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Conceptual Design of a Semi-Submersible
Floating Oil and Gas Production System for
Offshore Malaysia

Team Members:
Mr. HuiDong Bea
Mr. Aaron Gregg
Mr. Milton Hooks
Mr. Brian Riordan
Mr. Justin Russell
Mr. Miles Williams

OCEN 407 Design Class Team Malaysia i

Table of Contents________________________
Nomenclature...................................................................................................................... ii
List of Figures .................................................................................................................... iii
List of Tables ..................................................................................................................... iv
Executive Summary ............................................................................................................ 1
Overview......................................................................................................................... 1
Competencies.................................................................................................................. 1
General Arrangement and Overall Hull Design.......................................................... 2
Local and Global Loading .......................................................................................... 2
Weight, Buoyancy, and Stability ................................................................................ 3
Hydrodynamics of Motions and Loading ................................................................... 4
General Strength and Structural Design ..................................................................... 4
Mooring....................................................................................................................... 5
Cost ............................................................................................................................. 6
Regulatory Compliance .................................................................................................. 7
Acknowledgements............................................................................................................. 7
Introduction......................................................................................................................... 7
Objectives ......................................................................................................................... 10
Design Basis...................................................................................................................... 11
Design Methodology......................................................................................................... 12
General Arrangement and Overall Hull Design............................................................ 12
Wind and Current Loading ........................................................................................... 14
Local and Global Loading ............................................................................................ 15
Weight, Buoyancy and Stability ................................................................................... 21
Hydrodynamics of Motions and Loading ..................................................................... 31
Strength and Structural Design - General ..................................................................... 32
Mooring......................................................................................................................... 33
Cost ............................................................................................................................... 35
Results............................................................................................................................... 35
General Arrangement and Overall Hull Design............................................................ 35
Wind and Current Loading ........................................................................................... 35
Local and Global Loading ............................................................................................ 36
Weight, Buoyancy and Stability ................................................................................... 37
Hydrodynamics of Motions and Loading ..................................................................... 39
Strength and Structural Design - General ..................................................................... 42
Mooring......................................................................................................................... 43
Cost ............................................................................................................................... 47
Conclusion ........................................................................................................................ 48
References......................................................................................................................... 50
Appendix A: Hydrodynamics of Motions and Loading ................................................... 51
Appendix B: Ballasting..................................................................................................... 55
Appendix C: Environmental Loads .................................................................................. 58
Appendix D: StabCad ....................................................................................................... 61
Appendix E: GMOOR .................................................................................................... 117

OCEN 407 Design Class Team Malaysia ii

Nomenclature
ABS - American Bureau of Shipping
API - American Petroleum Institute
Area ratio - Ratio of the area between the righting moment curve, downflooding
angle, and heeling moment curve added to the area under the heeling
moment curve, righting moment curve, downflooding angle, and
inclination angle compared to the sum of the area under the heeling
moment curve, righting moment curve, downflooding angle, and
inclination angle and the area between the heeling moment curve and
righting moment curve
Down flooding angle - Static angle from the intersection of the vessel's centerline and the
waterline in calm water to the first opening that cannot be closed
watertight and through which downflooding can occur
Damaged condition - Condition in which ballast tanks are flooded to represent a damaged
situation
Heave - Vertical motion in water
Heeling arm - The horizontal distance between the buoyant force and the center of
gravity to move the vessel away from the upright position
Height coefficient - Characteristic of a structure in wind based on its height
Intact condition - Condition in which the semi-submersible is unharmed
MARPOL - Marine Pollution Convention
Pitch - Rotation around the transverse (x) axis
Righting arm - The horizontal distance between the buoyant force and the center of
gravity to bring the vessel to an upright position
Roll - Rotation around the longitudinal (y) axis
SCR - Steel Catenary Risers
Semi-taut mooring system - Mooring system in which the mooring lines are pulled in tension, but
leaving some slack
Shape coefficient - Characteristic of a structure in wind based on its shape
SOLAS - Safety of Life at Sea
Taut-leg mooring system - Mooring system in which the mooring lines are pulled in tension,
having no slack
USCG - United States Coast Guard
Yaw - Rotation about the vertical (z) axis

OCEN 407 Design Class Team Malaysia iii

List of Figures

Figure 1: Final Team Organization..................................................................................... 8
Figure 2: General Iterative Process..................................................................................... 9
Figure 3: Proposed Project Location................................................................................. 11
Figure 4: Original Basic Hull Design ............................................................................... 13
Figure 5: Modified Basic Hull Design.............................................................................. 13
Figure 6: Wind Height Divisions...................................................................................... 14
Figure 7: Semi-submersible Current Drag Coefficients ................................................... 15
Figure 8: ETABS Structural Model .................................................................................. 16
Figure 9: ETABS Model with Waterline .......................................................................... 16
Figure 10: Static Wave Model with Crest across Diagonal.............................................. 17
Figure 11: Static Wave Model with Trough across Diagonal........................................... 17
Figure 12: Pressures in Positive X, Y, and Z Directions without Wave........................... 18
Figure 13: Pressures in Negative X, Y, and Z Directions without Wave ......................... 18
Figure 14: Pressures in Positive X, Y, and Z Directions with Static Wave...................... 19
Figure 15: Pressures in Negative X, Y, and Z Directions with Static Wave .................... 19
Figure 16: Weight Distribution of Semi-Submersible Components................................. 20
Figure 17: Buoyancy Force............................................................................................... 20
Figure 18: Springs Attached at Nodes .............................................................................. 21
Figure 19: Column Diagram ............................................................................................. 21
Figure 20: Pontoon Diagram............................................................................................. 22
Figure 21: Node Diagram ................................................................................................. 22
Figure 22: Riser Placement and Ballasting Tank Diagram............................................... 24
Figure 23: Balancing Spreadsheet .................................................................................... 25
Figure 24: Auto Ballasting Example................................................................................. 26
Figure 25: Horizontally Centered Topside CG Float-on/Float-off Scenario .................... 27
Figure 26: Excessive topside CG offset for Float-on/Float-off ........................................ 28
Figure 27: Intact Stability Objective Diagram.................................................................. 29
Figure 28: Damaged Stability Objective Diagram............................................................ 30
Figure 29: Final Hull Dimensions..................................................................................... 35
Figure 30: Deflections and Displacements ....................................................................... 36
Figure 31: Reaction Forces ............................................................................................... 37
Figure 33: Shear Diagram................................................................................................. 42
Figure 34: Bending Moment Diagram.............................................................................. 42
Figure 35: 12 Leg Spread.................................................................................................. 43
Figure 36: 16 Leg Spread.................................................................................................. 44
Figure 37: Mooring Analysis Results ............................................................................... 44
Figure 38: Mooring Material and Equipment Costs ......................................................... 45
Figure 39: Chosen Mooring Design.................................................................................. 46

OCEN 407 Design Class Team Malaysia iv

List of Tables_____ ____________
Table 1: Metocean Data .................................................................................................... 12
Table 2: Total Environmental Loads ................................................................................ 36
Table 3: Deflections, Moments, and Stresses of Structure ............................................... 42

OCEN 407 Design Class Team Malaysia 1

Executive Summary
Overview
With the relatively low cost of energy and the rapid advancement of technology in the last

century, humankind has become highly dependent upon large quantities of inexpensive fuel for

everything from transportation of goods and people to the harvesting of crops for food to being

able see at night. As the population increases exponentially and demands for inexpensive energy

solutions continue to rise, the readily available reservoirs of energy rich fossil fuels have rapidly

declined. In order to meet these demands, fossil fuel exploration and production is forced into

more and more inhospitable conditions such as the extreme deep sea. The expenses associated

with fixed production platforms at this depth are no longer within a feasible range making a

floating production platform design a far more economical choice. One of the floating

production platform designs in use today and pursued by this conceptual design team is the semi-

submersible production platform. This conceptual design team was given the weights and

dimensions of the platform topsides required for production and drilling, and a location in the

South China Sea at a water depth of 5,500 ft. A highly iterative process of altering the key

dimensions of a four-column ring pontoon was conducted in order to meet applicable regulations

and client requests while minimizing costs and satisfying the chosen competencies. The design

team was divided into subgroups that would tackle different components of design based on

individual talent, desire and schedule availability and given a 3 month deadline.

