkhead Hotspot stress at bulkhead near bottom plate Hotspot stress at bulkhead near cover plate Nominal stress at mid-height location of vertical bulkhead Hotspot stress at bulkhead near bottom plate
Primary box Primary box Secondary box Secondary box Secondary box
Fig. 15. Vertical stress at the mid-height of the primary-box bulkhead (Point A) vs. the applied pressure for CCS#2. 8
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Fig. 16. Vertical stress at the mid-height of the primary-box bulkhead (Point A) vs. the applied pressure for CCS#3.
FE analysis. The estimated capacities are presented in Tables 10–12 for CCS#2, CCS#3, and CCS#4. The underlined data are the capacities of #7 and #8 obtained via the failure effect analysis of the horizontal plates, i.e., the bottom plate of the primary box and the cover plate of the secondary box. In the previous section, the FE model using four 2D node shell ele ments could yield reasonable stress values in hotspots such as check points B, C, and E. Therefore, A and D were the likely strength criteria for
the crushing-failure mode. In this study, point D was used to define the failure criteria for crushing. Figs. 47–51 show reference stresses at each check point for CCS#2 under the critical failure crushing pressure determined by Table 10 Maximum crushing capacity by bending strength on the horizontal plates for CCS#2. Figs. 52–54 correspond to CCS#3, and Figs. 55–58 correspond to CCS#4. Table 13 presents the design reference stresses at check point D for
Fig. 17. Vertical stress at the mid-height of the primary-box bulkhead (Point A) vs. the applied pressure for CCS#4. 9
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Fig. 18. Vertical stress at the hotspot of the primary-box bulkhead (Point B) vs. the applied pressure for CCS#2.
Fig. 19. Vertical stress at the hotspot of the primary-box bulkhead (Point B) vs. the applied pressure for CCS#3.
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Fig. 20. Vertical stress at the hotspot of the primary-box bulkhead (Point B) vs. the applied pressure for CCS#4.
Fig. 21. Vertical stress at the hotspot of the secondary-box bulkhead (Point C) vs. the applied pressure for CCS#2.
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Fig. 22. Vertical stress at the hotspot of the secondary-box bulkhead (Point C) vs. the applied pressure for CCS#3.
Fig. 23. Vertical stress at the hotspot of the secondary-box bulkhead (Point C) vs. the applied pressure for CCS#4.
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Fig. 24. Vertical stress at the mid-height of the secondary-box bulkhead (Point D) vs. the applied pressure for CCS#2.
Fig. 25. Vertical stress at the mid-height of the secondary-box bulkhead (Point D) vs. the applied pressure for CCS#3.
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Fig. 26. Vertical stress at the mid-height of the secondary-box bulkhead (Point D) vs. the applied pressure for CCS#4.
Fig. 27. Vertical stress output at the hotspot of the secondary-box bulkhead (Point D) vs. the applied pressure for CCS#2.
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Fig. 28. Vertical stress output at the hotspot of the secondary-box bulkhead (Point E) vs. the applied pressure for CCS#3.
Fig. 29. Vertical stress output at the hotspot of the secondary-box bulkhead (Point E) vs. the applied pressure for CCS#4.
Table 6 Stress-response check points on the horizontal member, i.e., the bottom plate of the primary box and the cover plate of the secondary box. ID
Description
Location
F G H I
Strong direction of bending stress at bottom plate of primary box Weak direction of bending stress at bottom plate of primary box Strong direction of bending stress at cover plate of secondary box Weak direction of bending stress at cover plate of secondary box
Primary box Primary box Secondary box Secondary box
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Fig. 30. Locations of the stress-output checks at the bottom and cover plates of the primary and secondary boxes, respectively.
Fig. 31. Bending-stress contours at the bottom of the primary box along the strong direction.
Fig. 32. Bending-stress contours at the bottom of the primary box along the weak direction.
Fig. 33. Bending-stress contours on the cover plate of the secondary box along the strong direction.
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Fig. 34. Bending-stress contours on the cover plate of the secondary box along the weak direction.
each CCS under the critical design pressure for crushing failure. In this study, 39.1, 45.3, and 46.5 MPa were obtained as the design reference stresses at check point D for CCS#2, CCS#3, and CCS#4, respectively.
Table 7 Mean strength properties for 9-mm and 12-mm plywood laminates (DNVGL, 2016). Plywood thickness
9 mm
Temperature Mc,1 (Nmm/mm) Mc,2 (Nmm/mm)
20 � C 1100 760
6. Conclusion
12 mm 163 � C 935 650
20 � C 1850 1380
The ultimate crushing-failure strength of the NO96 CCS was evalu ated via nonlinear FE simulation. To define the nonlinear material properties related to the thickness behavior, the experimental results of a plywood-crushing test were used. They were calibrated to obtain the equivalent overall inelastic stiffness characteristics of the intersections of the bulkhead for the mesh density used in the FE model. Crushing deformation limited the redistribution of forces by the bending or shearing of the horizontal members, i.e., the bottom plate of
163 � C 1580 1180
Note: M1 denotes bending about the 2-axis, and M2 denotes bending about the 1axis.
Fig. 35. Output of the sectional moment along the strong direction at the bottom plate of the primary box (point F) for CCS#2.
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Fig. 36. Output of the sectional moment along the strong direction at the bottom plate of the primary box (point F) for CCS#3.
Fig. 37. Output of the sectional moment along the strong direction at the bottom plate of the primary box (point F) for CCS#4.
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Fig. 38. Output of the sectional moment along the weak direction at the bottom plate of the primary box (point G) for CCS#2.
Fig. 39. Output of the sectional moment along the weak direction at the bottom plate of the primary box (point G) for CCS#3.
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Fig. 40. Output of the sectional moment along the weak direction at the bottom plate of the primary box (point G) for CCS#4.
Fig. 41. Output of the sectional moment along the strong direction on the cover plate of the secondary box (point H) for CCS#2.
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Fig. 42. Output of the sectional moment along the strong direction on the cover plate of the secondary box (point H) for CCS#3.
Fig. 43. Output of the sectional moment along the strong direction on the cover plate of the secondary box (point H) for CCS#4.
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Fig. 44. Output of the sectional moment along the weak direction on the cover plate of the secondary box (point I) for CCS#2.
Fig. 45. Output of the sectional moment along the weak direction on the cover plate of the secondary box (point I) for CCS#3.
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Fig. 46. Output of the sectional moment along the weak direction on the cover plate of the secondary box (point I) for CCS#4.
Table 8 Dependence of the level of consequence on the failure mode for determining the crushing capacity.
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Table 9 Failure effect analysis of the horizontal plates for the bending mode.
