Windfloat

JOURNAL OF RENEWABLE AND SUSTAINABLE ENERGY 2, 033104 ͑2010͒

WindFloat: A floating foundation for offshore wind turbines
Dominique Roddier,1,a͒ Christian Cermelli,2 Alexia Aubault,2 and
Alla Weinstein1
1
Principle Power, Inc., Seattle, Washington, USA
2
Marine Innovation and Technology, 2610 Marin Ave., Berkeley, California 94708, USA
͑Received 8 January 2010; accepted 2 May 2010; published online 15 June 2010͒

This manuscript summarizes the feasibility study conducted for the WindFloat tech-
nology. The WindFloat is a three-legged floating foundation for multimegawatt
offshore wind turbines. It is designed to accommodate a wind turbine, 5 MW or
larger, on one of the columns of the hull with minimal modifications to the nacelle
and rotor. Potential redesign of the tower and of the turbine control software can be
expected. Technologies for floating foundations for offshore wind turbines are
evolving. It is agreed by most experts that the offshore wind industry will see a
significant increase in activity in the near future. Fixed offshore turbines are limited
in water depth to ϳ30– 50 m. Market transition to deeper waters is inevitable,
provided that suitable technologies can be developed. Despite the increase in com-
plexity, a floating foundation offers the following distinct advantages: Flexibility in
site location; access to superior wind resources further offshore; ability to locate in
coastal regions with limited shallow continental shelf; ability to locate further off-
shore to eliminate visual impacts; an integrated hull, without a need to redesign the
transition piece between the tower and the submerged structure for every project;
simplified offshore installation procedures. Anchors are significantly cheaper to
install than fixed foundations and large diameter towers. This paper focuses first on
the design basis for wind turbine floating foundations and explores the require-
ments that must be addressed by design teams in this new field. It shows that the
design of the hull for a large wind turbine must draw on the synergies with oil and
gas offshore platform technology, while accounting for the different design require-
ments and functionality of the wind turbine. This paper describes next the hydro-
dynamic analysis of the hull, as well as ongoing work consisting of coupling hull
hydrodynamics with wind turbine aerodynamic forces. Three main approaches are
presented: The numerical hydrodynamic model of the platform and its mooring
system; wave tank testing of a scale model of the platform with simplified aerody-
namic simulation of the wind turbine; FAST, an aeroservoelastic software package
for wind turbine analysis with the ability to be coupled to the hydrodynamic model.
Finally, this paper focuses on the structural engineering that was performed as part
of the feasibility study conducted for qualification of the technology. Specifically,
the preliminary scantling is described and the strength and fatigue analysis meth-
odologies are explained, focusing on the following aspects: The coupling between
the wind turbine and the hull and the interface between the hydrodynamic loading
and the structural response. © 2010 American Institute of Physics.
͓doi:10.1063/1.3435339͔

I. INTRODUCTION
Currently, there are a number of offshore wind turbine floating foundation concepts in various
stages of development. They fall into three main categories: Spars, tension leg platforms ͑TLPs͒,

a͒
Author to whom correspondence should be addressed. Electronic mail: dominique.roddier@marineitech.com. Tel.: 510-
200-0530 ext 101.

1941-7012/2010/2͑3͒/033104/34/$30.00 2, 033104-1 © 2010 American Institute of Physics