Competencies
The following 8 competencies were chosen to analyze and address in the conceptual design:

1. General Arrangement and Overall Hull Design
2. Wind and Current Loading
3. Local and Global Loading
4. Weight, Buoyancy and Stability


5. Hydrodynamics of Motions and Loading
6. General Strength and Structural Design
7. Mooring
8. Cost

General Arrangement and Overall Hull Design
(Methodology pg. 12, Results pg.35)

The final hull design was a square ring pontoon with outer lengths of 270 ft., a height of 40 ft.

and a width of 55 ft. The ring pontoon is topped with four square columns with a length of 50 ft.

to a side and height of 80 ft. A square topside deck is supported by the four columns with the

corners of the topside deck placed at the geographical centers of column tops.

Wind and Current Loading

The provided metocean data for the Malaysia site includes the wave height, period, wind speed

and current speed for both the operational and survival design condition. The operational and

survival design storms are 10 year and 100 year storm conditions respectively for the design

location. Shape and height coefficients taken from ABS were applied to the constant topside

shapes, the final exposed column shapes (with zero shielding) and the final component elevations

once the final freeboard was calculated. Wind, current and mean wave drift were applied

concurrently in the bow, beam and quartering directions resulting in maximum environmental

loads for operational and survival conditions to be 473.3 kips and 603.1 kips from the quartering

direction.

Local and Global Loading

An analysis of the semi-submersible structure was performed using a three-dimensional analysis

program for building systems, ETABS, to determine the deflections, moments, and stresses

acting on the topside, columns, and pontoon of the hull. Two types of analyses were viewed: one


with the structure floating at a constant draft and the other with a static wave with the crest of the

wave acting on two of the columns opposite from each other. This wave, modeled as per ABS

code 3-3-A2/3.7, produced maximum moments, stresses, and deflections. Then, the weights of

the hull, topside, ballast tanks filled with water, and nodes were added, along with the buoyancy

acting on the bottom of the pontoon. Springs were added to the joint of the columns and pontoon

to represent the mooring lines. Wind, wave, and current forces were not modeled for the

structural analysis, thus reaction forces for the springs in the x and y directions are zero when the

analysis was run. In addition the weights of all of the components are of equal and opposite

force to the buoyancy resulting in zero net global displacement in the z direction.

Weight, Buoyancy, and Stability

Weight approximations of 9, 12 and 14 lbs/ft3 were used for columns, pontoons and nodes

respectively. An additional 48-62.5% was added for marine systems, brackets, paint, and design

allowances. The hull buoyancy was simply the weight of the water displaced by the pontoons,

nodes and the part of the columns that were submerged either during float-on, float-off,

operational or survival conditions. Physical centers used for small angle stability and input into

StabCad for large angle stability were: KG = 70 ft., KB = 29.9 ft., BM = 63.4 ft., KM = 93.2 ft.,

GM = 23.1 ft., with a draft of 85 ft. and a freeboard of 35 ft. All down flooding angles and area

ratios for intact stability were in excess of the ABS recommended values. Two consecutive

tanks were flooded all the way around semi-submersible to analyze the damaged stability

response. The distance between down flooding angle and the first righting arm/ heeling arm

intercept is greater than the recommended 7 degree minimum in the worst case damaged case

scenario. Six tanks in each pontoon are floodable to reach operational and survival drafts as well


as to offset a permanent topside center of gravity that is as far from the geographical center of

gravity as (47, 47) (ft).

Hydrodynamics of Motions and Loading

Using closed-form equations, the heave and pitch/roll period were determined to be 21.8 25.5

seconds respectively. Heave and pitch/roll response was found using response amplitude

operators found from a hull of similar dimensions for the degree headings of 0, 22.5, 45, 67.5

and 90. The maximum heave response was found to be from the 22.5 degree heading and was

+/-8.9 ft and +/-12.6 ft for the operating 10 year and survival 100 year storm conditions

respectively. The maximum pitch/roll response was found to be from the 0 degree heading and

was +/-2.5 degrees and +/-4.3 degrees for the operating 10 year and the survival 100 year storm

conditions respectively.

General Strength and Structural Design

As stated before, the structural analysis was performed using the program called ETABS. The

stick figure and loadings as previously mentioned were used. The stresses, for both wave cases,

that were created from the pressures, weights, and buoyancy were compared to guidelines from

ABS 3-2-3.3 through 3-2-3.11. The stresses for the topside and pontoon were the same while the

stresses in the columns were slightly less.

The structural analysis was run and the maximum values for the deflections, moments, and

stresses for the topside, columns, and pontoon came from the moment diagrams in the static

wave case. The maximum deflection of the hull occurs in the topside, which is 1.17 in., while

the columns contain the maximum moment at 13.8x106 kip-ft. The maximum stress, 29.92 ksi,


occurs in the pontoon, which is less than the calculated value, 29.94 ksi, as per ABS

recommendations.

Mooring
(Methodology: pg. 33, Results: pg.43)

The mooring design for the semi-submersible was created with the aid of the software program

GMOOR. Environmental and mooring line component data was manually entered into GMOOR

which constructed a model of the system. Twelve and sixteen leg mooring system designs were

tested with chain, polyester, and spiral wire components in both taut-leg and semi-taut leg

mooring system designs. The mooring lines at both the fairleads and anchors are composed of

chain to account for an increase of stresses at the fairleads and to fit chain jacks, as well as to

accommodate the situation where the mooring lines may lay on the ocean floor, not to get

sediment into the polyester or wire components. Wire as the middle component performed well

in both taut and semi-taut designs, providing the required in-water weight for the semi-taut

system for restoring force but added a significantly larger downward force on the hull. The

polyester middle component could also be adapted to work well in both taut and semi-taut leg

mooring systems. While this provided the required stretch for the restoring force in the taut-leg

configuration, it lacked the in-water weight to provide an adequate restoring force in the semi-

taut design. A dynamic analysis of the mooring system was conducted with concurrent

environmental forces applied in the direction of each mooring line to find the worst case intact

and damaged conditions. The results of these analyses were checked against API 2SK

recommended mooring guidelines. It was determined that the polyester line was ideal for a taut-

leg system, and the wire line was ideal for a semi-taut system. An increase of 0.02-0.04mm to

the diameter of all the steel components of the mooring line per design year is recommended to


account for corrosion as per API 2SK Stationkeeping Manual. The final design was dictated by

meeting all mooring objectives and pertinent regulations stipulated in API 2SK as well as

minimizing cost. The finalized system was a twelve leg taut system with 200 ft. of 2.75 in. K4

Studless chain at the anchor, 7,350 ft. of 3.94 in. polyester rope as the middle component and

300 ft. of 3.0 in. K4 Studless chain at the fairlead. The anchors chosen for this design are 11.36

tons Stevmanta vertically loaded drag embedment anchors built by Vryhof. The Stevmanta was

chosen because it performs ideally for taut leg systems and it is much more cost effective than

using suction piles. SEPLAs (Suction Embedded Plate Anchors) were also considered for the

anchors, but were not included in the final design due to unknown installation costs, which are

expected to be significantly higher than Stevmanta placement.

Cost
(Methodology: pg. 35, Results: pg.47)

The total hull cost was approximated by steel weight at $12,000 per short ton of steel. Three

weight estimates were used for cost approximation purposes: 12 lbs/ft3 for pontoons, 9 lbs/ft3 for

columns and 14 lbs/ft3 for nodes. The final semi-submersible design resulted in a steel weight of

21,902 (kips) or 10,951 (tons) resulting in a net hull cost of $184 Million (USD). Using

Stevmanta drag embedment anchors projected to be $51,000 (USD) apiece; total anchor cost is

$612,000 (USD). With a 12-leg taut mooring system design as previously mentioned, line and

associated equipment costs are estimated to be $192 Million (USD). Total hull and mooring

capital would then cost $7.8 Million (USD).


