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composite skin hybrid ship hull, Part II: Manufacturing and sagging testing
Composites: Part A 38 (2007) 1763–1772 www.elsevier.com/locate/compositesa
Steel truss/composite skin hybrid ship hull. Part II: Manufacturing and sagging testing William J. Maroun a, Jun Cao a, Joachim L. Grenestedt a
b,*
Department of Mechanical Engineering and Mechanics, Lehigh University, 19 Memorial Drive West, Bethlehem, PA 18015, USA b Department of Mechanical Engineering and Mechanics, ATLSS Center Faculty Associate, Lehigh University, 19 Memorial Drive West, Bethlehem, PA 18015, USA Received 10 April 2006; received in revised form 1 November 2006; accepted 3 November 2006
Abstract A six meter subscale steel truss/composite skin hybrid ship hull model was developed based on an optimized design of a full scale model. The subscale model was finite element analyzed, manufactured and tested under sagging loads. It was made of a welded stainless steel truss to which 60 composite sandwich panels were bonded. The model was loaded to 36% above the design load, at which point there was substantial yielding and residual deformation of the steel truss. However, there was no indication of damage in any of the composite sandwich panels, nor in the bonds between the panels and the steel truss. 2006 Elsevier Ltd. All rights reserved. Keywords: Hybrid ship hulls; C. Finite element analysis; D. Mechanical testing; E. Joints/joining
1. Introduction A number of hybrid ship hull concepts have recently been studied with the aim to combine the attributes of both steel and composites [1–4]. The present study concentrated on hybrid ship hulls consisting of a steel truss closed out with very large composite sandwich panels. A short section showing the structural principle is presented in Fig. 1. The steel truss was designed to carry the bending moment, whereas the composite skins were designed mainly to carry shear and water pressure loads. A 142 m ship hull, similar to a destroyer in terms of size, weight and speed, was designed, finite element analyzed and optimized prior to the current research [5]. At present, a simplified model with a 1:35 scale cross-section was developed and analyzed based on the optimized full scale model [5]. This subscale model was subsequently manufactured. Joints between the steel truss and the composite panels were investigated experimentally. A 6*
Corresponding author. Tel.: +1 610 758 4129; fax: +1 610 758 6224. E-mail address: [email protected] (J.L. Grenestedt).
1359-835X/$ - see front matter 2006 Elsevier Ltd. All rights reserved. doi:10.1016/j.compositesa.2006.11.003
point bending test fixture was developed and forces were applied to simulate sagging loads. The FE model of the subscale hybrid hull was analyzed, using the loads applied during the test. Results from FE analyses were compared with the experiment. There are naturally a number of items which could not be correctly scaled, including fiber and fabric geometry, material toughness over strength, bond line thickness, etc. Some efforts aimed at reducing these effects were attempted by using a light fabric with thin tows (few filaments in the yarns), and thin bond lines. 2. Joints between composite panel and steel truss For the hybrid ship hull it is envisioned that the composite panels would first be fabricated and trimmed to size and then attached to the steel truss. Attaching the large panels to the steel truss may most readily be done by adhesive bonding. In order to reduce the severe elastic stiffness mismatch between the steel and the composite adherends, the edge of the steel can be modified to reduce its effective stiffness. Various ways to achieve this have been investigated, including perforating the steel [6,7], cutting fingers in the
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4.5mm Ø3.5mm
2.5mm
2.5mm
Ø3mm 4.5 mm Ø2.5mm
4.5mm Ø2.75mm 25mm Perforation type A
1.5mm
1.5mm
4.5mm
6.4mm
steel edge [8,9], etc. In current research, joints with either perforations or fingers cut in the steel adherend were experimentally investigated. The goal was to develop a joint suitable for the small scale hybrid ship hull that is simple to manufacture but still light and strong. The reasons for developing a small scale joint, rather than developing a full scale joint and scaling it down, were mainly cost and simplicity. Coupon specimens of various joints were made by manufacturing composite sandwich plates and steel adherends and bonding these together. The cured specimens were mechanically tested to failure under tension. A typical specimen loaded in a tensile test machine is shown in Fig. 2. The sandwich plates were made by vacuum infusing vinyl ester resin into thin glass fiber skins on each side of a PVC foam core. The resin used was Derakane 8084 vinyl ester and the reinforcement was Hexcel 7725 2/2 twill glass
6.4mm
Fig. 1. Short section showing the principle of the hybrid hull. Only one sandwich panel, the deck, is shown.
fiber weave. Two layers of the glass fabric were put on a release agent coated steel mold, on top of which a 12.7 mm thick Divinycell H200 foam piece with 45 beveled edges was placed. One layer of glass fiber was laid on top, such that the inner and outer skins came together outside the beveled edge of the foam. A local reinforcing glass fiber strip was laid over the beveled edges on the inner (beveled) skin. The whole assembly was then vacuum bagged to the mold and infused with vinyl ester resin. After demolding, the panels were trimmed with a waterjet cutter. The steel adherends were made of 2.7 mm thick AL6XN stainless steel. One end of the steel adherend was either perforated or cut with fingers, using a waterjet cutter. Two different perforation configurations and three different finger configurations were designed as shown in Fig. 3. In Finger type B, both the tips of the fingers and the roots between the fingers were nominally 0.95 mm wide. The tips and roots were then rounded (R = 0.4 mm). The distance between fingers (center to center) was 4 mm and the fingers were 10 mm long. Finger type C was identical except the fingers were 6 mm long. Instead of cutting individual steel adherend to the required specimen width (25 mm) before bonding to the composite sandwich adherend, a big steel plate cut with the designed shapes of many specimens was grit blasted, solvent cleaned and bonded to the
90mm
1764
25mm Perforation type B
4mm
4mm
90mm
Fig. 2. Coupon test specimen for joint development loaded in tensile test machine.
Finger type A
Finger type B
25mm
R0.4mm
25mm
6mm
10mm
2.5mm R0.4mm
25mm
R0.75mm
10mm
2.5mm 1.25mm
Finger type C
Fig. 3. Five different designs of the steel adherends: two perforation configurations and three finger configurations.
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The test results also showed that specimens with reinforcement strips on both the inner and the outer skins had considerably lower failure loads than specimens with reinforcement strips only on the inner skin. It appears that the reinforcement strips on the outside created a more unbalanced specimen and thus lead to more bending, which reduced the overall strength. Based on these tests, the tapered finger joint design and the SIA E2119 adhesive were chosen for bonding the composite sandwich panels to the steel truss in the small scale ship hull model. The reinforcement strips would be applied only on the inner side of the sandwich panels.
composite sandwich plate. After cure, the plate was cut into individual 25 mm wide specimens. The overlap of the steel and composite skin was 20 mm. A total of 56 specimens were manufactured and two comparison tests were performed. An MTI modified 44 kN (10,000 lbs) Instron universal test machine was used for all tests. The specimens were tested under quasi static tensile load at a speed of 1 mm/min. The most common failure mode was skin debonding from the foam core some distance from the joint. Five different joints types (Fig. 3) were first compared by using specimens with the same fiber glass skins (outlined above), foam and adhesive (Hysol 9360). The specimens had the following strengths (failure load/width): Perforation type A: 143.0 ± 0.3 N/mm; Perforation type B: 151.1 ± 6.6 N/mm; Finger type A: 154.8 ± 9.9 N/mm; Finger type B: 164.2 ± 10.6 N/mm; Finger type C: 163.2 ± 13.8 N/mm. Specimens with finger type B and finger type C had similar failure loads and were stronger than specimens with other joint designs. Three different adhesives, Hysol 9360, Hysol 9430 and SIA E2119, were then compared by using specimens having the same joint design (Finger type C in Fig. 3), same fiber glass skins and same foam. When manufacturing these specimens, local reinforcement glass fiber strips were applied on both the inner (beveled) and the outer (flat) skins, which differed from the specimens used in the earlier testing. The specimens had the following strengths (failure load/width): specimens using Hysol 9360: 78.9 ± 6.5 N/ mm; specimens using Hysol 9430: 87.1 ± 8.4 N/mm; specimens using SIA E2119: 84.0 ± 6.0 N/mm. Specimens using Hysol 9430 adhesive had the highest failure load, followed by specimens bonded with SIA E2119. Considering the manufacturing convenience and the large amount of adhesive required, SIA E2119 adhesive mixed with a mixing gun was preferred over the hand-mixed Hysol 9430 adhesive.