033104-2 Roddier et al. J. Renewable Sustainable Energy 2, 033104 ͑2010͒

and semisubmersible/hybrid systems. A barge-type support structure has been studied1 but is not
included in this discussion due to its significant angular motions that hinder its commercial de-
velopment. In general terms, spar type has better heave performance than semisubmersibles due to
its deep draft and reduced vertical wave-exciting forces, but it has more pitch and roll motions
since the water plane area contribution to stability is reduced. TLPs have very good heave and
angular motions, but the complexity and cost of the mooring installation, the change in tendon
tension due to tidal variations, and the structural frequency coupling between the mast and the
mooring system are three major hurdles for such systems. When comparing floater types, wave
and wind-induced motions are not the only elements of performance to consider. Economics play
a significant role. It is, therefore, important to carefully study the fabrication, installation, com-
missioning, and ease of access for maintenance methodologies.2,3
Even though there have been a few visionary papers on the topic of floating wind turbines,
significant research and development efforts only started at the turn of this century.4 In the U.S.,
researchers from NREL and MIT started a significant R&D effort5 with the development of
coupled hydroaerotools,6–8 while model test campaigns were performed at Marintek in Norway on
a spar hull,9 the first version of the HyWind spar concept. The use of a semisubmersible hull as a
floating foundation was proposed independently by Fulton et al.10 and Zambrano et al.11 The latter
paper’s proposed design was a MiniFloat hull, the predecessor of the presented WindFloat
design.12
Over the past few years, academic interest in floating foundations for offshore wind turbines
has reached industry, and a significant amount of funding has been allocated to prototype devel-
opment. Leading the effort, shown in Fig. 1 from top left to bottom right, are the Statoil Norsk-
Hydro Hywind spar, ͑top left͒, the Blue H TLP recent prototype ͑top right͒, the SWAY spar/TLP
hybrid ͑bottom left͒, and the Force Technology WindSea semi submersible ͑bottom right͒.
The WindFloat hull is semisubmersible fitted with heave plates. Extensive technical qualifi-
cation of the hull has been performed over the past 5 years by Marine Innovation & Technology.
Multiple studies have been performed on the MiniFloat—the trademark of the original hull
name—and are published in permanent literature.13–15. These include model tests, hydrodynamic
and structural studies, along with specific tasks based on oil and gas and other industry require-
ments. The work described herein is based on the learning from those previous studies.
The WindFloat system described in this paper aims at enabling floating offshore wind tech-
nology by providing both technical and economical solutions. Its intent is to provide acceptable
static and dynamic motions for the operation of large wind turbines while limiting expensive
offshore installation and maintenance procedures.16–18
The challenges associated with design and operations of floating wind turbines are significant.
A floater supporting a large payload ͑wind turbine and nacelle͒ with large aerodynamic loads high
above the water surface challenges basic naval architecture principles due to the raised center of
gravity and large overturning moment. The static and dynamic stability criteria are difficult to
achieve especially in the context of offshore wind energy production where economics requires the
hull weight to be minimal.19,20
The following fundamental aspects must be addressed to design such system: ͑1͒ The influ-
ence of the turbine on the floater and ͑2͒ the influence of the floater motions on the turbine
performance. A large body of work has been published on the hydrodynamics of floating plat-
forms; see Refs. 21 and 22 for comprehensive overviews. Hydrodynamics of a minimal floating
platform with similar substructure was discussed by Cermelli and Roddier.23 Wind loads on float-
ing structures discussed in the above references are normally computed using a simple relation
between the apparent wind speed and loading based on empirical drag coefficients or results from
wind-tunnel tests. In the case of a floating offshore wind turbine, wind load components generated
by the turbine and their effects on platform motion are significant and may lead to coupling
effects, which cannot be accounted for using conventional methods.
The following methodology is applied in this paper, with increasing level of refinement of the
coupling effects between the wind turbine and platform motion. In the first step, consisting of
global sizing of the floater, coupling between the turbine and floater is accounted for using the

033104-3 WindFloat: Offshore floating wind J. Renewable Sustainable Energy 2, 033104 ͑2010͒

FIG. 1. HyWind ͑spar͒, blue H ͑tension leg͒, SWAY ͑tension leg/spar͒, and WindSea ͑semisubmersible͒.

following approximation: The wind thrust is determined by assuming that the base of the turbine
is fixed and it is applied as force and overturning moment at the base of the mast. This approach
is further described in Ref. 11.
The second step involves time-domain simulations of the hydrodynamic response of the
platform using TIMEFLOAT software. The software was modified to compute wind turbine loads
based on an equivalent drag model, which provides suitable wind thrust at the hub, and also
generates aerodynamic damping. Gyroscopic effects due to the gyration of the rotor coupled with
platform rotations are also included. This model is relatively simple to implement numerically, and
could also be adapted to an experimental setup in order to verify the platform motion predictions
during wave tank testing of a small-scale model. Results obtained at the UC Berkeley ship-model
testing facility are presented. This model does not account for turbine flexibilities and the various
control systems installed on large wind turbines, which have the ability to pitch the rotor blades
resulting in variable thrust and torque, in order to keep the rotor speed constant and the tower
stable, despite variable wind velocities.
In the third and most advanced step, the aeroservoelastic calculation software FAST developed
at the National Renewable Energy Laboratory ͑NREL͒5,8,18 was coupled with the hull hydrody-
namic software TIMEFLOAT to compute the platform motion and wind turbine loads including the
effects of turbine dynamics and the effect of platform motion on the resulting aerodynamic forces.
This offers the ability to compute simultaneously the effects of the mooring system, water-


