Citation
Preliminary Design and Nonlinear Numerical Analysis of an Inflatable Open-Ocean Aquaculture Cage

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Title:
Preliminary Design and Nonlinear Numerical Analysis of an Inflatable Open-Ocean Aquaculture Cage
Creator:
SUHEY, JEFFREY ( Author, Primary )
Copyright Date:
2008

Subjects

Subjects / Keywords:
Bending ( jstor )
Boundary conditions ( jstor )
Diameters ( jstor )
Geometry ( jstor )
Internal pressure ( jstor )
Modeling ( jstor )
Moduli of elasticity ( jstor )
Simulations ( jstor )
Stiffness ( jstor )
Structural deflection ( jstor )

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University of Florida
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University of Florida
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Copyright Jeffrey Suhey. Permission granted to University of Florida to digitize and display this item for non-profit research and educational purposes. Any reuse of this item in excess of fair use or other copyright exemptions requires permission of the copyright holder.
Embargo Date:
8/7/2004

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Full Text












PRELIMINARY DESIGN AND NONLINEAR NUMERICAL ANALYSIS OF AN
INFLATABLE OPEN-OCEAN AQUACULTURE CAGE















By

JEFFREY SUHEY


A THESIS PRESENTED TO THE GRADUATE SCHOOL
OF THE UNIVERSITY OF FLORIDA IN PARTIAL FULFILLMENT
OF THE REQUIREMENTS FOR THE DEGREE OF
MASTER OF SCIENCE

UNIVERSITY OF FLORIDA


2004

































Copyright 2004

by

Jeffrey Suhey

































This work is dedicated to anyone who strives to challenge himself.















ACKNOWLEDGMENTS

I would like to express my gratitude to my advisor, Dr. Christopher Niezrecki, for

his guidance and wisdom. Working on this project has provided me with an opportunity

to advance my education and potential. My sincerest thanks go out to Dr. Nam-Ho Kim

for his enthusiasm and expertise. I would also like to thank Dr. Ashok Kumar for serving

on my defense committee. My thanks go to Dr. Ifju and the members of the

Experimental Stress Analysis Lab for their assistance. Srikant Ranj an, Vann Chesney,

and JeffLeismer were a great help with my ANSYS installation and licensing issues. I

also owe a debt of gratitude to Anne Baumstarck, the Baumstarck family, and Erik

Mueller for providing me with temporary lodgings while in Gainesville. And special

thanks go to God, my family, and all my friends for their continued support and

encouragement.
















TABLE OF CONTENTS

page

A C K N O W L E D G M E N T S ................................................................................................. iv

L IST O F T A B L E S ......... ............................. ......... ... .............. .. vii

LIST OF FIGURES ............. ........ .......... ..... ................... viii

ABSTRACT .............. ..................... .......... .............. xii

CHAPTER

1 IN TR OD U CTION ............................................... .. ......................... ..

1.1 A quaculture Introduction.. .............................................................. ............... 1
1.2 Inflatable Structure Introduction....................................... .......................... 4
1.3 Inflatable Structure Analyses..................... ............................. 12
1.3.1 Modified Conventional Beam Theory Approach............................... 13
1.3.2 N onlinear A approach ............................................................................14
1.3.3 L inear Shell M ethod ............................ ........................ .... ............... 14
1.3.4 Finite Element Analysis Approach........................ ... ............... 15
1.3.5 Selected A approach ............ .... ...................................... ........ .... ......... 16

2 THEORETICAL DEVELOPMENT ........................................ ...................... 17

2.1 Fluid A nalysis........................................17
2.2 Non-wrinkled Inflatable Theory ................ ... ...................... 19
2.3 Case Specific Non-wrinkled Inflatable Theory .......................................... 21
2.3.1 B ending M om ent ............................................. .. ........................... 21
2.3.2 V ertical D election .............................................................................. 22
2.3.3 W rinkle L ength ............ .... .................................... ...... .......................23
2.4 Conclusion ................................. ............................... ........ 25

3 EXPERIMENTAL PROCEDURE AND RESULTS........................ 26

3.1 Experim mental Setup and Procedure.................................... ....................... 27
3.1.1 Test Sam ple Orientation ....................................................................... 27
3.1.2 Sample Shapes and Correction for Fiber Discontinuity ..........................28
3.1.3 Experim ental Procedure ........................................ ........................ 29



v









3 .2 S am ple Sh ap e C h oice ........................................ ............................................32
3.3 D ata A analysis and Results ............................................................................. 33

4 SIMULATION DEVELOPMENT ................................... .................................... 39

4.1 Verification Model Development and Results ............................................. 40
4.1.1 A applied Loading Conditions ........................................... ............... 41
4.1.2 W ater C current D rag ................................................ ......... ............... 42
4.1.3 Element Selection and Properties...........................................................44
4.1.4 V erification M odel Results...................................................................... 48
4.1.5 V erification M odel Conclusion ...................................... ............... 52
4 .2 C age Sim ulation P process ........................................................... .....................52
4.2.1 Flow D esign M odels........................................................ ............... 53
4.2.2 Elem ent C choice .................................... ................. ......... 53
4.2.3 Loading Conditions ............................................................................. 53
4.3 Top-D ow n Flow C age M odel .........................................................................55
4.3.1 A applied L loading C conditions ........................................... .....................56
4 .3.2 B oundary C onditions...................................................................... .. .... 56
4 .4 Side-F low C age M odel .............................................................. .....................57
4.4.1 A applied L loading C conditions ........................................... .....................58
4.4.2 B oundary C onditions........................................................ ............... 58

5 SIM U L A TIO N R E SU L T S ........................................ ...................... .....................63

5.1 Sim ulation Strategy .......................................... ............ ...... ........63
5 .2 F ailu re C criteria .............................................................................. 6 4
5.3 Stress Results ................................................................... .........66
5.4 G eom etry D eform ation R results ........................................ ........................ 67
5.5 Sim ulation D election R results ........................................ ......................... 71
5.5.1 Qualitative D election Results................................................................. 71
5.5.2 Maximum Deflection Results ...................................... ...............73
5.6 C age Size R results .............................. .... ...................... .. ...... .... ..... ...... 74
5.7 V velocity R results ......... .. ................ .. .............. ........ .. ...... 77

6 C O N C L U SIO N ......... ......................................................................... ........ .. ..... .. 78

APPENDIX

A AN SY S SCRIPT COD E .................................................. ............................... 80

B EX PERIM EN TAL D A TA ......................................................................... .... .... 102

L IST O F R E FE R E N C E S ........................................................................ ................... 104

BIOGRAPHICAL SKETCH ........................................................... ........ 106
















LIST OF TABLES


Table pge

3.1 Experimental elastic modulus results for three test sample shapes..........................33

3.2 Summary of experimental and adjusted material properties...............................37

5.1 Manufacturer data for Vinyl-Flow commercial drainage tubing ...........................64

5.2 Maximum cage deflections for the top-flow orientation.............. ............ 74

5.3 Maximum cage deflections for the side-flow orientation .....................................74
















LIST OF FIGURES


Figure pge

1.1 Example of a class 1 cage design (Loverich and Forster, 2000).............................3

1.2 Example of a class 2 cage design (Loverich and Forster, 2000).............................3

1.3 Example of a class 3 cage design (Loverich and Forster, 2000).............................4

1.4 Example of a class 4 cage design (Loverich and Forster, 2000).............................4

1.5 'Millennium Arches' in Stockholm, Sweden: entirely self-supporting inflated
structure. (L indstrand, 2000) ........................................ ............... ............... 6

1.6 'Millennium Arches' alternate view (Lindstrand, 2000) .......................................6

1.7 Section of 'Millennium Arches' before installation (Lindstrand, 2000)....................7

1.8 Hawkmoor self-supporting inflated 'Temprodome': shown with three units linked
together, each with dimensions (6 x 4 x 2.8m). (Hawkmoor, 2003).....................7

1.9 Exterior of 'Court TV' inflated dome. 20 ft tall by 40 ft diameter. For scale, note
the person in the doorway. (Promotional Design Group, 2001) .............................7

1.10 Interior of 'Court TV' dome showing inflated arch columns. (Promotional Design
G ro u p 2 0 0 1) ..............................................................................................................8

1.11 Thin-film inflated torus used as a support structure for optical reflector
(S R S 2 0 0 0 ) ................................................................................................................8

1.12 Inflated thin film with reflective coating to be used as a light-weight deployable
antenna, 5 m eter diam eter (SRS, 2000)................................... ....................... 9

1.13 Inflatable Antenna Experiemnt (IAE): 14 meter diameter deployed antenna with
inflated support beams (Domheim, 1999) ....... ........ ................................... 9

1.14 IAE Antenna during inflation. Stowed volume is one-tenth of deployed volume.
(Domheim, 1999) .......................... ....................... 10

1.15 IAE Reflector (14m diameter). For scale, note the person circled in the upper
right (L'Garde, 1996) .............. .. ........ .. ...... ......... ......... 10









1.16 Goodyear Corporation inflatable truss radar antenna (Jenkins, 2001)...................11

1.17 L'Garde inflatable solar array (Jenkins, 2001).................................... ............... 11

1.18 ILC Dover, Inc. modular split blanket solar array. Shown deployed and stowed
(Jen k in s, 2 0 0 1) ......... ................................................ ............................... 1 1

1.19 Planar deployment solar shade (Jenkins, 2001) ...............................................12

1.20 Inflatable Goodyear Inflatoplane (Jenkins, 2001)..............................................12

2 .1 F low on a vertical tube ............................................................................. ..... .......18

2.2 Flow on an inclined tube ...................................................................... 18

2.4 Drag loading assumed as a distributed force .........................................................22

2.5 Free body diagram of beam section .............................................. ............... 22

3.1 Vinyl-Flow internal fiber layer orientations...................... ..................... 26

3.2 Manufacturer information for Kuriyama Vinyl-Flow series.................................27

3.3 Test sample orientations with respect to continuous tubing ..................................27

3.4 Sample shape B (shown flat and folded)........ ...................... .............. 29

3.5 Sam ple shape C .......................................................................29

3.6 Test equipment: MTI screw-drive tensile test machine (MTI, 2004) ...................31

3.7 Test equipment: Interface 1220AF-25k load cell............ .............................31

3.8 Test equipment: Curtis 30k self-cinching grip..................................................... 31

3.9 Example of sample locked in grips before pull test ...........................................32

3.10 Experimental data: five hoop orientation pull samples.......................................35

3.11 Experimental data: five longitudinal orientation pull samples.............................35

3.12 Experimental data: Five hoop-orientation pull samples grouped together ............36

3.13 Experimental data: Five longitudinal-orientation pull samples grouped together..36

3.14 Tested sam ple w ith grip slip m arks................................. ........................ .. ......... 38

4.1 Cage prototype inflated to -12ft diameter, 9ft height............................................39

4.2 Cage prototype deflated to ~70cm ........................................ ....................... 40









4.3 Tubing union to PV C connector........................................ ........................... 40

4.5 Detail of boundary conditions for verification model...........................................41

4.6 Steps in representing internal pressure as pre-tension forces................................42

4.7 Representative projected area around one surface node ........................................43

4.9 Verification model parameter study: EFS effect on beam deflection ...................46

4.10 Verification model parameter study: Added stiffness effect on beam deflection ...46

4.11 Verification model parameter study: Added stiffness effect on longitudinal beam
store ss ............ ............ .. .......... ...... .............................................. . 4 7

4.12 Verification model parameter study: Poisson's ratio effect on beam deflection ....47

4.13 Verification model parameter study: Poisson's ratio effect on longitudinal beam
store ss ............ ............ .. .......... ...... .............................................. . 4 8

4.14 Deflection results for the verification model with length of 5m...........................50

4.15 Deflection results for the verification model with length of 10m length.................50

4.16 Deflection results for verification model with 0.5 m/s flow velocity ....................51

4.17 Deviation in deflection results for verification model with 0.5 m/s flow velocity ..51

4.18 Longitudinal bending stress results for verification model with 0.5 m/s flow
v elo city ......... ..... ............. .................................... ...........................52

4.19 Two cage simulation flow orientations ........................................ ............... 53

4.20 Control volume analysis around a rigid 45'-connector containing pressurized
internal fluid ............. .......... .... ................... ............ ............ 54

4.21 Free body diagram of rigid cap with algebraically summed pretension force
components ............... ..... .. .......... ........ .. ........ ........ ......... 55

4.22 Force components applied to finite element model of rigid connector....................55

4.23 Top-flow boundary condition locations ....................................... ............... 57

4.24 Cage center tube components ........... .. ......... ................... 59

4.25 Expanded view of representative internal cable connections.............................60

4.26 Expanded View of representative mooring cable connections .............................60









4.27 D im tensions of anchor location........................................... .......................... 61

4.28 Side-flow applied loading locations...................................................................... 61

4.29 Side-flow boundary condition locations ...................................... ............... 62

4.30 Expanded view of center tube boundary conditions .............................................62

5.3 Maximum stresses in each simulation and ultimate material strength...................67

5.4 Top-Flow deformed geometry at the wrinkle point: View 1...............................68

5.5 Top-Flow deformed geometry at the wrinkle point: View 2..................................68

5.6 Top-Flow deformed geometry at the wrinkle point: View 3................................69

5.7 Side-Flow deformed geometry at the wrinkle point: View 1 ...............................69

5.8 Side-Flow deformed geometry at the wrinkle point: View 2...............................70

5.9 Representative rotational effect occurring in side-flow simulation .......................70

5.10 Cage x-diameter net deflection for VF-800 tested at flow speed of 1 knot,
top -flow ......... ..... ............. ..................................... ...........................72

5.11 Maximum z-deflection for VF-800 cage tested at flow speed of 1 knot,
top -flow ......... ..... ............. ..................................... ...........................72

5.12 Cage y-diameter net deflection for VF-500 tested at flow speed of 1 knot,
sid e -flo w ...................................................................... 7 3

5.13 Top-flow maximum cage diameters for each tubing material grouped by internal
pressure ................ ..................................... ........................... 75

5.14 Side-flow maximum cage diameters for each tubing material grouped by internal
pressure ................ ..................................... ........................... 76

5.15 Summary of maximum possible cage size for each material in two flow cases
subject to one knot flow velocity VF Series number corresponds to diameter,
working pressure, and thickness in Table 5.1 .................................. ............... 76















Abstract of Thesis Presented to the Graduate School
of the University of Florida in Partial Fulfillment of the
Requirements for the Degree of Master of Science

PRELIMINARY DESIGN AND NONLINEAR ANALYSIS OF AN OPEN-OCEAN
AQUACULTURE CAGE

By

Jeffrey Suhey

August 2004

Chair: Dr. Christopher Niezrecki
Major Department: Mechanical and Aerospace Engineering

A nonlinear finite element analysis is used to predict the static performance of a

novel aquaculture cage constructed from inflated-beam members. Simulations are

performed to determine the cage maximum deflections, stresses in the inflated

components, and ultimately the maximum possible cage size. Experimental tests are

performed on a commercially available anisotropic fabric material to determine

longitudinal and hoop elastic moduli.

Fluid drag forces on the cage are applied to the inflated members as a constant

distributed force that is dependent on the flow speed of the current. Simulations are

performed up to 1.6 knots of water current. Two flow cases based on the orientation of

the cage with respect to water current are considered. Drag on the surrounding cage nets

is excluded from the preliminary design analysis. Within this study, cage members were

pressurized from 30 to 80 psi, depending on the material.









Important results include successfully merging a linear and nonlinear element to

represent tension-only behavior of an inflated fabric tube. Verification was performed by

comparing a simply-supported distributed-load inflated beam finite element model with a

modified traditional beam analysis. Maximum possible cage sizes are found at constant

flow velocity for different materials with varying internal pressure, wall thickness, and

tube diameter. Effects of changing flow velocity are found for a specific material case in

both flow directions. The results indicate that the largest sized cage for a 1 knot (0.51

m/s) current has a cage diameter of 39 m for the material used in this study.














CHAPTER 1
INTRODUCTION

The current trend in new inflated structure applications lends itself to studying an

underwater application. In this case, a preliminary engineering design analysis of an

open-ocean aquaculture cage is presented. This chapter provides a brief introduction to

aquaculture and describes recent developments in inflatable structures.

1.1 Aquaculture Introduction

Aquaculture can be described using the following definition: "Aquaculture is

defined as the propagation and rearing of aquatic organisms in controlled or selected

aquatic environments for any commercial, recreational, or public purpose" (Department

of Commerce, 2002). Aquaculture is an agricultural approach that provides an alternative

to the other two common methods of obtaining fish; commercial fishing and fish farming

with large permanent structures in the ocean or on land (Loverich and Forster, 2000).

"Aquaculture is expected to play an increasingly important role in meeting the global

demand for fisheries products as the world population continues to expand and fish

stocks approach their biological limits" (Fredriksson et al., 2000). The problem is

making the use of aquaculture cages economical. A growing need for improved cage

designs will hopefully stimulate engineering research to overcome the obstacles now

facing aquaculture.

Several factors comprise aquaculture cage design. Costs of manufacturing,

installation and maintenance, cage stability, durability, and predictability, all play an

important role in the design process. One aspect specifically important to aquaculture









cage design is the divergence volume. This can be defined as a maximum potential

geometric deformation of the cage. Large divergence volumes can cause high fish

mortality rates and decrease diver safety. These problems can lead to large economic

drawbacks, making the divergence volume a key factor on which to base a design. A

variety of cages have been designed to accommodate individual design advantages and

disadvantages. These designs have been grouped into four classes that are in current use

throughout the aquaculture community.

The class 1 cage is the most common in use today and is based on forces due to

gravity; weight and buoyancy. The cage typically comprises a floating frame with a

weighted net hanging below. This simple design has several problems, including

unpredictable failures and high fish mortality. The major disadvantage to this cage

design is the large divergence volume. Figure 1.1 shows a typical class 1 cage design.

The class 2 cage has anchored rigid truss members used to hold the shape of the

netting. Due to the taught netting, there is a significant improvement in the divergence

volume compared to the class 1 cage. Similarly, class 3 cages maintain taught netting for

this same improvement. Class 3 cages generally have a central truss with a rigid ring

connected by ropes and covered in netting. This design allows the cage to maintain its

shape without relying on anchors. The deformation of the class 3 cage is governed by the

rigidity of its components. It achieves a smaller divergence volume than the class 1 and 2

by providing more supports over the span of the net. Figures 1.2 and 1.3 show examples

of typical class 2 and 3 cages, respectively.

The Class 4 cage is a completely rigid self-supported frame covered in netting.

Although seemingly ideal due to the very low divergence volume, the costs of material,







3


construction, and installation labor are much higher than of the previous three classes of

cage design. Figure 1.4 shows an example of a typical class 4 cage design.






.Figue E a c 1 c i i a
:,. ?,, :.... ...... : ", : ... I.; .1 .. ,, .. :.,


^ 1~ .,'. I 1




Figure 1.1 Example of a class 1 cage design (Loverich and Forster, 2000)


Figure 1.2 Example of a class 2 cage design (Loverich and Forster, 2000)





















I.- .
N


Figure 1.3 Example of a class 3 cage design (Loverich and Forster, 2000)


Figure 1.4 Example of a class 4 cage design (Loverich and Forster, 2000)

1.2 Inflatable Structure Introduction

Several applications exist for inflatable structures. Applications for inflatable

structures have been used successfully in the past. Some familiar applications that are

still in use include vehicle tires, surface watercraft (inflated boats, pontoons, hovercraft,

etc), and aircraft (blimps, balloons, etc).

Within the past few decades, new inflatable applications are being explored and

gaining acceptance. Inflatable land structures are completely self-supported, requiring no

solid structural members. These are currently in use in military and architectural designs

for tents, hangars, roofs, and small buildings. These applications take advantage of the


^ ^-.









convenience of inflatable structures to save assembly time and travel weight. "...today,

it's well within the bounds of possibility to find a 300 ft long building housing a concert

venue, exhibition hall, museum, or space-age science display, with not a solid supporting

member in sight" (Lindstrand, 2000). Figures 1.5 through 1.10 show examples of

inflated land structures, with no rigid support members, supported only by internal

pressure. "The [Millennium Arches] whole building is 328ft(100m) long, 59ft(18m)

wide, and 56ft(17m) high. The main arches are 38ft(11.5m) high, and 59ft(18m) wide

while the end arches are 41ft(12.5m) long and the center arch is 164ft(50m) long"

(Lindstrand, 2000).

The aerospace field is working with inflatables in gossamer structures for space

applications such as sunshields, antennas, solar sails, habitats, and structural booms.

Space applications take advantage of several features of inflatable structures. "They

[inflatable structures] offer large potential reductions in stowed volume, cost, and often

weight.... Giant space structures such as 1,000-ft. antennas or solar sails may not even be

possible with mechanical deployment, but may be doable with inflatable design"

(Domheim, 1999). Figures 1.11 through 1.15 show examples of space applications of

inflatable structures.




























A1N W T.-. "2Ki,
Figure 1.5 'Millennium Arches' in Stockholm, Sweden: entirely self-supporting inflated
structure. (Lindstrand, 2000)


Figure 1.6 'Millennium Arches' alternate view (Lindstrand, 2000)























Figure 1.7 Section of 'Millennium Arches' before installation (Lindstrand, 2000)


.'... .I...
,n"


Figure 1.8 Hawkmoor self-supporting inflated 'Temprodome': shown with three units
linked together, each with dimensions (6 x 4 x 2.8m). (Hawkmoor, 2003)


Figure 1.9 Exterior of 'Court TV' inflated dome. 20 ft tall by 40 ft diameter. For scale,
note the person in the doorway. (Promotional Design Group, 2001)


'..'.'.'. .".



















Figure 1.10 Interior of 'Court TV' dome showing inflated arch columns. (Promotional
Design Group, 2001)


Figure 1.11 Thin-film inflated torus used as a support structure for optical reflector (SRS,
2000)





























Figure 1.12 Inflated thin film with reflective coating to be used as a light-weight
deployable antenna, 5 meter diameter (SRS, 2000)


Figure 1.13 Inflatable Antenna Experiemnt (IAE): 14 meter diameter deployed antenna
with inflated support beams (Domheim, 1999)































Figure 1.14 IAE Antenna during inflation.


Stowed volume is one-tenth of deployed
(Dornheim, 1999)


Figure 1.15 IAE Reflector (14m diameter). For scale, note the person circled in the upper
right (L'Garde, 1996)






















Figure 1.16 Goodyear Corporation inflatable truss radar antenna (Jenkins, 2001)


Figure 1.17 L'Garde inflatable solar array (Jenkins, 2001)


Figure 1.18 ILC Dover, Inc. modular split blanket solar array. Shown deployed and
stowed (Jenkins, 2001)























Figure 1.19 Planar deployment solar shade (Jenkins, 2001)











Figure 1.20 Inflatable Goodyear Inflatoplane (Jenkins, 2001)

1.3 Inflatable Structure Analyses

This section describes various forms of modeling inflatable structures. A specific

emphasis is placed on determining deflection and internal stress behavior of inflatable

structures subject to cantilever bending. Also of interest is analysis of the application of

fabrics for the inflated material. The basic concept of an inflatable structure is a closed

end pressure vessel subject to some applied loading. The internal pressure translates to a

force applied over the capped end in the axial direction. This force then creates an axial

stress that is able to resist bending and axial stress caused by the applied loading

conditions. Chapter 2 more fully explains the derivation of the inflatable theory.









1.3.1 Modified Conventional Beam Theory Approach

Experimental data has shown that inflatable beams do not behave according to

traditional beam theory. With modification to the traditional beam theory, an accurate

model can be made. One study analyzed inflatable beam deflections and stresses for

loads between incipient buckling, where bending stress equals axial stress due to

pressure, and final collapse. It was shown that initially bending occurred with no

wrinkling. When wrinkling occurred, a slack region and a taught region were present. A

formulation for the stress was developed showing zero axial stress in the slack region and

a portion of the max stress carried in the taught region. By equating expressions for

maximum axial stress, a formulation for the relationship between internal pressure and

applied tip force was determined. The curvature was determined and integrated twice to

find the deflection due to bending. The deflection due to shear was also determined

noting that it is negligible when the length of the beam is much larger than the radius of

the cross section. The total deflection was then determined as the algebraic sum of the

deflections due to bending and shear. Similar results were obtained for a beam subject to

distributed load bending (Comer & Levy, 1963). This analysis shows when wrinkles

begin to form and where they form. The wrinkled membrane is shown to have some

rigidity. The effect of the wrinkling on the bending strength of the beam is determined.

Another study applies the modified traditional beam theory approach from Comer

and Levy for fabric materials. An experimentally determined modulus of membrane of

the fabric material was used in place of Young's modulus. This modulus of membrane is

defined as the slope of the experimental stress resultant vs. engineering strain plot, where

stress resultant is the force per width of a pull-test sample. Using an approach similar to

that of Comer and Levy, the curvature was integrated numerically to determine tip









deflection behavior. Experimental results agreed for aspect ratios of l/d>6 (Main et al.,

1994).

1.3.2 Nonlinear Approach

Since inflatable beams behave nonlinearly, a nonlinear approach, although

complex, is the most logical method. Douglas uses the theory of incremental

deformations to find the Cauchy-Green deformation tensors which can be related to

determine the Lagrangian and Eulerian Strains. These lead to a stress-deformation

relation. This is all then applied to an inflated beam incorporating the pressure in as a

stress. The variation in beam stiffness is plotted as a function of changing internal

pressure. Unlike a linear approach, this thorough method takes into account the changes

in geometry and changes in material properties (Douglas, 1968).

Another study included an explanation of the geometry of fabrics and the

interaction between fibers. The general relationship between stress and strain for a fabric

material is ultimately shown to be a nonlinear trend (Bulson, 1973).

1.3.3 Linear Shell Method

Another method for analyzing inflatable structures is to use a linear shell approach.

One study develops a free body diagram of a width of flexed strip of the loaded material.

Static equilibrium of forces is then applied, including the applied load and the skin

tensions on the width of the material. First, a vertical force summation is made with the

application of a small angle approximation. Then, a horizontal force summation is

developed and combined with the vertical summation canceling the skin tension terms.

The result is an expression analogous to the equation for shearing angular deflection of a

solid beam. By comparing these two equations, the internal pressure is found to be









analogous to an effective shear modulus. This is used to determine the effect of pressure

on the shearing deflection (Topping, 1964).

Another source applies continuum mechanics to membranes to obtain several

general relations. It defines the kinematics, deformation, strain, strain rate, and stress for

membranes. Balance laws are used to derive general mass balance, momentum balance,

and energy balance relations (Jenkins, 2000, pp. 49-64). To model wrinkling in beams, a

tension field model is developed with an approach for application to linear finite element

method (Jenkins, 2000, pp 103-105).

1.3.4 Finite Element Analysis Approach

Another approach is to apply finite element analysis to inflated structures. Finite

element analysis can yield accurate results for problems with complex analytical

solutions. One study compares results from the modified traditional beam theory with

finite element results. An assumption of Brazier's effect is used here. This states that as

the tube deforms due to bending and the cross-section becomes flatter, the bending

stiffness of the entire beam decreases. This is most likely due to geometry changes that

affect the area moment of inertia.

This study formulates finite elements of geometrically nonlinear motion of

membrane. These allow for changes in length of the elements to account for stretching of

the material in the ends but not in the cylinder. Triangle elements are used to model the

circular end caps and quad elements are used to model the cylindrical tubing. A pressure

relationship is used to apply a force to each element. FEA results match theoretical

results from the modified traditional beam approach.

An analysis of tubes containing multiple pockets is conducted as well. The

behavior of these multi-cellular inflated beams is shown graphically. By increasing the









number of cells, the pressure can be varied throughout the length of the tube and thereby

optimize the design (Sakamoto and Natori, 2001).