Regulatory Compliance
Considerations were made towards American Bureau of Shipping (ABS), American Petroleum
Institute (API), International Convention for the Safety of Life at Sea (SOLAS), United States
Coast Guard (USCG) and the Marine Pollution Convention (MARPOL) in all applicable areas of
chosen conceptual design competencies.

Acknowledgements
It should be noted that a significant amount of help in understanding the problems and
approaches to solving offshore issues was provided by engineering mentors currently working in
the offshore industry. This list of engineers included: Rod King of Lloyds Register, Peter Noble
of ConocoPhillips, Philip Poll of Houston Offshore Engineering and Kent Longridge of
InterMoor.

Introduction
The conceptual design team decided to analyze the aforementioned competencies from the
standpoint of a four-column ring pontoon semi-submersible platform. Regard was made to given
topside dimensions and weight, required risers and their respective tensions, water depth and
Metocean data associated with the reservoir location. Engineers of varied backgrounds assisted
in the design process by giving guest presentations during classroom meetings, answering some
of the questions presented by the different design teams and helping with approximation
methods.

The teams first meeting resulted in the election of Brian Riordan as the team captain who then
broke the rest of the members into two-man teams to tackle what was thought at the time the
most crucial and most likely governing factors to the design: ABS/API regulatory compliance
research, StabCad familiarization, ETABS (structural analysis program) familiarization, and
automation of generating StabCad input points from prototype spreadsheet. After two weeks of
meetings with only small accomplishments in some StabCad familiarization and a Matlab
program to quickly develop a StabCad model based on dimension changes a strategic
restructuring was required to re-identify and prioritize the order in which analysis would take
place. The two-man teams also lead to a lack of accountability in analysis so everyone was
given a “point man” or lead status for different aspects of design with a partner assigned to work
on the project at the same time to offer aid. Regularly scheduled team meetings reviewed
progress from each lead to keep everyone on track. This structuring seemed to work for a few
weeks before it became apparent that each lead had developed his own semi-submersible that
met his immediate goals, whether it was large angle stability, hydrodynamic properties or
structural strength constraints, but did not meet all of them. The final restructuring kept the point
man positions previously assigned but had them report immediately to the Team Captain any
changes they made or problems they encountered and almost all inter-team communication was
then facilitated through the team captain in the following organizational structure seen in Figure
1: Final Team Organization.


Figure 1: Final Team Organization

Team members worked jointly with each other to help solve certain issues within their specific
assignments and tried to keep designs dynamic so changes in an interdependent competency
were easily updated in their design during the iterative process.

All other responsibilities were absorbed by the team captain and cost was determined jointly by
the Mooring lead and the Team Captain- combining costs of anchors and mooring with the
structural steel costs of the hull. Once spreadsheets and programs such as GMOOR, StabCad
and ETABS were understood, a spreadsheet would take the input changes to any dimensional
changes, provide hydrodynamic responses, freeboards, drafts and ballasting requirements. If all
of these past inspection, a batch file would run a Matlab program converting the X,Y,Z points
generated in the spreadsheet to a StabCad compatible input file for large angle stability analysis,
and other dimensions and area moments of inertia were also given to ETABS for structural
analysis. The topside dimensions (with the exception of a horizontal center of gravity) remained
constant, so a change in draft and exposed areas of columns were all that were needed to
recalculated concurrent environmental forces. The maximum concurrent forces were then used
in GMOOR to determine mooring line sizes.

To aid in iteration cycle speed, the acceptable column width range was determined by a
hydrodynamic analysis, and a minimum area moment of inertia used in ETABS while a
maximum projected area used for environmental forces. With a freeboard set to a constant, all
dimension iterations were then compared to the maximum bending and shear stresses for the
weakest structural design and with the largest environmental loads. If the tested dimensions past
hydrodynamic inspection and a quick comparison to worst case scenarios for structural and
environmental conditions, then it would definitely pass an in-depth analysis and was worth
pursuing. The general iteration process can be seen in Figure 2: General Iterative Process.


Figure 2: General Iterative Process


Objectives
The conceptual design is to satisfy the following criterion:
• Drafts & Freeboard
o A maximum float-on float-off draft of 29.52 ft. determined by Dockwise
boat capacity.
o A minimum operational freeboard (air gap) of 35 ft. based on metocean
data.
• Hydrodynamic responses
o A minimum natural heave and pitch/roll period of 20 seconds based on
client requests for optimum operation of topside separation tanks and
risers.
o A maximum of +/-16 ft heave response based recommendations from
HOE
o A maximum 4◦ pitch/roll for 10 year storm conditions and a maximum 10◦
pitch/roll for 100 year storm conditions based on HOE recommendations.
• Large Angle Intact Stability
o Righting Arm should be greater than 2 times the heeling arm at the down
flooding angle.
o Area ratio for a non-ship shaped floating structure should be greater than
1.3.
• Large Angle Damaged Stability
o There should be a distance greater than 7 degrees between the first
intercept of the righting and heeling arm and the down flooding angle.
• Structural Stress Allowances
o Maximum topside stress of 29.94 ksi
o Maximum column stress of 27.8 ksi
o Maximum pontoon stress of 29.94 ksi
• Mooring
o Dynamic intact mooring analysis factor of safety should be greater than
1.67 as per API codes.
o Dynamic damaged mooring analysis factor of safety should be greater
than 1.25 as per API codes.
o Pretension stress for semi-taut and taut-leg mooring systems should not
exceed 20% of the breaking strength.
o Platform offset for intact mooring system during operations should not
exceed 5% of the water depth.
o Platform offset for intact mooring system during survival mode should not
o Platform offset for damaged mooring system should not exceed 7% of the
water depth.
o Platform must be able to move 150 feet off center in any direction for
drilling purposes.


• Cost
o Cost should be minimized while meeting all other objectives.

Design Basis
The design site is in 5,500 ft. water depth in the South China Sea off the coast of Malaysia.
Singapore has fabrication capacity for both platform topsides and semi-submersible hull only
585 miles from the proposed platform location making Singapore the ideal construction site as
seen Figure 3.

Figure 3: Proposed Project Location

Even though the relatively short distance makes a wet-towing option economical, the ring-
pontoon design considered is not as hydrodynamic as a two-pontoon hull, so designing the hull
for the dry-tow option left greater transportation versatility in the event that a wet tow was too
expensive. A maximum draft for the dry-tow option is 29.52 ft.

The metocean data for the proposed design site was given as follows in Table 1.


Table 1: Metocean Data
Location Malaysia
Max Design
Condition
Operating Extreme
Environmental
10 Year 100 Year
Component
Wind:
Wind Spectrum API RP 2A API RP 2A
Speed, Vw (1 hour31.1 knots 36.9 knots
mean at 10 m)
Sea Waves:
Wave Spectrum JONSWAP3 JONSWAP3
Significant Wave
16.1 ft. 20.7 ft.
Height, Hs
Maximum Wave
30.8 ft. 39.4 ft.
Height, Hmax
Peak Period, Tp 12.7 sec. 13.1 sec.
Peakedness
1.7 1.9
Parameter, γ
Surface Speed, Vc 2.3 knots 2.5 knots

The environment is fairly benign with insignificant swells and a 100 year significant wave height
of about 21 ft. Assuming linear wave theory applies, ABS suggests adding 5 ft. to the 100 year
significant wave height to determine a minimum freeboard or air gap, putting the minimum air
gap at 26 ft.

Riser fatigue and topside separation tank effectiveness are a function of natural heave and
pitch/roll period. As such, natural heave and pitch/roll period were designed to be higher than 20
seconds as per client requests.

HOE recommends heave response to be within +/- 16 ft. for both 10 year and 100 year storm
conditions. In addition, ABS stipulates an acceptable 10 year storm pitch/roll to be a maximum
of 4 degrees and a 100 year storm pitch/roll to be a maximum of 10 degrees.

For large angle intact stability analysis, ABS recommends that the righting arm should be at least
twice the heeling arm at the down flooding angle. Since the semi-submersible is not considered
ship-shape, the required minimum area ratio is 1.3. In a damaged stability condition, ABS
recommends a minimum distance between the first intercept of the righting and heeling arm to
be at least 7 degrees from the down flooding angle.