entrapment plates, as well as all wind-induced loads on the turbine. The methodology is similar to
that of Jonkman1 but coupling with TIMEFLOAT allows accurate modeling of the nonlinear viscous
forces generated by the water-entrapment plates.
To address the influence of the floater motion on turbine performance, a study was performed
in which floater motions determined using the approach presented in this paper were applied at the
base of the mast and turbine performance was evaluated. The MSC.ADAMS with the ADAMS-TO-
AERODYN interface software allows for motion time series input, similar to earthquake loading.
The resulting forces in the various components of the turbine were compared to the case of a fixed
base. Results of this study will be published shortly.
As part of the design qualification process, a global structural analysis must be performed and
structural sizing and reinforcement of the components of the WindFloat were achieved. The
structural assessment of the design necessitates the use of a methodology and design criteria that
account for the specificities of the structure. Large wind forces and hydrodynamic loading need to
be accounted for accurately. In the absence of full-scale experience, the foundation is designed
according to a combination of recommendations for offshore oil and gas platforms, and for fixed
offshore wind turbines. To ensure that the design is sufficiently conservative, an extensive numeri-
cal analysis is carried out on all novel parts of the structure, such as the truss connecting the
columns together, and the turbine tower and its interface with the hull. In a later phase of the
project, structural optimizations of the platform will be carried out to reduce overall steel weight.
A review of the available design standards for the WindFloat is presented briefly, along with
a summary of the main characteristics of the platform and preliminary scantling of the columns.
Sections XVI and XVII of the present paper focuses on the design of the truss and tower with
finite-element analysis using the full description of environmental loads on the platform from
hydrodynamic analysis. Strength and fatigue analyses are performed. The design of the tower is of
particular interest since it is at the interface between the floater and the wind turbine.
Space does not permit a complete description of the system, in particular, wall thicknesses in
various parts of the structure. The intent of this paper is to not provide specific results for a given
geometry, but rather to expose practical methodologies that can be used for design, while includ-
ing all significant hydrodynamic and aerodynamic loading contributions.

II. STANDARDS

There are presently no standards specific to floating offshore wind turbines. There are, how-
ever, rules and guidelines for offshore floaters and for offshore fixed wind turbines. Saiga et al.24
had a very useful discussion on the various design guidelines. In the scope of this preliminary
work, the following documents provided sufficient information for the framework of the project.
We note that the IEC standards are very similar to those of DNV and Germanischer Lloyd. The
latter were used for this work.

A. Hull and mooring
• American Bureau of Shipping ͑ABS͒
• Guide for Building and Classing Floating Production Installations, 2004
• Rules for Building and Classing Mobile Offshore Drilling Units, 2006
• American Petroleum Institute ͑API͒
• API RP 2SK, Recommended Practice for Design and Analysis of Stationkeeping Systems for
Floating Structures, 2005
• API RP 2SM, Recommended Practice for Design, Manufacture, Installation, and Mainte-
nance of Synthetic Fiber Ropes for Offshore Mooring, 2001
• API RP 2A-WSD Recommended Practice for Planning, Designing and Constructing Fixed
Offshore Platforms—Working Stress Design, 22nd edition


B. Safety
• International Maritime Organization ͑IMO͒
• IMO International Convention for the Safety of Life at Sea ͑SOLAS͒, 1974

C. Offshore turbine
• Germanischer Lloyd ͑GL͒
• Guideline for the Certification of Offshore Wind Turbines, 2005
An alternative set of design codes published by Det Norske Veritas ͑DNV͒ will be considered
in the next phase of work. These include:
• DNV-OS-C101 Design of Offshore Steel Structures, General ͑LRFD method͒, April 2004
͓October 2007͔
• DNV-OS-C103 Structural Design of Column Stabilized Units ͑LRFD method͒, April 2004
͓October 2007͔
• DNV-OS-C201 Structural Design of Offshore Units ͑WSD method͒, April 2005 ͓April 2008͔
• DNV-OS-C301 Stability and Watertight Integrity, January 2001 ͓April 2007͔
• DNV-OS-C401 Fabrication and Testing of Offshore Structures, April 2004 ͓October 2007͔
• DNV-RP-A203, Qualification Procedures for New Technology. Sept. 2001
• DNV-OS-J101 Design of Offshore Wind Turbine Structures, October 2007
• DNV-OS-J102 Design and Manufacture of Wind Turbine Blades, Offshore and Onshore
Wind Turbines, October 2006

