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Cryogenic reliability of composite insulation panels for liquefied natural gas (LNG) ships
Composite Structures 94 (2012) 462–468
Contents lists available at SciVerse ScienceDirect
Composite Structures journal homepage: www.elsevier.com/locate/compstruct
Cryogenic reliability of composite insulation panels for liquefied natural gas (LNG) ships Young Ho Yu a, Bu Gi Kim a, Dai Gil Lee b,⇑ a School of Mechanical Aerospace & Systems Engineering, Korea Advanced Institute of Science and Technology, ME3221, Guseong-dong, Yuseong-gu, Daejeon 305-701, Republic of Korea b KAIST Institute for Design of Complex Systems (KIDCS), Guseong-dong, Yuseong-gu, Daejeon 305-701, Republic of Korea
a r t i c l e
i n f o
Article history: Available online 27 August 2011 Keywords: LNG ship Composite insulation panel Secondary barrier Crack retardation Glass fabric reinforcement
a b s t r a c t The major carrier of liquefied natural gas (LNG) is LNG ships, whose containment system is composed of dual barriers and composite insulation panels. The LNG containment system should have cryogenic reliability and high thermal insulation performance for safe and efficient transportation of LNG. The secondary barrier composed of adhesive bonded aluminum strips should keep tightness for 15 days, when the welded stainless primary barrier fails. However, cracks are generated in the composite insulation panels due to the local stress concentration and the brittleness of insulation materials at the cryogenic temperature of 163 °C. If cracks generated in the insulation panel propagate into the secondary barrier, LNG leakage problem might occur, which is a remaining concern in ship building industries. In this study, crack retardation capability in the composite insulation panel was investigated with glass fabric reinforcement. Finite element analysis was conducted to estimate the thermal stress at the cryogenic temperature and a new experimental method was developed to investigate the failure of secondary barrier of composite insulation panel. From the experimental results, it was found that the glass fabric reinforcement was effective to retard the crack propagation into the aluminum secondary barrier from the polyurethane insulation foam at the cryogenic temperature. Ó 2011 Elsevier Ltd. All rights reserved.
1. Introduction As the consumption of liquefied natural gas (LNG) has been increased much specially due to nuclear power plant disaster in Japan, the LNG ships have been spotlighted as major carriers of LNG [1–4]. Since LNG is normally carried in ships at 163 °C, the functional requirements of LNG ship are to have cryogenic reliability due to thermal cyclic stresses as well as to have high thermal insulation performance for safe and efficient transportation of LNG [5,6]. For efficient transportation, recently, the membrane type LNG containment systems are usually employed [7]. The membrane type LNG containment system has octagonal pillar shape with high capacity (larger than 39,000 m3), which is composed of dual barriers and composite insulation panels as shown in Fig. 1 [8,9]. The secondary barrier composed of adhesively bonded metal strips should keep tightness for 15 days when the welded stainless primary barrier fails. The insulation foam should have very low thermal conductivity for high thermal insulation performance, which limits the density of foam less than 120 kg/ ⇑ Corresponding author. E-mail addresses: [email protected] (Y.H. Yu), [email protected] (B.G. Kim), [email protected] (D.G. Lee). URLs: http://scs.kaist.ac.kr (Y.H. Yu), http://kidcs.kaist.ac.kr (D.G. Lee). 0263-8223/$ - see front matter Ó 2011 Elsevier Ltd. All rights reserved. doi:10.1016/j.compstruct.2011.08.009
m3. Since the foam density is not high enough for structural purpose and brittle at cryogenic temperatures, the reliability of conventional containment systems are not fully guaranteed due to the crack propagation in insulation foams, which is a remaining concern in ship building industries [10]. The composite insulation panels are composed of aluminum strip of 0.3 mm thickness, plywood and fiber reinforced polyurethane foam as shown in Fig. 1. The size of the insulation panel is 1 m (Width) 3 m (Length) 0.27 m (Height). The aluminum strip secondary barrier and polyurethane foam are adhesively bonded with polyurethane adhesive as shown in Fig. 2. During the operation of the cryogenic containment system (CCS) of LNG ships at the cryogenic temperature, cracks are generated in the composite insulation panels by the local stress concentrations due to the cyclic thermal stresses and the low fracture toughness of polyurethane foam at the cryogenic temperature [11–13]. In order to retard the crack generation, the polyurethane foam (PUF) is reinforced with glass fibers as shown in Fig. 2. However, the uniform glass fiber reinforcement with high volume fraction in PUF is not only possible in the mass production, but also it increases the thermal conductivity of polyurethane foam, which decreases the efficiency of the LNG transportation. When the cracks are generated, they propagate into the aluminum secondary barrier as shown in Fig. 2, which will induce the leakage of LNG.
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LNG
Corner insulation panel Flat insulation panel
Containment system
Containment system
LNG
Primary barrier
Composite insulation panel
Stainless steel
Plywood
Polyurethane foam
Aluminum
Secondary barrier
Fig. 3. Photograph of the large scale mock-up to investigate the cryogenic reliability of LNG containment system.
Corner panel
Fig. 1. Schematic diagrams of LNG containment system composed of dual barriers and composite insulation panels.
Flat panel Polyurethane foam Plywood
Aluminum
Stainless steel
LNG
(a)
Primary barrier
Composite insulation panel
Crack
Plywood
Stainless steel
Aluminum Section A
Polyurethane foam
Aluminum
Secondary barrier Polyurethane adhesive
(b) Aluminum barrier Polyurethane adhesive
Local elongation due to the crack propagation
Tension due to thermal shrinkage
Fig. 4. Composite insulation panels used in mock-up test: (a) schematic diagram of flat and corner of composite insulation panels in mock-up test and (b) photograph and schematic diagram of crack on the aluminum barrier generated on section A during mock-up test.
insulation panels retarded the crack propagation into the aluminum secondary barrier, which improved the cryogenic reliability of composite insulation panel. 2. Experimental Glass fiber
Crack
Polyurethane foam
Fig. 2. Schematic diagrams of crack propagation from the polyurethane foam into aluminum secondary barrier due to the local elongation with thermal shrinkage of composite insulation panels.
In this study, the crack generation was investigated with finite element analysis. Also a new cryogenic experimental method was developed to investigate the cryogenic reliability of the composite insulation panels. For the crack retardation from the insulation panel into the aluminum secondary barrier, the adhesive between the aluminum barrier and the polyurethane foam was reinforced with glass fabric. From the experimental results, it was found that the glass fabric reinforced adhesive of composite
The large scale mock-up has been usually constructed to investigate the cryogenic reliability of the CCS with various types of composite insulation panels as shown in Fig. 3 [14]. The composite insulation panels can be classified into the flat panel and corner panel as shown in Fig. 4a. From the cryogenic experimental results with large scale mock-up with composite insulation panels, cracks were detected on the aluminum secondary barrier of the section A as shown in Fig. 4b. Although the fully annealed aluminum strip of 1000 series has high failure strain of 34%, local strain concentrations in the aluminum strip due to the local failures of polyurethane foam, led to the crack propagations in the aluminum secondary barrier as shown in Fig. 4b. The crack propagation into the aluminum secondary barrier will cause the safety problems such as LNG leakage and explosion.