Table 10 Maximum crushing capacity by bending strength on the horizontal plates for CCS#2. Load case Full patch Patch 1 Patch 2 Patch 3 Patch 4
Bottom plate of the primary box
Cover plate of the secondary box
Strong direction
Weak direction
Strong direction
Weak direction
19.36 bar 25.35 bar 22.66 bar 25.74 bar 19.05 bar
27.59 bar 40.71 bar 48.88 bar 41.35 bar 28.00 bar
20.77 bar 30.06 bar 30.65 bar 30.46 bar 21.09 bar
37.26 bar 54.62 bar 41.32 bar 57.48 bar 41.73 bar
Table 11 Maximum crushing capacity by bending strength on the horizontal plates for CCS#3. Load case Full patch Patch 1 Patch 2 Patch 3 Patch 4
Bottom plate of the primary box
Cover plate of the secondary box
Strong direction
Weak direction
Strong direction
Weak direction
14.74 bar 50.77 bar 26.04 bar 28.78 bar 35.57 bar
31.60 bar N.A. N.A. 58.69 bar 48.02 bar
16.95 bar N.A. N.A. 68.22 bar 52.31 bar
33.09 bar N.A. N.A. 64.75 bar N.A.
Table 12 Maximum crushing capacity by bending strength on the horizontal plates for CCS#4. Load case Full patch Patch 1 Patch 2 Patch 3 Patch 4
Bottom plate of the primary box
Cover plate of the secondary box
Strong direction
Weak direction
Strong direction
Weak direction
16.01 bar 20.79 bar 18.99 bar 21.00 bar 15.71 bar
33.91 bar 51.17 bar N.A. N.A. 33.98 bar
17.81 bar 23.98 bar 22.88 bar 23.18 bar 16.88 bar
45.14 bar 56.51 bar 41.38 bar N.A. 44.13 bar
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Fig. 47. Reference stresses at each check point under 27.59 bar for the critical bending failure in the uniform-pressure case of CCS#2 determined by Table 10.
Fig. 48. Reference stresses at each check point under 40.71 bar for the critical bending failure in the Patch 1 case of CCS#2 determined by Table 10.
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Fig. 49. Reference stresses at each check point under 41.32 bar for the critical bending failure in the Patch 2 case of CCS#2 determined by Table 10.
Fig. 50. Reference stresses at each check point under 41.35 bar for the critical bending failure in the Patch 3 case of CCS#2 determined by Table 10.
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Fig. 51. Reference stresses at each check point under 28.00 bar for the critical bending failure in the Patch 4 case of CCS#2 determined by Table 10.
Fig. 52. Reference stresses at each check point under 31.60 bar for the critical bending failure in the uniform-pressure case of CCS#3 determined by Table 11.
27
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Ocean Engineering 188 (2019) 106224
Fig. 53. Reference stresses at each check point under 64.75 bar for the critical bending failure in the Patch 3 case of CCS#3 determined by Table 11.
Fig. 54. Reference stresses at each check point under 52.31 bar for the critical bending failure in the Patch 4 case of CCS#3 determined by Table 11.
28
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Ocean Engineering 188 (2019) 106224
Fig. 55. Reference stresses at each check point under 33.91 bar for the critical bending failure in the uniform-pressure case of CCS#4 determined by Table 12.
Fig. 56. Reference stresses at each check point under 51.17 bar for the critical bending failure in the Patch 1 case of CCS#4 determined by Table 12.
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Y.IL. Park
Fig. 57. Reference stresses at each check point under 41.38 bar for the critical bending failure in the Patch 2 case of CCS#4 determined by Table 12.
Fig. 58. Reference stresses at each check point under 33.98 bar for the critical bending failure in the Patch 4 case of CCS#4 determined by Table 12.
Table 13 Reference stresses at check point D of each CCS under the critical design pressure for crushing failure. Critical pressure of bending failure on the horizontal members [bar] Uniform pressure Patch 1 Patch 2 Patch 3 Patch 4
Reference stress at check point D for the critical design pressure of crushing failure [MPa]
CCS#2
CCS#3
CCS#4
CCS#2
CCS#3
CCS#4
27.6 40.7 41.3 41.4 28.0
31.6
33.9 51.2 41.4
39.1 42.3 37.7 42.8 39.5
45.3
46.5 51.5 37.1
64.8 52.3
34.0
30
53.3 38.8
46.7
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the primary box and the cover plate of the secondary box. Therefore, the load capacity of the bulkhead intersections was evaluated via a bendingcapacity assessment. The failure criteria were defined through failure effect analyses. The design crushing-failure reference stresses at the middle of the bulkhead of the secondary box were proposed as 39.1, 45.3, and 46.5 MPa for CCS#2, CCS#3, and CCS#4, and the design values are applicable to the NO96 CCS FE model using 2D shell elements to define the ultimate crushing strength with the following inequality.