3. Design of small scale ship hull specimen A 6 m subscale model was developed for structural testing purposes. The cross-section of the model was obtained by scaling the optimized full-sized ship by 1:35. In order to obtain a feasible load distribution using a few discrete loads, the model was made longer than a true scale model would be. The transverse structural dimensions were all scaled by 1:35. If overhead bars are used for quantities of the scale model, and n = 35 is the scale factor, then the scaled loads are M ; n3 F F ¼ 2; n P ¼ P; M¼
ð1Þ ð2Þ ð3Þ
20 ˚
where M = Msl = 2.0 · 109 N m is the maximum slamming induced bending moment, F = Fsl = 5.1 · 107 N is the maximum slamming induced shear force, and P = Psl = 0.29 MPa is the maximum bottom slamming pressure used
Keel longeron Bottom panel
Bottom panel
Side longeron
Side longeron Bulkhead
Side panel
Side panel Bulkhead
Bulkhead Deck longeron
30.6 mm
Bulkhead
24.9 mm
364.3 mm
Bulkhead
Deck longeron
Deck panel
514.3 mm Fig. 4. Cross-section of 1:35 subscale hull model. This picture shows the keel up, as the model was mounted in the test frame during the sagging test.
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Table 1 Dimensions and design loads for the subscale model Specimen dimension
Length of Thickness Thickness skin Thickness Thickness Design bending moment M Design shear load F
1:35 hybrid ship hull of steel in truss of inner (beveled) composite
6.096 m 2 mm 0.25 mm
of outer (flat) composite skin of foam core
0.5 mm 12.7 mm 46,600 N m 41,600 N
for the design of the full scale ship [5]. Elastic moduli were not scaled. The cross-section of the 1:35 scale hull specimen is shown Fig. 4 and dimension of this small scale model and its design loads are listed in Table 1. 4. Small scale ship hull manufacturing A 6 m long ship hull was manufactured. It consisted of a welded non-magnetic AL-6XN stainless steel truss, which was closed out with 60 glass fiber reinforced vinyl ester skin/foam core sandwich panels (Fig. 5). Stainless steel AL-6XN by Allegheny Ludlum Corp., was used to make the steel truss. The stainless steel thickness was 2 mm (which differs from the previous test specimens). The fingers tips and the roots between the fingers were nominally 2.28 mm wide, and then rounded (R = 1 mm). The distance between fingers (center to center) was 8.28 mm and the height of the fingers was 6 mm. The composite panels consisted of glass fiber skins and a PVC foam core. The reinforcement was Hexcel 7725 2/2 twill glass fiber weave.
Divinycell PVC based foam cores with the thickness 12.7 mm were used for all panels. Four different core densities were used: H130, H160, H200, and H250. Higher density foams were used in water pressure loaded bottom panels and highly shear loaded side panels. The exact locations will be explained in a later paragraph. The resin used throughout was Derakane 8084 vinyl ester. SIA E2119 adhesive from Sovereign Specialty Chemicals was used for bonding the sandwich panels to the steel truss. 5. Steel truss manufacturing The stainless steel truss was made by laser and waterjet cutting the 2 mm thick steel. The edges of the steel to be joined to the sandwich panels were cut with fingers (Fig. 5). The 6 m steel truss was made in three separate sections, 1.5 m, 3 m, and 1.5 m long, respectively. In each section, the steel components were first tack welded together. The longerons and the keel were welded using the Short Arc Metal Inert Gas (MIG) welding process with a shielding gas consisting of 97% Argon and 3% CO2 and a 0.023 in. diameter Alloy 625 welding wire. The bulkheads were then welded to the longerons and the keel using the Tungsten Inert Gas (TIG) welding process with a 100% Argon shielding gas and alloy 625 filler wire. The three sections were finally joined using the TIG process. 6. Composite panels manufacturing The composite panels were made by vacuum infusing vinyl ester resin into thin glass fiber skins on each side of a Divinycell foam core. The foam cores were routed to the design dimension and the edges were beveled 45. The foam cores were slightly smaller than the openings in the truss frame, such that the two skins could come together completely where the panels were bonded to the frame. Divinycell H130 foam was used in all deck panels and Divinycell H250 foam was used in all bottom panels. For the side panels, Divinycell H200 foam was used in 14 panels at high shear loaded regions, and Divinycell H160 foam was used in 10 panels at low/no shear load areas (Fig. 6). The manufacturing was as follows. One layer of thin resin distribution medium (RDM) and one layer of peel Side panel with H160 foam Side panels with H200 foam
Deck panels
Side panels with H160 foam
Side panels with H200 foam Side panel with H160 foam
Fig. 5. Small scale ship hull specimen. The stainless steel truss being closed out with composite sandwich panels. Fingers can be seen at the edge of steel surrounding the sandwich panels.
Fig. 6. Arrangements of foam cores of varying density in the hull.
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ply were first placed on a flat steel mold. Two layers of glass fiber for the outer (flat) sandwich skin were then placed on the peel ply. All fibers were laid at ±45 with respect to the edges of the panel. The foam was placed on the fiber reinforcements with the beveled side up. One layer of glass fiber was laid on top of the foam. A reinforcement strip was laid along the four edges of the foam, also with the fibers at ±45. Another layer of peel ply and RDM were laid over the stack. A vacuum bag was placed on top to seal the assembly. The resin was mixed according to the specifications (1.5% methyl ethyl ketone peroxide (MEKP), 0.3% cobalt naphthenate (CoNap) of 6% concentration, and 0.05% dimethylaniline (DMA)), and was degassed prior to the infusion until the majority of the bubbles had vanished (approximately 5 min). The resin was drawn into the mold by vacuum. After the whole part had been infused, the
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resin port was closed off, but the vacuum port was left open until the resin had cured. After the resin had fully cured, the panels were demolded. Sixty panels were made and cut with a waterjet cutter to the required dimensions. 7. Panel bonding The bonding area on the steel truss was sanded with a 150 mm diameter 40 grit sanding disc on a heavy duty angle grinder, and then cleaned with trichloroethylene. The peel ply on the inner side of the composite panels was peeled off immediately prior to bonding and the bonding area on the panels was also cleaned with trichloroethylene. The composite panels were not sanded. SIA E2119A/B adhesive was mixed and applied to the bonding surfaces of both the steel truss and the composite panels. The panels were bonded to the outside of the truss.
2(T1+T2)
Load spreader beam
2839.2 mm T1+T2
T1+T2 403.6 mm
403.6 mm
Load spreader beam
T1
Load spreader beam
T2
T1
T2
T2
T1
T2
T1
Specimen
L1
2032 mm
L2
L1
L2
T1+T2
T1+T2
T (T1+T2) T2 -T2 -(T1+T2)
(T1+3T2)L1
M (T1+T2)L1
(T1+3T2)L1 (T1+T2)L1
Fig. 7. Schematic of the six point bending test showing the loading tree (top) and the resulting shear force (middle) and bending moment (bottom).