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Table 1 Properties of polyurethane foam used for insulation foam.
Modulus Ex, Ez (MPa) Modulus Ey (MPa) Shear modulus (Gxy, Gxy, Gxy) Poisson’s ratio
vxy vyz vzx
Coefficient of thermal expansion (105)
ax ay az
Failure strain at 170 °C Tensile strength at 170 °C (MPa)
x, z direction x, z direction
Table 2 Properties of aluminum used for secondary barrier. Modulus (MPa) Poisson’s ratio Coefficient of thermal expansion (105 m/m K) Failure strain at 170 °C Tensile strength at 170 °C (MPa)
y
70 0.3 2.4 34% 98
x
y-symmetry
y
x z -163°C
x-symmetry
y
y x
y x
y x
x
Fixed condition, 25°C
(a) Stainless steel Adhesives
Aluminum
Polyurethane foam
Section A
(b) Fig. 5. Finite element model of flat panel and corner panel: (a) boundary conditions with coordinates and (b) mesh configuration of finite element model with fine mesh.
2.1. Finite element analysis of the composite insulation panels The stress and strain distributions in the composite insulation panels at the cryogenic temperature were calculated with finite element analysis with ABAQUS 6.10 (SIMULIA, USA). The properties of polyurethane foam and aluminum are listed in Tables 1 and 2, where x, z directions are in-plane direction and y direction is out of plane direction as shown in Fig. 5a. The flat panel and corner panel were modeled with the boundary condition in Fig. 5a. Fig. 5b shows the mesh configuration with fine meshes on the section A, where the cracks were detected in the aluminum secondary barrier during the mock-up test.
20 °C 136 73 11 0.24 0.39 0.09 3.09 5.86 3.05 2.1% 3.8
170 °C 185 117 21 0.20 0.32 0.09
Since the cracks were propagated into the aluminum secondary barrier from the insulation foam in section A of a corner panel in the mock up test, the stress and strain distributions induced by the thermal shrinkage at the cryogenic temperature were estimated with finite element analysis. Since the aluminum strip can function as the secondary barrier even though it yields until its rupture, the strain criterion rather than the yield criterion was employed for the failure of aluminum barrier. From the calculated results as shown in Fig. 6a, the maximum strain of the aluminum strip was 5.9%, which was 5.8 times lower than the maximum failure strain of the aluminum of 34% at the cryogenic temperature. Fig. 6a shows the distribution of maximum strain along the aluminum barrier. Since the polyurethane foam was brittle at the cryogenic temperature, the maximum strain criterion was employed for the failure criterion of polyurethane foam as follows [15]:
ecx < ex < etx
ð1:1Þ
ecy < ey < ety
ð1:2Þ
ecz < ez < etz
ð1:3Þ
Cyz < cyz < Cyz
ð1:4Þ
Czx < czx < Czx
ð1:5Þ
Cxy < cxy < Cxy
ð1:6Þ
where eti and eci (i = x, y, z) represent the failure tensile strain and compressive strain respectively, Cyz, Czx and Cxy represent the failure shear strains in the planes yz, zx and xy, respectively. The calculated maximum strain ex in the polyurethane foam (PUF) was 3.2% at the cryogenic temperature as shown in Fig. 6b. Fig. 6b shows the strains along the insulation foam under the aluminum barrier, where the strain concentrations occurred at the cryogenic temperature. Since the failure strain of polyurethane foam in x-direction is 2.1% as in Table 1, the polyurethane foam failed at the cryogenic temperature and the cracks propagated through the aluminum secondary barrier although the strain of aluminum was much lower than that of its failure strain as illustrated in Fig. 2. Since the aluminum secondary barrier failed at the low strain due to the crack propagated from the adhesively bonded PUF as shown in Fig. 4, the strain in the aluminum secondary barrier was calculated by the finite element analysis to estimate the reference strain for the safety of the aluminum barrier in the composite insulation panel. In Fig. 7, the section A is pulled 0.935 (0.321 + 0.614) mm along the x-direction due to the thermal shrinkage of flat panel and the corner panel at the cryogenic temperature. Also, the section A should extend to accommodate the thermal shrinkage at the cryogenic temperature, whose magnitude DL can be calculated as follows:
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y x z
Path along aluminum barrier (secondary barrier) Distance
Section A
Path along polyurethane foam under the aluminum barrier
3.5
35
(%)
25 20
2.5 2.0
Maximum value: 5.9%
15 10 5 0
Maximum value: 3.2% (strain concentration)
3.0
Failure strain of aluminum barrier: 34%
Strain,
Strain (%)
30
1.5
Failure strain of polyurethane foam: 2.1%
1.0 0.5
0
5
10
15
20
25
30
35
40
0.0
0
5
10
15
20
25
Distance (mm)
Distance (mm)
(a)
(b)
30
35
40
Fig. 6. Strain distributions of the aluminum barrier and polyurethane foam: (a) strain of the aluminum along the path of aluminum barrier (secondary barrier) and (b) strain of polyurethane foam along the path of polyurethane foam under the aluminum barrier.
DL ¼ a L0 DT
ð2Þ
where L0 is original length, a is coefficient of thermal expansion, and DT is temperature change [16,17]. The original length of section A is 35 mm and temperature difference is 188 °C when the panel cools down from 25 °C to 163 °C. From Eq. (2), the elongation of section A due to thermal shrinkage at the cryogenic temperature is 0.203 mm. Therefore, the total elongation of section A at the cryogenic temperature of 163 °C can be calculated to be 1.138 mm by summing up 0.935 mm and 0.203 mm and the strain of section A is expected to be 3.3% with the gage length of 35 mm.
Corner panel Flat panel
2.2. Cryogenic experiments Original position
0.321 mm mm 0.321
35 mm Shrinkage of flat panel
0.614 mm
Shrinkage of corner panel
Fig. 7. Schematic diagram of extension of section A along x-direction at cryogenic temperature due to the thermal shrinkage of flat panel and corner panel at cryogenic temperature.