1.3.5 Selected Approach

The modified traditional beam theory will be used in conjunction with a nonlinear

finite element model for future analysis. The finite element approach solves complex

geometries and provides a visualization of deformed behavior. The modified traditional

beam theory will be used to verify the behavior of a simple finite element model.

Following verification, the finite element approach will then be used to model complex

cage geometries and interactions. Further theoretical development of the modified

traditional beam theory is presented in Chapter 2 and the finite element approach is

described in Chapter 4.















CHAPTER 2
THEORETICAL DEVELOPMENT

The development of the theory applied to analytical solutions used in later chapters

is described in Chapter 2. A standard fluid flow analysis is first presented, followed by

the traditional beam method applied to inflatable theory, which is chosen from the

methods presented in Chapter 1. The theory presented focuses on inflated beams before

and at the point of wrinkling. The model used for derivation is an internally pressurized

circular-cylindrical closed-end fabric tube. The internal pressure acting over the surface

area of the end caps translates to pretension forces in the fabric tube. The essence of the

inflatable theory is that bending causes a compressive stress on the underside of the tube

opposing the tensile stress caused by the pretension forces. Wrinkling occurs when the

compressive portion of bending stress exceeds the pretension stress. In some cases, stress

resultant, c*, is used to replace axial stress, c, throughout the analysis, as described in

Main's theory, to more accurately model fabric behavior (Main et al., 1994). This also

leads to using a resultant elastic modulus, E*, with units of force per unit length.

2.1 Fluid Analysis

The fluid analysis is used to determine the total drag force acting on a cylinder for a

given external flow condition. The following assumptions are made for the fluid

analysis:

* Smooth cylinder surface

* Constant temperature fluid and material properties at 200C






18

* Steady state flow

* Constant cross-sectional area

In this case, the drag force is caused by friction of fluid current flowing over the

tube. The drag force is a function of tube cross-sectional area, A, drag coefficient, CD,

external fluid density, p, and external flow velocity, V. Equation 2.1 defines the drag

force, Fdrag, on a cylinder based on applicable assumptions.

Fdr = CApV2 (2.1)
2

The cross-sectional area perpendicular to the flow direction is defined, A = Dh, with tube

diameter, D, and cross-sectional length perpendicular to flow, h. Figures 2.1 and 2.2

define h for a vertical tube and a tube inclined by angle, p, from the horizontal,

respectively.




Flow Direction h=L
1



Figure 2.1 Flow on a vertical tube



Flow Direction h = L sin (p

---- ^ /^


Figure 2.2 Flow on an inclined tube









The coefficient of drag, CD, is determined using the dimensionless parameter, Re, and the

approximated range taken from documented experimental data. Equations 2.2 and 2.3

show the Reynold's number relation and the approximated value of CD based on the

expected flow velocity range.


Re pVD (2.2)


CD =1, for 10,000 < Re < 200,000 (2.3)

D represents the diameter of the tube. Fluid properties of the flowing current including

density and absolute viscosity, [t, are determined from documented values at a

temperature of 200 C (Fox & McDonald, 1999).

2.2 Non-wrinkled Inflatable Theory

Figure 2.3 shows the assumed stress distribution for a non-wrinkled inflated beam

subject to a bending moment.


Urn






-0


Figure 2.3 Stress distribution in tube subject to bending moment

Comer & Levy (1963) defined the stress distribution in this unwrinkled inflated circular-

cylindrical beam using Equation 2.4.

co (1 + cos ) O-m(1- cos 0)
S= (2.4)
2 2









As shown in Figure 2.3, Gm and ao are the maximum and minimum longitudinal stresses

in the circumferential stress distribution, respectively. Equation 2.4 can be rewritten as in

Equation 2.5, by diving through by the material thickness. Stress, C, is replaced with

stress resultant c*.

( (1+ cos0) o (1 cos))
a= + (2.5)
2 2

Equation 2.6 describes the bending moment, M(x), as a product of the tensile forces,

F(O), and the distance, h(O), from the neutral axis of bending. Force is then replaced with

the product of stress and area. The area is defined by integrating the fabric thickness, t,

over the tube circumference using tube radius, r, and angle, 0, from 0427n.


M(x) = F(O)h(O) = (O)h(O)rt dO (2.6)
8=0 o

Defining height, h(O) = r cosO, and stress resultant, u (0) = c(O)t, Equation 2.6 can be

rewritten using symmetry as Equation 2.7.


M(x)= -2J *r2 cos 0dO (2.7)
0

Substituting Equation 2.5 into Equation 2.7 and integrating yields Equation 2.8.

Sx 2M((x)28)
&,m O 0 (2.8)

According to Main (1994), the unwrinkled curvature, pc, is given by Equation 2.9, where

E* is the resultant elastic modulus.


(2.9)
Pc 2rE*









Substituting Equation 2.8 into Equation 2.9 and approximating curvature as d y/dx2,

where x and y are coordinates, yields Equation 2.10.

1 M(x) d2y
= (2.10)
pc itrE* dx2

This leads to the final relationship, Equation 2.11, which corresponds to the traditional

beam bending relation with I* representing the resultant area moment of inertia. Note the

units of resultant area moment of inertia are cubic which corresponds to using the stress

resultant, c*, and resultant elastic modulus, E*.

d y M(x) I*
d2y M (x) ,where I = r' by association (2.11)
dx2 E I

This differential equation describes the bending behavior of an inflated circular tube. It

will be integrated to determine deflection, y, for a specific set of boundary conditions.

This theoretical deflection solution is then used to verify simulation deflection results in

Chapter 4.

2.3 Case Specific Non-wrinkled Inflatable Theory

2.3.1 Bending Moment

Figure 2.4 shows the assumed drag force loading, w, and model boundary

conditions, with L representing the tube length. The specific bending moment, M(x), for

this situation is derived from free body diagrams of a cut section of the beam as shown in

Figure 2.5. Shear, V, and moment, M are shown as reaction forces. Variable x is used as

a local coordinate along the tube length. Equations 2.12 and 2.13 show the sum of

moments about point A assuming static equilibrium.









Fdrag
L


Internal pressure, P





Figure 2.4 Drag loading assumed as a distributed force


x x
2 2



wx
M
tA

4 V

wL
2

Figure 2.5 Free body diagram of beam section

+ MA = wL )x + (wx) +M = 0 (2.12)


M(x) = w(Lx- x2) (2.13)
2

2.3.2 Vertical Deflection

To solve for beam deflections, the specific moment equation, Equation 2.13, is

substituted into Equation 2.11 and integrated twice, yielding Equations 2.14 through 2.16.

d2y w
d y (Lx x2) (2.14)
dxC2 2E* mr3










dy w Lx2
dx 2E*Ir3 2

w Lx3 x4
2E*zr3 6 12

With boundary conditions at x = 0, y = 0, and at x

and C2, are found to be Equations 2.17 and 2.18

C, =0


+ C )


(2.15)



(2.16)


+ C C 2)


L, y = 0, constants of integration, Ci



(2.17)


(2.18)


Substituting Equations 2.17 and 2.18 into Equation 2.16 yields the vertical deflection

solution shown as Equation 2.19.


w Lx3 X4
y2E 6 12
2E* r3 6 12


SL3
12


(2.19)


2.3.3 Wrinkle Length

Wrinkle length is the length of the tube at which wrinkling first occurs. This

length depends on the moment necessary to initiate wrinkling. Comer and Levy first

determine the moment necessary to initiate wrinkling by equating expressions for om

(Comer & Levy, 1963). Equation 2.20 describes a force balance between the internal

pressure acting over the area of the circular end-plate and the longitudinal stress

integrated over the circumference of the tube.


pr2 = 2 o-*r dO


(2.20)


Substituting Equation 2.5 into Equation 2.20 and integrating yields Equation 2.21, the

first expression for om.









zpr(1 + cosO )
am co= (2.21)
2t[sin 00, + (r 0,)cos0 ]

Substituting Equation 2.5 into the balance of moments, Equation 2.7, yields the second

expression for om.

2M(1 + cos 0 )
"tr2 (2r 200 + sin 200 )

Combining Equations 2.21 and 2.22 by eliminating om and setting 00 = 0, yields Equation

2.23, the moment necessary to initiate wrinkling.


M = (2.23)
2

Equating the moment to initiate wrinkling, Equation 2.23, with the value of the maximum

moment in Equation 2.13 yields Equation 2.24.

L wL2 pair3
max M( )-w (2.24)
2 8 2

Solving for L and setting the load, w = F/L, gives the general expression for wrinkle

length, Equation 2.25. The wrinkle length is the length at which wrinkling first occurs in

an inflated tube subject to bending, for a given pressure and flow rate.


Lwr 4pR (2.25)
F

As expected, the wrinkle length will be reduced for high loading force, F, and

proportional to the internal pressure and the tube radius. Tube radius, r, is the largest

contributing factor to wrinkle length for these conditions.









For this specific application, the loading force, F, is the drag force, Fdrag. Since

drag force is a function of tube length, Equation 2.25 is further expanded using Equations

2.1 and 2.3 to obtain Equation 2.26.


Lw pr2 (2.26)
SpV2

Equation 2.26 now shows the wrinkle length as a function of the core contributing

variables for this application; internal pressure, tube radius, and external fluid velocity.

The wrinkle lengths in Equations 2.25 and 2.26 also depend on the boundary conditions

and are derived for the simply-supported conditions shown in Figure 2.4.

2.4 Conclusion

The equations derived in this chapter provide a theory for an inflated beam subject to

bending. The solution for vertical deflection, Equation 2.19, is used to verify the finite

element model developed in Chapters 4 and 5. The wrinkle length is used to verify the

final results of a verification model. Verifying a simple model is necessary to justify

extending the computer analysis to more complex geometries, not easily solved with

traditional analysis.















CHAPTER 3
EXPERIMENTAL PROCEDURE AND RESULTS

The setup, procedure, and results of the experiments performed to determine the

orthotropic material properties of a specific commercial drainage tubing is presented in

chapter 3. The Kuriyama Vinyl-Flow series is a composite comprised of a combination

of molded PVC layers and synthetic fibers. One fiber layer runs in the longitudinal

direction while two other layers are oriented at approximately +30 and -30O from the

longitudinal direction as shown in Figure 3.1. Figure 3.2 shows specifications given by

the Kuriyama for the different size models of Vinyl Flow. For this application, the

desired material properties are the elastic modulus and Poisson's ratio. These material

properties are determined experimentally and to be used in a numerical simulation as

described in Chapter 4.

+30



Longitudinal Direction



-30

Figure 3.1 Vinyl-Flow internal fiber layer orientations

















* P .*eri =l ip .,iS.e ihe .] 5:rnar.ei rnis featuring a bal-
anced polyester yarn spiral wrap, longitudinal strength
member, and homogeneous PVC tube and cover.
* Use for general purpose water discharge applications -
ideal for use as a drip irrigation supply line.
* Smooth tube provides low friction loss (see charts on
opposite side).
* Lays ; i ;iqr,I i n r ir rnr ii nj ; a I l ui ,-c r pressure.
* Hc.e ,Tia t.e u"ii,:r o e-, l nv r ji 'eaiirrg I.,r *.r,[. ii, -
tion applications.
* Homogeneous tube and cover construction eliminates
separation.
* Ultraviolet inhibitors reduce aging and weather checking.


Vinylflouw
general purpose water discharge hose
Nominal Hose Approximate Working Coil Approx.
Series Size ID Wall Thickness Pressure Length Weight
No. (in.) (in.) (in.) (psi) (ft) (Ibs.coil)
VF 150 11, 1.673 .0669 80 300 48
VF 200 2 2.165 .0669 80 300 69
VF 250 21/ 2.598 .0787 80 300 87
VF 300 3 3.130 .0787 70 300 117
VF 400 4 4.134 .0827 70 300 156
VF 500 5 5.039 .0866 40 300 204
VF 600 6 6.181 .0866 50 300 258
VF 800 8 8.169 .1063 45 100/300 130/390
VF 1000 10 10.118 .1181 35 100 181
VF 1200 12 12.126 .1181 30 100 195
VF 1400 14 14.134 .1181 30 100 262
VF 1600 16 16.142 .1181 30 100 310


Figure 3.2 Manufacturer information for Kuriyama Vinyl-Flow series



3.1 Experimental Setup and Procedure

3.1.1 Test Sample Orientation

Due to the orthotropic fiber orientation, a value for elastic modulus in both the


longitudinal and hoop directions is desired. To achieve this, samples were made to have


the fibers aligned in a specific orientation. Figure 3.3 shows the convenient coordinate

system chosen and the orientation of each sample. Also shown are Fx and Fy, the tensile


forces applied to each orientation sample.


Longitudinal, y Fy



Hoop,x K




Longitudinally Oriented Sample


Tubing Coordinate System


HoFx


Hoop Oriented Sample


Figure 3.3 Test sample orientations with respect to continuous tubing









3.1.2 Sample Shapes and Correction for Fiber Discontinuity

In order to accurately model the materials' performance, a tensile test sample shape

was needed. It was desired to simply use a continuous tube sample but it was not

possible to achieve this shape in the hoop direction. It was also inconvenient to test a

continuous tube for the large diameter tubes due to the grip size of the tensile test

machine. Since all the samples contain discontinuous fibers, it is necessary to determine

the numerical effect on elastic modulus. To do this, three different sample shapes are

tested and compared to see how the width and continuity of the fibers affect the elastic

modulus.

The first sample, shape A, is a completely continuous length of tubing with a

rectangular shape. The second sample, shape B, is the same as shape A, but differs only

with a single longitudinal cut running the entire length of the specimen. The cut severs

the continuity of the 30 cross fibers. Figure 3.4 shows the longitudinal cut on the

circular cross-section, representing shape B. In order to fit the sample completely within

the width of the test grips, samples A and B are folded twice along the longitudinal

direction as shown in Figure 3.4. The third sample, shape C, is cut from a template with

a shape mimicking a tensile test sample. The ends are designed to fit the width of the

grips of the testing device. The center section is designed to taper from the grips to a

narrow strip to control break location. Care was taken to keep the tapering smooth to

avoid stress concentrations in sharp covers. Figure 3.5 shows an example of shape C.

All samples are 23 cm long. Sample C is 2 cm wide at the narrowest section and samples

A and B have the same width when folded. Figures are shown with the same standard 5

34" pen for scale.

















Figure 3.4 Sample shape B (shown flat and folded)








Figure 3.5 Sample shape C

3.1.3 Experimental Procedure

An MTI 25k screw-driving tensile test machine was used with Curtis 30k self-

cinching jaw grips to perform the tensile tests. An Interface 1220AF-25k load cell was

used in conjunction with MTI software to measure force. After observing preliminary

tests, a test speed of 0.10 in/min was considered reasonable and chosen for all tests.

Figure 3.6 through 3.8 show examples of the test equipment. Figure 3.9 shows an

example of a specimen locked in the grips before pulling.

A sample is closed in the test grips and pulled at a constant velocity to the initial

breaking point. Two marks were made in the middle region of the sample oriented

perpendicular to the length. The length between marks and width (width for sample

shape C) were measured and recorded intermittently throughout the pull test with digital

calipers. The pull force on the specimen is measured and recorded simultaneously with

the length and width measurements. The instantaneous axial stress is calculated at each

point by dividing the force by the cross-sectional area. The cross-sectional area is

calculated as the product of the width and the thickness of the test strip. The material









width and thickness is assumed to be constant in Samples A and B. Only the thickness is

assumed constant in Sample C since the varying width was recorded. The instantaneous

axial strain, F, is calculated at each point using Equation 3.1.

s-s
c5- 5 (3.1)
80

The variable, 6, represents the recorded distance between marks and 60 represents the

original distance between marks. This references strain calculations to the original

unstressed length.

An attempt was made to experimentally determine the material Poisson's ratio by

using measurements of the strain in the width and the strain in the length during the

uniaxial pull tests. It was discovered, however, that since the cut samples do not have

stiffness in compression, and the width decreases with the sample under tension, the

width strain is not a valid measure to be used in determining Poisson's ratio. To

accurately describe Poisson's ratio, the material must be loaded in two directions at once.

Two possible solutions are to use a pull tester capable of pulling in two directions, or to

internally pressurize a uniaxial test specimen. Given the complexity of these tests,

Poisson's ratio is instead determined in chapter 4, using a parameter study in an ANSYS

simulation.









Figure 3.6 Test equipment: MTI screw-drive tensile test machine (MTI, 2004)


Figure 3.7 Test equipment: Interface 1220AF-25k load cell


Figure 3.8 Test equipment: Curtis 30k self-cinching grip
































Figure 3.9 Example of sample locked in grips before pull test

3.2 Sample Shape Choice

Using the analysis described in section 3.3, data from the sample shape tests is used

to determine material elastic modulus. Five specimens were tested for each sample

shape. Table 3.1 shows a summary of the experimentally determined elastic moduli for

each sample shape. It can be concluded that neither sample B nor sample C closely

represent the full tube structure. This is due to the breakage of the continuous hoop fibers

that form the structure of the fabric. Since breaking these fibers is unavoidable in hoop

orientation tests, the assumption was made to apply a multiplicative correction factor to

all tests made from cut samples. The correction factor is determined from a ratio of the

elastic modulus of the continuous tube, sample A, to the elastic modulus of the cut tube,

sample B. This correction factor of 1.29 is applied to all future experimental results,









including both the longitudinal and hoop oriented samples. Sample shape C is used for

all future testing for convenience and predictability.

Table 3.1 Experimental elastic modulus results for three test sample shapes

Sample Shape and Description Orientation Elastic Modulus (psi)


A Folded Continuous Tube Longitudinal 49009

B Folded Cut Tube Longitudinal 38148

C Strip from Template Longitudinal 36098


3.3 Data Analysis and Results

This section describes the analysis of the data collected and its application

towards the orthotropic material properties. Figures 3.10 and 3.11 show plots of axial

stress and axial length-strain for VF-600 material in hoop and longitudinal orientations,

respectively. Five samples per orientation were tested and are plotted with distinct linear

trends. In order to interpret one linear trend from all five tests, all the tests are grouped

together with one trend as shown in Figures 3.12 and 3.13. Combining five distinct

trends is justified by looking at the slope and y-intercepts of the trends. The trend from

each test has a distinct slope, but they are all close enough to consider within

experimental error. The y-intercepts between tests, however, are largely different. The

y-intercept of each trend line indicates an initial load present on the sample before any

length-strain is recorded. This is explained by the observed behavior of the test grips

during the test. The self-cinching grips initially sink in and tighten around the soft

material for a brief period of time before a length-strain is present in the entire sample.

This settling period is interpreted as a pre-loading on the sample by the load cell,

resulting in the y-intercept in the data trend. The presence of the y-intercept is a result of









the specific sample and test apparatus, so it is therefore assumed inconsequential to the

experimental material properties. Since the slope is the only valid material property

obtained from the data, all five groups can be justifiably analyzed with one trend. The

slope of the trend represents the elastic modulus for the corresponding sample

orientation.

Table 3.2 shows a summary of the experimental material property results. Shown

are the experimental and adjusted orthotropic elastic moduli for the hoop and longitudinal

orientations. Adjusted elastic moduli are based on the correction factor for breaking the

continuity of the fibers, determined in section 3.2. It is interesting to note that the elastic

modulus for the longitudinal orientation is much larger than for the hoop orientation since

the hoop stress is generally larger than the longitudinal stress. This is best attributed to

the fiber layer present in the longitudinal orientation. The material was most likely not

designed to be optimized as a pressure vessel, but rather designed for ease of the

manufacturing process.

















* Test 1
* Test 2
Test 3
Test 4
x Test 5


1800

1600

1400

1200

1000

800

600

400

200

0
O.C


4500

4000

3500

3000

2500

S2000

1500

1000

500

0
0.0000


0.0200 0.0400 0.0600 0.0800
Length Strain (inlin)


* Test 6
* Test 7
Test 8
Test 9
STest 10


0.1000 0.1200


Figure 3.11 Experimental data: five longitudinal orientation pull samples


000 0.0500 0.1000 0.1500 0.2000 0.250(
Length Strain (in/in)



Figure 3.10 Experimental data: five hoop orientation pull samples


------I


0


0












1800

1600

1400

1200

1000

800

600
-

400

200

0
0.0000


0.0500 0.1000 0.1500 0.2000
Length Strain (in./in.)


0.2500


Figure 3.12 Experimental data: Five hoop-orientation pull samples grouped together


0.0200 0.0400 0.0600
Length strain (inlin)


0.0800 0.1000 0.1200


Figure 3.13 Experimental data: Five longitudinal-orientation pull samples grouped
together


4500

4000

3500

3000

2500 -

2000 -

1500 -

1000 -

500

0
0.0000









Table 3.2 Summary of experimental and adjusted material properties

Matea / On Experimental Elastic Adjusted Elastic
Material/ Orientation .
Modulus (psi) Modulus (psi)

VF-600/ Hoop 5798.4 7479.9

VF-600 / Longitudinal 36098.0 46566.4



3.5 Errors and Accuracy

Unavoidable experimental errors occurred in all the tests. Due to the nature of the

cross-hatched fabric material, in-plane and out-of-plane shape distortions occurred in

samples when loaded under tension. The in-plane distortions can be described as uneven

shifting of the material in the pull direction. The strain measurements were largely

affected by the in-plane distortion because the measurement lines did not remain

perpendicular to the sample length. The out-of-plane distortions can be described as

cupping of the flat, rectangular cross-sectional area. This largely affects the width and, in

some cases, thickness measurements. Tiny slip marks were observed on some tested

samples where the self-cinching grips cut through or slipped on the slick outer material.

This would contribute additional error to the strain measurements. Figure 3.14 shows an

example of the lines visible on a tested pull sample that indicate slipping. In some cases,

manufacturing inconsistencies in the material were exposed when material samples

visibly broke in a location other than the expected thinnest portion of the sample. Human

error in preparing samples, taking length measurements, and reading the force gage also

contributed to experimental error. The wide scattering of the data points in the stress-

strain curve can also be attributed to the settling period described in section 3.3.









These experimental errors all contribute to the wide scattering of the data points

in the stress-strain curve. Due to the nature of the composite, flexible material and the

test equipment used, the experimental values are considered a valid representation of the

material.







Pull Direction


Grip slippage
Grip slippage


Normal Grip Wear


Figure 3.14 Tested sample with grip slip marks















CHAPTER 4
SIMULATION DEVELOPMENT

The process of developing a finite element model to represent an actual prototype

cage is described in Chapter 4. A simple verification model is first made to explore and

verify finite element membrane analysis for an inflatable tube subject to bending. A full

cage is then modeled to explore the interaction of all the components. Two flow

direction cases are modeled to see the effect of flow direction on stress and deflection in

the cage members. Figures 4.1 and 4.2 show a prototype cage inflated and deflated,

respectively. Figure 4.3 shows an example of the union between the tubing and a PVC

connector using a hose clamp.


Figure 4.1 Cage prototype inflated to -12ft diameter, 9ft height


























lgure 4.2 Cage prototype detlated to -/Ucm

4 Hose clamp
Flexible tubing PVC connector
o A


L I
Figure 4.3 Tubing union to PVC connector

4.1 Verification Model Development and Results

A verification model is designed to establish a basis for further simulations. The

verification model is a simply supported tubular beam with a distributed load on top.

Figure 4.4 shows the conceptual beam model and distributed loading. The hose clamps

are modeled by fixing the x and y displacements of all the nodes along the circumference

of each end. In order to avoid rigid body motion in the analysis, the tube is constrained in

the z direction. To avoid artificial boundary stress concentrations at the ends, the z-

displacements of only 2 nodes on the neutral axis of bending are fixed. Figure 4.5 shows

the applied boundary conditions and the coordinate system used.






















Figure 4.4 Conceptual boundary conditions and loading for the verification model


z-


Dz=
(one end of tube only)
Nodes from outer ring
falling on neutral axis
of bending


Figure 4.5 Detail of boundary conditions for verification model

4.1.1 Applied Loading Conditions

The internal pressure is broken up into pressure acting in the radial direction and

pressure acting longitudinally. The radial pressure is applied normal to the interior tube

surface. The longitudinal pressure is converted to nodal forces, Fi using steps shown in

Figure 4.6. Figures 4.6-A and 4.6-B show the internal pressure, Pi, applied over the area

of the circular cap, Acap. Assuming the cap is rigid, the applied pressure translates into an


I f I f 1 I I I f I









equivalent total axial force, Feq, acting on the cap as shown in Equation 4.2 and Figure

4.6-C. This total force is then distributed equally and applied directly to the number of

nodes, n, on the outer circular edge at the end of the tube as shown in Equation 4.3 and

Figure 4.6-D.


Feq = P, Acap

n
Feq = F,
1i=1

F, Feq
n


(4.1)


(4.2)


(4.3)


Rigid Cap



Acap =7r2




Pi




ntHF-


Fabric Tube


i=1 (D)

Figure 4.6 Steps in representing internal pressure as pre-tension forces

4.1.2 Water Current Drag

Water flowing over the outer surface of the tube creates a drag force. Fluid

assumptions and equations are derived in Chapter 2. A programming do-loop was


Rigid Cap


(A)


(C)









initially implemented into a script file to approximate a parabolic pressure flow

distribution by applying individual forces to nodes. Forces are determined from the

product of the local pressure and the projected area around the node. Figure 4.7 shows a

representative projected area for one node. Applying the pressure distribution to the

tension-only membrane tube in this manner yielded numeric instability. This numeric

instability and a need for only global effects of the drag leads to the assumption that local

pressure effects of the flow on the tube can be neglected. This simplifies the model by

approximating the total drag as an equally distributed load. Figure 4.8 shows an example

of the total drag force divided equally and applied to all the nodes along the length of a

typical tube. All nodal drag forces are applied in the same global coordinate direction.


Xi-1 Xi Xi+1
Figure 4.7 Representative projected area around one surface node














Equal forces_ _
applied at each
node around -
circumference







Figure 4.8 Representative application of forces to nodes around circumference

4.1.3 Element Selection and Properties

The geometry is created in ANSYS by extruding a circle along a linear path. This

area is then meshed with rectangular shell elements. The material to be modeled is a thin

fabric with orthotropic material properties. Since fabrics have tensile stiffness, but very

little compressive stiffness, a thin shell element with optional compressive stiffness is the

best choice. The Shell41 element with Keyopt(1) = 2 meets these requirements. It was

determined that the zero compressive stiffness caused a numerical instability in the

solution of a pressurized beam in drag which resulted in spikes in the deformed shape.

Initially, a small elastic foundation stiffness (EFS) was added to the Shell41 elements to

attempt to fix the instability. Successive ANSYS simulations varying the EFS showed a

large change in mean deflection, shown in Figure 4.9. This large deflection makes

choosing a small EFS value unacceptable and eliminates the use of the EFS in all models.

In another attempt to remedy the instability, equivalent-geometry Shell63

elements with small elastic modulus were merged to the Shell41 elements to add a small

compressive stiffness to the elements. Successive ANSYS simulations varying the added









stiffness of the shell63 elements showed an acceptable change in deflection and stress as

shown in Figures 4.10 and 4.11. A value of 10,000 Pa is chosen for the added stiffness.

Merging the shell63 elements is a preferable choice to fix the instability because the

added stiffness has a less interfering effect on the desired results. Implementing this

choice eliminates the spikes and smoothes out the deformed shape.

The isotropic material properties needed to solve for stress and deflection are

elastic modulus and Poisson's ratio. Since the material is orthotropic, hoop elastic

modulus, longitudinal elastic modulus, shear modulus, and Poisson's ratio are required.

Values for elastic modulus in the hoop and longitudinal directions are found

experimentally in Chapter 3. Since Poisson's ratio is not experimentally determined, a

value is chosen based on finite element model results. Simulations showed that varying

Poisson's ratio had negligible effect on longitudinal stress and deflection. Figure 4.12

shows maximum deflection plotted for various Poisson's ratios determined in an ANSYS

model of a cantilevered tube loaded with an end force. Figure 4.13 shows a similar plot

for maximum longitudinal stress using the same model. The small variation in stress and

deflection justifies assuming a reasonable value of 0.4 for Poisson's ratio.