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The composite panels were clamped for at least 8 h to achieve handling strength of the adhesive. Full cure was presumably achieved after 72 h at room temperature. 8. Load and test fixture A six point bending test was designed as shown in Fig. 7. This layout enabled the maximum bending moment and the maximum shear load according to the ABS regulations [10] to be reached simultaneously but not at the same location. As shown in Fig. 7, the maximum shear force is F max ¼ T 1 þ T 2
ð4Þ
and the maximum bending moment is M max ¼ T 1 L1 þ T 2 L1 þ T 2 L2 ¼ ðT 1 þ 3T 2 ÞL1 ;
ð5Þ
where T1 and T2 are the forces shown in Fig. 7. Given the F sl maximum shear force F max ¼ F ¼ 35 2 = 41,600 N and the sl = 46,600 N m maximum bending moment M max ¼ M ¼ M 353 (Table 1), and L1 = 508 mm and L2 = 2L1 where L1 is the length of each bay in the 12-bay specimen, the forces can be solved from Eqs. (4) and (5), T 1 ¼ 3F max =2 M max =ð2L1 Þ ¼ 16; 534 N; T 2 ¼ F max =2 þ M max =ð2L1 Þ ¼ 25; 066 N:
ð6Þ ð7Þ
The design load is 2(T1 + T2) = 83,200 N. A single hydraulic jack was used in conjunction with load spreader beams, as shown in Fig. 8. Each of the two shorter spreader beams was made of two C5 · 9 steel channels bolted back to back. The longer spreader beam was made of two C12 · 25 channels bolted together back to back. Loads were introduced to the specimens through eight thin steel loading arms bolted to brackets that were welded on the steel truss directly outside the bulkheads. The specimen was supported by four other arms bolted to floor mounts. The thin arms allowed the hull girder to deform freely with no moment introduced through the supports. 9. Strain gages, LVDT’s and data acquisition system The specimen was instrumented with 192 strain gages. Of these, 110 gages were mounted on the steel truss, 74 on composite panels, and eight on the four support arms. The strain gages mounted on the steel truss were Vishay CEA-06-250UN-350. Of the 110 gages on the steel truss, 30 were mounted on the longerons and 80 on the steel bulkheads. There were six strain gages on each of the five longerons. All gages on the steel longerons were oriented along the length direction of the hull. A total of 20 bulkheads were instrumented with stain gages, with four gages on each bulkhead. The gages on the bulkheads were placed along the length direction of the individual bulkhead members. The composite panels were instrumented with 74 Vishay CEA-06-500UW-350 strain gages. Of these 74 strain gages, 46 were mounted on bottom panels, 16 mounted on side
Fig. 8. Six point bending test fixture with test specimen.
panels and 12 on deck panels. Fig. 9 shows the arrangement and orientation of these strain gages. A total of 33 linear variable differential transformers (LVDT’s) from Macro Sensors were used to measure deformations of the specimen (Fig. 10). Sixteen of them were mounted diagonally on eight highly shear loaded side panels to measure the shear deformation. There were twelve LVDT’s mounted along the length of the hull to measure the axial deformation under bending. The remaining five LVDT’s were mounted vertically from the lab floor to the deck surface to measure the vertical deformation under bending. A CR9000 Measurement and Control System, made by Campbell Scientific Inc., was used for the data acquisition. CR9050 Analog Input Modules and 4WFB350 Terminal Input Modules were used to measure all signals from the strain gages, LVDT’s and the load cell. Test data was recorded by three personal computers.
10. Sagging test As shown in Fig. 8, the ship hull specimen was installed upside down and forces were applied to simulate sagging load. The design load was 83,200 N (18.7 kips). Before load was applied, all bolts between the four supporting arms and the lab floor were disconnected and the
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Strain gages on deck panels (in 0º)
Strain gages on side panels (in 45º)
6 Strain gages on bot tom panels, near keel longeron Strain gages on bottom panels (in 0º) Strain gages on bottom panels (in 45º) Fig. 9. Arrangements of strain gages on composite sandwich panels.
Fig. 10. Some of the LVDT’s on the hybrid ship hull specimen.
specimen was lifted by the hydraulic jack to allow the specimen to hang loosely. The strain gages were then zeroed. When hanging freely, the gravity induced bending moment in the hull is lower than when the hull is standing on the lower supports. The LVDT’s mounted on the hull were also zeroed (but not the five LVDT’s mounted between the hull and the lab floor). The specimen was then lowered until the four support arms fully supported the specimen. The lower supports where then connected. The jack was then further lowered until there was no load on the specimen. The weight of the load tree was still carried by the jack (through the load cell). The load cell and the five vertical LVDT’s mounted between the hull and the lab floor were then zeroed. The strain gages and the LVDT’s on the hull were not re-zeroed at this time. With this setup, the strains recorded when the load is zero are mainly due to the dead weight of the specimen. The sagging test commenced with loading the specimen to 8888 N (2000 lbf) in the first load step. It was then
unloaded and reloaded three times to this level before proceeding to the next load level. The load increment between each load step was 8888 N (2000 lbf), and at each load step the specimen was loaded and unloaded three times before moving to a new load level. Both the load and unload times were 2 min. Each time the specimen was loaded to a higher level, the specimen was carefully checked visually for damage. After the design load 83.2 kN (18.7 kips) was reached, the specimen was carefully checked. No failure was observed in any of the 60 composite sandwich panels or in the bonds between the panels and the steel truss. The test was continued with same load procedure, but was finally terminated when the load reached 112.9 kN (25.4 kips). This is approximately 36% above the design load. There was no indication of any damage in any of the 60 composite sandwich panels or in the adhesive bonds between the panels and the steel truss throughout the test, even though the maximum load was 36% higher than the design load. The test data of load–unload cycles clearly
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showed that there was plastic yielding of the steel truss near the design load and there was substantial yielding and residual deformation of the hull girder at the final load step. As described in [5], the hull was designed for a maximum von Mises stress of 95% of the yield stress of the steel when loaded to the design load.
11. Finite element analysis of the ship hull model Finite element analyses of the ship hull model were performed using the ANSYS finite element package [11]. An FE model of the subscale hull was created and analyzed, using the quasi static loads that were to be applied during the test. Similar to the full scale FE analysis [5], shell elements were used exclusively in the model. Material properties used for this model are given in Table 2. The orthotropic material properties of the glass fiber skins were measured using a vibrations identification method [12]. The dynamic response of a plate is a function of plate geometry, density, boundary conditions and elastic properties. The elastic properties can be determined if the Table 2 Material properties used in finite element analysis of the subscale model Steel
Young’s modulus Poisson’s ratio Density
195 GPa 0.33 8060 kg/m3
Fiber glass
Young’s modulus in 0 direction E11 Young’s modulus in 90 direction E22 Young’s modulus through the thickness E33 Shear modulus G12 Shear modulus G13 Shear modulus G23 Poisson’s ratio m12 Density
23.01 GPa 20.64 GPa 2 GPa 9.96 GPa 7.11 GPa 7.06 GPa 0.275 1800 kg/m3
Young’s modulus of H250 foam Poisson’s ratio of H250 foam Density of H250 foam Young’s modulus of H200 foam Poisson’s ratio of H200 foam Density of H200 foam Young’s modulus of H130 foam Poisson’s ratio of H130 foam Density of H130 foam
0.4 GPa 0.27 250 kg/m3 0.31 GPa 0.29 200 kg/m3 0.175 GPa 0.31 130 kg/m3
Foam
dynamic behavior of the plate is known. A non-destructive test procedure using a combined theoretical and experimental vibration method was applied. A composite plate was made by vacuum infusing vinyl ester resin into three layers of Hexcel 7725 2/2 twill glass fiber weave. The plate was then cut into several rectangular specimens with the dimensions 90 mm · 120 mm. The specimens were hung in light compliant rubber bands in a partial vacuum and excited by hitting them with a ball bearing ball. The partial vacuum reduced the inertia and damping effect of the surrounding air. The sound from the excited plate was recorded with a calibrated modal microphone and a National Instruments data acquisition system. The time response was FFT analyzed and the eigenfrequencies were determined. Finite element models of the test plates were created and analyzed with ANSYS. The OPTRIX multidisciplinary optimization code was used with ANSYS to automatically search for elastic properties that would minimize the difference between experimentally measured and numerically calculated eigenfrequencies. The material properties of the foam cores were taken from the Diab Divinycell H Grade Technical Manual [13]. In the FE model, H250 was used for the bottom panels, H200 for the side panels and H130 for the deck panels. Only linear elastic FE analyses were performed in the present study (Fig. 11). The results from linear elastic FE analysis are compared to the experimentally measured data below. 12. Discussion The strain gages on the steel longerons measured essentially the strain induced by the bending moment. Because of the distribution of the bending moment and the location of the neutral axis, the largest axial strains occurred at the highest bending moment region of the deck longerons. Fig. 12 shows the strain measured from strain gage 21, located on a deck longeron in the region of largest axial strain. The results from the linear elastic FE analyses matched the initial, linear part of the test results quite well. The difference in slope between the FE analysis and the experimentally measured data is around 5%. At the design load, the strain measured by strain gage 21 was around
Fig. 11. Finite element model of the subscale hybrid ship hull under the simulated sagging load of the test.