A new test method was developed to evaluate the failure strain of aluminum secondary barrier in the foam insulation panel. The effect of reinforcement was also investigated with the developed method. The configuration of the specimen is shown in Fig. 8, where the specimen has a symmetric shape to avoid bending at the cryogenic temperature due to the difference of coefficient of thermal expansion. The gage length of the specimen was 35 mm in order to simulate the section A as shown in Fig. 4. The tab length was 60 mm for the uniform stress distribution at the gage length according to the Saint–Venant’s principle [18,19]. Since the crack
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Y.H. Yu et al. / Composite Structures 94 (2012) 462–468 Table 3 Thickness and areal density of plain weave E-glass fabric.
155 60
35
35
Fabric A Fabric B Fabric C Steel tab
RPUF
Aluminum
Thickness (mm)
Areal density (kg/m2)
0.06 0.18 0.23
0.05 0.17 0.25
10
Crack (3 mm) 5
Table 4 Adhesive thickness and volume fraction of E-glass fabric. Without reinforcement
20
Curing pressure (MPa)
Glass fabric reinforcement Fabric A
0.001 Aluminum RPUF
Polyurethane adhesive
Crack (3 mm)
0.05 (unit: mm)
Fig. 8. Shape of specimen to investigate the failure strain of aluminum barrier with polyurethane foam.
Insulation chamber
Fan
Controller
Liquid nitrogen tank
Specimen
Material testing system
(a) Thermocouple
Specimen
0.1
Adhesive thickness (mm) Volume fraction (%)
0.50
0.54
–
3.4
Adhesive thickness (mm) Volume fraction (%)
0.24
0.36
–
5.1
Adhesive thickness (mm) Volume fraction (%)
0.17
0.22
–
8.2
Fabric B
Fabric C
0.61
0.64
10.5 0.47
13.5 0.33
19.3
15.4 0.49
20.3 0.35
28.0
strain of the specimens at the cryogenic temperature. The crosshead speed was 1 mm/min and the chamber was cooled down to 150 °C, which was the lowest temperature with the material testing system available in this work. The temperature of a specimen was measured with a thermocouple as shown in Fig. 9b. The failure strain was measured with the rupture of aluminum. The moment of failure was detected by the load displacement curve. Since the crack generated in the PUF propagated into the aluminum barrier, the PUF below the aluminum secondary barrier was reinforced with plain weave E-glass fabrics. Several types of glass fabrics listed in Table 3 were impregnated with polyurethane adhesive. Different curing pressures up to 0.1 MPa, the maximum pressure for the curing of larger panels, were applied to avoid wrinkling of the secondary aluminum barrier during the bonding process. 3. Results and discussion
Insulation chamber Liquid nitrogen tank
(b) Fig. 9. Experimental setup for cryogenic experiment: (a) schematic diagram of the experimental setup and (b) photograph of the experimental setup.
would propagate from the polyurethane foam to the aluminum barrier, a pre-crack of 3 mm depth was fabricated on the PUF at the center of specimen with the razor blade tapping method [20]. The aluminum strip was bonded on the PUF with the polyurethane adhesive, which was cured at 25 °C for 12 h. The tests were conducted with material testing system (INSTRON 4206, Instron, USA) with a thermal insulation chamber, as shown in Fig. 9a. The tensile load was applied to the specimen to measure the failure
The adhesive thickness and volume fraction of glass fiber were measured with respect to curing pressures and fabric types as listed in Table 4. The measured failure strains of the aluminum barriers without reinforcement were 3.0%, 3.9%, 3.7% when the curing pressures were 0.001 MPa, 0.05 MPa and 0.1 MPa, respectively as shown in Fig. 10. Since the average value of the failure strain without reinforcement was 3.5%, which was slightly larger than the strain value of 3.3% obtained from the finite element analysis, the cracks might start in the aluminum secondary barrier at the cryogenic temperature of 163 °C. Therefore, without reinforcement, the aluminum barrier could be failed due to the local elongation of panels and the aluminum strip itself at cryogenic temperature of 163 °C. For the specimens with the reinforcement of fabric type of A in Table 3, the failure strains of the aluminum barrier were 3.7%, 4.0%, 3.9% when the curing pressures were 0.001 MPa, 0.05 MPa and 0.1 MPa, respectively. The average value of the failure strain of
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Failure strain of aluminum barrier (%)
Y.H. Yu et al. / Composite Structures 94 (2012) 462–468
Large yielding area 18 16
Curing pressure 0.001 MPa
0.05 MPa
Aluminum Reinforced polyurethane adhesive
0.1 MPa
14 12 10 8
Local interfacial debonding
6 4
Fiber reinforced polyurethane foam
2 0
1
2
1: without reinforcement
3
4
2: with glass fabric reinforcement (fabric A)
(a) Aluminum
3: with glass fabric reinforcement (fabric B) 4: with glass fabric reinforcement (fabric C)
Fig. 10. Failure strain of aluminum barrier with respect to the curing pressure for the specimens reinforced with fabric A, fabric B and fabric C.
Fiber reinforced polyurethane foam
Small yielding area Aluminum Polyurethane adhesive
(b) Fiber reinforced polyurethane foam
(a) Aluminum
Fiber reinforced polyurethane foam
(b) Fig. 11. Failure of the aluminum secondary barrier by local elongation: (a) schematic diagram of failure mode and (b) photograph of failure mode.
specimen was 3.9%. For the specimens with the reinforcement of fabric type B in Table 3, the failure strains of the aluminum barrier were 3.7%, 4.0%, 13.1% when the curing pressures were 0.001 MPa, 0.05 MPa and 0.1 MPa, respectively. For the specimens with the reinforcement of fabric type C in Table 3, the failure strains of aluminum barrier were 14.0%, 14.6%, 14.3% when the curing pressures were 0.001 MPa, 0.05 MPa and 0.1 MPa, respectively. The average value of the failure strain of specimen was 14.3% with the reinforcement of fabric type C, which was 4.3 times larger than the maximum strain of 3.3% obtained from the finite element analysis. However, the reinforcement with fabric type B hardly improved the failure strain of the aluminum secondary barrier when the curing pressures were 0.001–0.05 MPa. From the results as shown in Fig. 10 and Table 4, it was found that the glass fabric reinforcement was much effective to increase the failure strain of the aluminum barrier on the PUF when the volume fraction of glass fabric in polyurethane adhesive was larger than 15.4%. Therefore, the reinforcement with the glass fabric type C in Table 4 has been em-
Fig. 12. Failure of aluminum barrier after local interfacial debonding: (a) schematic diagram of failure mode and (b) photograph of failure mode.