0.00000E+00


-2.00000E-03


-4.00000E-03


-6.00000E-03


o -8.00000E-03


-1.00000E-02


-1.20000E-02


-1.40000E-02


0 20000 40000 60000 80000 100000 120000 1400
EFS Value (Pa)

Figure 4.9 Verification model parameter study: EFS effect on beam deflection


0.00E+00


-1.00E-03


-2.00E-03


-3.00E-03


-4.00E-03


-5.00E-03


-6.00E-03


-7.00E-03


-8.00E-03


Figure 4.10


0 20000 40000 60000 80000 100000 120000 140000

Added Compressive Stiffness to Shell41 Elements (Pa)

Verification model parameter study: Added stiffness effect on beam
deflection


00












7.00E+06


6.00E+06 *** *


a 5.00E+06


S4.00E+06



0
-.00E+06


o
-J
1.00E+06


0.00E+00


* Maximum Stress
* Minimum Stress


0 20000 40000 60000 80000 100000 120000 140000

Added Compressive Stiffness to Shell41 Elements

Figure 4.11 Verification model parameter study: Added stiffness effect on longitudinal
beam stress


0.025

0.0225

0.02

0.0175

0.015

0.0125

0.01

0.0075

0.005

0.0025

0


0 0.05 0.1 0.15 0.2 0.25 0.3 0.35 0.4 0.45

Poisson's Ratio, vxy (v_hoop-longitudinal)


Figure 4.12 Verification model parameter study: Poisson's ratio effect on beam
deflection


-c
LU

,LL
4-
I-
C.
o.

t?




E

x
E


* *







48



6.00E+06

5.00E+06
*

S4.00E+06
S* Maximum Stress
3.00E+06 m Minimum Stress


2.00E+06


S1.00E+06


0.00E+00
0 0.05 0.1 0.15 0.2 0.25 0.3 0.35 0.4

Poisson's Ratio, vxy (v_hoop-longitudinal)

Figure 4.13 Verification model parameter study: Poisson's ratio effect on longitudinal
beam stress

4.1.4 Verification Model Results

To verify the model, simulation results are compared to the stress and deflection

theoretical solutions derived in Chapter 2. Since flow velocity is proportional to the drag

force acting on the beam, flow velocity is used to compare models. The simulations are

conducted until the beam reaches failure. A study is done for 5m and 10m tubes. Results

from the 5m long tube are shown in Figure 4.14 which correspond to a constant average

deviation of 3.22% between theoretical and simulation. Figure 4.15 shows similar results

for a 10m tube with a constant average deviation between theoretical and simulation to be

0.567%. These results show that the deflection results between theoretical and simulation

stay below an acceptable 5% deviation. Results also show that the deviation decreases as

tube length increases.

A length study was also performed to see how maximum deflection changed with

length. Simulations are performed up to an approximately 10% deflection:length ratio.









Figure 4.16 shows a parabolic trend in the simulation and theoretical results. Figure 4.17

plots the deviation between simulation and theoretical results, showing deviation

decreasing exponentially as length increases. Deviation drops below 10% above a tube

length of approximately 3.5m. The large deviation occurring in tubes shorter than 3.5m

is believed to be due to the Saint-Venant's principle (Hibbeler, 1997), applicable at the

fixed end boundary condition. In this case, the simulated cap restricts radial expansion in

the fabric tube, which affects the maximum deflection.

A stress simulation is also performed to verify the model behavior. Figure 4.18

shows longitudinal stress results for the verification model. The maximum stress

increases, while the minimum stress decreases as expected in bending. A rigid body

motion error is received when the minimum longitudinal stress decreases to

approximately zero, verifying the tension-only behavior of the model. This defines

failure by wrinkling in this model.













0.00E+00

-5.00E-02

-1.00E-01

E -1.50E-01
C
.2 -2.00E-01

4 -2.50E-01
a
S-3.00E-01

S-3.50E-01

-4.00E-01

-4.50E-01

-5.00E-01


Flowing Water Speed (mph)



Figure 4.14 Deflection results for the verification model with length of 5m


* Simulation
* Theoretical


0 0.2 0.4 0.6 0.8 1 1.2 1.4 1.6
Flowing Water Speed (mph)


Figure 4.15 Deflection results for the verification model with length of 10m length


S







I


* Simulation
* Theoretical


S


0.00E+00

-2.00E-01

-4.00E-01

E
- -6.00E-01
C
a
0
o -8.00E-01

9 -1.00E+00
>-

2 -1.20E+00

-1.40E+00

-1.60E+00

-1.80E+00













0.00E+00

-1.00E-01

-2.00E-01

-3.00E-01

-4.00E-01

-5.00E-01

-6.00E-01

-7.00E-01

-8.00E-01

-9.00E-01

-1.00E+00


0 1 2 3 4 5 6

Tube Length (m)


II


7 8 9


Figure 4.16 Deflection results for verification model with 0.5 m/s flow velocity


140.00%


120.00%


= 100.00%
0


G 80.00%
c

S60.00%


O 40.00%


20.00%


0.00%


0 1 2 3 4 5 6
Tube Length (m)


7 8 9


Figure 4.17 Deviation in deflection results for verification model with 0.5 m/s flow
velocity


* Simulation
* Theoretical


A











A



A

A
A A A A AAA A A.. .











1.40E+07

1.20E+07

1.00E+07
0-
S8.00E+06

n Maximum Stress
S6.00E+06 -
~m Minimum Stress

a 4.00E+06

3 2.00E+06

0.OOE+00

-2.00E+06
0 2 4 6 8 10 12 14
Tube Length (m)

Figure 4.18 Longitudinal bending stress results for verification model with 0.5 m/s flow
velocity

4.1.5 Verification Model Conclusion

The results from the verification model show the merged shell41 and shell63

elements acceptably model the tension-only tubing material. The Saint-Venant's

principle effect is specific to the boundary conditions and geometry of the verification

model and is not applied elsewhere. The drag loading and boundary condition

applications are acceptable and accurately model the actual conditions.

4.2 Cage Simulation Process

The actual tubes, when assembled as a cage, have rigid end caps with hose clamps

constraining each end. Based on the geometry of the cage, the tube can flex and extend

longitudinally. Internally, the tube contains pressurized air to maintain inflation. The

external loading on the tube is a distributed drag force caused by a cross-flowing water

current.









4.2.1 Flow Design Models

With infinite possible current flow-directions for an underwater cage, the number

of modeled choices is limited to cages subject to top-down flow and side-flow. Figure

4.19 defines the global coordinate system with side-flow in the positive x-direction and

top-flow in the negative y-direction. Modeling the entire cage was necessary because

unlike the single beam verification model, the boundary conditions that exist at the end of

each tube in the full cage structure are complex and interactive among the cage members.










\ -, / \



Top-Flow Side-Flow
Figure 4.19 Two cage simulation flow orientations

4.2.2 Element Choice

All tubes are meshed with merged tension-only shell41 and shell63 elements in

the same stacked element configuration as the verification model determined in Section

4.1.3. The tubular connectors in the cage are rigid relative to the tubing, so shell63

elements are used. LinklO elements model the tension-only cables that tie together

certain connectors throughout the cage.

4.2.3 Loading Conditions

Internal pressure in the fabric tubes acts in the radial and longitudinal directions.

Pressure acting in the radial direction is applied using the surface element pressure









command (sfe) to only the shell41 elements making up the fabric tubes. Since the tubes

are not actually capped, as in the verification model, a control volume analysis of the

fluid in the rigid connector is used to determine the net effect of the longitudinal pressure

in each tube. Figure 4.20 shows a control volume of fluid inside a typical 45'-connector.

Longitudinal pressure from each tube acts over the entrance area to the connector. The

product of the pressure and area is treated as an equivalent force, Fcap, similarly to

Equation 4.1 in the analysis of the verification model. Assuming the compressibility of

the fluid is negligible, and the fluid is static, the fluid and the connector are approximated

as one rigid body. Figure 4.21 shows Figure 4.20 translated into a free body diagram of a

typical 45'-connector modeled as a rigid body with net x and y component forces. Figure

4.22 shows a finite element model of the same typical 45-connector with the net

component cap forces divided and applied to all the nodes around the circumference of

the center of the connector. The same analysis and nodal force application is used for

each connector.

r - -------



fluid F
f. _^ ~F cap

F o ,,_

SFaap-y


F
Fcap-x

Figure 4.20 Control volume analysis around a rigid 45-connector containing pressurized
internal fluid










Fcap (1 cos 45)





cap (sin 45)

Figure 4.21 Free body diagram of rigid cap with algebraically summed pretension force
components











Forces act on nodes
around entire
circumference



Figure 4.22 Force components applied to finite element model of rigid connector

4.3 Top-Down Flow Cage Model

In the top-down flow model, the center tube is anchored at the top of the cage

with flow in the opposite direction. Small frictional drag occurring along the center tube

is neglected because it lies in the direction of flow. To save on computing time, the

center tube is eliminated from the model and replaced with a single node at the top and

bottom of the cage.

Although not immediately obvious, symmetry could be not be used to simplify

the top-down flow cage model analysis. Dividing the cage along the x, y, or both the x

and y axes does not work because the state of bending makes the z-direction boundary









conditions along the axis of symmetry unknown. The z-direction boundary conditions

also make it impossible to use an axis of symmetry along an axis 45 degrees from the x or

y axis. The cable element controlling the z-direction boundary condition lies along this

axis of symmetry and can not be split.

4.3.1 Applied Loading Conditions

Since external fluid flows over the entire cage, the total fluid drag force on the

cage is calculated based on its total cross-sectional area perpendicular to the direction of

flow. The total force is then divided equally and applied to all nodes of the eight caps

and eight tubes. Referring to the coordinate system defined in Figure 4.23, all drag forces

act in the negative z-direction for the top-down flow model.

4.3.2 Boundary Conditions

Boundary conditions are applied to simulate the conditions acting on the cage as

simply as possible while minimizing solution time. The two locations labeled A in

Figure 4.23 describe two single nodes at the top and bottom of the cage. These points are

placed to obtain a cage diameter-to-height ratio of 8:5. Both points are constrained to

zero deflection in the x,y, and z directions. Cables on the underside of the cage are slack

and do not affect the system, but were included for aesthetics along with the fixed node at

the bottom of the cage.

Locations labeled B and C in Figure 4.23 describe all nodes around the

circumference of a tube at the center of its length. Nodes at location B are constrained to

zero deflection in the y-direction. Nodes at location C are constrained to zero deflection

in the x-direction. Constraining locations B and C prevents rotational rigid body motion

while allowing the individual tubes and cage to expand symmetrically.











Z x A C




B B


(

A C

Figure 4.23 Top-flow boundary condition locations

4.4 Side-Flow Cage Model

The side-flow cage starts with the same octagon model as the top-down flow

model but includes a model of the center tube, all the cables, and a different anchoring

system. Z-direction pretension forces from internal pressure are additionally included at

the ends of the center tube in this model. The rigid connectors are pressurized to the

same internal pressure as the connecting tubing because successive analysis showed an

improvement in the smoothness of the deformation solution around the connections.

Symmetry could also not be used in the side-flow cage model analysis. The x-

axis would be the best choice for axis of symmetry, but the center tube, lying in the z-

direction, would have to be split longitudinally. Also, the x-direction boundary

conditions are unknown on the tubes along the axis of symmetry because the drag flow is

causing bending at those points.

Figure 4.24 shows the components of the center tube as two flexible membranes

and three straight, rigid connectors. The total center tube length is determined by the 8:5

cage diameter-to-height ratio. Figure 4.25 shows linkl0 elements connecting nodes in









the top, bottom, and center connectors of the center tube to nodes in the top and bottom

of the 45-degree connectors in the octagon. Cable elements are merged to single nodes

on the rigid connector symmetrically as shown. Four tension-only cables anchor the cage

from a single external point. Figure 4.26 shows an example of the linkl0 mooring cable

elements merged to one top and bottom node of a 45-degree connector.

Figure 4.27 shows the anchor point location in the global coordinate system,

where L is the length of an individual tube in the octagon. Setting the anchor point x-

location as a function of tube length maintains an anchor distance from the cage as the

cage size is increased. This location is chosen to be reasonably close, while minimizing

the y-tension component in the anchor cables, which affects the deformation of the cage.

4.4.1 Applied Loading Conditions

Drag flow for the side-flow case is approximated as a series of forces applied to

specific cage members. Locations A, B, and C in Figure 4.28 describe flexible

membrane tubes to which drag is applied. The magnitude of drag on locations A, B, and

C is individually based on the respective cross-sectional area. Small frictional drag is

neglected along location D in Figure 4.28 because the members are parallel to the flow

direction. The total drag force for each individual member is divided equally and applied

to the nodes corresponding to the respective location. Referring to the coordinate system

defined in Figure 4.23, all drag forces act in the positive x-direction for the side-flow

model.

4.4.2 Boundary Conditions

The location labeled A in Figure 4.29 describes the single-node anchor point fixed

to zero deflection in the x, y and z directions. To allow symmetric expansion and prevent

rotational rigid body motion, constraints are placed in two more locations in the cage.









Locations labeled B and C in Figure 4.29 describe all nodes around the circumference of

a tube at the center of its length. The nodes at location B are constrained to zero

deflection in the y-direction, while nodes at location C, further illustrated in Figure 4.30,

are fixed to zero deflection in the z-direction.


Flexible Membrane Tubing


Rigid Connectors


Figure 4.24 Cage center tube components








































Figure 4.25 Expanded view of representative internal cable connections


Mooring cables


Flexible
membrane
Rigid tubing
connector


Figure 4.26 Expanded View of representative mooring cable connections











E 3.4L


Figure 4.27 Dimensions of anchor location


Figure 4.28 Side-flow applied loading locations


'NA
























Figure 4.29 Side-flow boundary condition locations


Flexible
membrane
tubing


Rigid
/ connector



Displacement
boundary conditions
(nodes of entire
circumference)


S I I e e
Figure 4.30 Expanded view of center tube boundary conditions















CHAPTER 5
SIMULATION RESULTS

The design criteria and results from the ANSYS cage simulations are presented in

Chapter 5. The goal of the study is to determine the maximum cage size possible before

failure. Twelve commercially available flexible tubes are tested under top and side flow

conditions. The material yielding the largest cage is chosen and subject to an additional

velocity simulation.

5.1 Simulation Strategy

ANSYS script files are used to parameterize the variables involved in the

construction and analysis of each cage. Independent variables were kept constant for all

analyses run in this chapter. Independent variables include:

* External fluid temperature, viscosity, density
* Flexible membrane orthotropic elastic modulii
* Rigid connector elastic modulus
* Flexible membrane and rigid connector Poisson's ratios
* Flexible membrane shear modulus
* Cable constant, EA
* Rigid Connector geometry

Cable length is defined by the cage geometry. Flexible membrane material

properties are explained in Chapter 3, while reasonable values are chosen for external

fluid, connector, and cable properties. Dependent variables affecting the cage geometry

and loading are:

* Flexible membrane internal pressure
* Flexible membrane tube length
* Flexible membrane tube diameter









* Flexible membrane thickness
* External fluid velocity

Internal pressure causes the pretension forces in the flexible membrane tubes. The

pretension forces are also based on the area of the tube opening and so are functions of

the tube diameter. Drag on a given member is a direct function of tube length and tube

diameter. Drag is also a function of the flowing fluid velocity squared. Three of the five

independent variables are combined using data from commercially available materials.

Table 5.1 presents Vinyl-Flow data including tube diameter, thickness, and working

pressure for a given series number. Simulation results make convenient the grouping of

VF-Series' by pressure to show approximately constant pressure trends. The three

pressure range groups are indicated in Table 5.1. The following simulations vary Vinyl-

Flow series number, tube length, and external fluid velocity to determine effects on cage

performance.

Table 5.1 Manufacturer data for Vinyl-Flow commercial drainage tubing

Nominal Hose Approximate Working Coil Approx. Pressure
Series Size ID Wall Thickness Pressure Length Weight Group
No. (in.) (in.) (in.) (psi) (ft.) (lbs.lcoil) (psi)
VF 150 111, 1.673 .0669 80 300 48
VF 200 2 2.165 .0669 80 300 69 70-80
VF 250 2'/ 2.598 .0787 80 300 87
VF 300 3 3.130 .0787 70 300 117
VF 400 4 4.134 .0827 70 300 156
VF 500 5 5.039 .0866 40 300 204
VF 600 6 6.181 .0866 50 300 258 40-50
VF 800 8 8.169 .1063 45 100/300 130/390
VF1000 10 10.118 .1181 35 100 181
VF1200 12 12.126 .1181 30 100 195 30-35
VF1400 14 14.134 .1181 30 100 262
VF1600 16 16.142 .1181 30 100 310

5.2 Failure Criteria

The failure criteria for both flow cases are the points when wrinkling occurs

anywhere in the fabric tubing and when stress exceeds the ultimate strength of the









material. Wrinkling is defined here as a compressive stress. Cage wrinkle diameter is

defined as the maximum cage diameter before any flexible tube wrinkles due to a

compressive stress less than or equal to zero. It is important to consider both the

longitudinal and hoop stresses in the analysis.

Wrinkling is expected to occur in the longitudinal direction because the combined

loading of internal pressure and bending affect the longitudinal stress. Tubes are not

expected to wrinkle in the hoop direction because the only loading affecting the hoop

stress, internal pressure, creates a tensile hoop stress.

The maximum stress will occur in the longitudinal direction because the combined

loading of internal pressure and bending affect the longitudinal stress. The hoop stress is

affected only by the internal pressure loading. Using the manufacturer's recommended

working pressure ensures the hoop stress will not exceed the ultimate strength of the

material.

Figure 5.2 illustrates a representative case of compressive failure behavior for VF-

800 tested at a water speed of 1 knot in top-flow. Maximum and minimum longitudinal

stresses are shown for various cage sizes. As cage size is increased, maximum stress

increases, while minimum stress decreases. Pretension in the tubes keeps both stresses

positive, meaning they are always in tension. When the cage size increases to the point

that the minimum stress drops below zero, the model fails because the tension-only

elements can not support compressive stress. This is then interpreted by the program as

an unconstrained model and results in a rigid body motion error message. Note that the

compressive stress could go very slightly negative without failure because of the small

compressive strength provided by the merged shell63 elements comprising the tube.









Subsequent top-flow simulation failures are judged by the rigid body motion error

message.

The side-flow simulations do not show a smooth trend in the longitudinal stress

down to zero as in Figure 5.2, but show a sudden drop-off resulting in a rigid body

motion error message. This behavior can be described as a buckling of the structural

tubes. The stress behavior is due to changes in the higher order bending occurring in the

tubes as cage size changes. Section 5.4 illustrates and further explains this behavior.

The side-flow cage failure is therefore also determined by the rigid body motion error

message in the simulation.


1.20E+07
Maximum Stress
1.00E+07 Minimum Stress

S8.00E+06

S6.00E+06

I 4.00E+06

3 2.00E+06


0.00E+00
0 5 10 15 20 25 30
Cage Diameter (m)

Figure 5.2 Simulation results: Bending stresses in a VF-800 top-flow cage

5.3 Stress Results

Figure 5.3 illustrates the maximum longitudinal stresses occurring in each

simulation staying below the experimentally determined ultimate strength of the material,

2.6e7 Pa. This ensures the material strength is not exceeded in any simulation.












3.00E+07


2.50E+07
STop-Flow
Maximum Stress
S2.00E+07
m Side-Flow
g Maximum Stress
c 1.50E+07
S-Ultimate Material
o Strength
o- 1.OOE+07
0
-J

5.00E+06


O.OOE+00



Vinyl-Flow Series Number

Figure 5.3 Maximum stresses in each simulation and ultimate material strength

5.4 Geometry Deformation Results

Figures 5.4 through 5.6 show various views of a typical deformed geometry of a


cage subject to top-flow at the critical wrinkle point. All eight tubes bend in the negative

z-direction with maximum z-deflection occurring at half the tube length. Bending plus


the Poisson effect cause the cage diameter to decrease symmetrically throughout the


cage.


Figures 5.7 through 5.8 show two views of a typical deformed geometry of a cage


subject to side-flow at the critical wrinkle point. Note that the net cage x-diameter


increases, while the net cage y-diameter decreases. Internal reaction forces occurring


through the cable connections cause internal bending moments around the corner


connectors of the octagon. Deformation effects of these internal bending moments are


visible in the deformed curvature around the connectors and account for multiple-mode


bending visible in several of the beams. Figure 5.9 illustrates the rotational effect on a









representative rigid connector. The different rotational effects occurring throughout the

cage contribute to the multiple bending modes and subsequent buckling stress behavior

leading to failure.


Figure 5.4 Top-Flow deformed geometry at the wrinkle point: View 1


Undeformed
Deformed


Figure 5.5 Top-Flow deformed geometry at the wrinkle point: View 2











Undeformed



Deformed



Figure 5.6 Top-Flow deformed geometry at the wrinkle point: View 3
Undeformed Deformed


Figure 5.7 Side-Flow deformed geometry at the wrinkle point: View 1


;;











Deformed


Figure 5.8 Side-Flow deformed geometry at the wrinkle point: View 2


Undeformed







Deformed






Note
rotational
effect


Figure 5.9 Representative rotational effect occurring in side-flow simulation









5.5 Simulation Deflection Results

Deflection results are useful in predicting cage behavior. Presented here are

qualitative results for a specific material case and maximum deflection results for each

material case. Numeric sign associated with deflection is related to the previously

defined cage coordinate system. For diameter changes, a negative change indicates a

decrease in diameter, while a positive change indicates an increase.

5.5.1 Qualitative Deflection Results

Figure 5.10 shows deflection data of the net cage x-diameter deflection for a

representative case of the VF-800 tested in top-flow at a water speed of one knot at

various cage diameters. Deflection results are identical for the cage y-diameter based on

the symmetry of the geometry and loading. Cage diameter change has a qualitatively

linear trend for the top-flow cage loading. Figure 5.11 shows data of the maximum z-

deflection for the same representative case. The data represents the maximum bending

deflection occurring in each tube and has an expected qualitatively decreasing parabolic

trend as cage size is increased.

Figure 5.12 shows cage y-diameter deflection data for the VF-500 representative

case, but in side-flow. As seen in the geometry deformation of the side flow cage, the

cage y-diameter decreases drastically. No clear trend is visible in the smaller x-diameter

deflection. This is because the multiple bending modes largely affect the x-diameter

cross members and occur differently in each cage size. X-diameter cage deflections must

be considered on a case by case basis for side-flow.

















S-0.5
0 C
,m


.




S-21
.)


-2.5
Ca



0 5 10 15 20 25 30

Cage Diameter (m)


Figure 5.10 Cage x-diameter net deflection for VF-800 tested at flow speed of 1 knot,
top-flow



0-



-1



-2 -



-3 -



-4



-5
0 5 10 15 20 25 30

Cage Diameter (m)


Figure 5.11 Maximum z-deflection for VF-800 cage tested at flow speed of 1 knot, top-
flow











0.0
S5 10 15

-0.5







E
-1.0 -






> -2.0
01


Z -2.5-


-3.0
Cage Diameter (m)

Figure 5.12 Cage y-diameter net deflection for VF-500 tested at flow speed of 1 knot,
side-flow

5.5.2 Maximum Deflection Results

Two tables are shown to present deflection behavior with respect to the original

cage size. Data are summarized in table format because in each case, the maximum

deflections occur at different cage sizes which correspond to different drag loadings.

With different drag loadings, each case can not be directly compared to determine a

graphical trend. Tube diameter and working pressure are included in the tables for

comparison between cases. Table 5.2 shows the maximum diameter change and

maximum z-deflection with respect to original cage diameter for the top-flow case.

Diameter deflections are between 4-10%, while z-deflections are between 10-18%. Table

5.3 shows the maximum change in the x and y diameters with respect to original cage

diameter for the side-flow case. X-diameter deflections are small and between 2-4%,

while y-deflections are much larger, at 14-22%.











Table 5.2 Maximum cage deflections for the top-flow orientation

Tube diameter Working Maximum cage Ratio of change in cage x- Ratio of z-deflection to
Series number diameter to original cage
(in.) pressure (psi) diameter (m) diameter original cage diameter
diameter

VF 150 1.673 80 7.2 -4.93% -10.78%
VF 200 2.165 80 9.6 -6.49% -14.98%
VF 250 2.598 80 10.8 -6.49% -13.53%
VF 300 3.130 70 13.2 -7.02% -15.83%
VF 400 4.134 70 16.8 -8.71% -18.78%
VF 500 5.039 40 14.4 -5.80% -10.57%
VF 600 6.181 50 20.4 -8.83% -17.37%
VF 800 8.169 45 26.4 -8.75% -17.47%
VF 1000 10.118 35 27.6 -7.78% -13.82%
VF 1200 12.126 30 30 -7.60% -13.59%
VF 1400 14.134 30 34.8 -8.85% -15.76%
VF 1600 16.142 30 39.6 -10.10% -17.92%


Table 5.3 Maximum ca e deflections for the side-flow orientation
Sers Te d r Working Ratio of net x diameter Ratio of net y diameter
Series Te dia r pressure aimmae change to original cage change to original cage
number (in.) diameter (m)
(psi) diameter diameter
VF 150 1.673 80 14.4 2.498% -17.346%
VF 200 2.165 80 18 2.600% -18.633%
VF 250 2.598 80 20.4 2.598% -18.216%
VF 300 3.130 70 21.6 2.681% -18.361%
VF 400 4.134 70 27.6 3.138% -21.377%
VF 500 5.039 40 18 2.569% -14.222%
VF 600 6.181 50 21.6 2.731% -17.667%
VF 800 8.169 45 25.2 3.150% -18.619%
VF 1000 10.118 35 24 2.672% -15.442%
VF 1200 12.126 30 25.2 2.779% -14.976%
VF 1400 14.134 30 28.8 3.262% -17.278%
VF 1600 16.142 30 32.4 3.760% -19.605%



5.6 Cage Size Results

Maximum possible cage size is determined by the point at which the cage wrinkles.


Figures 5.13 and 5.14 show the maximum cage diameter results by tube diameter for the


top-flow and side-flow cases respectively. Slight pressure effects are noted between the


pressure groups of the top-flow case, but are neglected to form one linear increasing


trend. The trend generally shows that, independent of pressure, as diameter increases, the


maximum possible cage diameter increases for top-flow conditions. The side-flow case,


however, has significant pressure effects visible in Figure 5.14. Each pressure group is


plotted with an increasing linear trend, but a distinct slope and y-intercept. Within each










pressure group, as diameter increases, the maximum possible cage diameter increases.

Also evident is that as pressure increases, the slope of the constant pressure trends

becomes steeper. This indicates that for a given diameter, operating at higher pressures

increases the maximum possible cage size.

Since the flow situations are idealized and true water conditions vary no matter

how the cage is anchored, choosing the material for the largest cage is based on the

results of both flow cases. Figure 5.15 presents the data from figures 5.13 and 5.14 as a

summary of the top-flow and side-flow maximum cage diameters possible for all the

material models tested. The lesser cage maximum diameter of the two flow cases is the

limiting factor for each material. Refer to Table 5.1 for model specific manufacturer data

including tube diameter, thickness, and working pressure for each case.