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Fig. 12. Load–strain curves obtained from testing and linear elastic finite element analysis. Gage 21 was located at the middle of a deck longeron where the bending moment was the highest. The load is the global load applied by the hydraulic jack.
Fig. 14. Load–strain curves obtained from testing and linear elastic finite element analysis. Gage 141 and gage 175 were located on each side of a deck panel in the region of highest bending moment. The load is the global load applied by the hydraulic jack. The differing strains indicate bending of the panel.
Fig. 13. Load–strain curves obtained from testing and linear elastic finite element analysis. Gage 168 was located on one of the side panels where the shear load was the highest. The load is the global load applied by the hydraulic jack.
Throughout the test, there was no indication of any damage in any of the 60 composite sandwich panels or in the adhesive bonds between the panels and the steel truss. The highest strain measured in the sandwich panels was in the side panels located in the highest shear load regions. Fig. 13 shows the load–strain curve from such a panel. The load–strain curve is a straight line, indicating absence of any damage. The linear elastic finite element analysis matched the test results very well. The difference in the slope was less than 8%. In deck panels and bottom panels located in the region with the highest bending moment, there was residual strain after unloading (Fig. 14). The non-linearity seen in these gages started at a strain of approximately 0.1%, while damage in the composites was not expected to start before substantially higher strains. The residual strains were caused by yielding of the steel longerons, rather than damage in the composite panels. In the most highly compression loaded deck panels, strains measured on the outer skins (gage 175) were much higher than strains measure on the inner skins (gage 141) at high loads (Fig. 14), indicating significant bending of the sandwich. This was expected since all panels were highly non-symmetric. The outer skins were flat, whereas the inner skins followed the beveled cores to join the outer skins at the edges of the panels. However, there was no major difference in strain between the inner and outer skins in the shear loaded side panels or the tensile loaded (bottom) panels.
0.2%, which is the yield strain of virgin AL-6XN steel. However, the load and unload curves showed that plastic yielding initiated well before the strain reached 0.2%. Plastic yielding of the deck longerons started when the strain reached approximately 0.1%. One explanation for this lower yield strain of the steel is that the residual stresses from the welding drastically reduced the ‘‘effective’’ yield stress of the steel longeron. This was recently verified by uniaxial tests of welded AL-6XN stainless steel specimens [14]. Each bulkhead/frame was made of five channels. Four strain gages were mounted on each of 20 such channels, located in the most highly shear loaded regions of the hull. Two of these strain gages were mounted on the front and two on the rear of a channel. These 80 strain gages were installed to monitor local bending of the bulkhead. The maximum absolute value of strain in any strain gage on a bulkhead was less than 0.06%, even when the load reached 36% above the design load. There was no non-linearity perceptible in the load–strain curves, indicating that no yielding occurred in any of the bulkheads. The measured strains confirmed that the bulkheads deformed in an ‘‘S’’ shape, as also predicted by the FE analyses.
13. Conclusions The 1:35 scale hybrid ship hull model, consisting of a steel truss closed out with composite sandwich panels, proved to be easy to manufacture, be light, and have a ductile behavior not expected from an all-composite structure. The ductility may not be correctly scaled due to inherent material length parameters (toughness over strength squared, fiber diameter, etc.). The specimen was loaded to 36% above the design load. The maximum von Mises
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stress in the steel was supposed to be limited to approximately 95% of the yield strength at the design load; however, the test showed that yielding started considerably earlier. The joints proved to hold up very well, with no indication of failure throughout the test. Neither was there any damage in the 60 sandwich panels. There was very good agreement between a relatively simple finite element model and the 6 m ship hull specimen tested. The only discrepancy was the low yield strain of the steel in the test specimen. Residual stresses caused by welding of the truss have been identified as the reason [14]. Acknowledgement This work was supported by ONR Grant N00014-03-10597, with Dr. Roshdy Barsoum as Program Manager. References [1] Barsoum R. The best of both worlds: hybrid ship hulls use composites and steel. AMPTIAC Quart 2003;7(3):55–61. [2] Barsoum R. Hybrid ship hull. US Patent 6,386,131; 14 May 2002. [3] Cao J, Grenestedt JL. Design and testing of joints for composite sandwich/steel hybrid ship hulls. Compos Part A: Appl Sci Manuf 2004;35:1091–105. [4] Thompson L, Walls J, Caccese V. Design and analysis of a hybrid composite/metal structural system for underwater lifting bodies.
[5]
[6]
[7]
[8]
[9]
[10]
[11] [12]
[13] [14]
University of Maine Department of Mechanical Engineering. Report No. UM-MACH-RPT-01-08, June 2005. Cao J, Maroun WJ, Grenestedt JL, Steel truss/composite skin hybrid ship hull, Part I: design and analysis, in press, doi:10.1016/ j.compositesa.2006.11.004. Unde´n H, Ridder S-O. Load-introducing armature as component part of a laminated structural element. US Patent 4,673,606; 13 February 1985. Melograna JD, Grenestedt JL. Improving joints between composites and steel using perforations. Compos Part A: Appl Sci Manuf 2002;33(9):1253–61. Dvorak GJ, Zhang J, Canyurt O. Adhesive tongue-and-groove joints for thick composite laminates. Compos Sci Technol 2001;61(9): 1123–42. Grenestedt JL, Melograna JD, Maroun WJ. Adhesive tongue-andgroove joints for thin composite laminates. Compos Part A: Appl Sci Manuf 2003;34(2):119–24. ABS. Guide for building and classing high-speed naval craft 2003, Part 3 Hull construction and equipment, American Bureau of Shipping; 2003. ANSYS, Version 6.1, Ansys Inc., Southpointe, 275 Technology Drive, Canonsburg, PA, 15317; 2002. Hutapea P, Grenestedt JL. Effect of temperature on elastic properties of woven-glass epoxy composites for printed circuit board. J Electron Mater 2003;32(4):221–7. Divinycell Technical Manual H-Grade, DIAB, Box 201, SE-312 22 Laholm, Sweden; 2003. Cao J, Grenestedt JL, Maroun WJ. Testing and analysis of a 6-m steel truss/composite skin hybrid ship hull model. Mar Struct 2006;19:23–32.