DNVGL-CG-0158, 2016. "Sloshing analysis of LNG membrane tanks". Class Guideline. Dobashi, H., Usami, A., 2012. Dynamic amplification factor of NO96 insulation structures of membrane system. In: Proc. of the 2012 International Offshore and Polar Engineering Conference, pp. 525–529. Gaztransport & Technigaz, Production, 2017. http://www.gtt.fr/en/technologies-servic es/our-technologies. Hwang, J.O., Chun, S.E., Joh, K.H., Combos, P., Lauzon, J.D., White, N., Kim, M.S., Park, J.B., Lee, J.M., 2014. Direct assessment of structural capacity against sloshing using dynamic nonlinear FE analysis. In: Proc. of the 2014 International Ocean and Polar Engineering Conference, pp. 169–179. International Gas Union, 2017. IGU world LNG report. IGU 78–94, 1-53. Kim, M.H., Lee, S.M., Lee, J.M., Noh, B.J., Kim, W.S., 2010. Fatigue strength assessment of MARK-III type LNG cargo containment system. Ocean Eng. 37 (Issues 14–15), 1243–1252. Lee, D.J., Shin, S.B., 2014. Numerical model for dynamic behavior of PUF in membrane type LNG cargo containment system. In: Proc. of the 2014 International Ocean and Polar Engineering Conference, pp. 157–162. Lee, S.J., Yang, Y.S., Kim, S.C., Lee, J.H., 2011. Strength assessment procedure of LNG CCS under sloshing load based on the direct approach. In: Proc. of the 2011 International Offshore and Polar Engineering Conference, pp. 183–190. Lee, S.G., Kim, J.K., Nguyen, H.A., Nam, J.H., 2012. Structural safety assessment of LNGC MARK III membrane type CCS under sloshing impact loading. In: Proc. of the 2012 International Offshore and Polar Engineering Conference, pp. 487–494. Lloyd’s Register, May 2009. ShipRight, “Sloshing Assessment Guidance Document for Membrane Tank LNG Operations”, Version 2.0. LNG World Shipping, 2017. LNGC Fleet in Service and on Order. http://www.lngworldsh ipping.com/news/view,gtt -plans-mark-v-rethink_49315. Nho, I.S., Kim, S.C., Jang, B.S., Lee, J.H., 2012. Parametric investigation on the simplified triangular impulse of sloshing pressure and categorization of the structural response on the Mark III LNG CCS. In: Prof. of the 2012 International Offshore and Polar Engineering Conference, pp. 495–501. Park, Y.I., Lee, J.H., 2018. Buckling strength of GTT NO96 LNG Carrier cargo containment system. Ocean Eng. 154, 43–58. Pillon, B., Marhem, M., Leclere, G., Canler, G., 2009. Numerical approach for structural assessment of LNG containment systems. In: Proc. of the 2009 International Offshore and Polar Engineering Conference, pp. 175–182. Ryu, M.C., Jung, J.H., Kim, Y.S., Kim, Y.I., 2016. Sloshing design load prediction of a membrane type LNG cargo containment system with two-row tank arrangement in offshore application. Int. J. Naval Architect. Ocean Eng. 8 (6), 537–553. Wang, B., Han, S.K., Kim, Y.S., Park, Y.I., Shin, Y., 2009. Sloshing model tests and strength assessment of the NO 96 containment system. In: Proc. of the 2009 International Offshore and Polar Engineering Conference, pp. 261–268. Wang, B., Shin, Y., Wang, X., 2012. Reliability-based sloshing assessment of containment systems in LNGCs and FLNGs. In: Prof. of the 2012 International Offshore and Polar Engineering Conference, pp. 502–509. Yoo, M.J., Lee, S.H., Kim, S.C., Lee, J.H., Nho, I.S., 2011. Characteristics of dynamic response of MARK III LNG containment subjected to idealized triangular sloshing impact. In: Proc. of the 2011 International Offshore and Polar Engineering Conference, pp. 177–182.
σ av � σ av c where
σ av : nominal stress at mid-height of bulkhead of secondary box σ av c : design crushing strength (39.1 MPa for CCS#2, 45.3 MPa for CCS#3, and 46.5 MPa for CCS#4)
Design acceptance criteria for the crushing-failure evaluation for three CCSs were proposed. Future studies are needed to define the detailed CCS design. - Investigation of the crushing failure at the LNG operating temperature - A dynamic assessment to consider the design sloshing impact loads is necessary. References Arswendy, A., Moan, Torgeir, 2015. “Strength and stiffness assessment of an LNG containment system – crushing and buckling failure analysis of plywood components”. Eng. Fail. Anal. 48, 247–258. Chun, M.S., Kim, M.H., Kim, W.S., Kim, S.H., Lee, J.M., 2009. Experimental investigation on the impact behavior of membrane-type LNG carrier insulation system. J. Loss Prev. Process. Ind. 22 (Issue 6), 901–907. Chun, S.E., Hwang, J.O., Chun, M.S., Lee, J.M., Suh, Y.S., Hwangbo, S.M., White, Nigel, Wang, Z.H., 2011. Direct assessment of structural capacity against sloshing loads using nonlinear dynamic FE analysis including hull structural interactions. In: Proc. of the 2011 International Offshore and Polar Engineering Conference, pp. 191–199. Digital Times, 2018. LNG Cargo Ship Orders Rise. http://www.dt.co.kr/contents.html? article_no¼201801080210063200500.
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Ultimate crushing strength criteria for GTT NO96 LNG carrier cargo containment system under sloshing load Young IL. Park Department of Naval Architecture and Offshore Engineering, Dong-A University, Republic of Korea
A R T I C L E I N F O
A B S T R A C T
Keywords: LNG carrier cargo containment system (CCS) Finite-element method Failure mode Crushing failure Sloshing
The ultimate crushing strength of a Gaztransport and Technigaz NO96-type liquefied natural gas cargo containment system (NO96 CCS) was investigated through experiments and numerical simulations as one of the possible failure modes. As crushing failure is directly related to material failure along the through-thickness direction of the horizontal members, i.e., the bottom plate of the primary box and the top plate of the second ary box, small-scale experimental crushing test results were applied. Instead of using the test results directly, static failure capacity estimates were obtained via nonlinear finite-element analyses, which accounted for the crushing of the horizontal members and hence also the bending that this introduced in the horizontal members. The crushing-failure capacities were determined by the calculated bending moment in the two horizontal members. The calculated capacity will be used to select a design criterion using the vertical stress at the midheight of the bulkhead panel of the secondary box as an early design objective.
1. Introduction Since the Paris Agreement on climate change, the global liquefied natural gas (LNG) market has grown by 4%–6% annually. According to a report from the International Gas Union (IGU), the global LNG market is growing steadily, with an annual increase of 6.2%. The Asia–Pacific region accounts for 72.4% of the world’s total imports, and the largest importers of LNG have shifted from Japan and Korea to China, India, and the Southeast Asian countries. In particular, the Chinese government aims to increase its share of the global LNG from 6% to 15% by 2030. Moreover, the development of the Floating Storage Regasification Unit (FSRU) is expected to increase the import volumes of small LNGimporting countries and create new importers. The FSRU is considered to be easier and less expensive than LNG land terminals (IGU, 2017). It has lower initial investment cost than land terminals. Additionally, land terminals involve complicated licensing procedures from site selection to construction. With global market trends, technological development, and recovering international oil prices, demand for LNG is increasing, leading to increased orders in shipbuilding, such as LNG carriers and FSRUs (Digital Times, 2018). As LNG is transported in the liquid state at a cryogenic temperature of 163 � C, a cargo hold is installed inside the LNG carrier. The LNG cargo containment system (CCS) is generally classified as an independent-type or membrane-type cargo hold. As shown in Fig. 1, the
membrane-type cargo hold is strong (LNG World Shipping, 2017). The Gaztransport and Technigaz (GTT) NO96 and Mark III are representative membrane-type LNG CCSs (see Fig. 2). The Mark III is a PUF-type cargo hold with horizontal plywood on top of the primary barrier and at the bottom of the secondary barrier, and the R-PUF has it in its interior. There is a triplex between the barriers, and the primary and secondary barriers are fixed with an adhesive. The NO96 is a boxtype cargo hold with a box structure through the plywood, and it uses perlite or glass wool (GW) for the interior. There is invar steel between the barriers, and each barrier is fixed to the hull with a binding device. The NO96 LO3, NO96 LO3þ, Mark III Flex, and Mark V have been modified to reduce the boil-off gas (GTT, 2017). A major advantage of a membrane-type LNG CCS, such as the GTT MARK III and NO96, compared with an independent-type CCS is the flexibility for increasing the size of the cargo tank. However, the risk of sloshing damage to the membranes is an important factor to consider when selecting a membrane-type LNG CCS, particularly for offshore applications, e.g., LNG FPSO and FSRU. To assess the risk of sloshing damage and determine the feasibility of a membrane-type LNG CCS, several factors must be considered. For instance, adequate design capacities and parameters must be considered for each CCS and each design choice applied. Daewoo Shipbuilding & Marine Engineering Co., Ltd. developed a two-row tank arrangement for the NO96 CCS for LNG offshore
E-mail address: [email protected]. https://doi.org/10.1016/j.oceaneng.2019.106224 Received 15 November 2018; Received in revised form 15 July 2019; Accepted 16 July 2019 Available online 8 August 2019 0029-8018/© 2019 Elsevier Ltd. All rights reserved.