45

40

35

30
slight pressure effect
E



E
E 15 70-80 psi
m 40-50 psi
2 10 30-35 psi

5

0
0 2 4 6 8 10 12 14 16 18
Tube Diameter (in)

Figure 5.13 Top-flow maximum cage diameters for each tubing material grouped by
internal pressure

















25

20

S15 70-80 psi
E m 40-50 psi
0 10
E 30-35 psi
5 -
0 i i 15 -- ^------------- 70 8 p i --



0
0 2 4 6 8 10 12 14 16 18
Tube Diameter (in)

Figure 5.14 Side-flow maximum cage diameters for each tubing material grouped by
internal pressure


E
30

0 25
a
20
c 20
o
E
I 15
E

I 10


Vinyl-Flow Series Number


Figure 5.15 Summary of maximum possible cage size for each material in two flow cases
subject to one knot flow velocity VF Series number corresponds to
diameter, working pressure, and thickness in Table 5.1









5.7 Velocity Results

VF-1600 is chosen because it has the largest limiting cage size of the materials

tested. The maximum possible VF-1600 cage diameters for top-flow and side-flow are

averaged and then divided by two to obtain a cage diameter able to withstand higher than

one knot flow velocity. This cage diameter, 19.2 m, is used for a velocity test on the

cage. The cage withstands the same maximum flow speed of 1.6 knots in both top and

side flow cases before wrinkling occurs. In top-flow, cage diameter decreases 10.183%,

while in side-flow, cage x-diameter increases 4.027%, and cage y-diameter decreases

17.792%. These deflections are reasonable for an inflated flexible membrane cage of this

size and correspond to the expected deformation results shown in Section 5.4.














CHAPTER 6
CONCLUSION

The experimental work performed gave preliminary values of the orthotropic

material properties. More advanced test equipment is necessary to model orthotropic

fiber tubing to find the directional elastic moduli and Poisson's ratio. Large-deformation

strain gages would improve the accuracy in recording the strain used in determining the

elastic modulus. A two direction tensile tester would be necessary to experimentally

determine Poisson's ratio for the inflated-tube loading case. The values determined were

sufficient for the preliminary design and modeling of the orthotropic fiber material.

The analysis presented here focused on the inflatable behavior of the structure. The

simulation included the inflated tube members, rigid PVC connectors, and cable

connectors. It is important to note that the fish netting is excluded from the simulation,

but is necessary for a complete analysis. The simulation effectively models the nonlinear

behavior of the orthotropic, fabric material by translating the inflatable theory into a finite

element model. Preliminary simulations successfully show expected inflatable behavior

in full-cage top and side flow models.

Preliminary findings warrant further research that models the external netting. Net

selection has several biological issues to be addressed such as net biofouling and

biologically safe mesh size. Net analysis will potentially require dynamic analysis to

model the hydrodynamic interaction between the structural element and the fluid

(Tsukrov et al., 2000). Since the cage divergence volume is an important issue, future

work should include at least an approximation of the nets. Further simulation work









should also include optimizing the inflatable material. The cage material's size and

stiffness is expected to offset the increased loading of the nets, making the inflatable

aquaculture cage practical and economical.

The side and top flow models behave differently, particularly with respect to

divergence volume. The internal pressure and tube diameter play the largest role in

deciding the maximum possible cage diameter prior to wrinkling. Instead of including a

safety factor in the final results, the failure criterion was set to the point at which

wrinkling occurred in any tube. Inflatable tubes do not actually fail at the point of

wrinkling, but can continue past this point (Main et al., 1994). Future simulations may be

able to push the failure criteria by modeling the stiffness behavior of the wrinkled

material.

This work successfully models one new application of inflatable structures.

Nonlinear finite element modeling is shown to be an effective tool in analyzing the

behavior of these structures. Inflatables in many applications have a great potential for

improving cost, weight, and convenience. This effort adds one step to the progression of

inflatable structures, and may encourage their further development and acceptance.





















APPENDIX A
ANSYS SCRIPT CODE

The following contains two ANSYS script files used in the design and analysis of

the inflated cage. A top flow orientation script is first presented, followed by a side flow

orientation script. Many variables used in the analysis are parameterized in the beginning

of each script to account for changing conditions.

Top-Flow

/COM, this code has the link10 elements and uses knode command to
/COM, connect them to the caps
/PREP7
/COM, change colors to white background
/RGB,INDEX,100,100,100, 0
/RGB,INDEX, 80, 80, 80,13
/RGB,INDEX, 60, 60, 60,14
/RGB,INDEX, 0, 0, 0,15
/COM, *******************************************
/COM, define geometric tube parameters *
/COM, *******************************************
/COM, inner diameter, units = inches
ID = 8.169
/COM, material thickness, units = inches
thicknessinch = 0.1063
/COM, units = m
length = 8
/COM, end cap radius of curvature determines cap length
caprad=ID/12
/COM, **************************************
/COM, define element parameters *
/COM, **************************************
/COM, number of nodes on quarter circumference
n=7
/COM, number of nodes along length
m=51
/COM, element depth
edepth = length/(m-1)
/COM, Elastic Modulus in the hoop direction (Pa)
Ehoop = 5.157e7
/COM, Elastic Modulus in the longitudinal direction (Pa)










Elong = 3.211e8
/COM, Major Poisson's Ratio (xy)
poisson = 0.4
/COM, Shear Modulus, Gxy, G HL, units?
shearmod = 1.91e7
/COM, Elastic Modulus for Artificial stiffness (small) units=Pa
smallmod = 10000
/COM, Elastic Modulus for Cap (large) units=Pa
capmod=9e9
/COM,*******************************************************
/COM,* define external fluid (water (20 C) parameters *
/COM,*******************************************************
/COM, appendix A-3, external fluid absolute viscosity, kg/m-s
visc = (1e-3)
/COM, external fluid density, kg/m^3
dens = (998)
/COM, ***************************************
/COM, define loading *
/COM, ***************************************
/COM, internal pressure, psi
prespsi = 45
/COM, flowing fluid velocity m/s
vel= (0.5144444)
/COM, **********************************************
/COM, calculate dependant variables, units *
/COM, **********************************************
/COM, angles in degrees
*afundeg
/COM, cap arc length (m)
caplength=0.3
/COM, convert pressure units
pressure = (6894.757*prespsi)
/COM, convert diameter to radius in meters
radius = (ID/2)*0.0254
diam=radius*2
/COM, convert to units = m
thickness = thicknessinch*0.0254
/COM, cross-sectional area ONE tube, units = m^2
tubecsarea = 2*radius*length
/COM, cross-sectional area ONE cap, units = m^2
capcsarea=((3.1417*caprad*diam)/4)
/COM, cross-sectional area ENTIRE cage, units = m^2
totcsarea=((8*tubecsarea)+(8*capcsarea))
/COM, total number of nodes in the entire octagon (cap+tube)
ntot=32*(n**2+m*n-m-3*n+2)
/COM, Reynolds number, dimensionless
Re = (dens*vel*2*radius)/visc
/COM, coefficient of drag, found from Figure 9.13 based on Re
cd= 1
/COM, total drag force, units = N
drag = 0.5*cd*totcsarea*dens*(vel**2)
/COM, force per node equivalent to the internal pressure...
/COM, ...acting on the cap, units = N
capForce = (pressure*3.1417*radius*radius)
/COM,*******************************************************
/COM,* Define Elements *











/COM,*******************************************************
/COM,
/COM, 1 Membrane Shell Element
/COM, 2 Added Stiffness Membrane Shell Element
/COM, 3 Cap Membrane Shell Element
/COM, 4 Cable Element
ET,1,SHELL41,2
ET,2,SHELL63
ET,3,SHELL63
ET,4,LINK10,,1,0
/COM, Real Constants
R,1,thickness
R,2,thickness
R,3,thickness*10
R,4,0.5.0.3
/COM, Young's Modulus in the hoop direction
MP,EX,1, Ehoop
MP,EX,2, smallmod
MP,EX,3, capmod
MP,EX,4, 9e7
/COM, Young's Modulus in the longitudinal direction
MP,EY,1, Elong
/COM, Major Poisson's Ratio (xy)
MP,PRXY, 1,poisson
MP,PRXY,2,poisson
MP,PRXY,3,poisson
/COM, Shear Modulus, Gxy, G HL
MP,GXY,1, shearmod
/COM,*******************************************************
/COM,* CAGE GEOMETRY *
/COM,*******************************************************
/COM, generate one tube and two 45 degree caps (upper right comer)
k,l,length/2,length/2+length*sin(45),0
k,2,length/2+length*cos(45),length/2,0
circle,l,caprad,,,,8
circle,2,caprad,,,,8
1,4,12
circle,4,radius, 12
circle, 12,radius,4
ADRAG,18,19,20,21,,,17
ADRAG,18,19,20,21,,,2
ADRAG,22,23,24,25,,,9
/COM, mesh one tube (upper right comer)
lsel,s,line,,17
lsel,a,line,,27
lsel,a,line,,28
lcl a line 30
lsel,a,line,,32
lesize,all,,,m-1
TYPE,1
MAT,1
REAL,1
amesh, 1,4,1
/COM, mesh two caps (upper right comer)
lsel,s,line,, 18,25,1
lesize,all,,,n-1










lsel,s,line,,35
1scl a lile 36
sel a line 18
lsel,a,line,,40
lsel,a,line,,43
lsel,a,line,,44
lsel,a,line,,46
lsel,a,line,,48
lesize,all,,,n-1
TYPE,3
MAT,3
REAL,3
amesh,5,12,1
/COM, reflect areas and mesh to create all four covers of octagon
arsym,X,all
arsym,y,all
/COM, generate 2 vertical and 2 horizontal tube areas (sides)
k,,length/2,length/2+length*sin(45)+caprad,0
k,,-length/2,length/2+length*sin(45)+caprad,0
k,,-length/2,-length/2-length*sin(45)-caprad,0
k,,length/2,-length/2-length*sin(45)-caprad,0
k,,length/2+length*cos(45)+caprad,length/2,0
k,,length/2+length*cos(45)+caprad,-length/2,0
k,,-length/2-length*cos(45)-caprad,length/2,0
k,,-length/2-length*cos(45)-caprad,-length/2,0
1,99,100
1,101,102
1,103,104
1,105,106
ADRAG,63,66,68,69,,,146
ADRAG,95,98,100,101,,,147
ADRAG,104,108,111,113,,,148
ADRAG,72,76,79,81,,,149
/COM, mesh four tubes (sides)
lsel,s,line,,51
selaline, 5
lsel,a,line,,55
lsel,a,line,,58
lcl a liic S3
lsel,a,line,,85
lsel,a,line,,87
lsel,a,line,,90
lsel,a,line,,115
lsel,a,line,,117
lsel,a,line,,119
lsel,a,line,,122
lsel,a,line,,146
lsel,a,line,,147
lsel,a,line,,148
lsel,a,line,,149
lsel,a,line,,151
lsel,a,line,,152
lsel,a,line,,154
lsel,a,line,,156
lsel,a,line,,159
lsel,a,line,,160










lsel,a,line,,162
lsel,a,line,,164
lsel,a,line,,167
lsel,a,line,,168
lsel,a,line,,170
lsel,a,line,,172
lsel,a,line,,175
lsel,a,line,,176
lsel,a,line,,178
lsel,a,line,,180
lesize,all,,,m-1
TYPE,1
MAT,1
REAL,1
amesh,49,64,1
/COM, re-generate tube areas (sides) and mesh with artificial stiffness
lsel,all
ADRAG,63,66,68,69,,,146
ADRAG,95,98,100,101,,,147
ADRAG,104,108,111,113,,,148
ADRAG,72,76,79,81,,,149
arsym,X,1,4,1,0,1,0
arsym,y,l,4,1,0,1,0
arsym,X, 13,16,1,0,1,0
arsym,y,13,16,1,0,1,0
TYPE,2
MAT,2
REAL,2
amesh,65,96,1
/COM, merge all nodes, elements, and keypoints
nsel,all
esel,all
nummrg, node, le-5
nummrg, kp
nummrg, elem, le-5
/COM, apply total drag force / total # nodes
nsel,all
f,all,fz,-drag/ntot
/COM, apply force boundary conditions
esc1l i' pc 1
sfe,all,,pres,,pressure
/COM, create cage cables
k,200,0,0,((2.4*length)*(5/16))
k,201,0,0,-((2.4*length)*(5/16))
TYPE,4
MAT,4
REAL,4
csys,1
nsel,s,loc,y,22.49999,22.50001
csys,0
f,all,fy,(capforce*(1-sin(45)))/(4*(n-1))
f,all,fx,(capforce*cos(45))/(4*(n- 1))
csys,1
nsel,r,loc,z,radius
knode,203,all
1,203,200,1











lmesh,26
nsel,s,loc,y,22.49999,22.50001
nsel,r,loc,z,-radius
knode,204,all
1,204,201,1
lmesh,29
csys,1
nsel,s,loc,y,67.49999, 67.50001
csys,0
f,all,fx,(capforce*(1-cos(45)))/(4*(n-1))
f,all,fy,(capforce*sin(45))/(4*(n- 1))
csys,1
nsel,r,loc,z,radius
knode,205,all
1,205,200,1
lmesh,31
nsel,s,loc,y,67.49999, 67.50001
nsel,r,loc,z,-radius
knode,206,all
1,206,201,1
lmesh,33
csys,1
nsel,s,loc,y,112.49999, 112.50001
csys,0
f,all,fx,-(capforce*(1-cos(45)))/(4*(n-1))
f,all,fy,(capforce*sin(45))/(4*(n- 1))
csys,1
nsel,r,loc,z,radius
knode,207,all
1,207,200,1
lmesh,70
nsel,s,loc,y,112.49999, 112.50001
nsel,r,loc,z,-radius
knode,208,all
1,208,201,1
lmesh,74
csys,1
nsel,s,loc,y, 157.49999,157.50001
csys,0
f,all,fy,(capforce*(1-sin(45)))/(4*(n-1))
f,all,fx,-(capforce*cos(45))/(4*(n- 1))
csys,1
nsel,r,loc,z,radius
knode,209,all
1,209,200,1
lmesh,77
nsel,s,loc,y, 157.49999,157.50001
nsel,r,loc,z,-radius
knode,210,all
1,210,201,1
lmesh,80
csys,1
nsel,s,loc,y,-22.49999, -22.50001
csys,0
f,all,fy,-(capforce*(1-sin(45)))/(4*(n-1))
f,all,fx,(capforce*cos(45))/(4*(n- 1))











csys,1
nsel,r,loc,z,radius
knode,211,all
1,211,200,1
Imesh, 102
nsel,s,loc,y,-22.49999,-22.50001
nsel,r,loc,z,-radius
knode,212,all
1,212,201,1
Imesh, 106
csys,1
nsel,s,loc,y,-67.49999,-67.50001
csys,0
f,all,fx,(capforce*(1-sin(45)))/(4*(n- 1))
f,all,fy,-(capforce*cos(45))/(4*(n- 1))
csys,1
nsel,r,loc,z,radius
knode,213,all
1,213,200,1
Imesh, 109
nsel,s,loc,y,-67.49999,-67.50001
nsel,r,loc,z,-radius
knode,214,all
1,214,201,1
Imesh, 112
csys,1
nsel,s,loc,y,-l 12.49999,-112.50001
csys,0
f,all,fx,-(capforce*(1-sin(45)))/(4*(n-1))
f,all,fy,-(capforce*cos(45))/(4*(n-1))
csys,1
nsel,r,loc,z,radius
knode,215,all
1,215,200,1
Imesh,134
nsel,s,loc,y,-l 12.49999,-112.50001
nsel,r,loc,z,-radius
knode,216,all
1,216,201,1
Imesh, 138
csys,1
nsel,s,loc,y,-157.49999,-157.50001
csys,0
f,all,fy,-(capforce*(1-sin(45)))/(4*(n-1))
f,all,fx,-(capforce*cos(45))/(4*(n- 1))
csys,1
nsel,r,loc,z,radius
knode,217,all
1,217,200,1
Imesh, 141
nsel,s,loc,y,-157.49999,-157.50001
nsel,r,loc,z,-radius
knode,218,all
1,218,201,1
Imesh, 144
csys,0







87


nsel,s,loc,x,0
d,all,ux,0
nsel,s,loc,y,O
d,all,uy,0
dk,200,ux,0
dk,200,uy,0
dk,200,uz,0
dk,201,ux,0
dk,201,uy,0
dk,201,uz,0
nsel,all
esel,all
nummrg, node
nummrg, elem
nummrg, kp
fini
/COM, ************************************
/COM, Solution *
/COM, ************************************
/solve
nsel,all
esel,all
autots,on
deltim, 1
solve
postall
cys,0
nsel,s,loc,x,-length/2,length/2
nsel,r,loc,y,0,length*5
esln,s,all
escl I pi 1
plesol,s,x
esel,all
nsel,all
plnsol,u,x
esel,all
nsel,all
plnsol,u,z


Side-Flow

/PREP7
/COM, *******************************************
/COM, define geometric tube parameters *
/COM, *******************************************
/COM, change colors to white background
/RGB,INDEX,100,100,100, 0
/RGB,INDEX, 80, 80, 80,13
/RGB,INDEX, 60, 60, 60,14
/RGB,INDEX, 0, 0, 0,15
/COM, inner diameter, units = inches
ID = 16.142
/COM, material thickness, units = inches
thicknessinch = .1181
/COM, units = m




Full Text

PAGE 1

PRELIMINARY DESIGN AND NONLINEAR NUMERICAL ANALYSIS OF AN INFLATABLE OPEN-OCEAN AQUACULTURE CAGE By JEFFREY SUHEY A THESIS PRESENTED TO THE GRADUATE SCHOOL OF THE UNIVERSITY OF FLORIDA IN PARTIAL FULFILLMENT OF THE REQUIREMENTS FOR THE DEGREE OF MASTER OF SCIENCE UNIVERSITY OF FLORIDA 2004

PAGE 2

Copyright 2004 by Jeffrey Suhey

PAGE 3

This work is dedicated to anyone who strives to challenge himself.

PAGE 4

ACKNOWLEDGMENTS I would like to express my gratitude to my advisor, Dr. Christopher Niezrecki, for his guidance and wisdom. Working on this project has provided me with an opportunity to advance my education and potential. My sincerest thanks go out to Dr. Nam-Ho Kim for his enthusiasm and expertise. I would also like to thank Dr. Ashok Kumar for serving on my defense committee. My thanks go to Dr. Ifju and the members of the Experimental Stress Analysis Lab for their assistance. Srikant Ranjan, Vann Chesney, and Jeff Leismer were a great help with my ANSYS installation and licensing issues. I also owe a debt of gratitude to Anne Baumstarck, the Baumstarck family, and Erik Mueller for providing me with temporary lodgings while in Gainesville. And special thanks go to God, my family, and all my friends for their continued support and encouragement. iv

PAGE 5

TABLE OF CONTENTS page ACKNOWLEDGMENTS.................................................................................................iv LIST OF TABLES............................................................................................................vii LIST OF FIGURES.........................................................................................................viii ABSTRACT......................................................................................................................xii CHAPTER 1 INTRODUCTION........................................................................................................1 1.1 Aquaculture Introduction........................................................................................1 1.2 Inflatable Structure Introduction.............................................................................4 1.3 Inflatable Structure Analyses................................................................................12 1.3.1 Modified Conventional Beam Theory Approach......................................13 1.3.2 Nonlinear Approach...................................................................................14 1.3.3 Linear Shell Method...................................................................................14 1.3.4 Finite Element Analysis Approach.............................................................15 1.3.5 Selected Approach......................................................................................16 2 THEORETICAL DEVELOPMENT..........................................................................17 2.1 Fluid Analysis.......................................................................................................17 2.2 Non-wrinkled Inflatable Theory...........................................................................19 2.3 Case Specific Non-wrinkled Inflatable Theory....................................................21 2.3.1 Bending Moment........................................................................................21 2.3.2 Vertical Deflection.....................................................................................22 2.3.3 Wrinkle Length...........................................................................................23 2.4 Conclusion............................................................................................................25 3 EXPERIMENTAL PROCEDURE AND RESULTS.................................................26 3.1 Experimental Setup and Procedure.......................................................................27 3.1.1 Test Sample Orientation.............................................................................27 3.1.2 Sample Shapes and Correction for Fiber Discontinuity............................28 3.1.3 Experimental Procedure............................................................................29 v

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3.2 Sample Shape Choice...........................................................................................32 3.3 Data Analysis and Results....................................................................................33 4 SIMULATION DEVELOPMENT.............................................................................39 4.1 Verification Model Development and Results.....................................................40 4.1.1 Applied Loading Conditions......................................................................41 4.1.2 Water Current Drag....................................................................................42 4.1.3 Element Selection and Properties...............................................................44 4.1.4 Verification Model Results.........................................................................48 4.1.5 Verification Model Conclusion..................................................................52 4.2 Cage Simulation Process......................................................................................52 4.2.1 Flow Design Models...................................................................................53 4.2.2 Element Choice..........................................................................................53 4.2.3 Loading Conditions....................................................................................53 4.3 Top-Down Flow Cage Model...............................................................................55 4.3.1 Applied Loading Conditions......................................................................56 4.3.2 Boundary Conditions..................................................................................56 4.4 Side-Flow Cage Model.........................................................................................57 4.4.1 Applied Loading Conditions......................................................................58 4.4.2 Boundary Conditions..................................................................................58 5 SIMULATION RESULTS.........................................................................................63 5.1 Simulation Strategy..............................................................................................63 5.2 Failure Criteria......................................................................................................64 5.3 Stress Results........................................................................................................66 5.4 Geometry Deformation Results............................................................................67 5.5 Simulation Deflection Results..............................................................................71 5.5.1 Qualitative Deflection Results....................................................................71 5.5.2 Maximum Deflection Results.....................................................................73 5.6 Cage Size Results.................................................................................................74 5.7 Velocity Results....................................................................................................77 6 CONCLUSION...........................................................................................................78 APPENDIX A ANSYS SCRIPT CODE.............................................................................................80 B EXPERIMENTAL DATA........................................................................................102 LIST OF REFERENCES.................................................................................................104 BIOGRAPHICAL SKETCH...........................................................................................106 vi

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LIST OF TABLES Table page 3.1 Experimental elastic modulus results for three test sample shapes..........................33 3.2 Summary of experimental and adjusted material properties....................................37 5.1 Manufacturer data for Vinyl-Flow commercial drainage tubing.............................64 5.2 Maximum cage deflections for the top-flow orientation..........................................74 5.3 Maximum cage deflections for the side-flow orientation........................................74 vii

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LIST OF FIGURES Figure page 1.1 Example of a class 1 cage design (Loverich and Forster, 2000)................................3 1.2 Example of a class 2 cage design (Loverich and Forster, 2000)................................3 1.3 Example of a class 3 cage design (Loverich and Forster, 2000)................................4 1.4 Example of a class 4 cage design (Loverich and Forster, 2000)................................4 1.5 Millennium Arches in Stockholm, Sweden: entirely self-supporting inflated structure. (Lindstrand, 2000).....................................................................................6 1.6 Millennium Arches alternate view (Lindstrand, 2000)...........................................6 1.7 Section of Millennium Arches before installation (Lindstrand, 2000)....................7 1.8 Hawkmoor self-supporting inflated Temprodome: shown with three units linked together, each with dimensions (6 x 4 x 2.8m). (Hawkmoor, 2003).........................7 1.9 Exterior of Court TV inflated dome. 20 ft tall by 40 ft diameter. For scale, note the person in the doorway. (Promotional Design Group, 2001)................................7 1.10 Interior of Court TV dome showing inflated arch columns. (Promotional Design Group, 2001)..............................................................................................................8 1.11 Thin-film inflated torus used as a support structure for optical reflector (SRS, 2000)................................................................................................................8 1.12 Inflated thin film with reflective coating to be used as a light-weight deployable antenna, 5 meter diameter (SRS, 2000)......................................................................9 1.13 Inflatable Antenna Experiemnt (IAE): 14 meter diameter deployed antenna with inflated support beams (Dornheim, 1999)..................................................................9 1.14 IAE Antenna during inflation. Stowed volume is one-tenth of deployed volume. (Dornheim, 1999).....................................................................................................10 1.15 IAE Reflector (14m diameter). For scale, note the person circled in the upper right (LGarde, 1996)...............................................................................................10 viii

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1.16 Goodyear Corporation inflatable truss radar antenna (Jenkins, 2001).....................11 1.17 LGarde inflatable solar array (Jenkins, 2001).........................................................11 1.18 ILC Dover, Inc. modular split blanket solar array. Shown deployed and stowed (Jenkins, 2001).........................................................................................................11 1.19 Planar deployment solar shade (Jenkins, 2001).......................................................12 1.20 Inflatable Goodyear Inflatoplane (Jenkins, 2001).................................................12 2.1 Flow on a vertical tube.............................................................................................18 2.2 Flow on an inclined tube..........................................................................................18 2.4 Drag loading assumed as a distributed force............................................................22 2.5 Free body diagram of beam section.........................................................................22 3.1 Vinyl-Flow internal fiber layer orientations.............................................................26 3.2 Manufacturer information for Kuriyama Vinyl-Flow series....................................27 3.3 Test sample orientations with respect to continuous tubing....................................27 3.4 Sample shape B (shown flat and folded)..................................................................29 3.5 Sample shape C........................................................................................................29 3.6 Test equipment: MTI screw-drive tensile test machine (MTI, 2004).....................31 3.7 Test equipment: Interface 1220AF-25k load cell....................................................31 3.8 Test equipment: Curtis 30k self-cinching grip........................................................31 3.9 Example of sample locked in grips before pull test.................................................32 3.10 Experimental data: five hoop orientation pull samples...........................................35 3.11 Experimental data: five longitudinal orientation pull samples................................35 3.12 Experimental data: Five hoop-orientation pull samples grouped together.............36 3.13 Experimental data: Five longitudinal-orientation pull samples grouped together..36 3.14 Tested sample with grip slip marks..........................................................................38 4.1 Cage prototype inflated to ~12ft diameter, 9ft height..............................................39 4.2 Cage prototype deflated to ~70cm...........................................................................40 ix

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4.3 Tubing union to PVC connector...............................................................................40 4.5 Detail of boundary conditions for verification model..............................................41 4.6 Steps in representing internal pressure as pre-tension forces...................................42 4.7 Representative projected area around one surface node..........................................43 4.9 Verification model parameter study: EFS effect on beam deflection.....................46 4.10 Verification model parameter study: Added stiffness effect on beam deflection...46 4.11 Verification model parameter study: Added stiffness effect on longitudinal beam stress.........................................................................................................................47 4.12 Verification model parameter study: Poissons ratio effect on beam deflection....47 4.13 Verification model parameter study: Poissons ratio effect on longitudinal beam stress.........................................................................................................................48 4.14 Deflection results for the verification model with length of 5m..............................50 4.15 Deflection results for the verification model with length of 10m length.................50 4.16 Deflection results for verification model with 0.5 m/s flow velocity......................51 4.17 Deviation in deflection results for verification model with 0.5 m/s flow velocity..51 4.18 Longitudinal bending stress results for verification model with 0.5 m/s flow velocity.....................................................................................................................52 4.19 Two cage simulation flow orientations....................................................................53 4.20 Control volume analysis around a rigid 45-connector containing pressurized internal fluid.............................................................................................................54 4.21 Free body diagram of rigid cap with algebraically summed pretension force components...............................................................................................................55 4.22 Force components applied to finite element model of rigid connector....................55 4.23 Top-flow boundary condition locations...................................................................57 4.24 Cage center tube components...................................................................................59 4.25 Expanded view of representative internal cable connections...................................60 4.26 Expanded View of representative mooring cable connections................................60 x