Steel truss/composite skin hybrid ship hull. Part II: Manufacturing and sagging testing William J. Maroun a, Jun Cao a, Joachim L. Grenestedt a
b,*
Department of Mechanical Engineering and Mechanics, Lehigh University, 19 Memorial Drive West, Bethlehem, PA 18015, USA b Department of Mechanical Engineering and Mechanics, ATLSS Center Faculty Associate, Lehigh University, 19 Memorial Drive West, Bethlehem, PA 18015, USA Received 10 April 2006; received in revised form 1 November 2006; accepted 3 November 2006
Abstract A six meter subscale steel truss/composite skin hybrid ship hull model was developed based on an optimized design of a full scale model. The subscale model was finite element analyzed, manufactured and tested under sagging loads. It was made of a welded stainless steel truss to which 60 composite sandwich panels were bonded. The model was loaded to 36% above the design load, at which point there was substantial yielding and residual deformation of the steel truss. However, there was no indication of damage in any of the composite sandwich panels, nor in the bonds between the panels and the steel truss. 2006 Elsevier Ltd. All rights reserved. Keywords: Hybrid ship hulls; C. Finite element analysis; D. Mechanical testing; E. Joints/joining
1. Introduction A number of hybrid ship hull concepts have recently been studied with the aim to combine the attributes of both steel and composites [1–4]. The present study concentrated on hybrid ship hulls consisting of a steel truss closed out with very large composite sandwich panels. A short section showing the structural principle is presented in Fig. 1. The steel truss was designed to carry the bending moment, whereas the composite skins were designed mainly to carry shear and water pressure loads. A 142 m ship hull, similar to a destroyer in terms of size, weight and speed, was designed, finite element analyzed and optimized prior to the current research [5]. At present, a simplified model with a 1:35 scale cross-section was developed and analyzed based on the optimized full scale model [5]. This subscale model was subsequently manufactured. Joints between the steel truss and the composite panels were investigated experimentally. A 6*
Corresponding author. Tel.: +1 610 758 4129; fax: +1 610 758 6224. E-mail address: [email protected] (J.L. Grenestedt).
1359-835X/$ - see front matter 2006 Elsevier Ltd. All rights reserved. doi:10.1016/j.compositesa.2006.11.003
point bending test fixture was developed and forces were applied to simulate sagging loads. The FE model of the subscale hybrid hull was analyzed, using the loads applied during the test. Results from FE analyses were compared with the experiment. There are naturally a number of items which could not be correctly scaled, including fiber and fabric geometry, material toughness over strength, bond line thickness, etc. Some efforts aimed at reducing these effects were attempted by using a light fabric with thin tows (few filaments in the yarns), and thin bond lines. 2. Joints between composite panel and steel truss For the hybrid ship hull it is envisioned that the composite panels would first be fabricated and trimmed to size and then attached to the steel truss. Attaching the large panels to the steel truss may most readily be done by adhesive bonding. In order to reduce the severe elastic stiffness mismatch between the steel and the composite adherends, the edge of the steel can be modified to reduce its effective stiffness. Various ways to achieve this have been investigated, including perforating the steel [6,7], cutting fingers in the
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4.5mm Ø3.5mm
2.5mm
2.5mm
Ø3mm 4.5 mm Ø2.5mm
4.5mm Ø2.75mm 25mm Perforation type A
1.5mm
1.5mm
4.5mm
6.4mm
steel edge [8,9], etc. In current research, joints with either perforations or fingers cut in the steel adherend were experimentally investigated. The goal was to develop a joint suitable for the small scale hybrid ship hull that is simple to manufacture but still light and strong. The reasons for developing a small scale joint, rather than developing a full scale joint and scaling it down, were mainly cost and simplicity. Coupon specimens of various joints were made by manufacturing composite sandwich plates and steel adherends and bonding these together. The cured specimens were mechanically tested to failure under tension. A typical specimen loaded in a tensile test machine is shown in Fig. 2. The sandwich plates were made by vacuum infusing vinyl ester resin into thin glass fiber skins on each side of a PVC foam core. The resin used was Derakane 8084 vinyl ester and the reinforcement was Hexcel 7725 2/2 twill glass
6.4mm
Fig. 1. Short section showing the principle of the hybrid hull. Only one sandwich panel, the deck, is shown.
fiber weave. Two layers of the glass fabric were put on a release agent coated steel mold, on top of which a 12.7 mm thick Divinycell H200 foam piece with 45 beveled edges was placed. One layer of glass fiber was laid on top, such that the inner and outer skins came together outside the beveled edge of the foam. A local reinforcing glass fiber strip was laid over the beveled edges on the inner (beveled) skin. The whole assembly was then vacuum bagged to the mold and infused with vinyl ester resin. After demolding, the panels were trimmed with a waterjet cutter. The steel adherends were made of 2.7 mm thick AL6XN stainless steel. One end of the steel adherend was either perforated or cut with fingers, using a waterjet cutter. Two different perforation configurations and three different finger configurations were designed as shown in Fig. 3. In Finger type B, both the tips of the fingers and the roots between the fingers were nominally 0.95 mm wide. The tips and roots were then rounded (R = 0.4 mm). The distance between fingers (center to center) was 4 mm and the fingers were 10 mm long. Finger type C was identical except the fingers were 6 mm long. Instead of cutting individual steel adherend to the required specimen width (25 mm) before bonding to the composite sandwich adherend, a big steel plate cut with the designed shapes of many specimens was grit blasted, solvent cleaned and bonded to the
90mm
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25mm Perforation type B
4mm
4mm
90mm
Fig. 2. Coupon test specimen for joint development loaded in tensile test machine.
Finger type A
Finger type B
25mm
R0.4mm
25mm
6mm
10mm
2.5mm R0.4mm
25mm
R0.75mm
10mm
2.5mm 1.25mm
Finger type C
Fig. 3. Five different designs of the steel adherends: two perforation configurations and three finger configurations.
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The test results also showed that specimens with reinforcement strips on both the inner and the outer skins had considerably lower failure loads than specimens with reinforcement strips only on the inner skin. It appears that the reinforcement strips on the outside created a more unbalanced specimen and thus lead to more bending, which reduced the overall strength. Based on these tests, the tapered finger joint design and the SIA E2119 adhesive were chosen for bonding the composite sandwich panels to the steel truss in the small scale ship hull model. The reinforcement strips would be applied only on the inner side of the sandwich panels.
composite sandwich plate. After cure, the plate was cut into individual 25 mm wide specimens. The overlap of the steel and composite skin was 20 mm. A total of 56 specimens were manufactured and two comparison tests were performed. An MTI modified 44 kN (10,000 lbs) Instron universal test machine was used for all tests. The specimens were tested under quasi static tensile load at a speed of 1 mm/min. The most common failure mode was skin debonding from the foam core some distance from the joint. Five different joints types (Fig. 3) were first compared by using specimens with the same fiber glass skins (outlined above), foam and adhesive (Hysol 9360). The specimens had the following strengths (failure load/width): Perforation type A: 143.0 ± 0.3 N/mm; Perforation type B: 151.1 ± 6.6 N/mm; Finger type A: 154.8 ± 9.9 N/mm; Finger type B: 164.2 ± 10.6 N/mm; Finger type C: 163.2 ± 13.8 N/mm. Specimens with finger type B and finger type C had similar failure loads and were stronger than specimens with other joint designs. Three different adhesives, Hysol 9360, Hysol 9430 and SIA E2119, were then compared by using specimens having the same joint design (Finger type C in Fig. 3), same fiber glass skins and same foam. When manufacturing these specimens, local reinforcement glass fiber strips were applied on both the inner (beveled) and the outer (flat) skins, which differed from the specimens used in the earlier testing. The specimens had the following strengths (failure load/width): specimens using Hysol 9360: 78.9 ± 6.5 N/ mm; specimens using Hysol 9430: 87.1 ± 8.4 N/mm; specimens using SIA E2119: 84.0 ± 6.0 N/mm. Specimens using Hysol 9430 adhesive had the highest failure load, followed by specimens bonded with SIA E2119. Considering the manufacturing convenience and the large amount of adhesive required, SIA E2119 adhesive mixed with a mixing gun was preferred over the hand-mixed Hysol 9430 adhesive.