Design Methodology
The original general hull design was as basic as could be: a square ring pontoon supporting four
square columns with outer column edges flush with the given topside dimensions of 220 ft. by
220 ft. as seen in Figure 4.


Figure 4: Original Basic Hull Design

After calculating small angle stability and response periods, it was difficult to find a set of
dimensions that met the objectives adequately so the pontoon lengths were increased, spreading
the distance between the columns so that the topside deck corners were set on the geographical
centers of the tops of each column as seen in Figure 5.

Figure 5: Modified Basic Hull Design

Acceptable small angle stability and physical center data seemed within reason and response
periods, freeboard and float-on and float-off drafts were also easily obtainable within their


accepted ranges. All following iterations were made with the general basic hull design as seen in
Figure 5.

An Excel spreadsheet was created to calculate all of the wind forces acting on the platform. The
excel spreadsheet can be seen in Appendix 1. To compute the wind force the platform above the
waterline had to first be broken up into 50 foot increments as seen in Figure 6.

Figure 6: Wind Height Divisions

The area of each of the different components in between each 50 foot increment was then
computed. For all the components in between 50 foot increments a height coefficient was
assigned. These coefficients were found from the ABS MODU manual. Also based on the type
of structure each component was (i.e. truss, flat hull, cylinder etc.) assigned a shape coefficient
found in the ABS MODU manual. The forces were computed for every 45 degrees around the
platform. It was determined that the greatest load would come when the wind impacted the rig on
its northeast corner. The following equation was used to determine the wind forces:

F = C S C H CW SV 2

Where CS is a coefficient which comes from the shape of the structure in consideration, CH is the
coefficient of height which is based off the height of the structure above the mean water line. To
determine the height coefficients the semi was broken up into 50 foot increments. The areas of
these increments were then computed which is the value S. Both the coefficient of shape and
height come from recommended figures provided by ABS. The CW coefficient is a standard
coefficient that accounts for several constant factors. V is the average wind speed for a one hour
time interval.

The current forces were calculated using the same spreadsheet listed in


Appendix C: Environmental Loads. Because the geometry of the hull is symmetrical there was
no need to compute the load for more than one side and one corner. Since there is more exposed
surface area when the current is flowing toward the semi at a 45 degree angle the force acting on
the corner of the hull will be greatest at the corners under this condition. The current force was
computed using the equation below which is similar to that of the wind force equation mentioned
above. Unlike the wind force however the projected area of the submerged area did not have to
broken down into smaller areas, however a shape coefficient was computed based on the radius
the corners of the columns and pontoons.

F = C S SV 2

In equation 8 all of the variables have the exact same meaning as those for equation 4 but the
coefficient of shape was calculated from API 2SK Figure C-2 which is shown below in Figure 7.

Figure 7: Semi-submersible Current Drag Coefficients(ABS MODU)

The mean wave drift force was calculated using the force coefficients given in the GMOOR
example file. The wave period must be matched to the value given in the file. Then the force is
the coefficient times the wave amplitude squared.

The program that is used for structural analysis of the semi-submersible is ETABS. This is a
three-dimensional analysis program for building systems. The model that is inputted into
ETABS is a stick figure version of the hull design. The length of the pontoons and topside come
from the distance from center to center of the columns, which is 220 feet. The height of the
model is the distance from the center of the pontoon, entire length of the column, to half of the
topside height, which is 118 feet. Equivalent cross-sectional areas for the beams, pontoon, and
columns were determined by moments of inertia. A rigid plate was placed on top of the columns


to represent the topside. This model, shown in Figure 8 was input into ETABS with forces
applied.

Figure 8: ETABS Structural Model

Two different types of analyses were performed on the structure. One analysis has the structure
submerged in a constant draft of 67.74 ft., shown in Figure 9.

Figure 9: ETABS Model with Waterline

The other analysis contains a static wave that causes the maximum deflection and moments. To
find the maximum deflection of the topside and pontoons, the maximum moment needs to be
determined. ABS code 3-3-A2/3.7 states that “A diagonal wave (30-60 degrees) from the bow
with a length of about one and half times the diagonal distance between pontoon ends will induce
the maximum design longitudinal shear force between the pontoons. In this load case, the
longitudinal forces on the two pontoons and columns are maximized and acting in the opposite
directions. Thus, the horizontal bracings will be subject to the maximum bending moment.” The
diagonal wave was modeled at 45 degrees from the bow, with a wavelength of 466.7 ft, and the
crest of the wave was placed across the diagonal of the semi-submersible, shown in Figure 10.


Figure 10: Static Wave Model with Crest across Diagonal

The same wave was entered with the trough across the diagonal of the structure; this can be seen
in
Figure 11. The two cases involving the static wave gave similar results. From this point
forward, only the case with the crest across the diagonal will be discussed.

Figure 11: Static Wave Model with Trough across Diagonal

The main forces on the hull are from hydrostatic pressures, which are more dominating than
wind, wave, and current forces. The total pressure is determined by the static and dynamic
pressure equations:

Ptotal = Pstatic + Pdynamic
Pstatic = ρ gz
Where: ρ = 1.99 slugs/ft3
g = 32.2 ft/sec2
z = water depth

Pdynamic = ρ gAe kz
Where: A = amplitude of wave
k = wave number = 2π/L


These forces were put into ETABS as two different load combinations, for display reasons,
shown in Figure 12, Figure 13, Figure 14, and Figure 15. The two former figures display the
structure in water, without static wave acting on it, while the two latter figures contain the static
wave with 100 year storm significant wave height and a wavelength 1.5 times the hull diagonal
length as specified by ABS.

Figure 12: Pressures in Positive X, Y, and Z Directions without Wave

Figure 13: Pressures in Negative X, Y, and Z Directions without Wave


Figure 14: Pressures in Positive X, Y, and Z Directions with Static Wave

Figure 15: Pressures in Negative X, Y, and Z Directions with Static Wave

All of the weights of the semi-submersible were then added to the model. These weights include
the topside, columns, pontoon, nodes, risers, and the mooring lines. The weight of the topside is
40,256 kips and was uniformly distributed across the rigid plate on top of the model. The weight
of the columns and pontoons, 4,753 kips and 54,756 kips, respectively, were also inputted as
uniformly distributed loads. The weights of the nodes, 4,159 kips each, were added to the
bottom of the columns. The eight risers, weighing an average of 444 kips each, are located on
opposite sides of the pontoon with four on each side. Located at the top of the columns are the
twelve mooring lines, weighting 68 kips each. This weight distribution can be seen in Figure 16.


Figure 16: Weight Distribution of Semi-Submersible Components

The buoyancy force counteracts the weights of the entire semi-submersible and is equal to
135,028 kips as seen in Figure 17.

Figure 17: Buoyancy Force

Springs were placed at the nodes between the columns and the pontoons to represent the mooring
lines, shown in Figure 18.


Figure 18: Springs Attached at Nodes

The spring stiffness is determined by:
k = F/x

Where: F = average force of mooring line (from GMOOR) at fairlead = 116.8 kips
x = maximum displacement of semi-submersible = 550 ft. (from watch circle
defined by ABS as 10% of the water depth)

Weight, Buoyancy and Stability
Weight approximations were given for each of the main shapes used in the basic hull design
were 9, 12 and 14 lbs/ft3 for columns, pontoon and nodes respectively. The weight of the
columns seen in Figure 20 is modeled as a point mass for stability at the geometrical center of
the column. The buoyancy is modeled as a point force at the geometric center of the submerged
portion of the column with a magnitude equal to the weight of displaced water.

Figure 19: Column Diagram


The weight of the pontoon seen in Figure 20 is modeled as a point mass in the center of each
pontoon- there being four individual pontoons in the ring. The buoyancy is modeled as a point
buoyancy in the geometrical center of the submerged portion of each of the pontoons in the ring.
The force applied at the buoyancy point is equal to the weight of displaced water. For
operational and survival conditions, the pontoon is completely submerged. Only in a float-on or
float-off scenario is the pontoon possibly only partially submerged.