ployed for the cryogenic reliability of the aluminum secondary barrier adhesively bonded to the PUF core in the composite sandwich insulation panels. The increase of the failure strain of the aluminum secondary barrier with reinforcement of glass fabric was due to the change of failure mode. For the cases at the low volume fraction of glass fabric in Fig. 10, the aluminum barrier was failed with less than 5% strain due to the both failures of the polyurethane adhesive and the PUF by the local strain concentration as shown in Fig. 11. However, for the cases in Fig. 10 with high fiber volume fraction, whose failure strain of the aluminum was higher than 10%, the local interfacial debonding between the aluminum barrier and the polyurethane adhesive was followed by the failure of polyurethane adhesive and the PUF. The local interfacial debonding caused the yielding of large area of the aluminum barrier before failure as shown in Fig. 12. The typical load displacement graphs of two different failure modes are shown in Fig. 13. For the failure mode with the local debonding, the load decreased suddenly when the PUF failed, but maintained some load carrying capability, which prevented the leakage of LNG until the final rupture of the aluminum secondary barrier. With the high volume fraction of glass fabric, the glass fabric could retard the crack propagation from the PUF to the aluminum barrier because the interface between the aluminum barrier and the adhesive layer could not transfer the extra load after debonding. Also, at the high volume fraction of glass fabric, the interfacial bonding strength between aluminum barrier and adhesive would decrease due to the direct contact between some of the glass fibers without adhesive and the aluminum barrier. Although the glass fabric reinforcement which is not fully impregnated with adhesive could decrease the bonding strength, the bonding strength was much higher than the strength of polyurethane foam of 3.8 MPa in Table 1. Therefore, the glass fabric reinforcement could be employed in order to increase the failure strain of the aluminum barrier adhesively bonded to the PUF of the composite sandwich insulation panels.
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8 Failure by local elongation
7
Failure after local interfacial debonding
Load (kN)
6
Failure of aluminum by local elongation (low failure strain of aluminum) Local interfacial debonding (high failure strain of aluminum)
5
Aluminum yielding with large area
4 3 2 1 0
funded by the Ministry of Education, Science and Technology (R31-2008-000-10045-0), Grant No. EEWS-2011-N01110017 from EEWS Research Project of the office of KAIST EEWS Initiative (EEWS: Energy, Environment, Water, and Sustainability) and BK21.
0
1
2
3
4
5
Displacement (mm) Fig. 13. Typical load–displacement curves for two different failure modes, failure by local elongation and failure after local interfacial debonding with glass fabric reinforcement at cryogenic temperature of 150 °C.
4. Conclusions In this study, the cryogenic reliability of composite thermal insulation panels of the cryogenic containment system (CCS) of liquefied natural gas (LNG) ships was investigated. It was found that the aluminum secondary barrier adhesively bonded to the polyurethane foam (PUF) failed by the propagation of the crack generated from the PUF under the aluminum secondary barrier. Although the aluminum barrier had a much high failure strain of 34% at the cryogenic temperature, the local strain concentration in the aluminum barrier bonded to the PUF reduced the failure strain of the aluminum to 3.3%. For the crack retardation, the aluminum secondary barrier was reinforced by the plain weave E-glass fabrics between the aluminum secondary barrier and the PUF. From the experiments, the failure strain of aluminum with polyurethane foam improved to 14.3% when the volume fraction of glass fabric was higher than 15.4% and the fabrics impregnated with adhesive were cured under the pressure of 0.001–0.1 MPa, which was 4.3 times higher than the failure strain of the aluminum at the cryogenic temperature. Therefore, the glass fabric reinforcement is one of the promising method to improve the cryogenic reliability of composite insulation panels. Acknowledgments This research was supported by WCU (World Class University) program through the National Research Foundation of Korea
References [1] Kim BC, Yoon SH, Lee DG. Pressure resistance of the corrugated stainless steel membranes of LNG carriers. Ocean Eng 2011;38:592–608. [2] Wegrzyn JE, Litzke WL, Gurevich M. DOE/BNL Liquid natural gas heavy vehicle program. Future transportation technology conference and exposition, costa mesa, California. SAE Technical Paper Series 981919, August (11–13);1998. [3] Wegrzyn JE, Gurevich M. Liquefied natural gas for trucks and buses. Government/industry meeting, Washington DC. SAE Technical Paper Series 2000-01-2210, June (19–21);2000. [4] Kim MH, Lee SM, Lee JM, Noh BJ, Kim WS. Fatigue strength assessment of mark-3 type LNG cargo containment system. Ocean Eng 2010;37:1243–52. [5] Kim JG, Yoon SH, Park SW, Lee DG. Optimum silane treatment for the adhesively bonded aluminum adherend at the cryogenic temperature. J Adhes Sci Technol 2010;24:775–87. [6] Harris FS. Safety features on LNG ships. Cryogenics 1993;8(33):772–7. [7] Junshiro I, Kiyokazu K, Hidetoshi M, Hidefumi I, Yoshihiro S. Building of advanced large sized membrane type LNG carrier. Mitsubishi Heavy Industries, Ltd., Technical Review 2004; 41(6): 1–7. [8] Kim BG, Lee DG. The design of an optical sensor arrangement for the detection of oil contamination in an adhesively bonded structure of a liquified natural gas (LNG) ship. Measur Sci Technol 2009;20:065204. [9] Kim BG, Lee DG. Leakage characteristics of the glass fabric composite barriers of LNG ships. Compos Struct 2008;86:27–36. [10] Shin YS, Kim JW, Wang B. First principle-based analysis procedure for strength assessment of membrane-type LNG containment system due to sloshing impact. In: World Maritime technology conference, London, UK; 2006. p.29– 30. [11] Gibson LJ, Ashby MF. Cellular solids-structure and properties. NY: Pergamon Press; 1988. p. 51–5. [12] Melcher RJ, Johnson WS. Mode 1 fracture toughness of an adhesively bonded composite–composite joint in a cryogenic temperature. Compos Sci Technol 2007;67(3):501–6. [13] Anderson TL. Fracture mechanics-fundamentals and applications. NY: CRC Press; 2005. p. 12–7. [14] Kim BG, Yoon SH, Kim JG, Bang CS, Lee DG. Large scale cryogenic composite insulation mock-up for liquefied natural gas (LNG) ship. In: 7th Asian– australian conference on composite materials, Taipei, Taiwan; 2010. [15] Lee DG, Suh NP. Axiomatic design and fabrication of composite structures. Oxford University Press; 2006. p. 165–67. [16] Callister Jr WD. Materials science and engineering-an introduction. NY: John wiley and sons, Inc.; 2000. p. 740. [17] Shackelford JF. Introduction to material science for engineers. NY: Pearson Education, Inc.; 2009. p. 225. [18] Beer FP, Johnston Jr ER, DeWolf JT. Mechanics of materials. NY: The McGrawHill Companies, Inc.; 2002. p. 106. [19] Gere JM. Mechanics of materials. NY: Thomson Learning, part of Thomson Corporation; 2006. p. 138–40. [20] Yoon SH, Kim BC, Lee KH, Lee DG. Improvement of the adhesive fracture toughness of bonded aluminum joints using E-glass fibers at cryogenic temperature. J Adhes Sci Technol 2010;24:429–44.