3. Design of small scale ship hull specimen A 6 m subscale model was developed for structural testing purposes. The cross-section of the model was obtained by scaling the optimized full-sized ship by 1:35. In order to obtain a feasible load distribution using a few discrete loads, the model was made longer than a true scale model would be. The transverse structural dimensions were all scaled by 1:35. If overhead bars are used for quantities of the scale model, and n = 35 is the scale factor, then the scaled loads are M ; n3 F F ¼ 2; n P ¼ P; M¼
ð1Þ ð2Þ ð3Þ
20 ˚
where M = Msl = 2.0 · 109 N m is the maximum slamming induced bending moment, F = Fsl = 5.1 · 107 N is the maximum slamming induced shear force, and P = Psl = 0.29 MPa is the maximum bottom slamming pressure used
Keel longeron Bottom panel
Bottom panel
Side longeron
Side longeron Bulkhead
Side panel
Side panel Bulkhead
Bulkhead Deck longeron
30.6 mm
Bulkhead
24.9 mm
364.3 mm
Bulkhead
Deck longeron
Deck panel
514.3 mm Fig. 4. Cross-section of 1:35 subscale hull model. This picture shows the keel up, as the model was mounted in the test frame during the sagging test.
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Table 1 Dimensions and design loads for the subscale model Specimen dimension
Length of Thickness Thickness skin Thickness Thickness Design bending moment M Design shear load F
1:35 hybrid ship hull of steel in truss of inner (beveled) composite
6.096 m 2 mm 0.25 mm
of outer (flat) composite skin of foam core
0.5 mm 12.7 mm 46,600 N m 41,600 N
for the design of the full scale ship [5]. Elastic moduli were not scaled. The cross-section of the 1:35 scale hull specimen is shown Fig. 4 and dimension of this small scale model and its design loads are listed in Table 1. 4. Small scale ship hull manufacturing A 6 m long ship hull was manufactured. It consisted of a welded non-magnetic AL-6XN stainless steel truss, which was closed out with 60 glass fiber reinforced vinyl ester skin/foam core sandwich panels (Fig. 5). Stainless steel AL-6XN by Allegheny Ludlum Corp., was used to make the steel truss. The stainless steel thickness was 2 mm (which differs from the previous test specimens). The fingers tips and the roots between the fingers were nominally 2.28 mm wide, and then rounded (R = 1 mm). The distance between fingers (center to center) was 8.28 mm and the height of the fingers was 6 mm. The composite panels consisted of glass fiber skins and a PVC foam core. The reinforcement was Hexcel 7725 2/2 twill glass fiber weave.
Divinycell PVC based foam cores with the thickness 12.7 mm were used for all panels. Four different core densities were used: H130, H160, H200, and H250. Higher density foams were used in water pressure loaded bottom panels and highly shear loaded side panels. The exact locations will be explained in a later paragraph. The resin used throughout was Derakane 8084 vinyl ester. SIA E2119 adhesive from Sovereign Specialty Chemicals was used for bonding the sandwich panels to the steel truss. 5. Steel truss manufacturing The stainless steel truss was made by laser and waterjet cutting the 2 mm thick steel. The edges of the steel to be joined to the sandwich panels were cut with fingers (Fig. 5). The 6 m steel truss was made in three separate sections, 1.5 m, 3 m, and 1.5 m long, respectively. In each section, the steel components were first tack welded together. The longerons and the keel were welded using the Short Arc Metal Inert Gas (MIG) welding process with a shielding gas consisting of 97% Argon and 3% CO2 and a 0.023 in. diameter Alloy 625 welding wire. The bulkheads were then welded to the longerons and the keel using the Tungsten Inert Gas (TIG) welding process with a 100% Argon shielding gas and alloy 625 filler wire. The three sections were finally joined using the TIG process. 6. Composite panels manufacturing The composite panels were made by vacuum infusing vinyl ester resin into thin glass fiber skins on each side of a Divinycell foam core. The foam cores were routed to the design dimension and the edges were beveled 45. The foam cores were slightly smaller than the openings in the truss frame, such that the two skins could come together completely where the panels were bonded to the frame. Divinycell H130 foam was used in all deck panels and Divinycell H250 foam was used in all bottom panels. For the side panels, Divinycell H200 foam was used in 14 panels at high shear loaded regions, and Divinycell H160 foam was used in 10 panels at low/no shear load areas (Fig. 6). The manufacturing was as follows. One layer of thin resin distribution medium (RDM) and one layer of peel Side panel with H160 foam Side panels with H200 foam
Deck panels
Side panels with H160 foam
Side panels with H200 foam Side panel with H160 foam
Fig. 5. Small scale ship hull specimen. The stainless steel truss being closed out with composite sandwich panels. Fingers can be seen at the edge of steel surrounding the sandwich panels.
Fig. 6. Arrangements of foam cores of varying density in the hull.
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ply were first placed on a flat steel mold. Two layers of glass fiber for the outer (flat) sandwich skin were then placed on the peel ply. All fibers were laid at ±45 with respect to the edges of the panel. The foam was placed on the fiber reinforcements with the beveled side up. One layer of glass fiber was laid on top of the foam. A reinforcement strip was laid along the four edges of the foam, also with the fibers at ±45. Another layer of peel ply and RDM were laid over the stack. A vacuum bag was placed on top to seal the assembly. The resin was mixed according to the specifications (1.5% methyl ethyl ketone peroxide (MEKP), 0.3% cobalt naphthenate (CoNap) of 6% concentration, and 0.05% dimethylaniline (DMA)), and was degassed prior to the infusion until the majority of the bubbles had vanished (approximately 5 min). The resin was drawn into the mold by vacuum. After the whole part had been infused, the
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resin port was closed off, but the vacuum port was left open until the resin had cured. After the resin had fully cured, the panels were demolded. Sixty panels were made and cut with a waterjet cutter to the required dimensions. 7. Panel bonding The bonding area on the steel truss was sanded with a 150 mm diameter 40 grit sanding disc on a heavy duty angle grinder, and then cleaned with trichloroethylene. The peel ply on the inner side of the composite panels was peeled off immediately prior to bonding and the bonding area on the panels was also cleaned with trichloroethylene. The composite panels were not sanded. SIA E2119A/B adhesive was mixed and applied to the bonding surfaces of both the steel truss and the composite panels. The panels were bonded to the outside of the truss.
2(T1+T2)
Load spreader beam
2839.2 mm T1+T2
T1+T2 403.6 mm
403.6 mm
Load spreader beam
T1
Load spreader beam
T2
T1
T2
T2
T1
T2
T1
Specimen
L1
2032 mm
L2
L1
L2
T1+T2
T1+T2
T (T1+T2) T2 -T2 -(T1+T2)
(T1+3T2)L1
M (T1+T2)L1
(T1+3T2)L1 (T1+T2)L1
Fig. 7. Schematic of the six point bending test showing the loading tree (top) and the resulting shear force (middle) and bending moment (bottom).