III. WINDFLOAT DESCRIPTION
The WindFloat technology consists of a column-stabilized offshore platform with water-
entrapment plates and an asymmetric mooring system. A wind turbine mast is positioned directly
above one of the stabilizing columns ͑see Fig. 2͒.
It is comprised of the following elements:
• Three columns, which provide buoyancy to support the turbine and stability from the water
plane inertia. These columns are commonly used elements in floating offshore platforms and
one may rely on standard industry criteria, such as the ABS rules for column-stabilized units
for their design. The external cylindrical shell is stiffened with regularly spaced ring girders
and vertical L-shape stringers to provide sufficient local and global buckling stiffness to the
column. Scantling of the structural elements of the hull aims to determine the thickness of
shells, girders, and webs, as well as the size of their stiffeners and flanges. Since deeper shells
are subject to larger pressure loads, the hull is divided horizontally into four sections that are
sized according to their largest head overflow. This helps reduce the amount of steel required
to build the columns. It is important to note that such rules have been designed to extremely
low failure rates for structures undergoing heavy operational burden, such as the Mobile
Drilling Units. Constraints include the ability to withstand collisions with supply vessels, the
ability to support heavy equipment including rotating machinery, and frequent moves over
large distances. These will undoubtedly result in overly conservative scantlings for offshore
renewable energy systems. Further studies will be aimed at minimizing structural weight
while ensuring sufficient robustness, and will require extensive use of reliability analysis.
• Horizontal plates at the bottom of the columns, which ͑1͒ increase the added mass, hence
shift the natural period away from the wave energy, and ͑2͒, increase the viscous damping in
roll, pitch, and heave. Stiffeners cantilevered from the bottom of the columns with bracing
tying these stiffeners back to the columns support the plates. The water-entrapment plates
provide additional hydrodynamic inertia to the structure due to the large amount of water
displaced as the platform moves. In addition, vortices generated at the edge of the plates
generate large damping forces that further impede platform motion. Structural design of the
water-entrapment plates at the keel had to be carried out numerically since design codes do
not provide specific guidelines for such components. The authors have performed finite-
element analysis of the heave plates for a variety of projects, including a minimal water-


FIG. 2. Detail of structural reinforcement of water-entrapment plate on WindFloat.

injection platform for deep water marginal oil and gas fields, which is similar in payload and
displacement, and whose water-entrapment plates have the same edge length and surface
area. The results described by Aubault et al. ͑2006͒ are used to determine the size of stiff-
eners and stringers on the water-entrapment plate, as illustrated in Fig. 3.
• Permanent water ballast, inside the bottom of the columns, to lower the platform to its target
operational draft, once installed. An active ballast system moves water from column to
column to compensate for the mean wind loading on the turbine. This movable ballast
compensates for significant changes in wind speed and directions. It aims at keeping the mast
vertical to improve the turbine performance. Up to 200 ton of ballast water can be transferred
in approximately 30 min using two independent flow paths with redundant pumping capa-
bility. The active ballast compartment is located in the upper half of each column. The


FIG. 3. WindFloat hull and turbine.

damage design case includes the possibility of all the active ballast water being in the worse
compartment.
• Six mooring lines, made of conventional components ͑drag-embedment anchors, chains,
shackles, fairleads, and chain jacks͒.
• An offshore wind turbine, with as little requalification that is possible from existing fixed
offshore turbines. The tower is made of a number of sections with tapered diameter and
constant wall thickness that are welded together. At its lower end, the turbine tower extends
into the column in order to maximize continuity of the structure, leading to minimized stress
concentration in critical areas of the structure where bending moments are highest ͑due to
wind-induced overturning moment͒ and where large tubulars connect to the other stabilizing
columns. The connection is located above the wave zone, with a clearance above the largest
wave crests. The tower diameter is smaller than the column. A heavily stiffened top of
column section is designed to carry the tower loads into the column shell. The yaw bearing
is installed at the top of the tower and keeps the turbine headed into the wind.

The WindFloat, in its described configuration in this paper, has dimensions listed in Table I.
We note that this is not a final design and that each specific wind farm, being subjected to different
wind and wave environments, will have variations from this configuration. It is also noted that the
present design has significant safety margins. Subsequent design work was performed by the

TABLE I. WindFloat main dimensions.