Figure 20: Pontoon Diagram

The nodes were modeled as the densest part of the hull and are the portion of the hull where the
columns intersect the ring pontoon as seen in Figure 21.

Figure 21: Node Diagram

Like the columns and pontoon, the nodes weight is modeled as a point mass located at the
geometric center of the node and the buoyancy is modeled as point buoyancy located at the
center of the submerged portion of the node. Like the pontoon, under both operational and
survival conditions the nodes are completely submerged and are only partially submerged in a
float-on or float-off scenario. The magnitude of force applied at the buoyancy point is equal to
the weight of the displaced volume of water.


Two conditions were considered in regard to given topside horizontal center of gravity: a given
CG of (0,0) (with the origin at the top, middle of the deck), or an offset CG. The ideal scenario
would be a topside with a given centered horizontal center of gravity. Conditions may arise in
fabrication resulting in a slight or even drastic movement of the CG from the topside center that
would have to be permanently offset in ballasting tanks. This offset may require a significant
amount of ballasting to offset, and thus put the float-on/float-off draft over the maximum draft
allowed. If production deadlines are not met, a scenario might occur where most of the topside
is moved out with the hull to the project location and another topside component is derived at a
later time. In this scenario, ballasting would be used to temporarily offset the asymmetrical
weight of topside during float-on/float-off operations and until the missing topside component is
delivered. These scenarios were kept in mind during the ballasting tank design.

Eight steel catenary risers (SCR) were required for production, all at different tensions. It was
initially thought that ballasting tanks would have to be used to counteract the offset CG, but it
was also thought possible to hang the SCRs as evenly as possible off opposite sides of the hull to
minimize CG shift. These risers were placed in the spreadsheet shown in the diagram in Figure
22.


Figure 22: Riser Placement and Ballasting Tank Diagram

Mooring weight and/or tension at fairleads had to be added as well to find the final vertical
center of gravity, draft and freeboard. The mooring tensions (found in a later part of the report)
are also shown at the corners of the diagram in Figure 22: Riser Placement and Ballasting Tank
Diagram.

Resultant center of gravity was calculated from transverse, lateral and vertical moments as seen
in Figure 23.


Figure 23: Balancing Spreadsheet

To satisfy damaged stability criteria recommended in ABS, 6 equally sized tanks were placed in
each of the four pontoons. Tank ballasting levels were determined by an automated Macro
imbedded in a spreadsheet that would first balance the hull by filling some of the tanks to


counteract any offsetting topside CG, then filling them all to reach a 35 ft. freeboard, then
consolidating the ballasting water into as few tanks as a simple algorithm could handle. A
demonstration of an auto-ballasting spreadsheet system can be seen in Figure 24.

Figure 24: Auto Ballasting Example

A float-on/float-off scenario for a (0,0) topside CG analysis simply involved removing riser and
mooring tension and emptying tanks. If the small angle stability was acceptable (GM greater
than 10 ft.) and the draft was less than the maximum allowed by Dockwise equipment (29.52 ft.)
then float-on/float-off would remain a viable option. An example of this scenario can be seen in
Figure 25.


Figure 25: Horizontally Centered Topside CG Float-on/Float-off Scenario

As is evident in Figure 25, the gravity to metacentric height is in excess of the target 10 ft. and
the draft is less than the maximum 29.52 ft. leaving a dry-tow an option for this design.

An offset topside CG for a float-on/float-off scenario was also considered and also seemed more
critical than for an operational situation. The only method the hull has to counteract an offset
topside CG is with ballasting. Any increase in ballasting weight, however, would increase the
already relatively deep draft for float-on/float-off and could quickly remove dry-towing from
transportation options. An example of this scenario is seen in Figure 26.


Figure 26: Excessive topside CG offset for Float-on/Float-off

The draft in Figure 26: Excessive topside CG offset for Float-on/Float-off is highlighted red
signifying an error that was triggered by excessive draft for the float-on/float-off condition.

For quick analysis the gravity to metacentric height (GM) was used to determine right away if
the proposed design dimensions would result in a stable hull or not. This stability analysis was
only valid for small angles of list, so StabCad was used for the large angle stability analysis. An
intact stability analysis was run first resulting in a graph similar to the example shown in Figure
27.


Figure 27: Intact Stability Objective Diagram

For an intact analysis, two basic requirements were to be met: one, that the area ratio is greater
⎛ Area[ A + B] ⎞
than 1.3 ⎜ ≥ 1.3 ⎟ and two, that the righting arm moment is at least 2 times greater
⎝ Area[C + B ] ⎠
than the heeling arm moment at the down flooding inclination angle for 100 knot winds.

A damaged case analysis stipulated that the worst-case scenario of any two consecutive flooded
tanks anywhere in the hull had to result in a minimum 7 degree distance between the first
intercept between the righting arm and heeling arm and the down flooding angle as seen in
Figure 28 in 50 knot winds.


Figure 28: Damaged Stability Objective Diagram

Once dimensions passed small angle stability, freeboard and hydrodynamics tests a spreadsheet
recorded the X, Y and Z positions of each key point used for a StabCad model. A batch file then
ran a Matlab program that took the Excel spreadsheet X, Y, and Z points and saved them as a .txt
file in a StabCad friendly format. The text file was saved as a new StabCad file which was then
run to check large angle stability for intact and damaged cases. A simple example of this
StabCad model can be seen in Figure 29. Down flooding points were placed at the tops of the
columns, and panels were defined according to where solid planes were in the model. Hull
volumes were defined differently than the topside volumes and shape coefficients were applied
to each of the topside components.

Figure 29: StabCad Model


One of the early objective conformity tests was in hydrodynamics, having both heave and
pitch/roll natural periods greater than 20 seconds. A virtual added mass approximation of 100%
the hull weight was used to determine heave period. Two formulas were used to perform these
tests:
2m
THeave = 2Π
ρgAwp

Where:
m = mass
ρ = density
g = gravity
Awp = Water plane area

I
TRoll = 2Π
Awp gGM
Where:
I = Second Moment of Inertia with respect to water plane
GM = distance from the center of gravity to metacenter

In addition to heave and pitch/roll periods, heave and pitch responses, were to be determined;
conforming to ABS recommendations. A simplified dynamic free-body diagram can be seen in
Figure 30.

F(t)

m

k c

Figure 30: Simplified example of heave motion

RAO (Response Amplitude Operator) data was found using a hull of similar dimensions in
GMOOR. RAO is defined as response amplitude of the structure per unit wave height or wave
amplitude. Figure 30 shows a simplified example of heave motion, the governing equation used
for calculation can be seen in the following equation.


•• •
m X + c X + kX = F ( t ) = Re { F0 eiwt }

Where:
m = Equivalent Virtual Mass of System (Real mass +Added mass)
c = Equivalent Viscous Damping Coefficient of System
k = Equivalent Stiffness of System
F(t) =External Force (i.e. Regular Wave Force)

RAO values in conjunction with JONSWAP data were used to develop maximum heave and
pitch/roll responses for different headings using the following equations.

RMS = ∆ω iΣ ( RAO 2 i S (ω ) )

RS = ∆ω i RAO 2 i S (ω )

Where:
ω is in rad/s
and S(ω) is the metocean data calculated using JONSWAP

The limits recommended by ABS are defined as +/-16 feet for all heave maximum responses,
+/-4 degrees for pitch/roll maximum response using 10 year storm data, and +/-10 degrees for
pitch/roll maximum response using 100 year storm data.

Strength and Structural Design - General
As previously stated, the program called ETABS was used to perform the structural analysis.
The model length of the pontoons and topside is 220 feet, while the height of the model is 118
feet. Equivalent cross-sectional areas for the beams, pontoon, and columns were determined by
moments of inertia. The thickness of the columns and pontoon is 1 inch, and modeled as 1.25 in.
to account for scantling, ribs, and stiffeners. A rigid plate was placed on top of the columns to
represent the topside. A stick figure was modeled to represent the semi-submersible, with the
previous stated pressure, weight, and buoyancy loadings. The stresses that are created from
these forces are compared to the ABS code 3-2-3.3 through 3-2-3.11. The values of the stress for
the topside and pontoon from ABS come from the following equations.