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applications to reduce the sloshing design load (Ryu et al., 2016). Various experimental and numerical studies have been performed on the GTT MARK III-type CCS (Kim et al., 2010; Chun et al., 2009). Arswendy and Moan (2015) conducted experimental and numerical studies to determine the buckling and crushing strengths for a T-shaped plywood specimen as a part of the NO96 CCS. Chun et al. (2011), Hwang et al. (2014), Lee and Shin (2014), Lee et al. (2012), Nho et al. (2012), Pillon et al. (2009), Yoo et al. (2011), and Wang et al. (2012) performed structural evaluations under sloshing impact loads to define design sloshing pressure and/or ultimate strength of a MARK III-type LNG CCS depending on several failure modes. Dobashi and Usami (2012) proposed a design dynamic amplification factor for a NO96-type CCS. Lee et al. (2011) and Wang et al. (2009) investigated the general structural behavior of a NO96-type CCS under sloshing impact via simulations. However, few studies have been performed on the NO96 CCS, particularly for early-design purposes. Park and Lee (2018) proposed design dynamic buckling strength capacities for the critical failure mode of a box-type LNG CCS, such as the NO96 CCS. DNVGL (2016) identifies four design failure modes of the NO96 CCS: shear, bending failures at the cover panel of the primary box, buckling failure of the bulkhead panel of the primary and secondary boxes, and crushing failure at the intersection of the primary and secondary box bulkheads. However, it applies a comparative approach to a 138000-m3 an LNG carrier, which is a proven design and target vessel. In this study, the failure modes of the NO96 CCS are classified, and criteria are developed for ultimate crushing failure via numerical sim ulations tuned through experiments based on risk. The results will aid the initial design of LNG cargo holds and improve currently used holds. First, the failure modes of the box-type cargo hold as proposed by classification societies are compared, and several failure modes are added. To confirm the additions and the new list, a load due to sloshing is selected for a static-load test of an independent box-type cargo hold. A nonlinear finite-element (FE) analysis is performed to validate the experimental results. The experiments and FE analysis are also performed for additional failure modes, and design criteria are presented.
In this study, the design capacity of the crushing of plates at the bulkhead intersections was investigated numerically. 3. Experimental tests for crushing-failure analysis of plywood components Arswendy and Moan (2015) performed experimental tests of the crushing and buckling failure modes of plywood, which is the major strengthening part of a box-type LNG CCS. Fig. 5 shows the test spec imen configurations that they used for the crushing test, where (a), (b), and (c) correspond to crushing tests of a plywood plate and plywood bulkhead, of a plywood plate and steel bulkhead, and of a steel plate and plywood bulkhead, respectively. As the major failure mode of crushing for a box-type NO96 CCS should be related to horizontal member failure and vertical bulkhead member failure, the test configuration (a) of Fig. 5 was selected in this study. Fig. 6 shows the relationship between the vertical load and displacement for the (a) configuration. 4. FE simulations Proper interpretation of the test results requires knowledge of the correlation between the loads applied to the NO96 CCS and those applied to the test specimen under which the local stress at the inter section of the plates is similar. Thus, it is necessary to determine whether there is any difference in the stress concentration between the test setup and the NO96 CCS during crushing failure. To investigate the nonlinear behavior related to crushing failure at the intersection of the vertical bulkheads of the primary and the sec ondary boxes, a nonlinear static FE analysis was performed using pre vious experimental results (Arswendy and Moan, 2015) for cases (a) and (b) of Fig. 5, as the main focus of the present study was the crushing failure of the horizontal member. For the FE simulations, the commercial FE code ABAQUS was employed, with the specifications presented in Table 2. The material stiffness depends on the design temperature, plywood grain direction (orthotropic material properties), and bending/mem brane characteristics, as shown in Table 3 (DNVGL, 2016). An elastic modulus of 205800 MPa and a Poisson’s ratio 0.3 were considered for steel part in case (b) of Fig. 5. The indentation of the vertical bulkhead plates into the horizontal plate was localized to small regions directly adjacent to the edges of the bulkhead plate. An accurate FE representation of this localized damage required an extremely fine element mesh. Instead, an elastoplastic ma terial model for the thickness behavior of the horizontal plates was calibrated to provide an equivalent overall inelastic stiffness charac teristic of the bulkhead intersections for the mesh density used in the model. The crushable-foam model available in ABAQUS was applied to allow for volumetric plastic compressibility of the material. This was necessary to obtain reasonable inelastic deformation modes along the thickness direction of the horizontal plates. The incompressible plastic behavior of the standard metal plasticity model resulted in the unphysical locking of the elements of the horizontal plates located above and beneath the vertical bulkhead plates; consequently, the inelastic deformation was forced to occur in the adjacent element. Because the crushable-foam model is only applicable to isotropic materials in ABAQUS applications, four-node shell elements with orthotropic material properties were embedded into the solid element region, as shown in Fig. 7. The parameters of this stiffness-equivalent elastoplastic crushablefoam model were determined via FE simulations of the experimental test (Arswendy and Moan, 2015). The through-thickness elastic modulus, material yield stress, and plastic tangent modulus were tuned such that the simulated load–displacement response fitted the test re sults, as shown in Table 4. A comparison of the test and simulation
2. NO96 CCS The GTT NO96 is a box-type membrane LNG CCS containing plywood panels for strength, GW or perlite for insulation, and other components, e.g., a securing device and a membrane. Various NO96 CCSs have been developed to increase the thickness of the internal bulkheads or increase the number of internal members to bear greater load, as shown in Fig. 3 and Table 1. In this study, the crushing strength was evaluated for CCS#2, #3, and #4. The DNVGL (2016) and Lloyd’s Register (2009) identify the following failure modes for the NO96 CCS as shown in Fig. 4. 1) Bending/shear (including combined shear and bending) failure of the cover plate of the primary box 2) Buckling failure of the vertical bulkheads (internal and external bulkheads) 3) Crushing failure A. Bulkhead intersections B. Resin rope 4) Failure related to the box-securing system A. Bulkheads of the corner border box—cleats at the 90� transverse corner zone B. Bulkheads of the corner border box—cleats at the 135� longitu dinal corner zone 5) Failure connected to the membrane tongue attachment 2
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somewhat Fig. 11 shows the load vs. deflection curves under out-plane pressure for the two aforementioned FE models. The outputs are identical. As the FE model with the 3D solid elements and two-dimensional (2D) shell embedded was implemented for the horizontal member of the CCS—1) bottom plate of the primary box and 2) top plate of the secondary box—the out-plane behavior was more critical than the inplane behavior. Therefore, it was found to be acceptable to apply these concepts to the FE model. Elastoplastic FE analyses of a flat NO96 CCS were performed. The Table 1 General specifications of the NO96 CCS for each case study. Fig. 1. Global trends of LNG carrier CCSs.