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4.27 Dimensions of anchor location.................................................................................61 4.28 Side-flow applied loading locations.........................................................................61 4.29 Side-flow boundary condition locations..................................................................62 4.30 Expanded view of center tube boundary conditions................................................62 5.3 Maximum stresses in each simulation and ultimate material strength.....................67 5.4 Top-Flow deformed geometry at the wrinkle point: View 1...................................68 5.5 Top-Flow deformed geometry at the wrinkle point: View 2...................................68 5.6 Top-Flow deformed geometry at the wrinkle point: View 3...................................69 5.7 Side-Flow deformed geometry at the wrinkle point: View 1..................................69 5.8 Side-Flow deformed geometry at the wrinkle point: View 2..................................70 5.9 Representative rotational effect occurring in side-flow simulation.........................70 5.10 Cage x-diameter net deflection for VF-800 tested at flow speed of 1 knot, top-flow....................................................................................................................72 5.11 Maximum z-deflection for VF-800 cage tested at flow speed of 1 knot, top-flow....................................................................................................................72 5.12 Cage y-diameter net deflection for VF-500 tested at flow speed of 1 knot, side-flow...................................................................................................................73 5.13 Top-flow maximum cage diameters for each tubing material grouped by internal pressure.....................................................................................................................75 5.14 Side-flow maximum cage diameters for each tubing material grouped by internal pressure.....................................................................................................................76 5.15 Summary of maximum possible cage size for each material in two flow cases subject to one knot flow velocity VF Series number corresponds to diameter, working pressure, and thickness in Table 5.1..........................................................76 xi

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Abstract of Thesis Presented to the Graduate School of the University of Florida in Partial Fulfillment of the Requirements for the Degree of Master of Science PRELIMINARY DESIGN AND NONLINEAR ANALYSIS OF AN OPEN-OCEAN AQUACULTURE CAGE By Jeffrey Suhey August 2004 Chair: Dr. Christopher Niezrecki Major Department: Mechanical and Aerospace Engineering A nonlinear finite element analysis is used to predict the static performance of a novel aquaculture cage constructed from inflated-beam members. Simulations are performed to determine the cage maximum deflections, stresses in the inflated components, and ultimately the maximum possible cage size. Experimental tests are performed on a commercially available anisotropic fabric material to determine longitudinal and hoop elastic moduli. Fluid drag forces on the cage are applied to the inflated members as a constant distributed force that is dependent on the flow speed of the current. Simulations are performed up to 1.6 knots of water current. Two flow cases based on the orientation of the cage with respect to water current are considered. Drag on the surrounding cage nets is excluded from the preliminary design analysis. Within this study, cage members were pressurized from 30 to 80 psi, depending on the material. xii

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Important results include successfully merging a linear and nonlinear element to represent tension-only behavior of an inflated fabric tube. Verification was performed by comparing a simply-supported distributed-load inflated beam finite element model with a modified traditional beam analysis. Maximum possible cage sizes are found at constant flow velocity for different materials with varying internal pressure, wall thickness, and tube diameter. Effects of changing flow velocity are found for a specific material case in both flow directions. The results indicate that the largest sized cage for a 1 knot (0.51 m/s) current has a cage diameter of 39 m for the material used in this study. xiii

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CHAPTER 1 INTRODUCTION The current trend in new inflated structure applications lends itself to studying an underwater application. In this case, a preliminary engineering design analysis of an open-ocean aquaculture cage is presented. This chapter provides a brief introduction to aquaculture and describes recent developments in inflatable structures. 1.1 Aquaculture Introduction Aquaculture can be described using the following definition: Aquaculture is defined as the propagation and rearing of aquatic organisms in controlled or selected aquatic environments for any commercial, recreational, or public purpose (Department of Commerce, 2002). Aquaculture is an agricultural approach that provides an alternative to the other two common methods of obtaining fish; commercial fishing and fish farming with large permanent structures in the ocean or on land (Loverich and Forster, 2000). Aquaculture is expected to play an increasingly important role in meeting the global demand for fisheries products as the world population continues to expand and fish stocks approach their biological limits (Fredriksson et al., 2000). The problem is making the use of aquaculture cages economical. A growing need for improved cage designs will hopefully stimulate engineering research to overcome the obstacles now facing aquaculture. Several factors comprise aquaculture cage design. Costs of manufacturing, installation and maintenance, cage stability, durability, and predictability, all play an important role in the design process. One aspect specifically important to aquaculture 1

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2 cage design is the divergence volume. This can be defined as a maximum potential geometric deformation of the cage. Large divergence volumes can cause high fish mortality rates and decrease diver safety. These problems can lead to large economic drawbacks, making the divergence volume a key factor on which to base a design. A variety of cages have been designed to accommodate individual design advantages and disadvantages. These designs have been grouped into four classes that are in current use throughout the aquaculture community. The class 1 cage is the most common in use today and is based on forces due to gravity; weight and buoyancy. The cage typically comprises a floating frame with a weighted net hanging below. This simple design has several problems, including unpredictable failures and high fish mortality. The major disadvantage to this cage design is the large divergence volume. Figure 1.1 shows a typical class 1 cage design. The class 2 cage has anchored rigid truss members used to hold the shape of the netting. Due to the taught netting, there is a significant improvement in the divergence volume compared to the class 1 cage. Similarly, class 3 cages maintain taught netting for this same improvement. Class 3 cages generally have a central truss with a rigid ring connected by ropes and covered in netting. This design allows the cage to maintain its shape without relying on anchors. The deformation of the class 3 cage is governed by the rigidity of its components. It achieves a smaller divergence volume than the class 1 and 2 by providing more supports over the span of the net. Figures 1.2 and 1.3 show examples of typical class 2 and 3 cages, respectively. The Class 4 cage is a completely rigid self-supported frame covered in netting. Although seemingly ideal due to the very low divergence volume, the costs of material,

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3 construction, and installation labor are much higher than of the previous three classes of cage design. Figure 1.4 shows an example of a typical class 4 cage design. Figure 1.1 Example of a class 1 cage design (Loverich and Forster, 2000) Figure 1.2 Example of a class 2 cage design (Loverich and Forster, 2000)

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4 Figure 1.3 Example of a class 3 cage design (Loverich and Forster, 2000) Figure 1.4 Example of a class 4 cage design (Loverich and Forster, 2000) 1.2 Inflatable Structure Introduction Several applications exist for inflatable structures. Applications for inflatable structures have been used successfully in the past. Some familiar applications that are still in use include vehicle tires, surface watercraft (inflated boats, pontoons, hovercraft, etc), and aircraft (blimps, balloons, etc). Within the past few decades, new inflatable applications are being explored and gaining acceptance. Inflatable land structures are completely self-supported, requiring no solid structural members. These are currently in use in military and architectural designs for tents, hangars, roofs, and small buildings. These applications take advantage of the

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5 convenience of inflatable structures to save assembly time and travel weight. today, its well within the bounds of possibility to find a 300 ft long building housing a concert venue, exhibition hall, museum, or space-age science display, with not a solid supporting member in sight (Lindstrand, 2000). Figures 1.5 through 1.10 show examples of inflated land structures, with no rigid support members, supported only by internal pressure. The [Millennium Arches] whole building is 328ft(100m) long 59ft(18m) wide, and 56ft(17m) high. The main arches are 38ft(11.5m) high, and 59ft(18m) wide while the end arches are 41ft(12.5m) long and the center arch is 164ft(50m) long (Lindstrand, 2000). The aerospace field is working with inflatables in gossamer structures for space applications such as sunshields, antennas, solar sails, habitats, and structural booms. Space applications take advantage of several features of inflatable structures. They [inflatable structures] offer large potential reductions in stowed volume, cost, and often weight.Giant space structures such as 1,000-ft. antennas or solar sails may not even be possible with mechanical deployment, but may be doable with inflatable design (Dornheim, 1999). Figures 1.11 through 1.15 show examples of space applications of inflatable structures.

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6 Figure 1.5 Millennium Arches in Stockholm, Sweden: entirely self-supporting inflated structure. (Lindstrand, 2000) Figure 1.6 Millennium Arches alternate view (Lindstrand, 2000)

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7 Figure 1.7 Section of Millennium Arches before installation (Lindstrand, 2000) Figure 1.8 Hawkmoor self-supporting inflated Temprodome: shown with three units linked together, each with dimensions (6 x 4 x 2.8m). (Hawkmoor, 2003) Figure 1.9 Exterior of Court TV inflated dome. 20 ft tall by 40 ft diameter. For scale, note the person in the doorway. (Promotional Design Group, 2001)

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8 Figure 1.10 Interior of Court TV dome showing inflated arch columns. (Promotional Design Group, 2001) Figure 1.11 Thin-film inflated torus used as a support structure for optical reflector (SRS, 2000)

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9 Figure 1.12 Inflated thin film with reflective coating to be used as a light-weight deployable antenna, 5 meter diameter (SRS, 2000) Figure 1.13 Inflatable Antenna Experiemnt (IAE): 14 meter diameter deployed antenna with inflated support beams (Dornheim, 1999)

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10 Figure 1.14 IAE Antenna during inflation. Stowed volume is one-tenth of deployed volume. (Dornheim, 1999) Figure 1.15 IAE Reflector (14m diameter). For scale, note the person circled in the upper right (LGarde, 1996)

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11 Figure 1.16 Goodyear Corporation inflatable truss radar antenna (Jenkins, 2001) Figure 1.17 LGarde inflatable solar array (Jenkins, 2001) Figure 1.18 ILC Dover, Inc. modular split blanket solar array. Shown deployed and stowed (Jenkins, 2001)

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12 Figure 1.19 Planar deployment solar shade (Jenkins, 2001) Figure 1.20 Inflatable Goodyear Inflatoplane (Jenkins, 2001) 1.3 Inflatable Structure Analyses This section describes various forms of modeling inflatable structures. A specific emphasis is placed on determining deflection and internal stress behavior of inflatable structures subject to cantilever bending. Also of interest is analysis of the application of fabrics for the inflated material. The basic concept of an inflatable structure is a closed end pressure vessel subject to some applied loading. The internal pressure translates to a force applied over the capped end in the axial direction. This force then creates an axial stress that is able to resist bending and axial stress caused by the applied loading conditions. Chapter 2 more fully explains the derivation of the inflatable theory.

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13 1.3.1 Modified Conventional Beam Theory Approach Experimental data has shown that inflatable beams do not behave according to traditional beam theory. With modification to the traditional beam theory, an accurate model can be made. One study analyzed inflatable beam deflections and stresses for loads between incipient buckling, where bending stress equals axial stress due to pressure, and final collapse. It was shown that initially bending occurred with no wrinkling. When wrinkling occurred, a slack region and a taught region were present. A formulation for the stress was developed showing zero axial stress in the slack region and a portion of the max stress carried in the taught region. By equating expressions for maximum axial stress, a formulation for the relationship between internal pressure and applied tip force was determined. The curvature was determined and integrated twice to find the deflection due to bending. The deflection due to shear was also determined noting that it is negligible when the length of the beam is much larger than the radius of the cross section. The total deflection was then determined as the algebraic sum of the deflections due to bending and shear. Similar results were obtained for a beam subject to distributed load bending (Comer & Levy, 1963). This analysis shows when wrinkles begin to form and where they form. The wrinkled membrane is shown to have some rigidity. The effect of the wrinkling on the bending strength of the beam is determined. Another study applies the modified traditional beam theory approach from Comer and Levy for fabric materials. An experimentally determined modulus of membrane of the fabric material was used in place of Youngs modulus. This modulus of membrane is defined as the slope of the experimental stress resultant vs. engineering strain plot, where stress resultant is the force per width of a pull-test sample. Using an approach similar to that of Comer and Levy, the curvature was integrated numerically to determine tip

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14 deflection behavior. Experimental results agreed for aspect ratios of l/d>6 (Main et al., 1994). 1.3.2 Nonlinear Approach Since inflatable beams behave nonlinearly, a nonlinear approach, although complex, is the most logical method. Douglas uses the theory of incremental deformations to find the Cauchy-Green deformation tensors which can be related to determine the Lagrangian and Eulerian Strains. These lead to a stress-deformation relation. This is all then applied to an inflated beam incorporating the pressure in as a stress. The variation in beam stiffness is plotted as a function of changing internal pressure. Unlike a linear approach, this thorough method takes into account the changes in geometry and changes in material properties (Douglas, 1968). Another study included an explanation of the geometry of fabrics and the interaction between fibers. The general relationship between stress and strain for a fabric material is ultimately shown to be a nonlinear trend (Bulson, 1973). 1.3.3 Linear Shell Method Another method for analyzing inflatable structures is to use a linear shell approach. One study develops a free body diagram of a width of flexed strip of the loaded material. Static equilibrium of forces is then applied, including the applied load and the skin tensions on the width of the material. First, a vertical force summation is made with the application of a small angle approximation. Then, a horizontal force summation is developed and combined with the vertical summation canceling the skin tension terms. The result is an expression analogous to the equation for shearing angular deflection of a solid beam. By comparing these two equations, the internal pressure is found to be

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15 analogous to an effective shear modulus. This is used to determine the effect of pressure on the shearing deflection (Topping, 1964). Another source applies continuum mechanics to membranes to obtain several general relations. It defines the kinematics, deformation, strain, strain rate, and stress for membranes. Balance laws are used to derive general mass balance, momentum balance, and energy balance relations (Jenkins, 2000, pp. 49-64). To model wrinkling in beams, a tension field model is developed with an approach for application to linear finite element method (Jenkins, 2000, pp 103-105). 1.3.4 Finite Element Analysis Approach Another approach is to apply finite element analysis to inflated structures. Finite element analysis can yield accurate results for problems with complex analytical solutions. One study compares results from the modified traditional beam theory with finite element results. An assumption of Braziers effect is used here. This states that as the tube deforms due to bending and the cross-section becomes flatter, the bending stiffness of the entire beam decreases. This is most likely due to geometry changes that affect the area moment of inertia. This study formulates finite elements of geometrically nonlinear motion of membrane. These allow for changes in length of the elements to account for stretching of the material in the ends but not in the cylinder. Triangle elements are used to model the circular end caps and quad elements are used to model the cylindrical tubing. A pressure relationship is used to apply a force to each element. FEA results match theoretical results from the modified traditional beam approach. An analysis of tubes containing multiple pockets is conducted as well. The behavior of these multi-cellular inflated beams is shown graphically. By increasing the

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16 number of cells, the pressure can be varied throughout the length of the tube and thereby optimize the design (Sakamoto and Natori, 2001). 1.3.5 Selected Approach The modified traditional beam theory will be used in conjunction with a nonlinear finite element model for future analysis. The finite element approach solves complex geometries and provides a visualization of deformed behavior. The modified traditional beam theory will be used to verify the behavior of a simple finite element model. Following verification, the finite element approach will then be used to model complex cage geometries and interactions. Further theoretical development of the modified traditional beam theory is presented in Chapter 2 and the finite element approach is described in Chapter 4.

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CHAPTER 2 THEORETICAL DEVELOPMENT The development of the theory applied to analytical solutions used in later chapters is described in Chapter 2. A standard fluid flow analysis is first presented, followed by the traditional beam method applied to inflatable theory, which is chosen from the methods presented in Chapter 1. The theory presented focuses on inflated beams before and at the point of wrinkling. The model used for derivation is an internally pressurized circular-cylindrical closed-end fabric tube. The internal pressure acting over the surface area of the end caps translates to pretension forces in the fabric tube. The essence of the inflatable theory is that bending causes a compressive stress on the underside of the tube opposing the tensile stress caused by the pretension forces. Wrinkling occurs when the compressive portion of bending stress exceeds the pretension stress. In some cases, stress resultant, *, is used to replace axial stress, throughout the analysis, as described in Mains theory, to more accurately model fabric behavior (Main et al., 1994). This also leads to using a resultant elastic modulus, E*, with units of force per unit length. 2.1 Fluid Analysis The fluid analysis is used to determine the total drag force acting on a cylinder for a given external flow condition. The following assumptions are made for the fluid analysis: Smooth cylinder surface Constant temperature fluid and material properties at 20C 17

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18 Steady state flow Constant cross-sectional area In this case, the drag force is caused by friction of fluid current flowing over the tube. The drag force is a function of tube cross-sectional area, A, drag coefficient, C D external fluid density, and external flow velocity, V. Equation 2.1 defines the drag force, F drag on a cylinder based on applicable assumptions. 221VACFDdrag (2.1) The cross-sectional area perpendicular to the flow direction is defined,, with tube diameter, D, and cross-sectional length perpendicular to flow, h. Figures 2.1 and 2.2 define h for a vertical tube and a tube inclined by angle, from the horizontal, respectively. DhA Lh Flow Direction Figure 2.1 Flow on a vertical tube sinLh Flow Direction Figure 2.2 Flow on an inclined tube

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19 The coefficient of drag, C D is determined using the dimensionless parameter, Re, and the approximated range taken from documented experimental data. Equations 2.2 and 2.3 show the Reynolds number relation and the approximated value of C D based on the expected flow velocity range. VDeR (2.2) 1 DC for 10 000,200Re000, (2.3) D represents the diameter of the tube. Fluid properties of the flowing current including density and absolute viscosity, are determined from documented values at a temperature of 20 C (Fox & McDonald, 1999). 2.2 Non-wrinkled Inflatable Theory Figure 2.3 shows the assumed stress distribution for a non-wrinkled inflated beam subject to a bending moment. 0 m o Figure 2.3 Stress distribution in tube subject to bending moment Comer & Levy (1963) defined the stress distribution in this unwrinkled inflated circular-cylindrical beam using Equation 2.4. 2)cos1(2)cos1(0 m (2.4)

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20 As shown in Figure 2.3, m and 0 are the maximum and minimum longitudinal stresses in the circumferential stress distribution, respectively. Equation 2.4 can be rewritten as in Equation 2.5, by diving through by the material thickness. Stress, is replaced with stress resultant *. 2)cos1(2)cos1(**0*m (2.5) Equation 2.6 describes the bending moment, M(x), as a product of the tensile forces, F(), and the distance, h(), from the neutral axis of bending. Force is then replaced with the product of stress and area. The area is defined by integrating the fabric thickness, t, over the tube circumference using tube radius, r, and angle, from 02. (2.6) 2020)()()()()(drthhFxM Defining height, cos)(rh and stress resultant, Equation 2.6 can be rewritten using symmetry as Equation 2.7. t)()(* (2.7) 02*cos2)(drxM Substituting Equation 2.5 into Equation 2.7 and integrating yields Equation 2.8. 2*0*)(2 r xMm (2.8) According to Main (1994), the unwrinkled curvature, c is given by Equation 2.9, where E* is the resultant elastic modulus. **0*21rEmc (2.9)

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21 Substituting Equation 2.8 into Equation 2.9 and approximating curvature as 22dxyd where x and y are coordinates, yields Equation 2.10. 22*3)(1dxydErxMc (2.10) This leads to the final relationship, Equation 2.11, which corresponds to the traditional beam bending relation with I* representing the resultant area moment of inertia. Note the units of resultant area moment of inertia are cubic which corresponds to using the stress resultant, *, and resultant elastic modulus, E*. **22)( I E xMdxyd where by association (2.11) 3*rI This differential equation describes the bending behavior of an inflated circular tube. It will be integrated to determine deflection, y, for a specific set of boundary conditions. This theoretical deflection solution is then used to verify simulation deflection results in Chapter 4. 2.3 Case Specific Non-wrinkled Inflatable Theory 2.3.1 Bending Moment Figure 2.4 shows the assumed drag force loading, w, and model boundary conditions, with L representing the tube length. The specific bending moment, M(x), for this situation is derived from free body diagrams of a cut section of the beam as shown in Figure 2.5. Shear, V, and moment, M are shown as reaction forces. Variable x is used as a local coordinate along the tube length. Equations 2.12 and 2.13 show the sum of moments about point A assuming static equilibrium.

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22 Internal pressure, P LFwdrag Figure 2.4 Drag loading assumed as a distributed force wx M V A 2wL 2x 2x Figure 2.5 Free body diagram of beam section 02)()2( MxwxxwLMA (2.12) )(2)(2xLxwxM (2.13) 2.3.2 Vertical Deflection To solve for beam deflections, the specific moment equation, Equation 2.13, is substituted into Equation 2.11 and integrated twice, yielding Equations 2.14 through 2.16. )(223*22xLx r E wdxyd (2.14)

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23 )32(21323*CxLx r E wdxdy (2.15) )126(221433*CxCxLx r E wy (2.16) With boundary conditions at x = 0, y = 0, and at x = L, y = 0, constants of integration, C 1 and C 2 are found to be Equations 2.17 and 2.18 01 C (2.17) 1232LC (2.18) Substituting Equations 2.17 and 2.18 into Equation 2.16 yields the vertical deflection solution shown as Equation 2.19. )12126(23433*LxLx r E wy (2.19) 2.3.3 Wrinkle Length Wrinkle length is the length of the tube at which wrinkling first occurs. This length depends on the moment necessary to initiate wrinkling. Comer and Levy first determine the moment necessary to initiate wrinkling by equating expressions for m (Comer & Levy, 1963). Equation 2.20 describes a force balance between the internal pressure acting over the area of the circular end-plate and the longitudinal stress integrated over the circumference of the tube. (2.20) 0*22drrp Substituting Equation 2.5 into Equation 2.20 and integrating yields Equation 2.21, the first expression for m

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24 ]cos)([sin2)cos1(0000 tprm (2.21) Substituting Equation 2.5 into the balance of moments, Equation 2.7, yields the second expression for m )2sin22()cos1(20020 trMm (2.22) Combining Equations 2.21 and 2.22 by eliminating m and setting 0 = 0, yields Equation 2.23, the moment necessary to initiate wrinkling. 23rpM (2.23) Equating the moment to initiate wrinkling, Equation 2.23, with the value of the maximum moment in Equation 2.13 yields Equation 2.24. 28)2(32maxrpwLLMM (2.24) Solving for L and setting the load, LFw gives the general expression for wrinkle length, Equation 2.25. The wrinkle length is the length at which wrinkling first occurs in an inflated tube subject to bending, for a given pressure and flow rate. FrpLwr34 (2.25) As expected, the wrinkle length will be reduced for high loading force, F, and proportional to the internal pressure and the tube radius. Tube radius, r, is the largest contributing factor to wrinkle length for these conditions.

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25 For this specific application, the loading force, F, is the drag force, F drag Since drag force is a function of tube length, Equation 2.25 is further expanded using Equations 2.1 and 2.3 to obtain Equation 2.26. 224VrpLwr (2.26) Equation 2.26 now shows the wrinkle length as a function of the core contributing variables for this application; internal pressure, tube radius, and external fluid velocity. The wrinkle lengths in Equations 2.25 and 2.26 also depend on the boundary conditions and are derived for the simply-supported conditions shown in Figure 2.4. 2.4 Conclusion The equations derived in this chapter provide a theory for an inflated beam subject to bending. The solution for vertical deflection, Equation 2.19, is used to verify the finite element model developed in Chapters 4 and 5. The wrinkle length is used to verify the final results of a verification model. Verifying a simple model is necessary to justify extending the computer analysis to more complex geometries, not easily solved with traditional analysis.

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CHAPTER 3 EXPERIMENTAL PROCEDURE AND RESULTS The setup, procedure, and results of the experiments performed to determine the orthotropic material properties of a specific commercial drainage tubing is presented in chapter 3. The Kuriyama Vinyl-Flow series is a composite comprised of a combination of molded PVC layers and synthetic fibers. One fiber layer runs in the longitudinal direction while two other layers are oriented at approximately +30 and -30 from the longitudinal direction as shown in Figure 3.1. Figure 3.2 shows specifications given by the Kuriyama for the different size models of Vinyl Flow. For this application, the desired material properties are the elastic modulus and Poissons ratio. These material properties are determined experimentally and to be used in a numerical simulation as described in Chapter 4. Longitudinal Direction +30 -30 Figure 3.1 Vinyl-Flow internal fiber layer orientations 26

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27 Figure 3.2 Manufacturer information for Kuriyama Vinyl-Flow series 3.1 Experimental Setup and Procedure 3.1.1 Test Sample Orientation Due to the orthotropic fiber orientation, a value for elastic modulus in both the longitudinal and hoop directions is desired. To achieve this, samples were made to have the fibers aligned in a specific orientation. Figure 3.3 shows the convenient coordinate system chosen and the orientation of each sample. Also shown are Fx and Fy, the tensile forces applied to each orientation sample. Longitudinal, y Hoop, x Longitudinally Oriented Sample Hoop Oriented Sample Tubing Coordinate System Fy Fx Figure 3.3 Test sample orientations with respect to continuous tubing

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28 3.1.2 Sample Shapes and Correction for Fiber Discontinuity In order to accurately model the materials performance, a tensile test sample shape was needed. It was desired to simply use a continuous tube sample but it was not possible to achieve this shape in the hoop direction. It was also inconvenient to test a continuous tube for the large diameter tubes due to the grip size of the tensile test machine. Since all the samples contain discontinuous fibers, it is necessary to determine the numerical effect on elastic modulus. To do this, three different sample shapes are tested and compared to see how the width and continuity of the fibers affect the elastic modulus. The first sample, shape A, is a completely continuous length of tubing with a rectangular shape. The second sample, shape B, is the same as shape A, but differs only with a single longitudinal cut running the entire length of the specimen. The cut severs the continuity of the 30 cross fibers. Figure 3.4 shows the longitudinal cut on the circular cross-section, representing shape B. In order to fit the sample completely within the width of the test grips, samples A and B are folded twice along the longitudinal direction as shown in Figure 3.4. The third sample, shape C, is cut from a template with a shape mimicking a tensile test sample. The ends are designed to fit the width of the grips of the testing device. The center section is designed to taper from the grips to a narrow strip to control break location. Care was taken to keep the tapering smooth to avoid stress concentrations in sharp corners. Figure 3.5 shows an example of shape C. All samples are 23 cm long. Sample C is 2 cm wide at the narrowest section and samples A and B have the same width when folded. Figures are shown with the same standard 5 pen for scale.

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29 Figure 3.4 Sample shape B (shown flat and folded) Figure 3.5 Sample shape C 3.1.3 Experimental Procedure An MTI 25k screw-driving tensile test machine was used with Curtis 30k self-cinching jaw grips to perform the tensile tests. An Interface 1220AF-25k load cell was used in conjunction with MTI software to measure force. After observing preliminary tests, a test speed of 0.10 in/min was considered reasonable and chosen for all tests. Figure 3.6 through 3.8 show examples of the test equipment. Figure 3.9 shows an example of a specimen locked in the grips before pulling. A sample is closed in the test grips and pulled at a constant velocity to the initial breaking point. Two marks were made in the middle region of the sample oriented perpendicular to the length. The length between marks and width (width for sample shape C) were measured and recorded intermittently throughout the pull test with digital calipers. The pull force on the specimen is measured and recorded simultaneously with the length and width measurements. The instantaneous axial stress is calculated at each point by dividing the force by the cross-sectional area. The cross-sectional area is calculated as the product of the width and the thickness of the test strip. The material

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30 width and thickness is assumed to be constant in Samples A and B. Only the thickness is assumed constant in Sample C since the varying width was recorded. The instantaneous axial strain, is calculated at each point using Equation 3.1. 00 (3.1) The variable, represents the recorded distance between marks and 0 represents the original distance between marks. This references strain calculations to the original unstressed length. An attempt was made to experimentally determine the material Poissons ratio by using measurements of the strain in the width and the strain in the length during the uniaxial pull tests. It was discovered, however, that since the cut samples do not have stiffness in compression, and the width decreases with the sample under tension, the width strain is not a valid measure to be used in determining Poissons ratio. To accurately describe Poissons ratio, the material must be loaded in two directions at once. Two possible solutions are to use a pull tester capable of pulling in two directions, or to internally pressurize a uniaxial test specimen. Given the complexity of these tests, Poissons ratio is instead determined in chapter 4, using a parameter study in an ANSYS simulation.