Contents lists available at SciVerse ScienceDirect
Composite Structures journal homepage: www.elsevier.com/locate/compstruct
Cryogenic reliability of composite insulation panels for liquefied natural gas (LNG) ships Young Ho Yu a, Bu Gi Kim a, Dai Gil Lee b,⇑ a School of Mechanical Aerospace & Systems Engineering, Korea Advanced Institute of Science and Technology, ME3221, Guseong-dong, Yuseong-gu, Daejeon 305-701, Republic of Korea b KAIST Institute for Design of Complex Systems (KIDCS), Guseong-dong, Yuseong-gu, Daejeon 305-701, Republic of Korea
a r t i c l e
i n f o
Article history: Available online 27 August 2011 Keywords: LNG ship Composite insulation panel Secondary barrier Crack retardation Glass fabric reinforcement
a b s t r a c t The major carrier of liquefied natural gas (LNG) is LNG ships, whose containment system is composed of dual barriers and composite insulation panels. The LNG containment system should have cryogenic reliability and high thermal insulation performance for safe and efficient transportation of LNG. The secondary barrier composed of adhesive bonded aluminum strips should keep tightness for 15 days, when the welded stainless primary barrier fails. However, cracks are generated in the composite insulation panels due to the local stress concentration and the brittleness of insulation materials at the cryogenic temperature of 163 °C. If cracks generated in the insulation panel propagate into the secondary barrier, LNG leakage problem might occur, which is a remaining concern in ship building industries. In this study, crack retardation capability in the composite insulation panel was investigated with glass fabric reinforcement. Finite element analysis was conducted to estimate the thermal stress at the cryogenic temperature and a new experimental method was developed to investigate the failure of secondary barrier of composite insulation panel. From the experimental results, it was found that the glass fabric reinforcement was effective to retard the crack propagation into the aluminum secondary barrier from the polyurethane insulation foam at the cryogenic temperature. Ó 2011 Elsevier Ltd. All rights reserved.
1. Introduction As the consumption of liquefied natural gas (LNG) has been increased much specially due to nuclear power plant disaster in Japan, the LNG ships have been spotlighted as major carriers of LNG [1–4]. Since LNG is normally carried in ships at 163 °C, the functional requirements of LNG ship are to have cryogenic reliability due to thermal cyclic stresses as well as to have high thermal insulation performance for safe and efficient transportation of LNG [5,6]. For efficient transportation, recently, the membrane type LNG containment systems are usually employed [7]. The membrane type LNG containment system has octagonal pillar shape with high capacity (larger than 39,000 m3), which is composed of dual barriers and composite insulation panels as shown in Fig. 1 [8,9]. The secondary barrier composed of adhesively bonded metal strips should keep tightness for 15 days when the welded stainless primary barrier fails. The insulation foam should have very low thermal conductivity for high thermal insulation performance, which limits the density of foam less than 120 kg/ ⇑ Corresponding author. E-mail addresses: [email protected] (Y.H. Yu), [email protected] (B.G. Kim), [email protected] (D.G. Lee). URLs: http://scs.kaist.ac.kr (Y.H. Yu), http://kidcs.kaist.ac.kr (D.G. Lee). 0263-8223/$ - see front matter Ó 2011 Elsevier Ltd. All rights reserved. doi:10.1016/j.compstruct.2011.08.009
m3. Since the foam density is not high enough for structural purpose and brittle at cryogenic temperatures, the reliability of conventional containment systems are not fully guaranteed due to the crack propagation in insulation foams, which is a remaining concern in ship building industries [10]. The composite insulation panels are composed of aluminum strip of 0.3 mm thickness, plywood and fiber reinforced polyurethane foam as shown in Fig. 1. The size of the insulation panel is 1 m (Width) 3 m (Length) 0.27 m (Height). The aluminum strip secondary barrier and polyurethane foam are adhesively bonded with polyurethane adhesive as shown in Fig. 2. During the operation of the cryogenic containment system (CCS) of LNG ships at the cryogenic temperature, cracks are generated in the composite insulation panels by the local stress concentrations due to the cyclic thermal stresses and the low fracture toughness of polyurethane foam at the cryogenic temperature [11–13]. In order to retard the crack generation, the polyurethane foam (PUF) is reinforced with glass fibers as shown in Fig. 2. However, the uniform glass fiber reinforcement with high volume fraction in PUF is not only possible in the mass production, but also it increases the thermal conductivity of polyurethane foam, which decreases the efficiency of the LNG transportation. When the cracks are generated, they propagate into the aluminum secondary barrier as shown in Fig. 2, which will induce the leakage of LNG.
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LNG
Corner insulation panel Flat insulation panel
Containment system
Containment system
LNG
Primary barrier
Composite insulation panel
Stainless steel
Plywood
Polyurethane foam
Aluminum
Secondary barrier
Fig. 3. Photograph of the large scale mock-up to investigate the cryogenic reliability of LNG containment system.
Corner panel
Fig. 1. Schematic diagrams of LNG containment system composed of dual barriers and composite insulation panels.
Flat panel Polyurethane foam Plywood
Aluminum
Stainless steel
LNG
(a)
Primary barrier
Composite insulation panel
Crack
Plywood
Stainless steel
Aluminum Section A
Polyurethane foam
Aluminum
Secondary barrier Polyurethane adhesive
(b) Aluminum barrier Polyurethane adhesive
Local elongation due to the crack propagation
Tension due to thermal shrinkage
Fig. 4. Composite insulation panels used in mock-up test: (a) schematic diagram of flat and corner of composite insulation panels in mock-up test and (b) photograph and schematic diagram of crack on the aluminum barrier generated on section A during mock-up test.
insulation panels retarded the crack propagation into the aluminum secondary barrier, which improved the cryogenic reliability of composite insulation panel. 2. Experimental Glass fiber
Crack
Polyurethane foam
Fig. 2. Schematic diagrams of crack propagation from the polyurethane foam into aluminum secondary barrier due to the local elongation with thermal shrinkage of composite insulation panels.
In this study, the crack generation was investigated with finite element analysis. Also a new cryogenic experimental method was developed to investigate the cryogenic reliability of the composite insulation panels. For the crack retardation from the insulation panel into the aluminum secondary barrier, the adhesive between the aluminum barrier and the polyurethane foam was reinforced with glass fabric. From the experimental results, it was found that the glass fabric reinforced adhesive of composite
The large scale mock-up has been usually constructed to investigate the cryogenic reliability of the CCS with various types of composite insulation panels as shown in Fig. 3 [14]. The composite insulation panels can be classified into the flat panel and corner panel as shown in Fig. 4a. From the cryogenic experimental results with large scale mock-up with composite insulation panels, cracks were detected on the aluminum secondary barrier of the section A as shown in Fig. 4b. Although the fully annealed aluminum strip of 1000 series has high failure strain of 34%, local strain concentrations in the aluminum strip due to the local failures of polyurethane foam, led to the crack propagations in the aluminum secondary barrier as shown in Fig. 4b. The crack propagation into the aluminum secondary barrier will cause the safety problems such as LNG leakage and explosion.