results is shown in Fig. 6. Fig. 8 shows a comparison of the test results of Arswendy and Moan (2015) with the numerical simulations. In this study, the nonlinear material properties for the through-thickness direction, which is for crushing failure on the horizontal member, were calibrated according to case (a) (plywood–plywood), and the same material properties were considered for case (b). The comparison between the experimental test of case (b) and the numerical simulation reveals somewhat conservative behavior in the simulation; however, this can be acceptable for design purposes. Because the crushable-foam model is only applicable to isotropic materials in ABAQUS applications, four-node shell elements with orthotropic material properties were embedded into the solid element region, as shown in Fig. 7. Several FE analyses were performed to confirm whether the struc tural behavior of the embedded shell in the solid model was practically reasonable. Two FE models were employed: orthotropic shell elements (left figure of Fig. 9) and orthotropic shell elements with threedimensional (3D) solid elements (right figure of Fig. 9). The elastic buckling under in-plane compressive loads was evalu ated, as shown in Fig. 10. The buckling strength in the right frame of Fig. 10 is 12% higher than that in the left frame, which could be
Case No.
Primary box Thickness of cover plate
Thickness of BHD
Secondary box Thickness of BHD
No. of combs
CCS#1 CCS#2 CCS#3 CCS#4
12 mm 12 mm � 2 12 mm � 2 12 mm � 2
9 12 12 15
9 12 12 15
0 0 1 2
Fig. 4. Failure modes considered for the NO96 CCS (DNVGL, 2016).
Fig. 2. Overview of GTT Mark III and NO96 LNG CCSs (source: http://www.gtt.fr/).
Fig. 3. Insulation box of the NO96 CCS for various reinforcement designs (source: www.gtt.fr). 3
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Fig. 5. Crushing-test configurations (Arswendy and Moan, 2015).
Fig. 6. Relationship between the vertical load and displacement for the test configuration shown in Fig. 5 (Arswendy and Moan, 2015).
Table 3 Stiffness properties of the integrated orthotropic material for plywood panels (DNVGL, 2016).
Table 2 Specifications of the FE analysis. Item
Notes
Software (solver)
ABAQUS/Standard (V.6.14)
For static-buckling calculation
Element type
4-node shell
Reduced integration for strength member, i.e., plywood panel Spring connection, i.e., stapling Dependent on operational temperature, i.e., 163 to 20 � C
Material type Boundary conditions Surface contact
2-node beam elements Homogeneous orthotropic type 4 edges—simply supported Kinematic hard contact option
Parameter
Temperature 20 � C
Em,1 [MPa] Em,2 [MPa] Em,3 [MPa] Gm,12 [MPa] Gm,13 [MPa] Gm,23 [MPa]
ν12 ν 13 ν 23
Eb,1 9 mm [MPa] Eb,2 9 mm [MPa] Eb,1 12 mm [MPa] Eb,2 12 mm [MPa] Density [ton/mm3] σ F [MPa]
4
9450 8000 820 790 325 260 0.1 0.1 0.1 10950 6550 10450 7000 6.8e-10 43.0
163 � C 13200 11200 1800 2900 700 550 0.1 0.1 0.1 15350 9150 14650 9800 6.8e-10 65.0
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Fig. 7. Schematic of the crushable foam for the FE model.
the horizontal members, i.e., the bottom plate of the primary box and the top plate of the secondary box. The shell elements were constrained to the solid element mesh such that the composite mesh represented the total stiffness of the plates. The boundary conditions of the FE model were imposed at the resin support locations as simple support. FE analyses for five cases with different loads were conducted (refer to Fig. 13). The main objective of this study was to develop strength criteria for the crushing behavior of the NO96 CCS. Therefore, the structural response was investigated for various loading conditions. The vertical stresses in the bulkheads of the boxes were considered to be representative of the forces transferred from the primary box to the secondary box through the bulkhead intersections. Therefore, the stress was chosen as a candidate for the acceptance criterion of the crushing failure. Fig. 14 and Table 5 show the stress check points, which could have been representative strength criteria for the five cases of the load. Figs. 15–29 show the stress behavior under the applied pressure at the five check points for CCS#2, CCS#3, and CCS#4. The stress
Table 4 Material properties for the calibrated crushable-foam model. Material properties for crushable foams Elastic Crushable foam Crushable-foam hardening
Elastic modulus [MPa] Poisson’s ratio Compressive yield stress ratio Hydrostatic yield stress ratio Yield stress [MPa] Uniaxial plastic strain
80 0.1 0.9 0.9 10–28 0.0–0.93
analyses considered the thickness inelastic behavior of the primary-box bottom plate and the secondary-box top plate using the properties of the calibrated material (refer to Table 4). The quarter-size FE models of the insulation boxes used in the ana lyses are shown in Fig. 12, for CCS#2, CCS#3, and CCS#4 with different numbers of combs at the secondary box. The model was built using eight 3D node brick elements combined with four 2D node shell elements for
Fig. 8. Comparison of the test results with the numerical simulation results for the simplified model (left: plywood–plywood, case (a); right: steel–plywood, case (b) (Arswendy and Moan, 2015)).
5
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Fig. 9. 2 FE models: left, orthotropic shell model; right, 3D solid element þ 2D shell model.
Fig. 10. Elastic buckling strength for the in-plane compressive load comparison (orthotropic shell model – left: 2.65 MPa, shell–solid combination).
Fig. 11. Out-plane pressure behavior for an out-plane uniform pressure.