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31 Figure 3.6 Test equipment: MTI screw-drive tensile test machine (MTI, 2004) Figure 3.7 Test equipment: Interface 1220AF-25k load cell Figure 3.8 Test equipment: Curtis 30k self-cinching grip

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32 Figure 3.9 Example of sample locked in grips before pull test 3.2 Sample Shape Choice Using the analysis described in section 3.3, data from the sample shape tests is used to determine material elastic modulus. Five specimens were tested for each sample shape. Table 3.1 shows a summary of the experimentally determined elastic moduli for each sample shape. It can be concluded that neither sample B nor sample C closely represent the full tube structure. This is due to the breakage of the continuous hoop fibers that form the structure of the fabric. Since breaking these fibers is unavoidable in hoop orientation tests, the assumption was made to apply a multiplicative correction factor to all tests made from cut samples. The correction factor is determined from a ratio of the elastic modulus of the continuous tube, sample A, to the elastic modulus of the cut tube, sample B. This correction factor of 1.29 is applied to all future experimental results,

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33 including both the longitudinal and hoop oriented samples. Sample shape C is used for all future testing for convenience and predictability. Table 3.1 Experimental elastic modulus results for three test sample shapes Sample Shape and Description Orientation Elastic Modulus (psi) A Folded Continuous Tube Longitudinal 49009 B Folded Cut Tube Longitudinal 38148 C Strip from Template Longitudinal 36098 3.3 Data Analysis and Results This section describes the analysis of the data collected and its application towards the orthotropic material properties. Figures 3.10 and 3.11 show plots of axial stress and axial length-strain for VF-600 material in hoop and longitudinal orientations, respectively. Five samples per orientation were tested and are plotted with distinct linear trends. In order to interpret one linear trend from all five tests, all the tests are grouped together with one trend as shown in Figures 3.12 and 3.13. Combining five distinct trends is justified by looking at the slope and y-intercepts of the trends. The trend from each test has a distinct slope, but they are all close enough to consider within experimental error. The y-intercepts between tests, however, are largely different. The y-intercept of each trend line indicates an initial load present on the sample before any length-strain is recorded. This is explained by the observed behavior of the test grips during the test. The self-cinching grips initially sink in and tighten around the soft material for a brief period of time before a length-strain is present in the entire sample. This settling period is interpreted as a pre-loading on the sample by the load cell, resulting in the y-intercept in the data trend. The presence of the y-intercept is a result of

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34 the specific sample and test apparatus, so it is therefore assumed inconsequential to the experimental material properties. Since the slope is the only valid material property obtained from the data, all five groups can be justifiably analyzed with one trend. The slope of the trend represents the elastic modulus for the corresponding sample orientation. Table 3.2 shows a summary of the experimental material property results. Shown are the experimental and adjusted orthotropic elastic moduli for the hoop and longitudinal orientations. Adjusted elastic moduli are based on the correction factor for breaking the continuity of the fibers, determined in section 3.2. It is interesting to note that the elastic modulus for the longitudinal orientation is much larger than for the hoop orientation since the hoop stress is generally larger than the longitudinal stress. This is best attributed to the fiber layer present in the longitudinal orientation. The material was most likely not designed to be optimized as a pressure vessel, but rather designed for ease of the manufacturing process.

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35 0200400600800100012001400160018000.00000.05000.10000.15000.20000.2500Length Strain (in/in)Axial Stress (psi) Test 1 Test 2 Test 3 Test 4 Test 5 Figure 3.10 Experimental data: five hoop orientation pull samples 0500100015002000250030003500400045000.00000.02000.04000.06000.08000.10000.1200Length Strain (in/in)Axial Stress (psi) Test 6 Test 7 Test 8 Test 9 Test 10 Figure 3.11 Experimental data: five longitudinal orientation pull samples

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36 y = 5798.4x + 477.20200400600800100012001400160018000.00000.05000.10000.15000.20000.2500Length Strain (in./in.)Axial Stress (psi) Figure 3.12 Experimental data: Five hoop-orientation pull samples grouped together y = 36098x + 248.180500100015002000250030003500400045000.00000.02000.04000.06000.08000.10000.1200Length strain (in/in)Axial Stress (psi) Figure 3.13 Experimental data: Five longitudinal-orientation pull samples grouped together

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37 Table 3.2 Summary of experimental and adjusted material properties Material / Orientation Experimental Elastic Modulus (psi) Adjusted Elastic Modulus (psi) VF-600 / Hoop 5798.4 7479.9 VF-600 / Longitudinal 36098.0 46566.4 3.5 Errors and Accuracy Unavoidable experimental errors occurred in all the tests. Due to the nature of the cross-hatched fabric material, in-plane and out-of-plane shape distortions occurred in samples when loaded under tension. The in-plane distortions can be described as uneven shifting of the material in the pull direction. The strain measurements were largely affected by the in-plane distortion because the measurement lines did not remain perpendicular to the sample length. The out-of-plane distortions can be described as cupping of the flat, rectangular cross-sectional area. This largely affects the width and, in some cases, thickness measurements. Tiny slip marks were observed on some tested samples where the self-cinching grips cut through or slipped on the slick outer material. This would contribute additional error to the strain measurements. Figure 3.14 shows an example of the lines visible on a tested pull sample that indicate slipping. In some cases, manufacturing inconsistencies in the material were exposed when material samples visibly broke in a location other than the expected thinnest portion of the sample. Human error in preparing samples, taking length measurements, and reading the force gage also contributed to experimental error. The wide scattering of the data points in the stress-strain curve can also be attributed to the settling period described in section 3.3.

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38 These experimental errors all contribute to the wide scattering of the data points in the stress-strain curve. Due to the nature of the composite, flexible material and the test equipment used, the experimental values are considered a valid representation of the material. Pull Direction Grip slippageNormal Grip Wear Figure 3.14 Tested sample with grip slip marks

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CHAPTER 4 SIMULATION DEVELOPMENT The process of developing a finite element model to represent an actual prototype cage is described in Chapter 4. A simple verification model is first made to explore and verify finite element membrane analysis for an inflatable tube subject to bending. A full cage is then modeled to explore the interaction of all the components. Two flow direction cases are modeled to see the effect of flow direction on stress and deflection in the cage members. Figures 4.1 and 4.2 show a prototype cage inflated and deflated, respectively. Figure 4.3 shows an example of the union between the tubing and a PVC connector using a hose clamp. Figure 4.1 Cage prototype inflated to ~12ft diameter, 9ft height 39

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40 Figure 4.2 Cage prototype deflated to ~70cm PVC connector Hose clamp Flexible tubing Figure 4.3 Tubing union to PVC connector 4.1 Verification Model Development and Results A verification model is designed to establish a basis for further simulations. The verification model is a simply supported tubular beam with a distributed load on top. Figure 4.4 shows the conceptual beam model and distributed loading. The hose clamps are modeled by fixing the x and y displacements of all the nodes along the circumference of each end. In order to avoid rigid body motion in the analysis, the tube is constrained in the z direction. To avoid artificial boundary stress concentrations at the ends, the z-displacements of only 2 nodes on the neutral axis of bending are fixed. Figure 4.5 shows the applied boundary conditions and the coordinate system used.

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41 x z Figure 4.4 Conceptual boundary conditions and loading for the verification model Dz = 0 (one end of tube only) Nodes from outer ring falling on neutral axis of bending Dx = Dy = 0 (Both ends of tube) All nodes on the outer ring. X Y Z Figure 4.5 Detail of boundary conditions for verification model 4.1.1 Applied Loading Conditions The internal pressure is broken up into pressure acting in the radial direction and pressure acting longitudinally. The radial pressure is applied normal to the interior tube surface. The longitudinal pressure is converted to nodal forces, F i using steps shown in Figure 4.6. Figures 4.6-A and 4.6-B show the internal pressure, P i applied over the area of the circular cap, A cap Assuming the cap is rigid, the applied pressure translates into an

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42 equivalent total axial force, F eq acting on the cap as shown in Equation 4.2 and Figure 4.6-C. This total force is then distributed equally and applied directly to the number of nodes, n, on the outer circular edge at the end of the tube as shown in Equation 4.3 and Figure 4.6-D. capieqAPF (4.1) (4.2) niieqFF1 nFFeqi (4.3) Rigid Cap Fabric Tube Rigid Cap Pi Feq niiF1 Acap = r2 (A) (B) (C) (D) Pi Figure 4.6 Steps in representing internal pressure as pre-tension forces 4.1.2 Water Current Drag Water flowing over the outer surface of the tube creates a drag force. Fluid assumptions and equations are derived in Chapter 2. A programming do-loop was

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43 initially implemented into a script file to approximate a parabolic pressure flow distribution by applying individual forces to nodes. Forces are determined from the product of the local pressure and the projected area around the node. Figure 4.7 shows a representative projected area for one node. Applying the pressure distribution to the tension-only membrane tube in this manner yielded numeric instability. This numeric instability and a need for only global effects of the drag leads to the assumption that local pressure effects of the flow on the tube can be neglected. This simplifies the model by approximating the total drag as an equally distributed load. Figure 4.8 shows an example of the total drag force divided equally and applied to all the nodes along the length of a typical tube. All nodal drag forces are applied in the same global coordinate direction. Figure 4.7 Representative projected area around one surface node

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44 Forces applied at each node around circumference. Equal forces applied at each node around circumference Figure 4.8 Representative application of forces to nodes around circumference 4.1.3 Element Selection and Properties The geometry is created in ANSYS by extruding a circle along a linear path. This area is then meshed with rectangular shell elements. The material to be modeled is a thin fabric with orthotropic material properties. Since fabrics have tensile stiffness, but very little compressive stiffness, a thin shell element with optional compressive stiffness is the best choice. The Shell41 element with Keyopt(1) = 2 meets these requirements. It was determined that the zero compressive stiffness caused a numerical instability in the solution of a pressurized beam in drag which resulted in spikes in the deformed shape. Initially, a small elastic foundation stiffness (EFS) was added to the Shell41 elements to attempt to fix the instability. Successive ANSYS simulations varying the EFS showed a large change in mean deflection, shown in Figure 4.9. This large deflection makes choosing a small EFS value unacceptable and eliminates the use of the EFS in all models. In another attempt to remedy the instability, equivalent-geometry Shell63 elements with small elastic modulus were merged to the Shell41 elements to add a small compressive stiffness to the elements. Successive ANSYS simulations varying the added

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45 stiffness of the shell63 elements showed an acceptable change in deflection and stress as shown in Figures 4.10 and 4.11. A value of 10,000 Pa is chosen for the added stiffness. Merging the shell63 elements is a preferable choice to fix the instability because the added stiffness has a less interfering effect on the desired results. Implementing this choice eliminates the spikes and smoothes out the deformed shape. The isotropic material properties needed to solve for stress and deflection are elastic modulus and Poissons ratio. Since the material is orthotropic, hoop elastic modulus, longitudinal elastic modulus, shear modulus, and Poissons ratio are required. Values for elastic modulus in the hoop and longitudinal directions are found experimentally in Chapter 3. Since Poissons ratio is not experimentally determined, a value is chosen based on finite element model results. Simulations showed that varying Poissons ratio had negligible effect on longitudinal stress and deflection. Figure 4.12 shows maximum deflection plotted for various Poissons ratios determined in an ANSYS model of a cantilevered tube loaded with an end force. Figure 4.13 shows a similar plot for maximum longitudinal stress using the same model. The small variation in stress and deflection justifies assuming a reasonable value of 0.4 for Poissons ratio.

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46 -1.40000E-02-1.20000E-02-1.00000E-02-8.00000E-03-6.00000E-03-4.00000E-03-2.00000E-030.00000E+00020000400006000080000100000120000140000EFS Value (Pa)Mean Deflection (m) Figure 4.9 Verification model parameter study: EFS effect on beam deflection -8.00E-03-7.00E-03-6.00E-03-5.00E-03-4.00E-03-3.00E-03-2.00E-03-1.00E-030.00E+00020000400006000080000100000120000140000Added Compressive Stiffness to Shell41 Elements (Pa)Mean Deflection (m) Figure 4.10 Verification model parameter study: Added stiffness effect on beam deflection

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47 0.00E+001.00E+062.00E+063.00E+064.00E+065.00E+066.00E+067.00E+06020000400006000080000100000120000140000Added Compressive Stiffness to Shell41 ElementsLongitudinal Stress (Pa) Maximum Stress Minimum Stress Figure 4.11 Verification model parameter study: Added stiffness effect on longitudinal beam stress 00.00250.0050.00750.010.01250.0150.01750.020.02250.02500.050.10.150.20.250.30.350.40.45Poisson's Ratio, vxy (v_hoop-longitudinal)Maximum Vertical Deflection of Free End (m) Figure 4.12 Verification model parameter study: Poissons ratio effect on beam deflection

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48 0.00E+001.00E+062.00E+063.00E+064.00E+065.00E+066.00E+0600.050.10.150.20.250.30.350.4Poisson's Ratio, vxy (v_hoop-longitudinal)Longitudinal Stress (Pa) Maximum Stress Minimum Stress Figure 4.13 Verification model parameter study: Poissons ratio effect on longitudinal beam stress 4.1.4 Verification Model Results To verify the model, simulation results are compared to the stress and deflection theoretical solutions derived in Chapter 2. Since flow velocity is proportional to the drag force acting on the beam, flow velocity is used to compare models. The simulations are conducted until the beam reaches failure. A study is done for 5m and 10m tubes. Results from the 5m long tube are shown in Figure 4.14 which correspond to a constant average deviation of 3.22% between theoretical and simulation. Figure 4.15 shows similar results for a 10m tube with a constant average deviation between theoretical and simulation to be 0.567%. These results show that the deflection results between theoretical and simulation stay below an acceptable 5% deviation. Results also show that the deviation decreases as tube length increases. A length study was also performed to see how maximum deflection changed with length. Simulations are performed up to an approximately 10% deflection:length ratio.

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49 Figure 4.16 shows a parabolic trend in the simulation and theoretical results. Figure 4.17 plots the deviation between simulation and theoretical results, showing deviation decreasing exponentially as length increases. Deviation drops below 10% above a tube length of approximately 3.5m. The large deviation occurring in tubes shorter than 3.5m is believed to be due to the Saint-Venants principle (Hibbeler, 1997), applicable at the fixed end boundary condition. In this case, the simulated cap restricts radial expansion in the fabric tube, which affects the maximum deflection. A stress simulation is also performed to verify the model behavior. Figure 4.18 shows longitudinal stress results for the verification model. The maximum stress increases, while the minimum stress decreases as expected in bending. A rigid body motion error is received when the minimum longitudinal stress decreases to approximately zero, verifying the tension-only behavior of the model. This defines failure by wrinkling in this model.

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50 -5.00E-01-4.50E-01-4.00E-01-3.50E-01-3.00E-01-2.50E-01-2.00E-01-1.50E-01-1.00E-01-5.00E-020.00E+0000.511.522.53Flowing Water Speed (mph)Average Y-Deflection (m) Simulation Theoretical Figure 4.14 Deflection results for the verification model with length of 5m -1.80E+00-1.60E+00-1.40E+00-1.20E+00-1.00E+00-8.00E-01-6.00E-01-4.00E-01-2.00E-010.00E+0000.20.40.60.811.21.41.6Flowing Water Speed (mph)Average Y-Deflection (m) Simulation Theoretical Figure 4.15 Deflection results for the verification model with length of 10m length

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51 -1.00E+00-9.00E-01-8.00E-01-7.00E-01-6.00E-01-5.00E-01-4.00E-01-3.00E-01-2.00E-01-1.00E-010.00E+00012345678910Tube Length (m)Mean y-deflection (m) Simulation Theoretical Figure 4.16 Deflection results for verification model with 0.5 m/s flow velocity 0.00%20.00%40.00%60.00%80.00%100.00%120.00%140.00%012345678910Tube Length (m)% Difference in Deflection Figure 4.17 Deviation in deflection results for verification model with 0.5 m/s flow velocity

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52 -2.00E+060.00E+002.00E+064.00E+066.00E+068.00E+061.00E+071.20E+071.40E+0702468101214Tube Length (m)Longitudinal Stress (Pa) Maximum Stress Minimum Stress Figure 4.18 Longitudinal bending stress results for verification model with 0.5 m/s flow velocity 4.1.5 Verification Model Conclusion The results from the verification model show the merged shell41 and shell63 elements acceptably model the tension-only tubing material. The Saint-Venants principle effect is specific to the boundary conditions and geometry of the verification model and is not applied elsewhere. The drag loading and boundary condition applications are acceptable and accurately model the actual conditions. 4.2 Cage Simulation Process The actual tubes, when assembled as a cage, have rigid end caps with hose clamps constraining each end. Based on the geometry of the cage, the tube can flex and extend longitudinally. Internally, the tube contains pressurized air to maintain inflation. The external loading on the tube is a distributed drag force caused by a cross-flowing water current.

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53 4.2.1 Flow Design Models With infinite possible current flow-directions for an underwater cage, the number of modeled choices is limited to cages subject to top-down flow and side-flow. Figure 4.19 defines the global coordinate system with side-flow in the positive x-direction and top-flow in the negative y-direction. Modeling the entire cage was necessary because unlike the single beam verification model, the boundary conditions that exist at the end of each tube in the full cage structure are complex and interactive among the cage members. Figure 4.19 Two cage simulation flow orientations 4.2.2 Element Choice All tubes are meshed with merged tension-only shell41 and shell63 elements in the same stacked element configuration as the verification model determined in Section 4.1.3. The tubular connectors in the cage are rigid relative to the tubing, so shell63 elements are used. Link10 elements model the tension-only cables that tie together certain connectors throughout the cage. 4.2.3 Loading Conditions Internal pressure in the fabric tubes acts in the radial and longitudinal directions. Pressure acting in the radial direction is applied using the surface element pressure

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54 command (sfe) to only the shell41 elements making up the fabric tubes. Since the tubes are not actually capped, as in the verification model, a control volume analysis of the fluid in the rigid connector is used to determine the net effect of the longitudinal pressure in each tube. Figure 4.20 shows a control volume of fluid inside a typical 45-connector. Longitudinal pressure from each tube acts over the entrance area to the connector. The product of the pressure and area is treated as an equivalent force, F cap similarly to Equation 4.1 in the analysis of the verification model. Assuming the compressibility of the fluid is negligible, and the fluid is static, the fluid and the connector are approximated as one rigid body. Figure 4.21 shows Figure 4.20 translated into a free body diagram of a typical 45-connector modeled as a rigid body with net x and y component forces. Figure 4.22 shows a finite element model of the same typical 45-connector with the net component cap forces divided and applied to all the nodes around the circumference of the center of the connector. The same analysis and nodal force application is used for each connector. Figure 4.20 Control volume analysis around a rigid 45-connector containing pressurized internal fluid

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55 Figure 4.21 Free body diagram of rigid cap with algebraically summed pretension force components Figure 4.22 Force components applied to finite element model of rigid connector 4.3 Top-Down Flow Cage Model In the top-down flow model, the center tube is anchored at the top of the cage with flow in the opposite direction. Small frictional drag occurring along the center tube is neglected because it lies in the direction of flow. To save on computing time, the center tube is eliminated from the model and replaced with a single node at the top and bottom of the cage. Although not immediately obvious, symmetry could be not be used to simplify the top-down flow cage model analysis. Dividing the cage along the x, y, or both the x and y axes does not work because the state of bending makes the z-direction boundary

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56 conditions along the axis of symmetry unknown. The z-direction boundary conditions also make it impossible to use an axis of symmetry along an axis 45 degrees from the x or y axis. The cable element controlling the z-direction boundary condition lies along this axis of symmetry and can not be split. 4.3.1 Applied Loading Conditions Since external fluid flows over the entire cage, the total fluid drag force on the cage is calculated based on its total cross-sectional area perpendicular to the direction of flow. The total force is then divided equally and applied to all nodes of the eight caps and eight tubes. Referring to the coordinate system defined in Figure 4.23, all drag forces act in the negative z-direction for the top-down flow model. 4.3.2 Boundary Conditions Boundary conditions are applied to simulate the conditions acting on the cage as simply as possible while minimizing solution time. The two locations labeled A in Figure 4.23 describe two single nodes at the top and bottom of the cage. These points are placed to obtain a cage diameter-to-height ratio of 8:5. Both points are constrained to zero deflection in the x,y, and z directions. Cables on the underside of the cage are slack and do not affect the system, but were included for aesthetics along with the fixed node at the bottom of the cage. Locations labeled B and C in Figure 4.23 describe all nodes around the circumference of a tube at the center of its length. Nodes at location B are constrained to zero deflection in the y-direction. Nodes at location C are constrained to zero deflection in the x-direction. Constraining locations B and C prevents rotational rigid body motion while allowing the individual tubes and cage to expand symmetrically.

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57 Figure 4.23 Top-flow boundary condition locations 4.4 Side-Flow Cage Model The side-flow cage starts with the same octagon model as the top-down flow model but includes a model of the center tube, all the cables, and a different anchoring system. Z-direction pretension forces from internal pressure are additionally included at the ends of the center tube in this model. The rigid connectors are pressurized to the same internal pressure as the connecting tubing because successive analysis showed an improvement in the smoothness of the deformation solution around the connections. Symmetry could also not be used in the side-flow cage model analysis. The x-axis would be the best choice for axis of symmetry, but the center tube, lying in the z-direction, would have to be split longitudinally. Also, the x-direction boundary conditions are unknown on the tubes along the axis of symmetry because the drag flow is causing bending at those points. Figure 4.24 shows the components of the center tube as two flexible membranes and three straight, rigid connectors. The total center tube length is determined by the 8:5 cage diameter-to-height ratio. Figure 4.25 shows link10 elements connecting nodes in

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58 the top, bottom, and center connectors of the center tube to nodes in the top and bottom of the 45-degree connectors in the octagon. Cable elements are merged to single nodes on the rigid connector symmetrically as shown. Four tension-only cables anchor the cage from a single external point. Figure 4.26 shows an example of the link10 mooring cable elements merged to one top and bottom node of a 45-degree connector. Figure 4.27 shows the anchor point location in the global coordinate system, where L is the length of an individual tube in the octagon. Setting the anchor point x-location as a function of tube length maintains an anchor distance from the cage as the cage size is increased. This location is chosen to be reasonably close, while minimizing the y-tension component in the anchor cables, which affects the deformation of the cage. 4.4.1 Applied Loading Conditions Drag flow for the side-flow case is approximated as a series of forces applied to specific cage members. Locations A, B, and C in Figure 4.28 describe flexible membrane tubes to which drag is applied. The magnitude of drag on locations A, B, and C is individually based on the respective cross-sectional area. Small frictional drag is neglected along location D in Figure 4.28 because the members are parallel to the flow direction. The total drag force for each individual member is divided equally and applied to the nodes corresponding to the respective location. Referring to the coordinate system defined in Figure 4.23, all drag forces act in the positive x-direction for the side-flow model. 4.4.2 Boundary Conditions The location labeled A in Figure 4.29 describes the single-node anchor point fixed to zero deflection in the x, y and z directions. To allow symmetric expansion and prevent rotational rigid body motion, constraints are placed in two more locations in the cage.

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59 Locations labeled B and C in Figure 4.29 describe all nodes around the circumference of a tube at the center of its length. The nodes at location B are constrained to zero deflection in the y-direction, while nodes at location C, further illustrated in Figure 4.30, are fixed to zero deflection in the z-direction. Flexible Membrane Tubing Rigid Connectors Figure 4.24 Cage center tube components

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60 Figure 4.25 Expanded view of representative internal cable connections Mooring cables Rigid connector Flexible membrane tubing Figure 4.26 Expanded View of representative mooring cable connections

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61 +X +Y 3.4L L/2 L/2 Figure 4.27 Dimensions of anchor location Figure 4.28 Side-flow applied loading locations

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62 B B A C Figure 4.29 Side-flow boundary condition locations Flexible membrane tubing Rigid connector Displacement boundary conditions (nodes of entire circumference) Figure 4.30 Expanded view of center tube boundary conditions

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CHAPTER 5 SIMULATION RESULTS The design criteria and results from the ANSYS cage simulations are presented in Chapter 5. The goal of the study is to determine the maximum cage size possible before failure. Twelve commercially available flexible tubes are tested under top and side flow conditions. The material yielding the largest cage is chosen and subject to an additional velocity simulation. 5.1 Simulation Strategy ANSYS script files are used to parameterize the variables involved in the construction and analysis of each cage. Independent variables were kept constant for all analyses run in this chapter. Independent variables include: External fluid temperature, viscosity, density Flexible membrane orthotropic elastic modulii Rigid connector elastic modulus Flexible membrane and rigid connector Poissons ratios Flexible membrane shear modulus Cable constant, EA Rigid Connector geometry Cable length is defined by the cage geometry. Flexible membrane material properties are explained in Chapter 3, while reasonable values are chosen for external fluid, connector, and cable properties. Dependent variables affecting the cage geometry and loading are: Flexible membrane internal pressure Flexible membrane tube length Flexible membrane tube diameter 63

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64 Flexible membrane thickness External fluid velocity Internal pressure causes the pretension forces in the flexible membrane tubes. The pretension forces are also based on the area of the tube opening and so are functions of the tube diameter. Drag on a given member is a direct function of tube length and tube diameter. Drag is also a function of the flowing fluid velocity squared. Three of the five independent variables are combined using data from commercially available materials. Table 5.1 presents Vinyl-Flow data including tube diameter, thickness, and working pressure for a given series number. Simulation results make convenient the grouping of VF-Series by pressure to show approximately constant pressure trends. The three pressure range groups are indicated in Table 5.1. The following simulations vary Vinyl-Flow series number, tube length, and external fluid velocity to determine effects on cage performance. Table 5.1 Manufacturer data for Vinyl-Flow commercial drainage tubing 5.2 Failure Criteria The failure criteria for both flow cases are the points when wrinkling occurs anywhere in the fabric tubing and when stress exceeds the ultimate strength of the

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65 material. Wrinkling is defined here as a compressive stress. Cage wrinkle diameter is defined as the maximum cage diameter before any flexible tube wrinkles due to a compressive stress less than or equal to zero. It is important to consider both the longitudinal and hoop stresses in the analysis. Wrinkling is expected to occur in the longitudinal direction because the combined loading of internal pressure and bending affect the longitudinal stress. Tubes are not expected to wrinkle in the hoop direction because the only loading affecting the hoop stress, internal pressure, creates a tensile hoop stress. The maximum stress will occur in the longitudinal direction because the combined loading of internal pressure and bending affect the longitudinal stress. The hoop stress is affected only by the internal pressure loading. Using the manufacturers recommended working pressure ensures the hoop stress will not exceed the ultimate strength of the material. Figure 5.2 illustrates a representative case of compressive failure behavior for VF-800 tested at a water speed of 1 knot in top-flow. Maximum and minimum longitudinal stresses are shown for various cage sizes. As cage size is increased, maximum stress increases, while minimum stress decreases. Pretension in the tubes keeps both stresses positive, meaning they are always in tension. When the cage size increases to the point that the minimum stress drops below zero, the model fails because the tension-only elements can not support compressive stress. This is then interpreted by the program as an unconstrained model and results in a rigid body motion error message. Note that the compressive stress could go very slightly negative without failure because of the small compressive strength provided by the merged shell63 elements comprising the tube.

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66 Subsequent top-flow simulation failures are judged by the rigid body motion error message. The side-flow simulations do not show a smooth trend in the longitudinal stress down to zero as in Figure 5.2, but show a sudden drop-off resulting in a rigid body motion error message. This behavior can be described as a buckling of the structural tubes. The stress behavior is due to changes in the higher order bending occurring in the tubes as cage size changes. Section 5.4 illustrates and further explains this behavior. The side-flow cage failure is therefore also determined by the rigid body motion error message in the simulation. 0.00E+002.00E+064.00E+066.00E+068.00E+061.00E+071.20E+07051015202530Cage Diameter (m)Longitududinal Tube Stress (Pa) Maximum Stress Minimum Stress Figure 5.2 Simulation results: Bending stresses in a VF-800 top-flow cage 5.3 Stress Results Figure 5.3 illustrates the maximum longitudinal stresses occurring in each simulation staying below the experimentally determined ultimate strength of the material, 2.6e7 Pa. This ensures the material strength is not exceeded in any simulation.