User-input hull dimensions

Column diameter 35 ft 10.7 m
Length of heave plate edge 45 ft 13.7 m
Column center to center 185 ft 56.4 m
Pontoon diameter 6 ft 1.8 m
Operating draft 75 ft 22.9 m
Airgap 35 ft 10.7 m
Bracing diameter 4 ft 1.2 m
Displacement 7833 st 7105 ton


FIG. 4. Turbine thrust vs wind speed.

authors since these initial studies indicate that the hull presented in this paper has the capability to
support the loading forces of what can be expected of typical wind turbines with rated power up
to 10 MW.
The stabilizing columns are spread out forming an equilateral triangle between the three
column centers. A boat landing is installed on one or two of these columns to access the structure.
The columns are interconnected with a truss structure composed of main beams connecting col-
umns and bracings connecting main beams to columns or other main beams.
Minimal deck space is required between the tops of the columns. Figure 2 shows a gangway
connecting one column to the next and is the main deck element. Additional areas may be used to
support secondary structures, such as auxiliary solar cells, and to provide access around the wind
turbine mast. The height of the deck is positioned such that the highest expected wave crests will
not damage deck equipment or the turbine blades. The structure is anchored to the seabed using
conventional mooring lines arranged in an asymmetrical fashion. The turbine supporting tower is
carrying more mooring lines than the other two.

IV. WIND TURBINE
The philosophy of the WindFloat is to accommodate turbines from different manufacturers. It
is therefore important to work with the turbine manufacturers and use their data to optimize the
design. Figure 4 shows a typical turbine thrust loading on the tower as a function of wind speed.
This is a very useful information, which is used to understand the mean force and the moment the
turbine will apply on the top of the column, and is a key driver to the sizing of the hull.
Figure 5 shows a typical turbine rated power as a function of wind speed. This information is
necessary to predict the total amount of electricity that the turbine will produce when it is linked
directly with the wind data for a specific site.
In the initial phase of this feasibility study, conservative assumptions were made to develop
the platform global sizing. It was assumed that the wind-induced thrust at the top of the mast could
be estimated based on a drag coefficient applied to the overall area covered by the rotor, i.e., a 413
ft ͑126 m͒ diameter disk. The selected drag coefficient was 1.2 for wind speeds up to 12 m/s and
0.4 thereafter up to 25 m/s wind. The turbine was assumed parked for higher wind speeds. This
model is conservative and has been being significantly improved since these studies. NREL
turbine code FAST has been integrated with MI&T’s floating body motion prediction code TIME-
FLOAT. Fully coupled simulations can be performed to better understand the influence of hull
motions on the turbine and vice versa.


FIG. 5. Turbine rated power vs wind speed.

The turbine and mast main specifics for a 5 MW turbine are listed in Table II. These numbers
are specific to a manufacturer but most large turbines of the same size are very similar with respect
to principal weights and dimensions.

V. ENVIRONMENTAL DATA
Currently, two concurrent sites are being evaluated for the WindFloat: First, the West coast of
the U.S., from Northern California to Washington; second, the Atlantic coast of Portugal. In both
cases, the wind resources are acceptable for a wind farm development and the wave conditions are
quite severe. This paper focuses on the WindFloat design performed for the Western U.S. site. A
detailed metocean analysis was performed for the site shown in Figure 6. 25 years of wind and
wave data from the National Oceanic and Atmospheric Administration ͑NOAA͒ buoy 46022 were
used for the analysis.

A. Geographical location of the wind farm
The WindFloat is envisioned to be located 15–20 km ͑10–12 miles͒ offshore so as to minimize
risks/nuisance to the general public, and to mitigate the view impact from the coastline. The water
depth is assumed to be 500 ft ͑ϳ150 m͒. The WindFloat is intended to be suitable for open ocean
locations with relatively harsh metocean conditions over a wide range of water depths, and most
likely will be cost efficient at or beyond 50 m water depths. In this design phase, the conditions
assumed are those of Northern California, as shown in Fig. 6. This location was chosen in an early
assessment based on the good wind resources and the geographical proximity of Humboldt Bay.
The metocean conditions north of the Eureka site ͑Oregon, Washington͒ will be typical of the

TABLE II. 5 MW turbine characteristics.