6
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Fig. 12. 1/4-sized FE model of the NO96 CCS (left: CCS#2; middle: CCS#3; right: CCS#4); arrow: grain direction.
responses show that the redistribution of forces away from the bulkhead intersections and toward the adjacent parts of the bulkhead edges occurred in the structure owing to the bending of the horizontal bottom plate of the primary box and the cover plate of the secondary box. Owing to the crushing deformation at the intersections of the pri mary and secondary bulkheads, loads were redistributed from the bulkhead intersection to the remainder of their edges from the bending capacity of the bottom plate of the primary box and the cover plate of the secondary box. Both plates (the bottom plate of the primary box and the cover plate of the secondary box) were bent along their two in-plane directions—the strong and weak directions—and the capacity and effi ciency of the plates to redistribute loads was dependent on them maintaining their bending strength. To define the ultimate bending
capacities of the horizontal members (the bottom plate of the primary box and the cover plate of the secondary box), four additional stress responses were observed, as shown in Table 6 and Fig. 30. The strong/weak directions of the bending-stress contours of the bottom plate of the primary box and the cover plate of the secondary box for a full patch load are shown in Figs. 31–34. The criteria for the ultimate bending strength in Table 7, depending on the plywood grain direction, i.e., the strong or weak direction, were applied. The critical design pressures for each loading case on CCS#2, CCS#3, and CCS#4 were determined by Table 7: 1100 and 760 Nmm/mm for the strong and weak directions of the bending capacities for the bottom plate of the primary box, respectively, and 1850 and 1380 Nmm/mm for
Fig. 13. Cases of the FE analysis, with different patch loads. 7
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the strong and weak directions of the bending capacities for the top plate of the secondary box, respectively, as shown in Figs. 35–46. From Figs. 35–46, the following conclusions are drawn. - The structure had a limited capacity to transfer the applied force. The most critical patch-load configurations occurred in the case of a uniform load. - The stress responses of the bulkheads of the primary box, the bulk heads of the secondary box, and the hotspots near the horizontal plates exhibited similar behavior. 5. Acceptance criteria for crushing failure Crushing strength should be measured in terms of the nominal stress in the vertical bulkhead of the NO96 CCS, as the basic designer conducts FE calculations using 2D shell-element modeling. Crushing deformation between the primary and secondary boxes leads to the failure of the horizontal plates, i.e., the bottom plate of the primary box and the cover plate of the secondary box. Loads applied on the cover plate of the primary box are redistributed from the intersection of the bulkheads to the remainder of their edges by the horizontal plates, i.e., the bottom plate of the primary box and the cover plate of the secondary box. Both the horizontal plates were bent along their two inplane directions, and the ability and efficiency of the plates to redis tribute the loads were dependent on the bending and shear integrity of the horizontal plates. Plate failure effect analysis was performed, as summarized in Table 8 and Table 9, to assess the conditions of the critical failure of the inter section of the bulkheads. Bending was a more critical failure mechanism than shear for the plates. Therefore, the total capacity of the bulkhead intersections was evaluated via a bending-capacity assessment. The failure effect analysis in Tables 8 and 9 indicates that #7 and #8 were critical effects. They represent the worst case of the acceptable consequence. The system was induced to significantly limit the redis tribution of forces to the edge of the bulkhead. Therefore, the acceptance criterion for the bulkhead intersections was determined by #7 and #8. Each capacity of the bending mode for each of the two plates was estimated according to the pressure–bending stress curves obtained via
Fig. 14. Stress-output check points in the primary and secondary boxes.
Table 5 Stress-response check points on the vertical bulkhead member. ID
Description
Location
A B C D E
Nominal stress at mid-height location of vertical bulkhead Hotspot stress at bulkhead near bottom plate Hotspot stress at bulkhead near cover plate Nominal stress at mid-height location of vertical bulkhead Hotspot stress at bulkhead near bottom plate
Primary box Primary box Secondary box Secondary box Secondary box
Fig. 15. Vertical stress at the mid-height of the primary-box bulkhead (Point A) vs. the applied pressure for CCS#2. 8
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Fig. 16. Vertical stress at the mid-height of the primary-box bulkhead (Point A) vs. the applied pressure for CCS#3.
FE analysis. The estimated capacities are presented in Tables 10–12 for CCS#2, CCS#3, and CCS#4. The underlined data are the capacities of #7 and #8 obtained via the failure effect analysis of the horizontal plates, i.e., the bottom plate of the primary box and the cover plate of the secondary box. In the previous section, the FE model using four 2D node shell ele ments could yield reasonable stress values in hotspots such as check points B, C, and E. Therefore, A and D were the likely strength criteria for
the crushing-failure mode. In this study, point D was used to define the failure criteria for crushing. Figs. 47–51 show reference stresses at each check point for CCS#2 under the critical failure crushing pressure determined by Table 10 Maximum crushing capacity by bending strength on the horizontal plates for CCS#2. Figs. 52–54 correspond to CCS#3, and Figs. 55–58 correspond to CCS#4. Table 13 presents the design reference stresses at check point D for
Fig. 17. Vertical stress at the mid-height of the primary-box bulkhead (Point A) vs. the applied pressure for CCS#4. 9
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Fig. 18. Vertical stress at the hotspot of the primary-box bulkhead (Point B) vs. the applied pressure for CCS#2.
Fig. 19. Vertical stress at the hotspot of the primary-box bulkhead (Point B) vs. the applied pressure for CCS#3.
10
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Fig. 20. Vertical stress at the hotspot of the primary-box bulkhead (Point B) vs. the applied pressure for CCS#4.
Fig. 21. Vertical stress at the hotspot of the secondary-box bulkhead (Point C) vs. the applied pressure for CCS#2.
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Fig. 22. Vertical stress at the hotspot of the secondary-box bulkhead (Point C) vs. the applied pressure for CCS#3.
Fig. 23. Vertical stress at the hotspot of the secondary-box bulkhead (Point C) vs. the applied pressure for CCS#4.
12
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Fig. 24. Vertical stress at the mid-height of the secondary-box bulkhead (Point D) vs. the applied pressure for CCS#2.
Fig. 25. Vertical stress at the mid-height of the secondary-box bulkhead (Point D) vs. the applied pressure for CCS#3.
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Fig. 26. Vertical stress at the mid-height of the secondary-box bulkhead (Point D) vs. the applied pressure for CCS#4.
Fig. 27. Vertical stress output at the hotspot of the secondary-box bulkhead (Point D) vs. the applied pressure for CCS#2.
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Fig. 28. Vertical stress output at the hotspot of the secondary-box bulkhead (Point E) vs. the applied pressure for CCS#3.
Fig. 29. Vertical stress output at the hotspot of the secondary-box bulkhead (Point E) vs. the applied pressure for CCS#4.