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Table 1 Properties of polyurethane foam used for insulation foam.
Modulus Ex, Ez (MPa) Modulus Ey (MPa) Shear modulus (Gxy, Gxy, Gxy) Poisson’s ratio
vxy vyz vzx
Coefficient of thermal expansion (105)
ax ay az
Failure strain at 170 °C Tensile strength at 170 °C (MPa)
x, z direction x, z direction
Table 2 Properties of aluminum used for secondary barrier. Modulus (MPa) Poisson’s ratio Coefficient of thermal expansion (105 m/m K) Failure strain at 170 °C Tensile strength at 170 °C (MPa)
y
70 0.3 2.4 34% 98
x
y-symmetry
y
x z -163°C
x-symmetry
y
y x
y x
y x
x
Fixed condition, 25°C
(a) Stainless steel Adhesives
Aluminum
Polyurethane foam
Section A
(b) Fig. 5. Finite element model of flat panel and corner panel: (a) boundary conditions with coordinates and (b) mesh configuration of finite element model with fine mesh.
2.1. Finite element analysis of the composite insulation panels The stress and strain distributions in the composite insulation panels at the cryogenic temperature were calculated with finite element analysis with ABAQUS 6.10 (SIMULIA, USA). The properties of polyurethane foam and aluminum are listed in Tables 1 and 2, where x, z directions are in-plane direction and y direction is out of plane direction as shown in Fig. 5a. The flat panel and corner panel were modeled with the boundary condition in Fig. 5a. Fig. 5b shows the mesh configuration with fine meshes on the section A, where the cracks were detected in the aluminum secondary barrier during the mock-up test.
20 °C 136 73 11 0.24 0.39 0.09 3.09 5.86 3.05 2.1% 3.8
170 °C 185 117 21 0.20 0.32 0.09
Since the cracks were propagated into the aluminum secondary barrier from the insulation foam in section A of a corner panel in the mock up test, the stress and strain distributions induced by the thermal shrinkage at the cryogenic temperature were estimated with finite element analysis. Since the aluminum strip can function as the secondary barrier even though it yields until its rupture, the strain criterion rather than the yield criterion was employed for the failure of aluminum barrier. From the calculated results as shown in Fig. 6a, the maximum strain of the aluminum strip was 5.9%, which was 5.8 times lower than the maximum failure strain of the aluminum of 34% at the cryogenic temperature. Fig. 6a shows the distribution of maximum strain along the aluminum barrier. Since the polyurethane foam was brittle at the cryogenic temperature, the maximum strain criterion was employed for the failure criterion of polyurethane foam as follows [15]:
ecx < ex < etx
ð1:1Þ
ecy < ey < ety
ð1:2Þ
ecz < ez < etz
ð1:3Þ
Cyz < cyz < Cyz
ð1:4Þ
Czx < czx < Czx
ð1:5Þ
Cxy < cxy < Cxy
ð1:6Þ
where eti and eci (i = x, y, z) represent the failure tensile strain and compressive strain respectively, Cyz, Czx and Cxy represent the failure shear strains in the planes yz, zx and xy, respectively. The calculated maximum strain ex in the polyurethane foam (PUF) was 3.2% at the cryogenic temperature as shown in Fig. 6b. Fig. 6b shows the strains along the insulation foam under the aluminum barrier, where the strain concentrations occurred at the cryogenic temperature. Since the failure strain of polyurethane foam in x-direction is 2.1% as in Table 1, the polyurethane foam failed at the cryogenic temperature and the cracks propagated through the aluminum secondary barrier although the strain of aluminum was much lower than that of its failure strain as illustrated in Fig. 2. Since the aluminum secondary barrier failed at the low strain due to the crack propagated from the adhesively bonded PUF as shown in Fig. 4, the strain in the aluminum secondary barrier was calculated by the finite element analysis to estimate the reference strain for the safety of the aluminum barrier in the composite insulation panel. In Fig. 7, the section A is pulled 0.935 (0.321 + 0.614) mm along the x-direction due to the thermal shrinkage of flat panel and the corner panel at the cryogenic temperature. Also, the section A should extend to accommodate the thermal shrinkage at the cryogenic temperature, whose magnitude DL can be calculated as follows:
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y x z
Path along aluminum barrier (secondary barrier) Distance
Section A
Path along polyurethane foam under the aluminum barrier
3.5
35
(%)
25 20
2.5 2.0
Maximum value: 5.9%
15 10 5 0
Maximum value: 3.2% (strain concentration)
3.0
Failure strain of aluminum barrier: 34%
Strain,
Strain (%)
30
1.5
Failure strain of polyurethane foam: 2.1%
1.0 0.5
0
5
10
15
20
25
30
35
40
0.0
0
5
10
15
20
25
Distance (mm)
Distance (mm)
(a)
(b)
30
35
40
Fig. 6. Strain distributions of the aluminum barrier and polyurethane foam: (a) strain of the aluminum along the path of aluminum barrier (secondary barrier) and (b) strain of polyurethane foam along the path of polyurethane foam under the aluminum barrier.
DL ¼ a L0 DT
ð2Þ
where L0 is original length, a is coefficient of thermal expansion, and DT is temperature change [16,17]. The original length of section A is 35 mm and temperature difference is 188 °C when the panel cools down from 25 °C to 163 °C. From Eq. (2), the elongation of section A due to thermal shrinkage at the cryogenic temperature is 0.203 mm. Therefore, the total elongation of section A at the cryogenic temperature of 163 °C can be calculated to be 1.138 mm by summing up 0.935 mm and 0.203 mm and the strain of section A is expected to be 3.3% with the gage length of 35 mm.
Corner panel Flat panel
2.2. Cryogenic experiments Original position
0.321 mm mm 0.321
35 mm Shrinkage of flat panel
0.614 mm
Shrinkage of corner panel
Fig. 7. Schematic diagram of extension of section A along x-direction at cryogenic temperature due to the thermal shrinkage of flat panel and corner panel at cryogenic temperature.