Rotor mass 121 st 135 mt
Nacelle mass 264 st 294 mt
Mast mass 383 st 425 mt
Mast diameter 26.25 ft 8 m
Rotor diameter 413.4 ft 126 m
Clearance between TOC and bottom of blade 16.4 ft 5 m


FIG. 6. WindFloat location and metocean data buoy.

Eureka site, but will most likely have slightly larger significant wave height ͑Hs͒ value. There are
a number of NOAA buoys that can be used to derive the exact extreme conditions and will be used
in the detail design of a specific project.

B. Operational and survival „extreme… conditions
From the wave data, three design sea states were defined. An operational case is shown in
Table III, an extreme sea state with a wind gust, as defined in GL design guidelines and shown in
Table IV and the 100 year storm shown in Table V. The extreme wave event assumes a 100 year
return period in keeping with common practice from the offshore industry. It is noted that offshore
wind turbine codes, such as Germanischer Lloyd “Guideline for the Certification of Offshore Wind
Turbines,” only require 50 year return period events to be considered for design. Although like-
lihood of failure of an offshore wind turbine foundation may be comparable to that of an offshore
platform, the consequences are far less severe because they are unmanned structures and do not
have the potential for large pollutions. In the context of this feasibility study, 100 year return
period events were considered for preliminary design. This offers an element of robustness, which
is useful since the design typically evolves significantly at this early stage. Once the project
feasibility has been demonstrated, a reliability study will be conducted to set the final criteria for
the design of an offshore wind farm, with the objective of minimizing the overall project cost.
The data were also processed to find out if there are any directional effects between the wind
and waves. It was remarked that the wind and waves are collinear when they are both coming from
the north; however, when the wind came from the south, the waves had a tendency to come from
the west. Hence directional criteria are shown in Table VI.

TABLE III. Operational metocean case.

Sea state Operational

Significant wave height 7.8 ft ͑2.4 m͒
Peak period 10 s
Wind speed at 10 m elevation 40 ft/s ͑12.2 m/s͒
Current speed 0.98 ft/s ͑0.3 m/s͒


TABLE IV. ECG.

Sea state ECG

Peak period 10 s
Wind speed at 10 m elevation 0 to 85 ft/s ͑25.9 m/s͒ in 10 s

VI. OPERATIONAL REQUIREMENTS
The operational requirements provided in this section are typical of an offshore floater. They
form the basis of the initial design to be carried out as development work progresses further.

A. WindFloat normal operation „anchored…
As a base case, the WindFloat is assumed to be permanently moored using a conventional
anchoring system made of a chain jack, chain and wire sections, and an anchor. That means the
WindFloat will not be disconnected in case of extreme weather conditions.
The main purpose of the WindFloat is to generate electricity from the wind turbine. Therefore,
the WindFloat should be designed to maximize the amount of time the turbine is operational. Since
existing turbines stop operating at 25 m/s wind speed, it is desirable for the wave-induced motions
in waves typical of those wind speeds not to interfere with this operational limit. It is anticipated
that the turbine may need to be strengthened to survive extreme storms in their parked positions
due to the additional inertial accelerations caused by the wave-induced motions.
A closed-loop active ballast system is designed to compensate for the mean wind force and
direction. Water needs to be moved between columns such that the mast remains vertical, hence
optimizing electricity production. It is not envisioned that this active ballast system compensates
for the dynamic motions of the floater, as it should have a response time of between 30 and 60
min. In rapidly changing wind conditions, including wind turbulence, pitching of the blades
͑reduction in thrust͒ is performed to help minimize the wind-induced trim if necessary. The
response time for this mode is of the order of minutes or less.

B. Storm conditions
The WindFloat is designed to withstand very significant storms without failure. Borrowing
from the requirements for oil and gas platforms, the WindFloat hull was designed for the 100 year
return storm at the site.
There are three separate regimes for the turbine that are wind speed dependent.
͑1͒ The blades are optimally pitched to maximize electricity production.
͑2͒ The blades are pitched as to minimize the loading on the blades, but the turbine keeps
spinning.
͑3͒ The rotor is not spinning and the turbine is either idling or locked down, in survival mode,
depending on the severity of the environment.

TABLE V. 100 year storm.

Sea state 100 year storm

Peak period 17 s
Wind speed at 10 m elevation 85 ft/s ͑25.9 m/s͒


TABLE VI. Directional extreme events.