Fy
F=
F .S .

Where: Fy = yield strength of steel
F.S. = factor of safety = 1.67

For the columns, the equation for the stress is:


Fcr
F=
F .S .

⎛ Fy 2 ⎞ ⎛ Kl ⎞ 2
Fcr = Fy − ⎜ 2 ⎟ ⎜ ⎟
⎜ 4π E ⎟ ⎝ r ⎠
⎝ ⎠

Where: Fcr = critical stress of steel
E = Modulus of Elasticity
K = effective length factor
r = radius of gyration

By using 50 ksi, common for many types of steels, for the yield strength of the steel used, the
stresses for the topside and pontoon are both 29.94 ksi. For the columns, the stress is calculated
to be 27.8 ksi.

Mooring
The purpose of a mooring system is to limit the semi-submersibles movement so that the
platform remains in place over the equipment on the sea floor. The mooring system does allow
for some movement of the platform so that the amount of tension in the mooring lines is reduced
which allows for smaller diameter mooring lines thus reducing material costs. Also if the
mooring system is to tight then if the platform suffers an impact from a large wave such as a
rouge wave it can cause a very sudden and dramatic acceleration of the platform that can throw
workers to the floor and damage equipment. However, there is a limit to how much the platform
can offset because in most cases with steel catenary risers an offset greater than 5 percent of the
water depths will cause increased fatigue on the risers or break them. Traditionally mooring lines
have consisted of a chain fairlead and ground chain and a wire rope in between the two.
However, for this design analysis it was decided to investigate the use a polyester rope as well at
steel stranded wire. The reason for this is that polyester is much lighter than wire and so
decreases the weight on the platform and has a much hire modulus of elasticity which allows for
more restoring force.

It was decided that a twelve leg taut and semi-taut mooring system should be analyzed along
with a sixteen leg taut and semi-taut. A catenary mooring system was not analyzed because at
5500 foot water depths such systems become very heavy and much more difficult to install. Due
to the limited amount of time available for this design a preliminary analysis was performed that
tested the viability of using polyester rope and stranded wire for taut and semi-taut systems. The
actual material properties used for the polyester rope were those provided by InterMoor and for
the stranded wire the material properties of steel jacketed spiral stranded wire were used, these
properties were also provided by InterMoor.

After preliminary tests, by comparing the polyester rope and stranded wire designs with similar
break strengths, it was found that for a taut mooring configuration, the polyester outperformed


the stranded wire. The wire in fact did not pass API design guidelines because the minimum
factor of safety for the stranded wire configuration was only 1.1. The API guidelines used for
comparison are listed later in this section. The reason that the wire had much lower factors of
safety is because the wire is less elastic than the polyester, initiating fatigue sooner. To
compensate for this, the wire diameter and length were increased until they passed the
guidelines. As a result, the wire mooring configuration was 4.5 times heavier than the polyester
system. Although wire is less expensive per pound than polyester, the cost of materials for the
two designs is still similar. However, when a larger diameter is used for the line, it is more
difficult to install. When testing the semi-taut design, it was discovered that the polyester
allowed the offset of the platform to be greater than the API guidelines. There are two reasons
for this occurrence: elasticity of the polyester and that it is almost neutrally buoyant in water.
However, the semi-taut configuration needs the weight in the mooring lines to help pull the
platform back to its proper position. Due to these features of the polyester line, it was decided
that further analysis for taut legs would only be performed with polyester rope and chain
configurations. For semi-taut configurations only stranded wire and chain configurations would
be considered.

In order to test the four designs, a software program called GMOOR was used. GMOOR is a
software package that allows the user to input the water depth, environmental force coefficients,
which are computed from the environmental loads computed earlier in the report. The location
of the semi-submersible’s fairleads is also entered into the GMOOR input file along with RAO
data. A separate input file is created for the actual mooring lines to be used. This file contains
the physical characteristics of the different mooring lines used, such as weight, design break
strength, and stretch properties. This file also includes the length and diameter of each element
in the mooring line, the depth of the anchor, and the lateral distance from the fairlead to the
anchor. Once these files are created, GMOOR combines them and constructs a diagram of what
the design will look like and analyzes the tensions acting on the mooring lines and the anchor.
GMOOR then allows the user to enter environmental forces and the angle at which they act on
the platform. GMOOR assumes that anchors can never fail, allowing the user to break lines in
order to check the damaged performance of the mooring system. Four designs were fully tested
and optimized using GMOOR. The testing included running a batch file that dynamically
analyzed the environmental forces as being collinear and tested them at 22.5 degree intervals
around the platform. At each of these intervals, GMOOR would test the mooring system with all
lines intact and then break each line sequentially and then reset it and test the next line. By
analyzing the mooring system this way, it was possible to determine not only the response when
the 1st, 2nd or 3rd most loaded line is broken, but the response when any line is broken.
Therefore, GMOOR was able to find the worse case scenario. The optimization was performed
by shortening the line lengths and decreasing the diameters to minimize material costs while
making sure not to sacrifice the safety of the system. The mooring designs were tested against
guidelines created by the American Petroleum Institute (API) and the American Bureau of
Shipping (ABS) which are:

• Maximum Offset for Intact Operational Condition 5% of Water Depth
• Maximum Offset for Damaged/Survival Condition 7% of Water Depth
• Minimum Factor of Safety for Intact Condition 1.67 for Dynamically Analysis
• Minimum Factor of Safety for Damaged Condition 1.25 for Dynamically Analysis


• Polyester Rope/Jacketed Spiral must not touch seafloor

Ideally, this mooring analysis would include analyzing the various environmental loads coming
from the actual average directions and other various angles to the platform rather than examining
them only as collinear. Such situations can cause unexpected responses. However, this was
beyond the scope of the design because the actual current, wind and wave directions were
unknown since there are no obtainable historical records.

Cost
The total cost of this design is approximately $7.2 million, which includes materials and
equipment but not installation costs.

Results

Final dimensions are seen in Figure 31. These dimensions were larger than originally expected,
being governed by down flooding angle rather than the expected freeboard.

Figure 31: Final Hull Dimensions

After completing the excel spreadsheet to calculate the wind, current and wave loads the excel
spreadsheet added up the sum of all the forces to calculate the total environmental load acting on
the platform. Table 2 shows the result of these calculations, for both operational and survival
conditions.


Table 2: Total Environmental Loads
Operational Conditions (kips)
North South West East
381.6 370.0 354.4 349.7
NE NW SE SW
470.2 473.3 462.5 465.6
Survival Conditions (kips)
North South West East
488.6 472.4 450.5 443.8
NE NW SE SW
598.6 603.1 587.8 592.2

ETABS was run with pressures, weights, buoyancy, and spring forces on the model. Figure 32
displays an exaggerated model of the deflections and displacements of the topside, pontoons, and
columns.

Figure 32: Deflections and Displacements

Since the wind, wave, and current forces were not modeled for the structural analysis, the
reaction forces for the springs in the X and Y directions are zero, and the weights of all of the
components are cancelled out from buoyancy to have no reaction in the Z direction, as shown in
Figure 33.


Figure 33: Reaction Forces

Weight, Buoyancy and Stability
Finalized dimensions resulted in the float-on/float-off (labeled as tow out/ loading) draft,
operating draft and optional survival draft seen in Figure 34.

Figure 34: Final Drafts

Auto-ballasting could handle a permanent CG offset of 47 ft in both lateral and transverse
directions for operational and survival conditions as seen in Figure 35.


Figure 35: Maximum Permanent Offset for Operational and Survival Conditions

Due to draft constraints, and offset topside CG is likely more crucial for float-on/float-off
conditions. A maximum acceptable topside CG shift of 10 ft in both lateral and transverse
directions is seen in Figure 36.