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67 0.00E+005.00E+061.00E+071.50E+072.00E+072.50E+073.00E+07VF 150VF 200VF 250VF 300VF 400VF 500VF 600VF 800VF 1000VF 1200VF 1400VF 1600Vinyl-Flow Series NumberLongitudinal Stress (Pa) Top-FlowMaximum Stress Side-FlowMaximum Stress Ultimate MaterialStrength Figure 5.3 Maximum stresses in each simulation and ultimate material strength 5.4 Geometry Deformation Results Figures 5.4 through 5.6 show various views of a typical deformed geometry of a cage subject to top-flow at the critical wrinkle point. All eight tubes bend in the negative z-direction with maximum z-deflection occurring at half the tube length. Bending plus the Poisson effect cause the cage diameter to decrease symmetrically throughout the cage. Figures 5.7 through 5.8 show two views of a typical deformed geometry of a cage subject to side-flow at the critical wrinkle point. Note that the net cage x-diameter increases, while the net cage y-diameter decreases. Internal reaction forces occurring through the cable connections cause internal bending moments around the corner connectors of the octagon. Deformation effects of these internal bending moments are visible in the deformed curvature around the connectors and account for multiple-mode bending visible in several of the beams. Figure 5.9 illustrates the rotational effect on a

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68 representative rigid connector. The different rotational effects occurring throughout the cage contribute to the multiple bending modes and subsequent buckling stress behavior leading to failure. Figure 5.4 Top-Flow deformed geometry at the wrinkle point: View 1 X Y Deformed Undeformed Figure 5.5 Top-Flow deformed geometry at the wrinkle point: View 2

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69 Deformed Undeformed Figure 5.6 Top-Flow deformed geometry at the wrinkle point: View 3 Deformed Undeformed Figure 5.7 Side-Flow deformed geometry at the wrinkle point: View 1

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70 Deformed Undeformed Figure 5.8 Side-Flow deformed geometry at the wrinkle point: View 2 Undeformed Deformed Note rotational effect Figure 5.9 Representative rotational effect occurring in side-flow simulation

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71 5.5 Simulation Deflection Results Deflection results are useful in predicting cage behavior. Presented here are qualitative results for a specific material case and maximum deflection results for each material case. Numeric sign associated with deflection is related to the previously defined cage coordinate system. For diameter changes, a negative change indicates a decrease in diameter, while a positive change indicates an increase. 5.5.1 Qualitative Deflection Results Figure 5.10 shows deflection data of the net cage x-diameter deflection for a representative case of the VF-800 tested in top-flow at a water speed of one knot at various cage diameters. Deflection results are identical for the cage y-diameter based on the symmetry of the geometry and loading. Cage diameter change has a qualitatively linear trend for the top-flow cage loading. Figure 5.11 shows data of the maximum z-deflection for the same representative case. The data represents the maximum bending deflection occurring in each tube and has an expected qualitatively decreasing parabolic trend as cage size is increased. Figure 5.12 shows cage y-diameter deflection data for the VF-500 representative case, but in side-flow. As seen in the geometry deformation of the side flow cage, the cage y-diameter decreases drastically. No clear trend is visible in the smaller x-diameter deflection. This is because the multiple bending modes largely affect the x-diameter cross members and occur differently in each cage size. X-diameter cage deflections must be considered on a case by case basis for side-flow.

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72 -2.5-2-1.5-1-0.50051015202530Cage Diameter (m)Net Cage x-diameter Deflection (m) Figure 5.10 Cage x-diameter net deflection for VF-800 tested at flow speed of 1 knot, top-flow -5-4-3-2-10051015202530Cage Diameter (m) Figure 5.11 Maximum z-deflection for VF-800 cage tested at flow speed of 1 knot, top-flow

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73 -3.0-2.5-2.0-1.5-1.0-0.50.005101520Cage Diameter (m)Net Cage y-diameter Deflection (m) Figure 5.12 Cage y-diameter net deflection for VF-500 tested at flow speed of 1 knot, side-flow 5.5.2 Maximum Deflection Results Two tables are shown to present deflection behavior with respect to the original cage size. Data are summarized in table format because in each case, the maximum deflections occur at different cage sizes which correspond to different drag loadings. With different drag loadings, each case can not be directly compared to determine a graphical trend. Tube diameter and working pressure are included in the tables for comparison between cases. Table 5.2 shows the maximum diameter change and maximum z-deflection with respect to original cage diameter for the top-flow case. Diameter deflections are between 4-10%, while z-deflections are between 10-18%. Table 5.3 shows the maximum change in the x and y diameters with respect to original cage diameter for the side-flow case. X-diameter deflections are small and between 2-4%, while y-deflections are much larger, at 14-22%.

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74 Table 5.2 Maximum cage deflections for the top-flow orientation Series numberTube diameter (in.)Working pressure (psi)Maximum cage diameter (m)Ratio of change in cage x-diameter to original cage diameterRatio of z-deflection to original cage diameterVF 1501.673807.2-4.93%-10.78%VF 2002.165809.6-6.49%-14.98%VF 2502.5988010.8-6.49%-13.53%VF 3003.1307013.2-7.02%-15.83%VF 4004.1347016.8-8.71%-18.78%VF 5005.0394014.4-5.80%-10.57%VF 6006.1815020.4-8.83%-17.37%VF 8008.1694526.4-8.75%-17.47%VF 100010.1183527.6-7.78%-13.82%VF 120012.1263030-7.60%-13.59%VF 140014.1343034.8-8.85%-15.76%VF 160016.1423039.6-10.10%-17.92% Table 5.3 Maximum cage deflections for the side-flow orientation Series numberTube diameter (in.)Working pressure (psi)Maximum cage diameter (m)Ratio of net x diameter change to original cage diameterRatio of net y diameter change to original cage diameterVF 1501.6738014.42.498%-17.346%VF 2002.16580182.600%-18.633%VF 2502.5988020.42.598%-18.216%VF 3003.1307021.62.681%-18.361%VF 4004.1347027.63.138%-21.377%VF 5005.03940182.569%-14.222%VF 6006.1815021.62.731%-17.667%VF 8008.1694525.23.150%-18.619%VF 100010.11835242.672%-15.442%VF 120012.1263025.22.779%-14.976%VF 140014.1343028.83.262%-17.278%VF 160016.1423032.43.760%-19.605% 5.6 Cage Size Results Maximum possible cage size is determined by the point at which the cage wrinkles. Figures 5.13 and 5.14 show the maximum cage diameter results by tube diameter for the top-flow and side-flow cases respectively. Slight pressure effects are noted between the pressure groups of the top-flow case, but are neglected to form one linear increasing trend. The trend generally shows that, independent of pressure, as diameter increases, the maximum possible cage diameter increases for top-flow conditions. The side-flow case, however, has significant pressure effects visible in Figure 5.14. Each pressure group is plotted with an increasing linear trend, but a distinct slope and y-intercept. Within each

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75 pressure group, as diameter increases, the maximum possible cage diameter increases. Also evident is that as pressure increases, the slope of the constant pressure trends becomes steeper. This indicates that for a given diameter, operating at higher pressures increases the maximum possible cage size. Since the flow situations are idealized and true water conditions vary no matter how the cage is anchored, choosing the material for the largest cage is based on the results of both flow cases. Figure 5.15 presents the data from figures 5.13 and 5.14 as a summary of the top-flow and side-flow maximum cage diameters possible for all the material models tested. The lesser cage maximum diameter of the two flow cases is the limiting factor for each material. Refer to Table 5.1 for model specific manufacturer data including tube diameter, thickness, and working pressure for each case. 051015202530354045024681012141618Tube Diameter (in)Maximum Cage Diameter (m) 70-80 psi 40-50 psi 30-35 psi slight pressure effectslight pressure effect Figure 5.13 Top-flow maximum cage diameters for each tubing material grouped by internal pressure

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76 05101520253035024681012141618Tube Diameter (in)Maximum Cage Diameter (m) 70-80 psi 40-50 psi 30-35 psi Figure 5.14 Side-flow maximum cage diameters for each tubing material grouped by internal pressure 051015202530354045VF 150VF 200VF 250VF 300VF 400VF 500VF 600VF 800VF 1000VF 1200VF 1400VF 1600Vinyl-Flow Series NumberMaximum Cage Diameter (m) Side-Flow Top-Flow Figure 5.15 Summary of maximum possible cage size for each material in two flow cases subject to one knot flow velocity VF Series number corresponds to diameter, working pressure, and thickness in Table 5.1

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77 5.7 Velocity Results VF-1600 is chosen because it has the largest limiting cage size of the materials tested. The maximum possible VF-1600 cage diameters for top-flow and side-flow are averaged and then divided by two to obtain a cage diameter able to withstand higher than one knot flow velocity. This cage diameter, 19.2 m, is used for a velocity test on the cage. The cage withstands the same maximum flow speed of 1.6 knots in both top and side flow cases before wrinkling occurs. In top-flow, cage diameter decreases 10.183%, while in side-flow, cage x-diameter increases 4.027%, and cage y-diameter decreases 17.792%. These deflections are reasonable for an inflated flexible membrane cage of this size and correspond to the expected deformation results shown in Section 5.4.

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CHAPTER 6 CONCLUSION The experimental work performed gave preliminary values of the orthotropic material properties. More advanced test equipment is necessary to model orthotropic fiber tubing to find the directional elastic moduli and Poissons ratio. Large-deformation strain gages would improve the accuracy in recording the strain used in determining the elastic modulus. A two direction tensile tester would be necessary to experimentally determine Poissons ratio for the inflated-tube loading case. The values determined were sufficient for the preliminary design and modeling of the orthotropic fiber material. The analysis presented here focused on the inflatable behavior of the structure. The simulation included the inflated tube members, rigid PVC connectors, and cable connectors. It is important to note that the fish netting is excluded from the simulation, but is necessary for a complete analysis. The simulation effectively models the nonlinear behavior of the orthotropic, fabric material by translating the inflatable theory into a finite element model. Preliminary simulations successfully show expected inflatable behavior in full-cage top and side flow models. Preliminary findings warrant further research that models the external netting. Net selection has several biological issues to be addressed such as net biofouling and biologically safe mesh size. Net analysis will potentially require dynamic analysis to model the hydrodynamic interaction between the structural element and the fluid (Tsukrov et al., 2000). Since the cage divergence volume is an important issue, future work should include at least an approximation of the nets. Further simulation work 78

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79 should also include optimizing the inflatable material. The cage materials size and stiffness is expected to offset the increased loading of the nets, making the inflatable aquaculture cage practical and economical. The side and top flow models behave differently, particularly with respect to divergence volume. The internal pressure and tube diameter play the largest role in deciding the maximum possible cage diameter prior to wrinkling. Instead of including a safety factor in the final results, the failure criterion was set to the point at which wrinkling occurred in any tube. Inflatable tubes do not actually fail at the point of wrinkling, but can continue past this point (Main et al., 1994). Future simulations may be able to push the failure criteria by modeling the stiffness behavior of the wrinkled material. This work successfully models one new application of inflatable structures. Nonlinear finite element modeling is shown to be an effective tool in analyzing the behavior of these structures. Inflatables in many applications have a great potential for improving cost, weight, and convenience. This effort adds one step to the progression of inflatable structures, and may encourage their further development and acceptance.

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APPENDIX A ANSYS SCRIPT CODE The following contains two ANSYS script files used in the design and analysis of the inflated cage. A top flow orientation script is first presented, followed by a side flow orientation script. Many variables used in the analysis are parameterized in the beginning of each script to account for changing conditions. Top-Flow /COM, this code has the link10 elements and uses knode command to /COM, connect them to the caps /PREP7 /COM, change colors to white background /RGB,INDEX,100,100,100, 0 /RGB,INDEX, 80, 80, 80,13 /RGB,INDEX, 60, 60, 60,14 /RGB,INDEX, 0, 0, 0,15 /COM, ******************************************* /COM, define geometric tube parameters /COM, ******************************************* /COM, inner diameter, units = inches ID = 8.169 /COM, material thickness, units = inches thicknessinch = 0.1063 /COM, units = m length = 8 /COM, end cap radius of curvature determines cap length caprad=ID/12 /COM, ************************************** /COM, define element parameters /COM, ************************************** /COM, number of nodes on quarter circumference n = 7 /COM, number of nodes along length m = 51 /COM, element depth edepth = length/(m-1) /COM, Elastic Modulus in the hoop direction (Pa) Ehoop = 5.157e7 /COM, Elastic Modulus in the longitudinal direction (Pa) 80

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81 Elong = 3.211e8 /COM, Major Poisson's Ratio (xy) poisson = 0.4 /COM, Shear Modulus, Gxy, G_HL, units? shearmod = 1.91e7 /COM, Elastic Modulus for Artificial stiffness (small) units=Pa smallmod = 10000 /COM, Elastic Modulus for Cap (large) units=Pa capmod=9e9 /COM,******************************************************* /COM,* define external fluid (water @20 C) parameters /COM,******************************************************* /COM, appendix A-3, external fluid absolute viscosity, kg/m-s visc = (1e-3) /COM, external fluid density, kg/m^3 dens = (998) /COM, *************************************** /COM, define loading /COM, *************************************** /COM, internal pressure, psi prespsi = 45 /COM, flowing fluid velocity m/s vel = (0.5144444) /COM, ********************************************** /COM, calculate dependant variables, units /COM, ********************************************** /COM, angles in degrees *afun,deg /COM, cap arc length (m) caplength=0.3 /COM, convert pressure units pressure = (6894.757*prespsi) /COM, convert diameter to radius in meters radius = (ID/2)*0.0254 diam=radius*2 /COM, convert to units = m thickness = thicknessinch*0.0254 /COM, cross-sectional area ONE tube, units = m^2 tubecsarea = 2*radius*length /COM, cross-sectional area ONE cap, units = m^2 capcsarea=((3.1417*caprad*diam)/4) /COM, cross-sectional area ENTIRE cage, units = m^2 totcsarea=((8*tubecsarea)+(8*capcsarea)) /COM, total number of nodes in the entire octagon (cap+tube) ntot=32*(n**2+m*n-m-3*n+2) /COM, Reynolds number, dimensionless Re = (dens*vel*2*radius)/visc /COM, coefficient of drag, found from Figure 9.13 based on Re cd = 1 /COM, total drag force, units = N drag = 0.5*cd*totcsarea*dens*(vel**2) /COM, force per node equivalent to the internal pressure... /COM, ...acting on the cap, units = N capForce = (pressure*3.1417*radius*radius) /COM,******************************************************* /COM,* Define Elements

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82 /COM,******************************************************* /COM, /COM, 1 Membrane Shell Element /COM, 2 Added Stiffness Membrane Shell Element /COM, 3 Cap Membrane Shell Element /COM, 4 Cable Element ET,1,SHELL41,2 ET,2,SHELL63 ET,3,SHELL63 ET,4,LINK10,,1,0 /COM, Real Constants R,1,thickness R,2,thickness R,3,thickness*10 R,4,0.5.0.3 /COM, Young's Modulus in the hoop direction MP,EX,1, Ehoop MP,EX,2, smallmod MP,EX,3, capmod MP,EX,4, 9e7 /COM, Young's Modulus in the longitudinal direction MP,EY,1, Elong /COM, Major Poisson's Ratio (xy) MP,PRXY,1,poisson MP,PRXY,2,poisson MP,PRXY,3,poisson /COM, Shear Modulus, Gxy, G_HL MP,GXY,1, shearmod /COM,******************************************************* /COM,* CAGE GEOMETRY /COM,******************************************************* /COM, generate one tube and two 45 degree caps (upper right corner) k,1,length/2,length/2+length*sin(45),0 k,2,length/2+length*cos(45),length/2,0 circle,1,caprad,,,,8 circle,2,caprad,,,,8 l,4,12 circle,4,radius,12 circle,12,radius,4 ADRAG,18,19,20,21,,,17 ADRAG,18,19,20,21,,,2 ADRAG,22,23,24,25,,,9 /COM, mesh one tube (upper right corner) lsel,s,line,,17 lsel,a,line,,27 lsel,a,line,,28 lsel,a,line,,30 lsel,a,line,,32 lesize,all,,,m-1 TYPE,1 MAT,1 REAL,1 amesh,1,4,1 /COM, mesh two caps (upper right corner) lsel,s,line,,18,25,1 lesize,all,,,n-1

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83 lsel,s,line,,35 lsel,a,line,,36 lsel,a,line,,38 lsel,a,line,,40 lsel,a,line,,43 lsel,a,line,,44 lsel,a,line,,46 lsel,a,line,,48 lesize,all,,,n-1 TYPE,3 MAT,3 REAL,3 amesh,5,12,1 /COM, reflect areas and mesh to create all four corners of octagon arsym,X,all arsym,y,all /COM, generate 2 vertical and 2 horizontal tube areas (sides) k,,length/2,length/2+length*sin(45)+caprad,0 k,,-length/2,length/2+length*sin(45)+caprad,0 k,,-length/2,-length/2-length*sin(45)-caprad,0 k,,length/2,-length/2-length*sin(45)-caprad,0 k,,length/2+length*cos(45)+caprad,length/2,0 k,,length/2+length*cos(45)+caprad,-length/2,0 k,,-length/2-length*cos(45)-caprad,length/2,0 k,,-length/2-length*cos(45)-caprad,-length/2,0 l,99,100 l,101,102 l,103,104 l,105,106 ADRAG,63,66,68,69,,,146 ADRAG,95,98,100,101,,,147 ADRAG,104,108,111,113,,,148 ADRAG,72,76,79,81,,,149 /COM, mesh four tubes (sides) lsel,s,line,,51 lsel,a,line,,53 lsel,a,line,,55 lsel,a,line,,58 lsel,a,line,,83 lsel,a,line,,85 lsel,a,line,,87 lsel,a,line,,90 lsel,a,line,,115 lsel,a,line,,117 lsel,a,line,,119 lsel,a,line,,122 lsel,a,line,,146 lsel,a,line,,147 lsel,a,line,,148 lsel,a,line,,149 lsel,a,line,,151 lsel,a,line,,152 lsel,a,line,,154 lsel,a,line,,156 lsel,a,line,,159 lsel,a,line,,160

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84 lsel,a,line,,162 lsel,a,line,,164 lsel,a,line,,167 lsel,a,line,,168 lsel,a,line,,170 lsel,a,line,,172 lsel,a,line,,175 lsel,a,line,,176 lsel,a,line,,178 lsel,a,line,,180 lesize,all,,,m-1 TYPE,1 MAT,1 REAL,1 amesh,49,64,1 /COM, re-generate tube areas (sides) and mesh with artificial stiffness lsel,all ADRAG,63,66,68,69,,,146 ADRAG,95,98,100,101,,,147 ADRAG,104,108,111,113,,,148 ADRAG,72,76,79,81,,,149 arsym,X,1,4,1,0,1,0 arsym,y,1,4,1,0,1,0 arsym,X,13,16,1,0,1,0 arsym,y,13,16,1,0,1,0 TYPE,2 MAT,2 REAL,2 amesh,65,96,1 /COM, merge all nodes, elements, and keypoints nsel,all esel,all nummrg, node, 1e-5 nummrg, kp nummrg, elem, 1e-5 /COM, apply total drag force / total # nodes nsel,all f,all,fz,-drag/ntot /COM, apply force boundary conditions esel,s,type,,1 sfe,all,,pres,,pressure /COM, create cage cables k,200,0,0,((2.4*length)*(5/16)) k,201,0,0,-((2.4*length)*(5/16)) TYPE,4 MAT,4 REAL,4 csys,1 nsel,s,loc,y,22.49999,22.50001 csys,0 f,all,fy,(capforce*(1-sin(45)))/(4*(n-1)) f,all,fx,(capforce*cos(45))/(4*(n-1)) csys,1 nsel,r,loc,z,radius knode,203,all l,203,200,1

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85 lmesh,26 nsel,s,loc,y,22.49999,22.50001 nsel,r,loc,z,-radius knode,204,all l,204,201,1 lmesh,29 csys,1 nsel,s,loc,y,67.49999, 67.50001 csys,0 f,all,fx,(capforce*(1-cos(45)))/(4*(n-1)) f,all,fy,(capforce*sin(45))/(4*(n-1)) csys,1 nsel,r,loc,z,radius knode,205,all l,205,200,1 lmesh,31 nsel,s,loc,y,67.49999, 67.50001 nsel,r,loc,z,-radius knode,206,all l,206,201,1 lmesh,33 csys,1 nsel,s,loc,y,112.49999, 112.50001 csys,0 f,all,fx,-(capforce*(1-cos(45)))/(4*(n-1)) f,all,fy,(capforce*sin(45))/(4*(n-1)) csys,1 nsel,r,loc,z,radius knode,207,all l,207,200,1 lmesh,70 nsel,s,loc,y,112.49999, 112.50001 nsel,r,loc,z,-radius knode,208,all l,208,201,1 lmesh,74 csys,1 nsel,s,loc,y,157.49999,157.50001 csys,0 f,all,fy,(capforce*(1-sin(45)))/(4*(n-1)) f,all,fx,-(capforce*cos(45))/(4*(n-1)) csys,1 nsel,r,loc,z,radius knode,209,all l,209,200,1 lmesh,77 nsel,s,loc,y,157.49999,157.50001 nsel,r,loc,z,-radius knode,210,all l,210,201,1 lmesh,80 csys,1 nsel,s,loc,y,-22.49999, -22.50001 csys,0 f,all,fy,-(capforce*(1-sin(45)))/(4*(n-1)) f,all,fx,(capforce*cos(45))/(4*(n-1))

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86 csys,1 nsel,r,loc,z,radius knode,211,all l,211,200,1 lmesh,102 nsel,s,loc,y,-22.49999,-22.50001 nsel,r,loc,z,-radius knode,212,all l,212,201,1 lmesh,106 csys,1 nsel,s,loc,y,-67.49999,-67.50001 csys,0 f,all,fx,(capforce*(1-sin(45)))/(4*(n-1)) f,all,fy,-(capforce*cos(45))/(4*(n-1)) csys,1 nsel,r,loc,z,radius knode,213,all l,213,200,1 lmesh,109 nsel,s,loc,y,-67.49999,-67.50001 nsel,r,loc,z,-radius knode,214,all l,214,201,1 lmesh,112 csys,1 nsel,s,loc,y,-112.49999,-112.50001 csys,0 f,all,fx,-(capforce*(1-sin(45)))/(4*(n-1)) f,all,fy,-(capforce*cos(45))/(4*(n-1)) csys,1 nsel,r,loc,z,radius knode,215,all l,215,200,1 lmesh,134 nsel,s,loc,y,-112.49999,-112.50001 nsel,r,loc,z,-radius knode,216,all l,216,201,1 lmesh,138 csys,1 nsel,s,loc,y,-157.49999,-157.50001 csys,0 f,all,fy,-(capforce*(1-sin(45)))/(4*(n-1)) f,all,fx,-(capforce*cos(45))/(4*(n-1)) csys,1 nsel,r,loc,z,radius knode,217,all l,217,200,1 lmesh,141 nsel,s,loc,y,-157.49999,-157.50001 nsel,r,loc,z,-radius knode,218,all l,218,201,1 lmesh,144 csys,0

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87 nsel,s,loc,x,0 d,all,ux,0 nsel,s,loc,y,0 d,all,uy,0 dk,200,ux,0 dk,200,uy,0 dk,200,uz,0 dk,201,ux,0 dk,201,uy,0 dk,201,uz,0 nsel,all esel,all nummrg, node nummrg, elem nummrg, kp fini /COM, ************************************ /COM, Solution /COM, ************************************ /solve nsel,all esel,all autots,on deltim,1 solve /post1 cys,0 nsel,s,loc,x,-length/2,length/2 nsel,r,loc,y,0,length*5 esln,s,all esel,r,type,,1 plesol,s,x esel,all nsel,all plnsol,u,x esel,all nsel,all plnsol,u,z Side-Flow /PREP7 /COM, ******************************************* /COM, define geometric tube parameters /COM, ******************************************* /COM, change colors to white background /RGB,INDEX,100,100,100, 0 /RGB,INDEX, 80, 80, 80,13 /RGB,INDEX, 60, 60, 60,14 /RGB,INDEX, 0, 0, 0,15 /COM, inner diameter, units = inches ID = 16.142 /COM, material thickness, units = inches thicknessinch = .1181 /COM, units = m

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88 length = 8 /COM, end cap radius of curvature determines cap length caprad=ID/12 /COM, ************************************** /COM, define element parameters /COM, ************************************** /COM, number of nodes on quarter circumference n = 7 /COM, number of nodes along length m = 81 /COM, element depth edepth = length/(m-1) /COM, Elastic Modulus in the hoop direction (Pa) Ehoop = 5.157e7 /COM, Elastic Modulus in the longitudinal direction (Pa) Elong = 3.211e8 /COM, Major Poisson's Ratio (xy) poisson = 0.4 /COM, Shear Modulus, Gxy, G_HL, units? shearmod = 1.91e7 /COM, Elastic Modulus for Artificial stiffness (small) units=Pa smallmod = 10000 /COM, Elastic Modulus for Cap (large) units=Pa capmod=3e9 /COM,******************************************************* /COM,* define external fluid (water @20 C) parameters /COM,******************************************************* /COM, appendix A-3, external fluid absolute viscosity, kg/m-s visc = (1e-3) /COM, external fluid density, kg/m^3 dens = (998) /COM, *************************************** /COM, define loading /COM, *************************************** /COM, internal pressure, psi prespsi = 45 /COM, flowing fluid velocity m/s vel = (0.5144444) /COM, ********************************************** /COM, calculate dependant variables, units /COM, ********************************************** /COM, angles in degrees *afun,deg /COM, cap arc length (m) caplength=0.3 /COM, convert pressure units pressure = (6894.757*prespsi) /COM, convert diameter to radius in meters radius = (ID/2)*0.0254 diam=radius*2 /COM, convert to units = m thickness = thicknessinch*0.0254 /COM, cross-sectional area ONE tube, units = m^2 tubecsarea = 2*radius*length /COM, cross-sectional area ONE cap, units = m^2 capcsarea=((3.1417*caprad*diam)/4)

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89 /COM, cross-sectional area ENTIRE cage, units = m^2 totcsarea=((8*tubecsarea)+(8*capcsarea)) /COM, total number of nodes in the entire octagon (cap+tube) !ntot=32*(n**2+m*n-m-3*n+2) /COM, total number nodes in one tube ntot=((n-1)*m*4) /COM, Reynolds number, dimensionless Re = (dens*vel*2*radius)/visc /COM, coefficient of drag, found from Figure 9.13 based on Re cd = 1 /COM, total drag force on one tube, units = N drag = 0.5*cd*tubecsarea*dens*(vel**2) /COM, total drag force on half of the center tube, units = N dragcentertube = 0.5*cd*2*radius*(((2.4*length)*(5/16))-caplength/2)*dens*(vel**2) /COM, force per node equivalent to the internal pressure... /COM, ...acting on the cap, units = N !capForce = (pressure*3.1417*radius*radius)/(4*(n-1)) capForce = (pressure*3.1417*radius*radius) /COM,******************************************************* /COM,* Define Elements /COM,******************************************************* /COM, 1 Membrane Shell Element /COM, 2 Added Stiffness Membrane Shell Element /COM, 3 Cap Membrane Shell Element /COM, 4 Cable Element ET,1,SHELL41,2 !ET,1,SHELL63 ET,2,SHELL63 ET,3,SHELL63 ET,4,LINK10,,1,0 /COM, Real Constants R,1,thickness R,2,thickness R,3,thickness*3 R,4,0.5 /COM, Young's Modulus in the hoop direction MP,EX,1, Ehoop MP,EX,2, smallmod MP,EX,3, capmod MP,EX,4, 9e7 /COM, Young's Modulus in the longitudinal direction MP,EY,1, Elong /COM, Major Poisson's Ratio (xy) MP,PRXY,1,poisson MP,PRXY,2,poisson MP,PRXY,3,poisson /COM, Shear Modulus, Gxy, G_HL MP,GXY,1, shearmod /COM,******************************************************* /COM,* CAGE GEOMETRY /COM,******************************************************* /COM, generate one tube and two 45 degree caps (upper right corner) k,1,length/2,length/2+length*sin(45),0 k,2,length/2+length*cos(45),length/2,0 circle,1,caprad,,,,8 circle,2,caprad,,,,8