Table 6 Stress-response check points on the horizontal member, i.e., the bottom plate of the primary box and the cover plate of the secondary box. ID
Description
Location
F G H I
Strong direction of bending stress at bottom plate of primary box Weak direction of bending stress at bottom plate of primary box Strong direction of bending stress at cover plate of secondary box Weak direction of bending stress at cover plate of secondary box
Primary box Primary box Secondary box Secondary box
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Fig. 30. Locations of the stress-output checks at the bottom and cover plates of the primary and secondary boxes, respectively.
Fig. 31. Bending-stress contours at the bottom of the primary box along the strong direction.
Fig. 32. Bending-stress contours at the bottom of the primary box along the weak direction.
Fig. 33. Bending-stress contours on the cover plate of the secondary box along the strong direction.
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Fig. 34. Bending-stress contours on the cover plate of the secondary box along the weak direction.
each CCS under the critical design pressure for crushing failure. In this study, 39.1, 45.3, and 46.5 MPa were obtained as the design reference stresses at check point D for CCS#2, CCS#3, and CCS#4, respectively.
Table 7 Mean strength properties for 9-mm and 12-mm plywood laminates (DNVGL, 2016). Plywood thickness
9 mm
Temperature Mc,1 (Nmm/mm) Mc,2 (Nmm/mm)
20 � C 1100 760
6. Conclusion
12 mm 163 � C 935 650
20 � C 1850 1380
The ultimate crushing-failure strength of the NO96 CCS was evalu ated via nonlinear FE simulation. To define the nonlinear material properties related to the thickness behavior, the experimental results of a plywood-crushing test were used. They were calibrated to obtain the equivalent overall inelastic stiffness characteristics of the intersections of the bulkhead for the mesh density used in the FE model. Crushing deformation limited the redistribution of forces by the bending or shearing of the horizontal members, i.e., the bottom plate of
163 � C 1580 1180
Note: M1 denotes bending about the 2-axis, and M2 denotes bending about the 1axis.
Fig. 35. Output of the sectional moment along the strong direction at the bottom plate of the primary box (point F) for CCS#2.
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Fig. 36. Output of the sectional moment along the strong direction at the bottom plate of the primary box (point F) for CCS#3.
Fig. 37. Output of the sectional moment along the strong direction at the bottom plate of the primary box (point F) for CCS#4.
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Fig. 38. Output of the sectional moment along the weak direction at the bottom plate of the primary box (point G) for CCS#2.
Fig. 39. Output of the sectional moment along the weak direction at the bottom plate of the primary box (point G) for CCS#3.
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Fig. 40. Output of the sectional moment along the weak direction at the bottom plate of the primary box (point G) for CCS#4.
Fig. 41. Output of the sectional moment along the strong direction on the cover plate of the secondary box (point H) for CCS#2.
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Fig. 42. Output of the sectional moment along the strong direction on the cover plate of the secondary box (point H) for CCS#3.
Fig. 43. Output of the sectional moment along the strong direction on the cover plate of the secondary box (point H) for CCS#4.
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Fig. 44. Output of the sectional moment along the weak direction on the cover plate of the secondary box (point I) for CCS#2.
Fig. 45. Output of the sectional moment along the weak direction on the cover plate of the secondary box (point I) for CCS#3.
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Fig. 46. Output of the sectional moment along the weak direction on the cover plate of the secondary box (point I) for CCS#4.
Table 8 Dependence of the level of consequence on the failure mode for determining the crushing capacity.
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Table 9 Failure effect analysis of the horizontal plates for the bending mode.
Table 10 Maximum crushing capacity by bending strength on the horizontal plates for CCS#2. Load case Full patch Patch 1 Patch 2 Patch 3 Patch 4
Bottom plate of the primary box
Cover plate of the secondary box
Strong direction
Weak direction
Strong direction
Weak direction
19.36 bar 25.35 bar 22.66 bar 25.74 bar 19.05 bar
27.59 bar 40.71 bar 48.88 bar 41.35 bar 28.00 bar
20.77 bar 30.06 bar 30.65 bar 30.46 bar 21.09 bar
37.26 bar 54.62 bar 41.32 bar 57.48 bar 41.73 bar
Table 11 Maximum crushing capacity by bending strength on the horizontal plates for CCS#3. Load case Full patch Patch 1 Patch 2 Patch 3 Patch 4
Bottom plate of the primary box
Cover plate of the secondary box
Strong direction
Weak direction
Strong direction
Weak direction
14.74 bar 50.77 bar 26.04 bar 28.78 bar 35.57 bar
31.60 bar N.A. N.A. 58.69 bar 48.02 bar
16.95 bar N.A. N.A. 68.22 bar 52.31 bar
33.09 bar N.A. N.A. 64.75 bar N.A.
Table 12 Maximum crushing capacity by bending strength on the horizontal plates for CCS#4. Load case Full patch Patch 1 Patch 2 Patch 3 Patch 4
Bottom plate of the primary box
Cover plate of the secondary box
Strong direction
Weak direction
Strong direction
Weak direction
16.01 bar 20.79 bar 18.99 bar 21.00 bar 15.71 bar
33.91 bar 51.17 bar N.A. N.A. 33.98 bar
17.81 bar 23.98 bar 22.88 bar 23.18 bar 16.88 bar
45.14 bar 56.51 bar 41.38 bar N.A. 44.13 bar
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Fig. 47. Reference stresses at each check point under 27.59 bar for the critical bending failure in the uniform-pressure case of CCS#2 determined by Table 10.
Fig. 48. Reference stresses at each check point under 40.71 bar for the critical bending failure in the Patch 1 case of CCS#2 determined by Table 10.
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Fig. 49. Reference stresses at each check point under 41.32 bar for the critical bending failure in the Patch 2 case of CCS#2 determined by Table 10.
Fig. 50. Reference stresses at each check point under 41.35 bar for the critical bending failure in the Patch 3 case of CCS#2 determined by Table 10.
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Fig. 51. Reference stresses at each check point under 28.00 bar for the critical bending failure in the Patch 4 case of CCS#2 determined by Table 10.
Fig. 52. Reference stresses at each check point under 31.60 bar for the critical bending failure in the uniform-pressure case of CCS#3 determined by Table 11.
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Fig. 53. Reference stresses at each check point under 64.75 bar for the critical bending failure in the Patch 3 case of CCS#3 determined by Table 11.
Fig. 54. Reference stresses at each check point under 52.31 bar for the critical bending failure in the Patch 4 case of CCS#3 determined by Table 11.
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Fig. 55. Reference stresses at each check point under 33.91 bar for the critical bending failure in the uniform-pressure case of CCS#4 determined by Table 12.
Fig. 56. Reference stresses at each check point under 51.17 bar for the critical bending failure in the Patch 1 case of CCS#4 determined by Table 12.
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Fig. 57. Reference stresses at each check point under 41.38 bar for the critical bending failure in the Patch 2 case of CCS#4 determined by Table 12.