Figure 36: Maximum Permanent Offset for Float-on/Float-off Condition

Using previously mentioned equations in a spreadsheet that can be seen in Appendix A:
Hydrodynamics of Motions and Loading, for a hull running on the light side of weight
approximations- response periods were 21.8 and 25.3 seconds for heave and pitch/roll
respectively. Since the auto-ballasting Macro run in the ballasting spreadsheet kept the freeboard
at 35 ft., the area of the water plane was the same for a hull on the heavy side of weight
approximation as one running on the light side. Therefore, a hull running on the heavy side of
the weight approximation had response periods of 21.8 and 25.5 seconds for heave and pitch/roll
respectively. Both heavy and light cases were above the minimum 20 second natural period
objectives.

Table 3 shows maximum responses for heave and pitch/roll for each heading angle used. The
maximum heave response occurs at a 22.5 degree heading and the maximum response for
pitch/roll occurs at a 0 degree heading as can also be seen in bold type in Table 3.
Table 3: Maximum Responses
Heave Maximum Response (ft) Pitch/Roll Maximum Response (°)
RAO Heading 0° 22.5° 45° 67.5° 90° 0° 22.5° 45° 67.5° 90°
10 Year Storm 7.26 8.91 8.46 8.24 8.24 2.49 2.23 1.55 1.58 0.69
100 Year Storm 10.61 12.57 11.78 11.25 11.15 4.26 3.86 2.76 2.22 0.86


Figure 37 through Figure 40 show the RAO, JONSWAP and resultant response plots for each of
the maximum cases.

10 Year, 22.5 Degree, Spectral Analysis (Heave)

1.200

1.000
[RAO,JONSWAP]/peak, Response

0.800 RAO

0.600 JONSWAP
Response
0.400

0.200

0.000
0 0.5 1 1.5
-0.200
w (rad/s)

Figure 37: 10 year, 22.5° heading heave spectral analysis

100 Year, 22.5 Degree, Spectral Analysis (Heave)

1.200

1.000

0.800 RAO

0.600 JONSWAP
Response
0.400

0.200

0.000
0 0.5 1 1.5
-0.200
w (rad/s)

Figure 38: 100 year, 22.5° heading heave spectral analysis


10 Year, 0 Degree, Spectral Analysis (Pitch)

1.200
1.000
0.800 RAO
0.600 JONSWAP

0.400 Response

0.200
0.000
0 0.5 1 1.5
-0.200
w (rad/s)

Figure 39: 10 year, 0° heading pitch/roll spectral analysis

100 Year, 0 Degree, Spectral Analysis (Pitch)

1.200

1.000

0.800 RAO
0.600 JONSWAP
Response
0.400

0.200

0.000
0 0.5 1 1.5
-0.200
w (rad/s)

Figure 40: 100 year, 0° heading pitch/roll spectral analysis

The final max responses compared to the recommended maximum values can be seen in Table 4
and are well within the maximum values recommended by ABS.
Table 4: Actual Hull Response and Recommended Values
Actual Max Recommended Max
Heave 10 Year Conditions +/- 8.91 ft +/- 16 ft
Heave 100 Year Conditions +/- 12.57 ft +/- 16 ft
Pitch/Roll 10 Year Conditions +/- 2.49 ft +/- 4 ft
Pitch/Roll 100 Year Conditions +/- 4.26 ft +/- 10 ft


Strength and Structural Design - General
The shear and moment diagrams due to the applied forces are shown in Figure 41 and Figure 42,
respectively.

Figure 41: Shear Diagram

Figure 42: Bending Moment Diagram

From the shear and moment diagrams and tables, the deflections, moments, and stresses of the
topside beams, pontoons, and columns for both cases are shown in Table 5.
Table 5: Deflections, Moments, and Stresses of Structure
Calculated Stresses from ABS
Deflections Moments Stresses Guidelines
4.37x106 kip-ft
Constant

Topside 1.16 in. (downward) 29.51 ksi 29.94 ksi
Draft

6
Columns 0.04 in. (outward) 13.6x10 kip-ft 23.6 ksi 27.8 ksi
6
Pontoon 0.76 in. (upward) 5.30x10 kip-ft 29.76 ksi 29.94 ksi
6
Topside 1.17 in. (downward) 4.62x10 kip-ft 29.88 ksi 29.94 ksi
Wave
Static

Columns 0.05 in. (outward) 13.8x106 kip-ft 23.7 ksi 27.8 ksi
Pontoon 0.81 in. (upward) 5.59x106 kip-ft 29.92 ksi 29.94 ksi


The deflections are small enough to not cause any damage, and should not cause any motion
discomfort. The stresses caused by the forces acting on the hull fall within the ABS code. The
thickness of 1 inch chosen for the hull is a good value to give small deflections and to keep the
stresses within ABS recommendations, but could possibly be reduced to help with reducing
moments and stresses, along with cost efficiency. Also, the stresses are very close to the
recommended values, which could cause buckling in the beams.

Mooring
When all required data is entered into GMOOR it can generate a plan view or a 3D view of the
inputted mooring system. Figure 43 shows the plan view for the twelve leg taut and semi-taut
configurations, the lines are positioned to 30 degrees from horizontal with 15 degrees between
each line.

Figure 43: 12 Leg Spread

Figure 44 shows the plan view for the sixteen leg taut and semi-taut configurations, the lines are
positioned to 30 degrees from horizontal with 15 degrees between each line


Figure 44: 16 Leg Spread

Figure 45 shows the results of the tests ran in GMOOR for each mooring configuration. The
figure compares the minimum factor of safety for intact and damaged conditions under survival
storm conditions with the API guidelines. It also compares the offset of the platform as percent
of water depth for operational, survival and damaged survival conditions to the recommended
guidelines.

7.00
7% Max Offset
6.00
5% Max Offset
5.00
12 Leg Taut
4.00
1.67 Min Damaged FOS 1.25 Min Damaged FOS 12 Leg Semi -Taut

3.00 16 Leg Taut
16 Leg Semi -Taut
2.00

1.00

0.00
Min FOS Intact Min FOS Max Offset Max Offset Max Offset
Damaged Operational (% Survival (% WD) Damaged (%
WD) WD)

Figure 45: Mooring Analysis Results


For each of the designs the cost was computed using the different line costs as $3.20 per pound
for polyester mooring line, the jacketed spiral costs $2.07 per pound and the chain costs $1.40
per pound. The cost of equipment (chain jacks, fairleads etc.) was $450,000 per line. Figure 46
reflects the cost of each mooring configuration for line material costs and equipment costs,
however it does not represent the installation costs.

Mooring Line Costs

$14,000,000.00

$12,000,000.00
Total Cost in US Dollars

$10,000,000.00

$8,000,000.00

$6,000,000.00

$4,000,000.00

$2,000,000.00

$0.00
12 Leg Taut 12 Leg Semi-Taut 16 Leg Taut 16 Leg Semi-Taut
Design Type

Figure 46: Mooring Material and Equipment Costs

After careful comparison of the different mooring designs it was decided that the twelve leg taut
mooring system would be the best option. This conclusion was developed based on the fact that
the twelve leg taut design while not having the best factor of safeties, it does have values well
above the recommended minimum. Also the twelve leg taut design does have a good offset
distance it is loose enough to not cause the response to be to rigid while it is tight enough that it
doesn’t get too close to recommended max offset. The price of the mooring design also makes
this the most cost effective choice. Figure 28 is a 3D drawing of the mooring system created by
GMOOR. The 12 leg taut system consists of 200 feet of 2.75 inch diameter K4 Studless chain for
the anchor, 7350 feet of 3.94 inch Marlowpoly polyester rope, and 300 feet of 3.0 inch K4
Studless chain for the fairlead. The extra diameter of the chain at the fairlead is to compensate
for the extra fatigue generated at the fairlead. The diameters of these chains reflect the added
diameter for corrosion allowance as specified by API 2SK guidelines which is a range of 0.2 to
0.4 millimeters per service year. It was decided to be a little conservative in the corrosion
allowance so a value of 0.3 millimeter per service year was chosen resulting in an increased of
0.25 inches in diameter for the 20 design life in this project. The tests were performed without
the additional corrosion allowance as suggested by API 2SK. The 300 feet of chain at the
fairlead allows for the platform to offset in any direction up to 150 feet drilling purposes, at the
maximum of 150 feet it was found that the minimum factor of safety in the 10 year storm


condition is 2.63 intact and 1.94 damaged which meets API guidelines. The total cost of this
design is approximately $7.2 million, which includes materials and equipment but not
installation costs. Figure 47 shows the 3D rendering created by GMOOR of the chosen 12 leg
taut leg mooring system.