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The composite panels were clamped for at least 8 h to achieve handling strength of the adhesive. Full cure was presumably achieved after 72 h at room temperature. 8. Load and test fixture A six point bending test was designed as shown in Fig. 7. This layout enabled the maximum bending moment and the maximum shear load according to the ABS regulations [10] to be reached simultaneously but not at the same location. As shown in Fig. 7, the maximum shear force is F max ¼ T 1 þ T 2
ð4Þ
and the maximum bending moment is M max ¼ T 1 L1 þ T 2 L1 þ T 2 L2 ¼ ðT 1 þ 3T 2 ÞL1 ;
ð5Þ
where T1 and T2 are the forces shown in Fig. 7. Given the F sl maximum shear force F max ¼ F ¼ 35 2 = 41,600 N and the sl = 46,600 N m maximum bending moment M max ¼ M ¼ M 353 (Table 1), and L1 = 508 mm and L2 = 2L1 where L1 is the length of each bay in the 12-bay specimen, the forces can be solved from Eqs. (4) and (5), T 1 ¼ 3F max =2 M max =ð2L1 Þ ¼ 16; 534 N; T 2 ¼ F max =2 þ M max =ð2L1 Þ ¼ 25; 066 N:
ð6Þ ð7Þ
The design load is 2(T1 + T2) = 83,200 N. A single hydraulic jack was used in conjunction with load spreader beams, as shown in Fig. 8. Each of the two shorter spreader beams was made of two C5 · 9 steel channels bolted back to back. The longer spreader beam was made of two C12 · 25 channels bolted together back to back. Loads were introduced to the specimens through eight thin steel loading arms bolted to brackets that were welded on the steel truss directly outside the bulkheads. The specimen was supported by four other arms bolted to floor mounts. The thin arms allowed the hull girder to deform freely with no moment introduced through the supports. 9. Strain gages, LVDT’s and data acquisition system The specimen was instrumented with 192 strain gages. Of these, 110 gages were mounted on the steel truss, 74 on composite panels, and eight on the four support arms. The strain gages mounted on the steel truss were Vishay CEA-06-250UN-350. Of the 110 gages on the steel truss, 30 were mounted on the longerons and 80 on the steel bulkheads. There were six strain gages on each of the five longerons. All gages on the steel longerons were oriented along the length direction of the hull. A total of 20 bulkheads were instrumented with stain gages, with four gages on each bulkhead. The gages on the bulkheads were placed along the length direction of the individual bulkhead members. The composite panels were instrumented with 74 Vishay CEA-06-500UW-350 strain gages. Of these 74 strain gages, 46 were mounted on bottom panels, 16 mounted on side
Fig. 8. Six point bending test fixture with test specimen.
panels and 12 on deck panels. Fig. 9 shows the arrangement and orientation of these strain gages. A total of 33 linear variable differential transformers (LVDT’s) from Macro Sensors were used to measure deformations of the specimen (Fig. 10). Sixteen of them were mounted diagonally on eight highly shear loaded side panels to measure the shear deformation. There were twelve LVDT’s mounted along the length of the hull to measure the axial deformation under bending. The remaining five LVDT’s were mounted vertically from the lab floor to the deck surface to measure the vertical deformation under bending. A CR9000 Measurement and Control System, made by Campbell Scientific Inc., was used for the data acquisition. CR9050 Analog Input Modules and 4WFB350 Terminal Input Modules were used to measure all signals from the strain gages, LVDT’s and the load cell. Test data was recorded by three personal computers.
10. Sagging test As shown in Fig. 8, the ship hull specimen was installed upside down and forces were applied to simulate sagging load. The design load was 83,200 N (18.7 kips). Before load was applied, all bolts between the four supporting arms and the lab floor were disconnected and the
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Strain gages on deck panels (in 0º)
Strain gages on side panels (in 45º)
6 Strain gages on bot tom panels, near keel longeron Strain gages on bottom panels (in 0º) Strain gages on bottom panels (in 45º) Fig. 9. Arrangements of strain gages on composite sandwich panels.
Fig. 10. Some of the LVDT’s on the hybrid ship hull specimen.
specimen was lifted by the hydraulic jack to allow the specimen to hang loosely. The strain gages were then zeroed. When hanging freely, the gravity induced bending moment in the hull is lower than when the hull is standing on the lower supports. The LVDT’s mounted on the hull were also zeroed (but not the five LVDT’s mounted between the hull and the lab floor). The specimen was then lowered until the four support arms fully supported the specimen. The lower supports where then connected. The jack was then further lowered until there was no load on the specimen. The weight of the load tree was still carried by the jack (through the load cell). The load cell and the five vertical LVDT’s mounted between the hull and the lab floor were then zeroed. The strain gages and the LVDT’s on the hull were not re-zeroed at this time. With this setup, the strains recorded when the load is zero are mainly due to the dead weight of the specimen. The sagging test commenced with loading the specimen to 8888 N (2000 lbf) in the first load step. It was then
unloaded and reloaded three times to this level before proceeding to the next load level. The load increment between each load step was 8888 N (2000 lbf), and at each load step the specimen was loaded and unloaded three times before moving to a new load level. Both the load and unload times were 2 min. Each time the specimen was loaded to a higher level, the specimen was carefully checked visually for damage. After the design load 83.2 kN (18.7 kips) was reached, the specimen was carefully checked. No failure was observed in any of the 60 composite sandwich panels or in the bonds between the panels and the steel truss. The test was continued with same load procedure, but was finally terminated when the load reached 112.9 kN (25.4 kips). This is approximately 36% above the design load. There was no indication of any damage in any of the 60 composite sandwich panels or in the adhesive bonds between the panels and the steel truss throughout the test, even though the maximum load was 36% higher than the design load. The test data of load–unload cycles clearly
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showed that there was plastic yielding of the steel truss near the design load and there was substantial yielding and residual deformation of the hull girder at the final load step. As described in [5], the hull was designed for a maximum von Mises stress of 95% of the yield stress of the steel when loaded to the design load.
11. Finite element analysis of the ship hull model Finite element analyses of the ship hull model were performed using the ANSYS finite element package [11]. An FE model of the subscale hull was created and analyzed, using the quasi static loads that were to be applied during the test. Similar to the full scale FE analysis [5], shell elements were used exclusively in the model. Material properties used for this model are given in Table 2. The orthotropic material properties of the glass fiber skins were measured using a vibrations identification method [12]. The dynamic response of a plate is a function of plate geometry, density, boundary conditions and elastic properties. The elastic properties can be determined if the Table 2 Material properties used in finite element analysis of the subscale model Steel
Young’s modulus Poisson’s ratio Density
195 GPa 0.33 8060 kg/m3
Fiber glass
Young’s modulus in 0 direction E11 Young’s modulus in 90 direction E22 Young’s modulus through the thickness E33 Shear modulus G12 Shear modulus G13 Shear modulus G23 Poisson’s ratio m12 Density
23.01 GPa 20.64 GPa 2 GPa 9.96 GPa 7.11 GPa 7.06 GPa 0.275 1800 kg/m3
Young’s modulus of H250 foam Poisson’s ratio of H250 foam Density of H250 foam Young’s modulus of H200 foam Poisson’s ratio of H200 foam Density of H200 foam Young’s modulus of H130 foam Poisson’s ratio of H130 foam Density of H130 foam
0.4 GPa 0.27 250 kg/m3 0.31 GPa 0.29 200 kg/m3 0.175 GPa 0.31 130 kg/m3
Foam
dynamic behavior of the plate is known. A non-destructive test procedure using a combined theoretical and experimental vibration method was applied. A composite plate was made by vacuum infusing vinyl ester resin into three layers of Hexcel 7725 2/2 twill glass fiber weave. The plate was then cut into several rectangular specimens with the dimensions 90 mm · 120 mm. The specimens were hung in light compliant rubber bands in a partial vacuum and excited by hitting them with a ball bearing ball. The partial vacuum reduced the inertia and damping effect of the surrounding air. The sound from the excited plate was recorded with a calibrated modal microphone and a National Instruments data acquisition system. The time response was FFT analyzed and the eigenfrequencies were determined. Finite element models of the test plates were created and analyzed with ANSYS. The OPTRIX multidisciplinary optimization code was used with ANSYS to automatically search for elastic properties that would minimize the difference between experimentally measured and numerically calculated eigenfrequencies. The material properties of the foam cores were taken from the Diab Divinycell H Grade Technical Manual [13]. In the FE model, H250 was used for the bottom panels, H200 for the side panels and H130 for the deck panels. Only linear elastic FE analyses were performed in the present study (Fig. 11). The results from linear elastic FE analysis are compared to the experimentally measured data below. 12. Discussion The strain gages on the steel longerons measured essentially the strain induced by the bending moment. Because of the distribution of the bending moment and the location of the neutral axis, the largest axial strains occurred at the highest bending moment region of the deck longerons. Fig. 12 shows the strain measured from strain gage 21, located on a deck longeron in the region of largest axial strain. The results from the linear elastic FE analyses matched the initial, linear part of the test results quite well. The difference in slope between the FE analysis and the experimentally measured data is around 5%. At the design load, the strain measured by strain gage 21 was around
Fig. 11. Finite element model of the subscale hybrid ship hull under the simulated sagging load of the test.