Fig. 58. Reference stresses at each check point under 33.98 bar for the critical bending failure in the Patch 4 case of CCS#4 determined by Table 12.
Table 13 Reference stresses at check point D of each CCS under the critical design pressure for crushing failure. Critical pressure of bending failure on the horizontal members [bar] Uniform pressure Patch 1 Patch 2 Patch 3 Patch 4
Reference stress at check point D for the critical design pressure of crushing failure [MPa]
CCS#2
CCS#3
CCS#4
CCS#2
CCS#3
CCS#4
27.6 40.7 41.3 41.4 28.0
31.6
33.9 51.2 41.4
39.1 42.3 37.7 42.8 39.5
45.3
46.5 51.5 37.1
64.8 52.3
34.0
30
53.3 38.8
46.7
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the primary box and the cover plate of the secondary box. Therefore, the load capacity of the bulkhead intersections was evaluated via a bendingcapacity assessment. The failure criteria were defined through failure effect analyses. The design crushing-failure reference stresses at the middle of the bulkhead of the secondary box were proposed as 39.1, 45.3, and 46.5 MPa for CCS#2, CCS#3, and CCS#4, and the design values are applicable to the NO96 CCS FE model using 2D shell elements to define the ultimate crushing strength with the following inequality.
DNVGL-CG-0158, 2016. "Sloshing analysis of LNG membrane tanks". Class Guideline. Dobashi, H., Usami, A., 2012. Dynamic amplification factor of NO96 insulation structures of membrane system. In: Proc. of the 2012 International Offshore and Polar Engineering Conference, pp. 525–529. Gaztransport & Technigaz, Production, 2017. http://www.gtt.fr/en/technologies-servic es/our-technologies. Hwang, J.O., Chun, S.E., Joh, K.H., Combos, P., Lauzon, J.D., White, N., Kim, M.S., Park, J.B., Lee, J.M., 2014. Direct assessment of structural capacity against sloshing using dynamic nonlinear FE analysis. In: Proc. of the 2014 International Ocean and Polar Engineering Conference, pp. 169–179. International Gas Union, 2017. IGU world LNG report. IGU 78–94, 1-53. Kim, M.H., Lee, S.M., Lee, J.M., Noh, B.J., Kim, W.S., 2010. Fatigue strength assessment of MARK-III type LNG cargo containment system. Ocean Eng. 37 (Issues 14–15), 1243–1252. Lee, D.J., Shin, S.B., 2014. Numerical model for dynamic behavior of PUF in membrane type LNG cargo containment system. In: Proc. of the 2014 International Ocean and Polar Engineering Conference, pp. 157–162. Lee, S.J., Yang, Y.S., Kim, S.C., Lee, J.H., 2011. Strength assessment procedure of LNG CCS under sloshing load based on the direct approach. In: Proc. of the 2011 International Offshore and Polar Engineering Conference, pp. 183–190. Lee, S.G., Kim, J.K., Nguyen, H.A., Nam, J.H., 2012. Structural safety assessment of LNGC MARK III membrane type CCS under sloshing impact loading. In: Proc. of the 2012 International Offshore and Polar Engineering Conference, pp. 487–494. Lloyd’s Register, May 2009. ShipRight, “Sloshing Assessment Guidance Document for Membrane Tank LNG Operations”, Version 2.0. LNG World Shipping, 2017. LNGC Fleet in Service and on Order. http://www.lngworldsh ipping.com/news/view,gtt -plans-mark-v-rethink_49315. Nho, I.S., Kim, S.C., Jang, B.S., Lee, J.H., 2012. Parametric investigation on the simplified triangular impulse of sloshing pressure and categorization of the structural response on the Mark III LNG CCS. In: Prof. of the 2012 International Offshore and Polar Engineering Conference, pp. 495–501. Park, Y.I., Lee, J.H., 2018. Buckling strength of GTT NO96 LNG Carrier cargo containment system. Ocean Eng. 154, 43–58. Pillon, B., Marhem, M., Leclere, G., Canler, G., 2009. Numerical approach for structural assessment of LNG containment systems. In: Proc. of the 2009 International Offshore and Polar Engineering Conference, pp. 175–182. Ryu, M.C., Jung, J.H., Kim, Y.S., Kim, Y.I., 2016. Sloshing design load prediction of a membrane type LNG cargo containment system with two-row tank arrangement in offshore application. Int. J. Naval Architect. Ocean Eng. 8 (6), 537–553. Wang, B., Han, S.K., Kim, Y.S., Park, Y.I., Shin, Y., 2009. Sloshing model tests and strength assessment of the NO 96 containment system. In: Proc. of the 2009 International Offshore and Polar Engineering Conference, pp. 261–268. Wang, B., Shin, Y., Wang, X., 2012. Reliability-based sloshing assessment of containment systems in LNGCs and FLNGs. In: Prof. of the 2012 International Offshore and Polar Engineering Conference, pp. 502–509. Yoo, M.J., Lee, S.H., Kim, S.C., Lee, J.H., Nho, I.S., 2011. Characteristics of dynamic response of MARK III LNG containment subjected to idealized triangular sloshing impact. In: Proc. of the 2011 International Offshore and Polar Engineering Conference, pp. 177–182.
σ av � σ av c where
σ av : nominal stress at mid-height of bulkhead of secondary box σ av c : design crushing strength (39.1 MPa for CCS#2, 45.3 MPa for CCS#3, and 46.5 MPa for CCS#4)
Design acceptance criteria for the crushing-failure evaluation for three CCSs were proposed. Future studies are needed to define the detailed CCS design. - Investigation of the crushing failure at the LNG operating temperature - A dynamic assessment to consider the design sloshing impact loads is necessary. References Arswendy, A., Moan, Torgeir, 2015. “Strength and stiffness assessment of an LNG containment system – crushing and buckling failure analysis of plywood components”. Eng. Fail. Anal. 48, 247–258. Chun, M.S., Kim, M.H., Kim, W.S., Kim, S.H., Lee, J.M., 2009. Experimental investigation on the impact behavior of membrane-type LNG carrier insulation system. J. Loss Prev. Process. Ind. 22 (Issue 6), 901–907. Chun, S.E., Hwang, J.O., Chun, M.S., Lee, J.M., Suh, Y.S., Hwangbo, S.M., White, Nigel, Wang, Z.H., 2011. Direct assessment of structural capacity against sloshing loads using nonlinear dynamic FE analysis including hull structural interactions. In: Proc. of the 2011 International Offshore and Polar Engineering Conference, pp. 191–199. Digital Times, 2018. LNG Cargo Ship Orders Rise. http://www.dt.co.kr/contents.html? article_no¼201801080210063200500.
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