Figure 47: Chosen 12-leg Taut Mooring Design

Figure 48 shows the final 12 leg taut mooring system design, the final component diameters
including corrosion allowances and component locations.

Figure 48: Final Mooring Design


After designing the mooring configuration the type and size of anchor needed to be determined.
The two options available for a taut leg mooring system are suction piles and vertically loaded
drag embedment anchors. For the vertically loaded drag embedment anchor the Vryhof
Stevmanta anchor was chosen because the information for sizing was available. Using the
Vryhof Anchor Manual and the tension at the anchor found from GMOOR for the chosen
mooring design the necessary size of the Stevmanta anchor was found to be 11.36 tons. For
designing the suction pile InterMoor volunteered to size the anchor base on the tension load
computed in GMOOR. The result was a suction pile 56 feet long with an outside diameter of 6
feet and a wall thickness of 1.65 in. and a top thickness of 5.5 in. The weight of this anchor is
75.5 tons. Because installation and manufacturing costs for both anchors was not available the
decision of which anchor to be used had to be based off the material costs. The steel cost of one
Stevmanta anchor is approximately $51,000 (USD) and the steel cost of one suction pile is
approximately $340,000 (USD). Based on the drastic difference in price the undeniably most
cost effective solution was the Stevmanta vertically loaded drag embedment anchor. The steel
cost for 12 of these anchors was only $612,000 (USD) approximately.

Cost
The final mooring line cost with associated equipment for the chosen 12-leg taut system can be
seen in Figure 49.

Mooring Line Costs

$14,000,000.00

$12,000,000.00
Total Cost in US Dollars

$10,000,000.00

$8,000,000.00

$6,000,000.00

$4,000,000.00

$2,000,000.00

$0.00
12 Leg Taut 12 Leg Semi-Taut 16 Leg Taut 16 Leg Semi-Taut
Design Type

Figure 49: Mooring Cost Summary

The Stevmanta anchors chosen were priced at $51,000 (USD) a piece, running a total capital cost
of $612,000 (USD). Final hull dimensions with approximate densities and $12,000 per ton, cost
approximately $184,000,000 (USD) capital. The total capital cost was approximately
$192,000,000 (USD).


Conclusion
All of the following objectives of the chosen competencies have been met in excess of minimum
recommendations stipulated by pertinent regulatory agencies.
• Drafts & Freeboard
o A maximum float-on float-off draft of 29.52 ft. determined by Dockwise
boat capacity.
o A minimum operational freeboard (air gap) of 35 ft. based on metocean
data.
• Hydrodynamic responses
o A minimum natural heave and pitch/roll period of 20 seconds based on
client requests for optimum operation of topside separation tanks and
risers.
o A maximum of +/-16 ft heave response based on ABS recommendations.
o A maximum 4◦ pitch/roll for 10 year storm conditions and a maximum 10◦
pitch/roll for 100 year storm conditions based on ABS recommendations.
• Large Angle Intact Stability
o Righting Arm should be greater than 2 X the heeling arm at the down
flooding angle.
o Area ratio for a non-ship shaped floating structure should be greater than
1.3.
• Large Angle Damaged Stability
o There should be a distance greater than 7 degrees between the first
intercept of the righting and heeling arm and the down flooding angle.
• Structural Stress Allowances
o Maximum topside stress of 29.94 ksi
o Maximum column stress of 27.8 ksi
o Maximum pontoon stress of 29.94 ksi
• Mooring
o Dynamic intact mooring analysis factor of safety should be greater than
1.67 as per API codes.
o Dynamic damaged mooring analysis factor of safety should be greater
than 1.25 as per API codes.
o Pretension stress for semi-taut and taut-leg mooring systems should not
exceed 20% of the breaking strength.
o Platform offset for intact mooring system during operations should not
o Platform offset for intact mooring system during survival mode should not
o Platform offset for damaged mooring system should not exceed 7% of the
water depth.
o Platform must be able to move 150 feet off center in any direction for
drilling purposes.
• Cost
o Cost should be minimized while meeting all other objectives.


For the eight aforementioned competencies:

1. General Arrangement and Overall Hull Design
2. Wind and Current Loading
3. Local and Global Loading
4. Weight, Buoyancy and Stability
5. Hydrodynamics of Motions and Loading
6. General Strength and Structural Design
7. Mooring
8. Cost

The general hull design had columns heights larger than would be expected. The governing
objective forcing the columns higher was the location of the modeled down flooding points
instead of the expected minimum freeboard. Further StabCad analysis is suggested to confirm
this. If indeed the down flooding angle is the governing factor, the down flooding points should
be moved up to the topside deck or above to reduce column height. A reduction of column
height in either case will allow a reduction in required pontoon and node sizes further,
significantly decreasing capital cost.

Sloshing analysis at the natural heave and pitch/roll periods for the tank size should be
completed to see how sloshing could effect hull motions before partially filled tanks are accepted
for long-term ballasting and trimming design. If sloshing proves to be an issue, tanks should be
filled as full as possible, or fitted with floating-top tanks.

A significant variable in cost for anchors is in installation. Installation costs could easily swing
the chosen anchor design from Stevmanta to SEPLAs. There is also a discrepancy between
industry review and product recommendations as to whether drag embedment anchors such as
Stevmantas are ideal for a taut-leg mooring system or not. Some suggest that they are ill-fitted
due to a lack of precision placement and the SEPLAs would be a better choice. Because of their
suction embedment, however, SEPLAs are expected to run a higher installation cost.
Clarification as to the precision placement capabilities of Stevmanta anchors and the associated
installation costs for the three different types of anchors analyzed (Suction Piles, Stevmanta and
SEPLAs) need to be completed before placement.


References
American Bureau of Shipping (ABS) Guide for Building and Classing Facilities on Offshore
Installations, Houston, TX. 2000

American Bureau of Shipping. Guide for Building and Classing Floating Production
Installations, Houston, TX. 2004

American Bureau of Shipping. Rules for Building and Classing Mobile Offshore Drilling Units,
Houston, TX. 2004

American Bureau of Shipping. Rules for Building and Classing Steel Vessels, Houston, TX.
2005.

American Petroleum Institute (API). Interim Guidance for Gulf of Mexico MODU Design
Practice- 2006 Hurricane Season (API RP95F), Washington D.C. 2006

American Petroleum Institute (API). Recommended Practice for Design and Analysis of Station
Keeping Systems for Floating Structures, First Edition. Washington D.C., June 1995. (API 2SK)

AutoDesk, Inc. AutoCAD Software, 2004

Computers and Structures, Inc., ETABS, Nonlinear Version 9.0.5, 2005

Global Maritime Consultancy, Ltd. Gmoor32, Version 9.406c, 2001

Gunter and Associates, Inc. StabCAD User’s Manual, 1999.


Appendix A: Hydrodynamics of Motions and Loading
The following spreadsheet was used to take given data, hull dimension changes and estimated
densities and weights to process some of the required hydrodynamic properties of the semi-
submersible. Design criterion being monitored per design iteration can be seen at the top of the
spreadsheet below.


Appendix B: Ballasting
The following spreadsheets give examples of design processes in which varied centers of topside
gravity location are tested versus freeboard, hydrodynamic responses, small angle stability,
maximum float-on draft and hull ability to counter-ballast offset center of gravity. The topside
center of gravity can be seen in the white cells on the left and a partially filled ballasting tank is
signified by the blue cells in the plan view of the tank arrangement diagram. Percentage that an
individual tank is filled is labeled on that individual tank and riser and mooring tensions are zero
in a float-on, float-off scenario. Also included in this appendix is the table used to calculate
moment arms used in ballasting.


Appendix C: Environmental Loads
The following spreadsheets were used in conjunction with shape and height coefficients as per
regulation to calculate concurrent wind, current and mean wave drift environmental loads on the
semi-submersible.

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