A new test method was developed to evaluate the failure strain of aluminum secondary barrier in the foam insulation panel. The effect of reinforcement was also investigated with the developed method. The configuration of the specimen is shown in Fig. 8, where the specimen has a symmetric shape to avoid bending at the cryogenic temperature due to the difference of coefficient of thermal expansion. The gage length of the specimen was 35 mm in order to simulate the section A as shown in Fig. 4. The tab length was 60 mm for the uniform stress distribution at the gage length according to the Saint–Venant’s principle [18,19]. Since the crack
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Y.H. Yu et al. / Composite Structures 94 (2012) 462–468 Table 3 Thickness and areal density of plain weave E-glass fabric.
155 60
35
35
Fabric A Fabric B Fabric C Steel tab
RPUF
Aluminum
Thickness (mm)
Areal density (kg/m2)
0.06 0.18 0.23
0.05 0.17 0.25
10
Crack (3 mm) 5
Table 4 Adhesive thickness and volume fraction of E-glass fabric. Without reinforcement
20
Curing pressure (MPa)
Glass fabric reinforcement Fabric A
0.001 Aluminum RPUF
Polyurethane adhesive
Crack (3 mm)
0.05 (unit: mm)
Fig. 8. Shape of specimen to investigate the failure strain of aluminum barrier with polyurethane foam.
Insulation chamber
Fan
Controller
Liquid nitrogen tank
Specimen
Material testing system
(a) Thermocouple
Specimen
0.1
Adhesive thickness (mm) Volume fraction (%)
0.50
0.54
–
3.4
Adhesive thickness (mm) Volume fraction (%)
0.24
0.36
–
5.1
Adhesive thickness (mm) Volume fraction (%)
0.17
0.22
–
8.2
Fabric B
Fabric C
0.61
0.64
10.5 0.47
13.5 0.33
19.3
15.4 0.49
20.3 0.35
28.0
strain of the specimens at the cryogenic temperature. The crosshead speed was 1 mm/min and the chamber was cooled down to 150 °C, which was the lowest temperature with the material testing system available in this work. The temperature of a specimen was measured with a thermocouple as shown in Fig. 9b. The failure strain was measured with the rupture of aluminum. The moment of failure was detected by the load displacement curve. Since the crack generated in the PUF propagated into the aluminum barrier, the PUF below the aluminum secondary barrier was reinforced with plain weave E-glass fabrics. Several types of glass fabrics listed in Table 3 were impregnated with polyurethane adhesive. Different curing pressures up to 0.1 MPa, the maximum pressure for the curing of larger panels, were applied to avoid wrinkling of the secondary aluminum barrier during the bonding process. 3. Results and discussion
Insulation chamber Liquid nitrogen tank
(b) Fig. 9. Experimental setup for cryogenic experiment: (a) schematic diagram of the experimental setup and (b) photograph of the experimental setup.
would propagate from the polyurethane foam to the aluminum barrier, a pre-crack of 3 mm depth was fabricated on the PUF at the center of specimen with the razor blade tapping method [20]. The aluminum strip was bonded on the PUF with the polyurethane adhesive, which was cured at 25 °C for 12 h. The tests were conducted with material testing system (INSTRON 4206, Instron, USA) with a thermal insulation chamber, as shown in Fig. 9a. The tensile load was applied to the specimen to measure the failure
The adhesive thickness and volume fraction of glass fiber were measured with respect to curing pressures and fabric types as listed in Table 4. The measured failure strains of the aluminum barriers without reinforcement were 3.0%, 3.9%, 3.7% when the curing pressures were 0.001 MPa, 0.05 MPa and 0.1 MPa, respectively as shown in Fig. 10. Since the average value of the failure strain without reinforcement was 3.5%, which was slightly larger than the strain value of 3.3% obtained from the finite element analysis, the cracks might start in the aluminum secondary barrier at the cryogenic temperature of 163 °C. Therefore, without reinforcement, the aluminum barrier could be failed due to the local elongation of panels and the aluminum strip itself at cryogenic temperature of 163 °C. For the specimens with the reinforcement of fabric type of A in Table 3, the failure strains of the aluminum barrier were 3.7%, 4.0%, 3.9% when the curing pressures were 0.001 MPa, 0.05 MPa and 0.1 MPa, respectively. The average value of the failure strain of
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Failure strain of aluminum barrier (%)
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Large yielding area 18 16
Curing pressure 0.001 MPa
0.05 MPa
Aluminum Reinforced polyurethane adhesive
0.1 MPa
14 12 10 8
Local interfacial debonding
6 4
Fiber reinforced polyurethane foam
2 0
1
2
1: without reinforcement
3
4
2: with glass fabric reinforcement (fabric A)
(a) Aluminum
3: with glass fabric reinforcement (fabric B) 4: with glass fabric reinforcement (fabric C)
Fig. 10. Failure strain of aluminum barrier with respect to the curing pressure for the specimens reinforced with fabric A, fabric B and fabric C.
Fiber reinforced polyurethane foam
Small yielding area Aluminum Polyurethane adhesive
(b) Fiber reinforced polyurethane foam
(a) Aluminum
Fiber reinforced polyurethane foam
(b) Fig. 11. Failure of the aluminum secondary barrier by local elongation: (a) schematic diagram of failure mode and (b) photograph of failure mode.
specimen was 3.9%. For the specimens with the reinforcement of fabric type B in Table 3, the failure strains of the aluminum barrier were 3.7%, 4.0%, 13.1% when the curing pressures were 0.001 MPa, 0.05 MPa and 0.1 MPa, respectively. For the specimens with the reinforcement of fabric type C in Table 3, the failure strains of aluminum barrier were 14.0%, 14.6%, 14.3% when the curing pressures were 0.001 MPa, 0.05 MPa and 0.1 MPa, respectively. The average value of the failure strain of specimen was 14.3% with the reinforcement of fabric type C, which was 4.3 times larger than the maximum strain of 3.3% obtained from the finite element analysis. However, the reinforcement with fabric type B hardly improved the failure strain of the aluminum secondary barrier when the curing pressures were 0.001–0.05 MPa. From the results as shown in Fig. 10 and Table 4, it was found that the glass fabric reinforcement was much effective to increase the failure strain of the aluminum barrier on the PUF when the volume fraction of glass fabric in polyurethane adhesive was larger than 15.4%. Therefore, the reinforcement with the glass fabric type C in Table 4 has been em-
Fig. 12. Failure of aluminum barrier after local interfacial debonding: (a) schematic diagram of failure mode and (b) photograph of failure mode.