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90 l,4,12 circle,4,radius,12 circle,12,radius,4 ADRAG,18,19,20,21,,,17 ADRAG,18,19,20,21,,,2 ADRAG,22,23,24,25,,,9 /COM, mesh one tube (upper right corner) lsel,s,line,,17 lsel,a,line,,27 lsel,a,line,,28 lsel,a,line,,30 lsel,a,line,,32 lesize,all,,,m-1 TYPE,1 MAT,1 REAL,1 amesh,1,4,1 /COM, mesh two caps (upper right corner) lsel,s,line,,18,25,1 lesize,all,,,n-1 lsel,s,line,,35 lsel,a,line,,36 lsel,a,line,,38 lsel,a,line,,40 lsel,a,line,,43 lsel,a,line,,44 lsel,a,line,,46 lsel,a,line,,48 lesize,all,,,n-1 TYPE,3 MAT,3 REAL,3 amesh,5,12,1 /COM, reflect areas and mesh to create all four corners of octagon arsym,X,all arsym,y,all /COM, generate 2 vertical and 2 horizontal tube areas (sides) k,,length/2,length/2+length*sin(45)+caprad,0 k,,-length/2,length/2+length*sin(45)+caprad,0 k,,-length/2,-length/2-length*sin(45)-caprad,0 k,,length/2,-length/2-length*sin(45)-caprad,0 k,,length/2+length*cos(45)+caprad,length/2,0 k,,length/2+length*cos(45)+caprad,-length/2,0 k,,-length/2-length*cos(45)-caprad,length/2,0 k,,-length/2-length*cos(45)-caprad,-length/2,0 l,99,100 l,101,102 l,103,104 l,105,106 ADRAG,63,66,68,69,,,146 ADRAG,95,98,100,101,,,147 ADRAG,104,108,111,113,,,148 ADRAG,72,76,79,81,,,149 /COM, mesh four tubes (sides) lsel,s,line,,51 lsel,a,line,,53

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91 lsel,a,line,,55 lsel,a,line,,58 lsel,a,line,,83 lsel,a,line,,85 lsel,a,line,,87 lsel,a,line,,90 lsel,a,line,,115 lsel,a,line,,117 lsel,a,line,,119 lsel,a,line,,122 lsel,a,line,,146 lsel,a,line,,147 lsel,a,line,,148 lsel,a,line,,149 lsel,a,line,,151 lsel,a,line,,152 lsel,a,line,,154 lsel,a,line,,156 lsel,a,line,,159 lsel,a,line,,160 lsel,a,line,,162 lsel,a,line,,164 lsel,a,line,,167 lsel,a,line,,168 lsel,a,line,,170 lsel,a,line,,172 lsel,a,line,,175 lsel,a,line,,176 lsel,a,line,,178 lsel,a,line,,180 lesize,all,,,m-1 asel,s,,,49,52,1 asel,a,,,53,56,1 asel,a,,,57,60,1 asel,a,,,61,64,1 TYPE,1 MAT,1 REAL,1 amesh,all /COM, re-generate tube areas (sides) and mesh with artificial stiffness lsel,all ADRAG,63,66,68,69,,,146 ADRAG,95,98,100,101,,,147 ADRAG,104,108,111,113,,,148 ADRAG,72,76,79,81,,,149 asel,s,,,1,4,1 arsym,X,all,,,,1 asel,s,,,1,4,1 arsym,y,all,,,,1 asel,s,,,13,16,1 arsym,X,all,,,,1 asel,s,,,13,16,1 arsym,y,all,,,,1 TYPE,2 MAT,2 REAL,2

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92 asel,s,,,65,96,1 amesh,all k,200,(-3.7*length) k,201,0,0,(((2.4*length)*(5/16))+caplength) k,202,0,0,((2.4*length)*(5/16)) k,203,0,0,caplength/2 k,204,0,0,0 l,201,202 l,202,203 l,203,204 circle,201,radius /com, center tube upper cap lsel,s,,,265,268,1 adrag,all,,,,,,262 lsel,s,line,,262 lsel,a,line,,270 lsel,a,line,,271 lsel,a,line,,273 lsel,a,line,,275 lesize,all,,,n-1 TYPE,3 MAT,3 REAL,3 asel,s,,,97,100,1 amesh,all /com, center tube circle,202,radius lsel,s,,,277,280,1 adrag,all,,,,,,263 lsel,s,line,,263 lsel,a,line,,282 lsel,a,line,,283 lsel,a,line,,285 lsel,a,line,,287 lesize,all,,,(m-1)/2 TYPE,1 MAT,1 REAL,1 asel,s,,,101,104,1 amesh,all circle,203,radius lsel,s,,,289,292,1 adrag,all,,,,,,263 lsel,s,line,,263 lsel,a,line,,294 lsel,a,line,,295 lsel,a,line,,297 lsel,a,line,,299 lesize,all,,,(m-1)/2 TYPE,2 MAT,2 REAL,2 amesh,105,108,1 circle,204,radius lsel,s,,,301,304,1 adrag,all,,,,,,264

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93 lsel,s,line,,264 lsel,a,line,,306 lsel,a,line,,307 lsel,a,line,,309 lsel,a,line,,311 lesize,all,,,(n-1)/2 TYPE,3 MAT,3 REAL,3 asel,s,,,109,112,1 amesh,all csys,0 asel,s,,,97,112,1 arsym,z,all csys,1 nsel,s,loc,y,22.49999,22.50001 csys,0 nsel,r,loc,y,radius*2,length*5 nsel,r,loc,z,radius knode,300,all csys,1 nsel,s,loc,y,22.49999,22.50001 csys,0 nsel,r,loc,y,radius*2,length*5 nsel,r,loc,z,-radius knode,301,all csys,1 nsel,s,loc,y,67.49999, 67.50001 csys,0 nsel,r,loc,y,radius*2,length*5 nsel,r,loc,z,radius knode,302,all csys,1 nsel,s,loc,y,67.49999, 67.50001 csys,0 nsel,r,loc,y,radius*2,length*5 nsel,r,loc,z,-radius knode,303,all csys,1 nsel,s,loc,y,112.49999, 112.50001 csys,0 nsel,r,loc,y,radius*2,length*5 nsel,r,loc,z,radius knode,304,all csys,1 nsel,s,loc,y,112.49999, 112.50001 csys,0 nsel,r,loc,y,radius*2,length*5 nsel,r,loc,z,-radius knode,305,all csys,1 nsel,s,loc,y,157.49999,157.50001 csys,0 nsel,r,loc,y,radius*2,length*5 nsel,r,loc,z,radius knode,306,all

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94 csys,1 nsel,s,loc,y,157.49999,157.50001 csys,0 nsel,r,loc,y,radius*2,length*5 nsel,r,loc,z,-radius knode,307,all csys,1 nsel,s,loc,y,-22.49999, -22.50001 csys,0 nsel,r,loc,y,-radius*2,-length*5 nsel,r,loc,z,radius knode,308,all csys,1 nsel,s,loc,y,-22.49999,-22.50001 csys,0 nsel,r,loc,y,-radius*2,-length*5 nsel,r,loc,z,-radius knode,309,all csys,1 nsel,s,loc,y,-67.49999,-67.50001 csys,0 nsel,r,loc,y,-radius*2,-length*5 nsel,r,loc,z,radius knode,310,all csys,1 nsel,s,loc,y,-67.49999,-67.50001 csys,0 nsel,r,loc,y,-radius*2,-length*5 nsel,r,loc,z,-radius knode,311,all csys,1 nsel,s,loc,y,-112.49999,-112.50001 csys,0 nsel,r,loc,y,-radius*2,-length*5 nsel,r,loc,z,radius knode,312,all csys,1 nsel,s,loc,y,-112.49999,-112.50001 csys,0 nsel,r,loc,y,-radius*2,-length*5 nsel,r,loc,z,-radius knode,313,all csys,1 nsel,s,loc,y,-157.49999,-157.50001 csys,0 nsel,r,loc,y,-radius*2,-length*5 nsel,r,loc,z,radius knode,314,all csys,1 nsel,s,loc,y,-157.49999,-157.50001 csys,0 nsel,r,loc,y,-radius*2,-length*5 nsel,r,loc,z,-radius knode,315,all /COM, ****************************************** /COM, create new keypoints

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95 /COM, ****************************************** /COM, center tube top csys,0 nsel,s,loc,x,radius nsel,r,loc,y,-radius*2,radius*2 nsel,r,loc,z,(((2.4*length)*(5/16))+(caplength/2)) knode,400,all csys,0 nsel,s,loc,x,-radius nsel,r,loc,y,-radius*2,radius*2 nsel,r,loc,z,(((2.4*length)*(5/16))+(caplength/2)) knode,401,all csys,0 nsel,s,loc,x,-radius*2,radius*2 nsel,r,loc,y,radius nsel,r,loc,z,(((2.4*length)*(5/16))+(caplength/2)) knode,402,all csys,0 nsel,s,loc,x,-radius*2,radius*2 nsel,r,loc,y,-radius nsel,r,loc,z,(((2.4*length)*(5/16))+(caplength/2)) knode,403,all /COM, center tube bottom csys,0 nsel,s,loc,x,radius nsel,r,loc,y,-radius*2,radius*2 nsel,r,loc,z,-(((2.4*length)*(5/16))+(caplength/2)) knode,600,all csys,0 nsel,s,loc,x,-radius nsel,r,loc,y,-radius*2,radius*2 nsel,r,loc,z,-(((2.4*length)*(5/16))+(caplength/2)) knode,601,all csys,0 nsel,s,loc,x,-radius*2,radius*2 nsel,r,loc,y,radius nsel,r,loc,z,-(((2.4*length)*(5/16))+(caplength/2)) knode,602,all csys,0 nsel,s,loc,x,-radius*2,radius*2 nsel,r,loc,y,-radius nsel,r,loc,z,-(((2.4*length)*(5/16))+(caplength/2)) knode,603,all /COM, center tube center csys,0 nsel,s,loc,x,radius nsel,r,loc,y,-radius*2,radius*2 nsel,r,loc,z,0 knode,500,all csys,0 nsel,s,loc,x,-radius nsel,r,loc,y,-radius*2,radius*2 nsel,r,loc,z,0 knode,501,all csys,0 nsel,s,loc,x,-radius*2,radius*2

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96 nsel,r,loc,y,radius nsel,r,loc,z,0 knode,502,all csys,0 nsel,s,loc,x,-radius*2,radius*2 nsel,r,loc,y,-radius nsel,r,loc,z,0 knode,503,all /COM, **************************************** /COM, create cable links /COM, **************************************** l,300,400,1 l,300,500,1 l,301,500,1 l,301,600,1 l,302,402,1 l,302,502,1 l,303,502,1 l,303,602,1 l,304,402,1 l,304,502,1 l,305,502,1 l,305,602,1 l,306,401,1 l,306,501,1 l,307,501,1 l,307,601,1 l,308,400,1 l,308,500,1 l,309,500,1 l,309,600,1 l,310,403,1 l,310,503,1 l,311,503,1 l,311,603,1 l,312,403,1 l,312,503,1 l,313,503,1 l,313,603,1 l,314,401,1 l,314,501,1 l,315,501,1 l,315,601,1 /com, tether support l,200,306,1 l,200,307,1 l,200,314,1 l,200,315,1 /com, select cage cables and mesh lsel,s,,,361,396,1 TYPE,4 MAT, 4 REAL,4 lmesh,all /com, merge nodes, elements, and keypoints in that order to increase accuracy nsel,all

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97 esel,all ksel,all nummrg,node,1e-5 nummrg,elem,1e-5 nummrg,kp,1e-5 ********************************************** ** ** ** apply boundary conditions ** ** ** ********************************************** nsel,all nplot esel,all eplot /com, apply internal pressure csys,0 nsel,s,loc,y,-1.2*radius,1.2*radius nsel,r,loc,x,-1.2*radius,1.2*radius esln,s esel,r,type,,1 sfe,all,,pres,,-pressure csys,0 nsel,s,loc,y,-1.2*radius,1.2*radius nsel,r,loc,x,-1.2*radius,1.2*radius nsel,inve esln,s esel,r,type,,1 sfe,all,,pres,,pressure csys,0 nsel,s,loc,y,-1.2*radius,1.2*radius nsel,r,loc,x,-1.2*radius,1.2*radius nsel,r,loc,z,0,-5*length esln,s esel,r,type,,3 sfe,all,,pres,,pressure csys,0 nsel,s,loc,y,-1.2*radius,1.2*radius nsel,r,loc,x,-1.2*radius,1.2*radius nsel,r,loc,z,0,5*length esln,s esel,r,type,,3 sfe,all,,pres,,pressure csys,0 nsel,s,loc,y,-1.2*radius,1.2*radius nsel,r,loc,x,-1.2*radius,1.2*radius nsel,inve esln,s esel,r,type,,3 sfe,all,,pres,,-pressure /com, apply drag csys,0 !nsel,s,loc,y,-length/2,length/2 nsel,s,loc,y,-((length/2)+(caprad*sin(22.5))),((length/2)+(caprad*sin(22.5))) nsel,r,loc,x,1.2*radius,5*length f,all,fx,(drag/ntot) csys,0

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98 !nsel,s,loc,y,-length/2,length/2 nsel,s,loc,y,-((length/2)+(caprad*sin(22.5))),((length/2)+(caprad*sin(22.5))) nsel,r,loc,x,-1.2*radius,-5*length f,all,fx,(drag/ntot) csys,1 nsel,s,loc,y,22.5,67.5 nsel,a,loc,y,112.5,157.5 !nsel,s,loc,y,30,60 !nsel,a,loc,y,120,150 csys,0 nsel,r,loc,y,1.2*radius,5*length !esln,s,all !esel,r,type,,1 !nsle,s,all f,all,fx,((drag*cos(45))/(ntot)) csys,1 nsel,s,loc,y,-22.5,-67.5 nsel,a,loc,y,-112.5,-157.5 !nsel,s,loc,y,-30,-60 !nsel,a,loc,y,-120,-150 csys,0 nsel,r,loc,y,-1.2*radius,-5*length f,all,fx,((drag*cos(45))/(ntot)) /com, fix X-displacement csys,1 nsel,s,loc,y,157.49999,157.50001 csys,0 nsel,r,loc,y,1.2*radius,5*length nsel,r,loc,z,radius !d,all,ux,0 csys,1 nsel,s,loc,y,157.49999,157.50001 csys,0 nsel,r,loc,y,1.2*radius,5*length nsel,r,loc,z,-radius csys,1 nsel,s,loc,y,-157.49999,-157.50001 csys,0 nsel,r,loc,y,-1.2*radius,-5*length nsel,r,loc,z,radius csys,1 nsel,s,loc,y,-157.49999,-157.50001 csys,0 nsel,r,loc,y,-1.2*radius,-5*length nsel,r,loc,z,-radius /com, fix Y-displacement csys,1 nsel,s,loc,y,-157.49999,-157.50001 csys,0 nsel,r,loc,y,-1.2*radius,-5*length nsel,r,loc,z,radius !d,all,uy,0 csys,1 nsel,s,loc,y,-157.49999,-157.50001 csys,0 nsel,r,loc,y,-1.2*radius,-5*length

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99 nsel,r,loc,z,-radius csys,1 nsel,s,loc,y,-22.49999,-22.50001 csys,0 nsel,r,loc,y,-1.2*radius,-5*length nsel,r,loc,z,radius csys,1 nsel,s,loc,y,-22.49999,-22.50001 csys,0 nsel,r,loc,y,-1.2*radius,-5*length nsel,r,loc,z,-radius csys,0 nsel,s,loc,y,0 nsel,r,loc,x,1.2*radius,5*length d,all,uy,0 csys,0 nsel,s,loc,y,0 nsel,r,loc,x,-1.2*radius,-5*length d,all,uy,0 ksel,s,,,200 nslk,s,all d,all,ux,0,,,,uy,uz /com, fix Z-displacement csys,0 nsel,s,loc,z,0 nsel,r,loc,y,-1.2*radius,1.2*radius nsel,r,loc,x,-1.2*radius,1.2*radius d,all,uz,0 2.49999,112.50001 /COM,****************************************************************** /COM,* apply cap forces and create keypoints /COM,****************************************************************** csys,1 nsel,s,loc,y,22.49999,22.50001 csys,0 nsel,r,loc,y,radius*2,length*5 f,all,fy,(capforce*(1-sin(45)))/(4*(n-1)) f,all,fx,(capforce*cos(45))/(4*(n-1)) csys,1 nsel,s,loc,y,67.49999, 67.50001 csys,0 nsel,r,loc,y,radius*2,length*5 f,all,fx,(capforce*(1-cos(45)))/(4*(n-1)) f,all,fy,(capforce*sin(45))/(4*(n-1)) csys,1 nsel,s,loc,y,112.49999, 112.50001 csys,0 nsel,r,loc,y,radius*2,length*5 f,all,fx,-(capforce*(1-cos(45)))/(4*(n-1)) f,all,fy,(capforce*sin(45))/(4*(n-1)) csys,1 nsel,s,loc,y,157.49999,157.50001 csys,0 nsel,r,loc,y,radius*2,length*5 f,all,fy,(capforce*(1-sin(45)))/(4*(n-1)) f,all,fx,-(capforce*cos(45))/(4*(n-1))

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100 csys,1 nsel,s,loc,y,-22.49999, -22.50001 csys,0 nsel,r,loc,y,-radius*2,-length*5 f,all,fy,-(capforce*(1-sin(45)))/(4*(n-1)) f,all,fx,(capforce*cos(45))/(4*(n-1)) csys,1 nsel,s,loc,y,-67.49999,-67.50001 csys,0 nsel,r,loc,y,-radius*2,-length*5 f,all,fx,(capforce*(1-sin(45)))/(4*(n-1)) f,all,fy,-(capforce*cos(45))/(4*(n-1)) csys,1 nsel,s,loc,y,-112.49999,-112.50001 csys,0 nsel,r,loc,y,-radius*2,-length*5 f,all,fx,-(capforce*(1-sin(45)))/(4*(n-1)) f,all,fy,-(capforce*cos(45))/(4*(n-1)) csys,1 nsel,s,loc,y,-157.49999,-157.50001 csys,0 nsel,r,loc,y,-radius*2,-length*5 f,all,fy,-(capforce*(1-sin(45)))/(4*(n-1)) f,all,fx,-(capforce*cos(45))/(4*(n-1)) /COM, center tube forces csys,0 nsel,s,loc,x,-radius*2,radius*2 nsel,r,loc,y,-radius*2,radius*2 nsel,r,loc,z,(((2.4*length)*(5/16))+caplength) f,all,fz,(capforce/(4*(n-1))) csys,0 nsel,s,loc,x,-radius*2,radius*2 nsel,r,loc,y,-radius*2,radius*2 nsel,r,loc,z,-(((2.4*length)*(5/16))+caplength) f,all,fz,-(capforce/(4*(n-1))) csys,0 nsel,s,loc,x,-radius*2,radius*2 nsel,r,loc,y,-radius*2,radius*2 nsel,r,loc,z,caplength/2,(((2.4*length)*(5/16))) esln,s,all esel,r,type,,1 nsle,s,all f,all,fx,dragcentertube/ntot csys,0 nsel,s,loc,x,-radius*2,radius*2 nsel,r,loc,y,-radius*2,radius*2 nsel,r,loc,z,-caplength/2,-(((2.4*length)*(5/16))) esln,s,all esel,r,type,,1 nsle,s,all f,all,fx,dragcentertube/ntot fini /COM, ************************************ /COM, Solution /COM, ************************************ /solve

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101 nsel,all esel,all autots,on deltim,1 solve /post1 nsel,all esel,all esel,r,type,,1 nsle,s,all plnsol,s,2 nsel,s,loc,y,-length/2,length/2 nsel,r,loc,x,1.2*radius,length*5 esln,s,all plnsol,u,x nsel,s,loc,y,-length/2,length/2 nsel,r,loc,x,-1.2*radius,-length*5 esln,s,all plnsol,u,x esel,all nsel,all plnsol,u,y nsel,s,loc,x,-1.2*radius,1.2*radius nsel,r,loc,y,-1.2*radius,1.2*radius esln,s,all nsle,s,all plnsol,u,x LOCAL,27,CART,0,0,0,45,0,0

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APPENDIX B EXPERIMENTAL DATA Longitudinal Orientation Pull Tests ForceWidthWidth StrainLengthLength StrainF/WF/(w*t)(lb.)(in.)(in.)(in. / in.)(lb./in.)(lb./in^2)00.75000.7870550.7445-0.00730.82050.042673.88869.1186Test 111000.7435-0.00870.82150.0438134.501582.3411500.7430-0.00930.85150.0820201.882375.1092500.7385-0.01530.87000.1055338.523982.63600.74450.7935750.7430-0.00200.82300.0372100.941187.554Test 121250.7375-0.00940.84550.0655169.491994.0182000.7315-0.01750.85150.0731273.413216.5982500.7300-0.01950.86200.0863342.474029.00900.74950.8040500.7405-0.01200.81450.013167.52794.3758Test 131100.7375-0.01600.84750.0541149.151754.7361600.7360-0.01800.85950.0690217.392557.5452250.7335-0.02130.87600.0896306.753608.8052700.913500.71100.8150750.71250.00210.84850.0411105.261238.39Test 141500.7070-0.00560.85350.0472212.162496.0482200.7030-0.01130.88250.0828312.943681.72750.7015-0.013400.72250.8120500.7190-0.00480.81500.003769.54818.1298Test 151100.7165-0.00830.83300.0259153.521806.1661600.7140-0.01180.85550.0536224.092636.3492150.7100-0.01730.87500.0776302.823562.5522600.8720 102

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103 Hoop Orientation Pull Tests ForceWidthWidth StrainLengthLength StrainF/WF/(w*t)(lb.)(in.)(in.)(in. / in.)(lb./in.)(lb./in^2)00.80900.8055410.7610-0.05930.85650.063353.88633.8Test 1570.7345-0.09210.89000.104977.60913.0700.7135-0.11800.91650.137898.111154.2800.6845-0.15390.94450.1726116.871375.0900.6695-0.17240.96050.1924134.431581.500.80950.8140300.7790-0.03770.82600.014738.51453.1400.7550-0.06730.84340.036152.98623.3Test 2500.7320-0.09570.86000.056568.31803.6600.7165-0.11490.88500.087283.74985.2700.6930-0.14390.91050.1186101.011188.4800.6800-0.16000.94200.1572117.651384.100.81550.8235300.7795-0.04410.86650.052238.49452.8500.7490-0.08150.84100.021366.76785.4Test 3600.7220-0.11470.88450.074183.10977.7700.7050-0.13550.89050.081499.291168.1800.6780-0.16860.91650.1129117.991388.200.72450.8245300.7065-0.02480.82550.001242.46499.6400.6915-0.04550.83800.016457.85680.5Test 4500.6655-0.08140.86400.047975.13883.9600.6520-0.10010.90600.098892.021082.6700.6260-0.13600.91150.1055111.821315.5800.6185-0.14630.92600.1231129.351521.700.73900.8170300.7105-0.03860.82350.008042.22496.8Test 5400.6870-0.07040.86450.058158.22685.0500.6695-0.09400.89550.096174.68878.6600.6400-0.13400.92800.135993.751102.9

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LIST OF REFERENCES Bulson, P, 1973, Design principles of pneumatic structures. The Structural Engineer, Vol 51, No 6, pp 209-215. Comer, R, and Levy, S, 1963, Deflections of an inflated circular-cylindrical cantilever beam. AIAA Journal, Vol 1, No 7, pp 1652-1655. Department of Commerce, 2002, http://www.nmfs.noaa.gov/trade/newgrant.htm, last accessed April 10, 2004. Dornheim, M, 1999, Inflatable structures taking to flight. Aviation Week and Space Technology, pp 60-62, http://www.lgarde.com/programs/iaearticle/awarticle.html, last accessed May 10, 2004. Douglas, W, 1969, Bending stiffness of an inflated cylindrical cantilever beam. AIAA Journal, Vol 7, No 7, pp 1248-1253. Fox, R, and McDonald, A, 1998, Introduction to Fluid Mechanics. New York: John Wiley & Sons, Inc. Fredriksson, D, Swift, M, Muller, E, Baldwin, K, and Celikkol, B, 2000, Open ocean aquaculture engineering: system design and physical modeling. Marine Technology Society Journal, Vol 34, pp 41-52. Hawkmoor "Temprodome," 2003, http://www.hawkmoor.com/tempro_dome.htm, last accessed April 10, 2004. Hibbeler, R, 1997, Mechanics of Materials Englewood Cliffs, NJ: Prentice Hall. Jenkins, C, 2001, Gossamer Spacecraft: Membrane and Inflatable Structures Technology for Space Applications. Reston, VA: American Institute of Aeronautics and Astronautics. Kuriyama of America, Inc, 2003, http://www.kuriyama.com/, last accessed May 10, 2004. LGarde, 2004, http://www.lgarde.com/programs/iae.html, last accessed April 10, 2004. Lindstrand, 2000, http://www.lindstrand.co.uk/inflatablebuildings/, last accessed April 10, 2004. 104

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105 Loverich G, and Forster J, 2000, Advances in offshore cage design using spar buoys. Marine Technology Society Journal, Vol 34, pp 18-28. Main, J, Peterson, S, and Strauss, A, 1994, Load-deflection behavior of space-based inflatable fabric beams. Journal of Aerospace Engineering, Vol 7, No 2, pp 225-238. Main, J, Peterson, S, and Strauss, A, 1995, Beam-type bending of space-based inflated membrane structures. Journal of Aerospace Engineering, Vol 8, No 2, pp 120-125. Measurements Technology Inc (MTI), 2004, http://www.mti-atlanta.com, last accessed May 10, 2004. Promotional Design Group, 2001, http://www.promotionaldesigngroup.com/inflatable_tents.htm, last accessed April 10, 2004. Sakamoto, H, Natori, M, 2001, Deflection of multi-cellular inflatable tubes for redundant space structures. AIAA/ASME/ASCE/AHS/ASC Structures, Structural Dynamics, and Materials Conference and Exhibit, 42nd, Seattle, WA, April 16-19. SRS Technologies, 2000, http://www.stg.srs.com/atd/antenna.htm, last accessed April 10, 2004. Topping, A, 1964, Shear deflections and buckling characteristics of inflated members. Journal of Aircraft, Vol 1, No 5, pp 289-293. Tsukrov, I, Ozbay, M, Fredriksson, D, Swift, M, Baldwin, K, Celikkol, B, 2000, Open ocean aquaculture engineering: numerical modeling. Marine Technology Society Journal, Vol 34, pp 29-40.

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BIOGRAPHICAL SKETCH Jeffrey Suhey was born in upstate New York in 1978 before he moved to Newtown, Pennsylvania, where he attended the Council Rock School District. In 1993, he moved to Jacksonville, Florida, where he attended Bishop Kenny High School. Following graduation, he enrolled in the University of Florida Department of Mechanical Engineering, where he obtained both a bachelors and masters degree, focusing on mechanical design. During the course of his studies, he has completed three internships in the Jacksonville area; first at Vistakon, Inc, a subsidiary of Johnson & Johnson that manufactures contact lenses, and later two internships at Unison Industries, a subsidiary of General Electric Engine Services, that designs and manufactures jet ignition leads and igniters. He plans to begin his career in the aerospace field as a design and stress analyst engineer. 106