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Fig. 12. Load–strain curves obtained from testing and linear elastic finite element analysis. Gage 21 was located at the middle of a deck longeron where the bending moment was the highest. The load is the global load applied by the hydraulic jack.
Fig. 14. Load–strain curves obtained from testing and linear elastic finite element analysis. Gage 141 and gage 175 were located on each side of a deck panel in the region of highest bending moment. The load is the global load applied by the hydraulic jack. The differing strains indicate bending of the panel.
Fig. 13. Load–strain curves obtained from testing and linear elastic finite element analysis. Gage 168 was located on one of the side panels where the shear load was the highest. The load is the global load applied by the hydraulic jack.
Throughout the test, there was no indication of any damage in any of the 60 composite sandwich panels or in the adhesive bonds between the panels and the steel truss. The highest strain measured in the sandwich panels was in the side panels located in the highest shear load regions. Fig. 13 shows the load–strain curve from such a panel. The load–strain curve is a straight line, indicating absence of any damage. The linear elastic finite element analysis matched the test results very well. The difference in the slope was less than 8%. In deck panels and bottom panels located in the region with the highest bending moment, there was residual strain after unloading (Fig. 14). The non-linearity seen in these gages started at a strain of approximately 0.1%, while damage in the composites was not expected to start before substantially higher strains. The residual strains were caused by yielding of the steel longerons, rather than damage in the composite panels. In the most highly compression loaded deck panels, strains measured on the outer skins (gage 175) were much higher than strains measure on the inner skins (gage 141) at high loads (Fig. 14), indicating significant bending of the sandwich. This was expected since all panels were highly non-symmetric. The outer skins were flat, whereas the inner skins followed the beveled cores to join the outer skins at the edges of the panels. However, there was no major difference in strain between the inner and outer skins in the shear loaded side panels or the tensile loaded (bottom) panels.
0.2%, which is the yield strain of virgin AL-6XN steel. However, the load and unload curves showed that plastic yielding initiated well before the strain reached 0.2%. Plastic yielding of the deck longerons started when the strain reached approximately 0.1%. One explanation for this lower yield strain of the steel is that the residual stresses from the welding drastically reduced the ‘‘effective’’ yield stress of the steel longeron. This was recently verified by uniaxial tests of welded AL-6XN stainless steel specimens [14]. Each bulkhead/frame was made of five channels. Four strain gages were mounted on each of 20 such channels, located in the most highly shear loaded regions of the hull. Two of these strain gages were mounted on the front and two on the rear of a channel. These 80 strain gages were installed to monitor local bending of the bulkhead. The maximum absolute value of strain in any strain gage on a bulkhead was less than 0.06%, even when the load reached 36% above the design load. There was no non-linearity perceptible in the load–strain curves, indicating that no yielding occurred in any of the bulkheads. The measured strains confirmed that the bulkheads deformed in an ‘‘S’’ shape, as also predicted by the FE analyses.
13. Conclusions The 1:35 scale hybrid ship hull model, consisting of a steel truss closed out with composite sandwich panels, proved to be easy to manufacture, be light, and have a ductile behavior not expected from an all-composite structure. The ductility may not be correctly scaled due to inherent material length parameters (toughness over strength squared, fiber diameter, etc.). The specimen was loaded to 36% above the design load. The maximum von Mises
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stress in the steel was supposed to be limited to approximately 95% of the yield strength at the design load; however, the test showed that yielding started considerably earlier. The joints proved to hold up very well, with no indication of failure throughout the test. Neither was there any damage in the 60 sandwich panels. There was very good agreement between a relatively simple finite element model and the 6 m ship hull specimen tested. The only discrepancy was the low yield strain of the steel in the test specimen. Residual stresses caused by welding of the truss have been identified as the reason [14]. Acknowledgement This work was supported by ONR Grant N00014-03-10597, with Dr. Roshdy Barsoum as Program Manager. References [1] Barsoum R. The best of both worlds: hybrid ship hulls use composites and steel. AMPTIAC Quart 2003;7(3):55–61. [2] Barsoum R. Hybrid ship hull. US Patent 6,386,131; 14 May 2002. [3] Cao J, Grenestedt JL. Design and testing of joints for composite sandwich/steel hybrid ship hulls. Compos Part A: Appl Sci Manuf 2004;35:1091–105. [4] Thompson L, Walls J, Caccese V. Design and analysis of a hybrid composite/metal structural system for underwater lifting bodies.
[5]
[6]
[7]
[8]
[9]
[10]
[11] [12]
[13] [14]
University of Maine Department of Mechanical Engineering. Report No. UM-MACH-RPT-01-08, June 2005. Cao J, Maroun WJ, Grenestedt JL, Steel truss/composite skin hybrid ship hull, Part I: design and analysis, in press, doi:10.1016/ j.compositesa.2006.11.004. Unde´n H, Ridder S-O. Load-introducing armature as component part of a laminated structural element. US Patent 4,673,606; 13 February 1985. Melograna JD, Grenestedt JL. Improving joints between composites and steel using perforations. Compos Part A: Appl Sci Manuf 2002;33(9):1253–61. Dvorak GJ, Zhang J, Canyurt O. Adhesive tongue-and-groove joints for thick composite laminates. Compos Sci Technol 2001;61(9): 1123–42. Grenestedt JL, Melograna JD, Maroun WJ. Adhesive tongue-andgroove joints for thin composite laminates. Compos Part A: Appl Sci Manuf 2003;34(2):119–24. ABS. Guide for building and classing high-speed naval craft 2003, Part 3 Hull construction and equipment, American Bureau of Shipping; 2003. ANSYS, Version 6.1, Ansys Inc., Southpointe, 275 Technology Drive, Canonsburg, PA, 15317; 2002. Hutapea P, Grenestedt JL. Effect of temperature on elastic properties of woven-glass epoxy composites for printed circuit board. J Electron Mater 2003;32(4):221–7. Divinycell Technical Manual H-Grade, DIAB, Box 201, SE-312 22 Laholm, Sweden; 2003. Cao J, Grenestedt JL, Maroun WJ. Testing and analysis of a 6-m steel truss/composite skin hybrid ship hull model. Mar Struct 2006;19:23–32.