ployed for the cryogenic reliability of the aluminum secondary barrier adhesively bonded to the PUF core in the composite sandwich insulation panels. The increase of the failure strain of the aluminum secondary barrier with reinforcement of glass fabric was due to the change of failure mode. For the cases at the low volume fraction of glass fabric in Fig. 10, the aluminum barrier was failed with less than 5% strain due to the both failures of the polyurethane adhesive and the PUF by the local strain concentration as shown in Fig. 11. However, for the cases in Fig. 10 with high fiber volume fraction, whose failure strain of the aluminum was higher than 10%, the local interfacial debonding between the aluminum barrier and the polyurethane adhesive was followed by the failure of polyurethane adhesive and the PUF. The local interfacial debonding caused the yielding of large area of the aluminum barrier before failure as shown in Fig. 12. The typical load displacement graphs of two different failure modes are shown in Fig. 13. For the failure mode with the local debonding, the load decreased suddenly when the PUF failed, but maintained some load carrying capability, which prevented the leakage of LNG until the final rupture of the aluminum secondary barrier. With the high volume fraction of glass fabric, the glass fabric could retard the crack propagation from the PUF to the aluminum barrier because the interface between the aluminum barrier and the adhesive layer could not transfer the extra load after debonding. Also, at the high volume fraction of glass fabric, the interfacial bonding strength between aluminum barrier and adhesive would decrease due to the direct contact between some of the glass fibers without adhesive and the aluminum barrier. Although the glass fabric reinforcement which is not fully impregnated with adhesive could decrease the bonding strength, the bonding strength was much higher than the strength of polyurethane foam of 3.8 MPa in Table 1. Therefore, the glass fabric reinforcement could be employed in order to increase the failure strain of the aluminum barrier adhesively bonded to the PUF of the composite sandwich insulation panels.
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8 Failure by local elongation
7
Failure after local interfacial debonding
Load (kN)
6
Failure of aluminum by local elongation (low failure strain of aluminum) Local interfacial debonding (high failure strain of aluminum)
5
Aluminum yielding with large area
4 3 2 1 0
funded by the Ministry of Education, Science and Technology (R31-2008-000-10045-0), Grant No. EEWS-2011-N01110017 from EEWS Research Project of the office of KAIST EEWS Initiative (EEWS: Energy, Environment, Water, and Sustainability) and BK21.
0
1
2
3
4
5
Displacement (mm) Fig. 13. Typical load–displacement curves for two different failure modes, failure by local elongation and failure after local interfacial debonding with glass fabric reinforcement at cryogenic temperature of 150 °C.
4. Conclusions In this study, the cryogenic reliability of composite thermal insulation panels of the cryogenic containment system (CCS) of liquefied natural gas (LNG) ships was investigated. It was found that the aluminum secondary barrier adhesively bonded to the polyurethane foam (PUF) failed by the propagation of the crack generated from the PUF under the aluminum secondary barrier. Although the aluminum barrier had a much high failure strain of 34% at the cryogenic temperature, the local strain concentration in the aluminum barrier bonded to the PUF reduced the failure strain of the aluminum to 3.3%. For the crack retardation, the aluminum secondary barrier was reinforced by the plain weave E-glass fabrics between the aluminum secondary barrier and the PUF. From the experiments, the failure strain of aluminum with polyurethane foam improved to 14.3% when the volume fraction of glass fabric was higher than 15.4% and the fabrics impregnated with adhesive were cured under the pressure of 0.001–0.1 MPa, which was 4.3 times higher than the failure strain of the aluminum at the cryogenic temperature. Therefore, the glass fabric reinforcement is one of the promising method to improve the cryogenic reliability of composite insulation panels. Acknowledgments This research was supported by WCU (World Class University) program through the National Research Foundation of Korea
References [1] Kim BC, Yoon SH, Lee DG. Pressure resistance of the corrugated stainless steel membranes of LNG carriers. Ocean Eng 2011;38:592–608. [2] Wegrzyn JE, Litzke WL, Gurevich M. DOE/BNL Liquid natural gas heavy vehicle program. Future transportation technology conference and exposition, costa mesa, California. SAE Technical Paper Series 981919, August (11–13);1998. [3] Wegrzyn JE, Gurevich M. Liquefied natural gas for trucks and buses. Government/industry meeting, Washington DC. SAE Technical Paper Series 2000-01-2210, June (19–21);2000. [4] Kim MH, Lee SM, Lee JM, Noh BJ, Kim WS. Fatigue strength assessment of mark-3 type LNG cargo containment system. Ocean Eng 2010;37:1243–52. [5] Kim JG, Yoon SH, Park SW, Lee DG. Optimum silane treatment for the adhesively bonded aluminum adherend at the cryogenic temperature. J Adhes Sci Technol 2010;24:775–87. [6] Harris FS. Safety features on LNG ships. Cryogenics 1993;8(33):772–7. [7] Junshiro I, Kiyokazu K, Hidetoshi M, Hidefumi I, Yoshihiro S. Building of advanced large sized membrane type LNG carrier. Mitsubishi Heavy Industries, Ltd., Technical Review 2004; 41(6): 1–7. [8] Kim BG, Lee DG. The design of an optical sensor arrangement for the detection of oil contamination in an adhesively bonded structure of a liquified natural gas (LNG) ship. Measur Sci Technol 2009;20:065204. [9] Kim BG, Lee DG. Leakage characteristics of the glass fabric composite barriers of LNG ships. Compos Struct 2008;86:27–36. [10] Shin YS, Kim JW, Wang B. First principle-based analysis procedure for strength assessment of membrane-type LNG containment system due to sloshing impact. In: World Maritime technology conference, London, UK; 2006. p.29– 30. [11] Gibson LJ, Ashby MF. Cellular solids-structure and properties. NY: Pergamon Press; 1988. p. 51–5. [12] Melcher RJ, Johnson WS. Mode 1 fracture toughness of an adhesively bonded composite–composite joint in a cryogenic temperature. Compos Sci Technol 2007;67(3):501–6. [13] Anderson TL. Fracture mechanics-fundamentals and applications. NY: CRC Press; 2005. p. 12–7. [14] Kim BG, Yoon SH, Kim JG, Bang CS, Lee DG. Large scale cryogenic composite insulation mock-up for liquefied natural gas (LNG) ship. In: 7th Asian– australian conference on composite materials, Taipei, Taiwan; 2010. [15] Lee DG, Suh NP. Axiomatic design and fabrication of composite structures. Oxford University Press; 2006. p. 165–67. [16] Callister Jr WD. Materials science and engineering-an introduction. NY: John wiley and sons, Inc.; 2000. p. 740. [17] Shackelford JF. Introduction to material science for engineers. NY: Pearson Education, Inc.; 2009. p. 225. [18] Beer FP, Johnston Jr ER, DeWolf JT. Mechanics of materials. NY: The McGrawHill Companies, Inc.; 2002. p. 106. [19] Gere JM. Mechanics of materials. NY: Thomson Learning, part of Thomson Corporation; 2006. p. 138–40. [20] Yoon SH, Kim BC, Lee KH, Lee DG. Improvement of the adhesive fracture toughness of bonded aluminum joints using E-glass fibers at cryogenic temperature. J Adhes Sci Technol 2010;24:429–44.