Wind dominated Wave dominated

Collinear case 1 Bi case 2 Collinear case 3 Bi case 4

Hs m 11.2 11.2 13.5 13.5
Tp s 16.67 16.67 19 19
Wave direction deg 0 270 0 270
Wind speed m/s 19.6 25.5 18.4 23.2
Wind direction deg 0 180 0 180
Current speed m/s 0.59 0.76 0.55 0.70
Current direction deg 0 180 0 180

This is typical of large wind turbines. However, as the platform moves in large waves, one
must recognize that regime 3 may occur sooner than expected due to the WindFloat wave re-
sponse. As part of the turbine qualification work, a specific turbine operational envelope must be
defined.

C. Emergency operations
The philosophy behind the emergency shut down system is to preserve the structure and
minimize the loss of equipment. Since the platform is normally unmanned, both automated and
remote shut down procedures must be in place.
The following points are a nonexhaustive list of key actions that should trigger a series of
checks and possible shutdown of the turbine.
• Failure of the active ballast system, noted by either a large mean pitch that does not diminish,
coupled with an abnormal power requirement of the pumps.
• Water leaks in a column, noted by a heel of the platform into that column, which cannot be
compensated by the functioning active ballast system.
• Large accelerations measured in the turbine, which would induce stresses above the design
threshold.
• Inability for the turbine to rotate into the wind, noted by a discrepancy between the measured
wind direction and the turbine heading.
• Power failure.
• Loss of communication between the WindFloat and the remote operator.

There should be enough backup power available on the WindFloat to complete an emergency
shutdown procedure and keep emergency and safety systems, such as navigation lights, opera-
tional until maintenance can be performed.

VII. FABRICATION, INSTALLATION, AND COMMISSIONING REQUIREMENTS
There are very strong synergies between the WindFloat hull and the MiniFloat oil and gas
platform in terms of fabrication, installation, and commissioning. The MiniFloat design philoso-
phy is to optimize the economics by reducing cost in all phases of the project. The same philoso-
phy is applied here and design decisions are made after clearly understanding their impact to all
stages of the process.

A. Fabrication: Quayside
The mast and turbine are fully integrated with the platform at quayside during fabrication. The
platform is then towed to its installation site using a tugboat. Due to its exceptional stability
performance, this operation can be conducted with minimal restrictions on weather conditions.


Unlike fixed offshore wind foundations, there is no requirement in lifting the turbine at the
offshore installation site, which was proven to be difficult and costly. Such heavy lift operations
for 5 MW turbines have been performed from floating heavy lift vessels in summer in the North
Sea but have been limited to 2 ft seas and hence, almost impossible off the Northern California
coast. With the proposed WindFloat floating platform, integration of the mast, turbine, and plat-
form is performed at quayside, and on-site operations consist only of deploying mooring lines and
connecting to the platform. In the case of an unexpected failure of the wind turbine, the installation
sequence can be reversed and the platform towed back to a port for repairs.
The fabrication site should meet the following requirements.
• The structure should be designed to minimize welding at the assembly yard, by providing
large preassembled cylindrical sections of the columns, which can be efficiently fabricated in
a workshop using automatic welding machines.
• It should be in the vicinity of a waterway, deep enough to allow for the WindFloat to be
towed, at transit draft to the open ocean. The WindFloat is designed to be stable at its transit
draft. Temporary buoyancy may be attached to the column carrying the turbine to accommo-
date the depth of the channel.
• The mast, nacelle, and turbine should be installed at quayside. This implies the use of a large
crane.
• The means of loading out the hull from the integration site into the water should be consid-
ered early on when considering specific yards. Possible solutions are single lift from a heavy
lift crane, dry dock/graving dock, or submersible barges.

B. Installation: Transit
The transit phase studies should address the following points.
• The platform is towed after precommissioning to avoid the large cost and risk of placing the
tower and turbine onto a floater in open water.
• If a buoyancy module is needed to get out of the fabrication yard, then it should be removed
as soon as practical and the platform can be ballasted down to be even keel, with approxi-
mately 50 ft ͑15 m͒ draft.
• The transit route should be as short as possible, which means that the location of the fabri-
cation yard is project specific. This is important especially since an offshore wind farm will
be comprised of multiple WindFloat units and each hull has to be towed.
• Proper selection of the installation vessel is fundamental to project economics. The benefits
of using the same vessel with the ability to perform: ͑1͒ The mooring installation, ͑2͒ the
towing of the WindFloat platforms, and ͑3͒ the power cable installation could be significant.

C. Installation: Commissioning
It is important to minimize the offshore commissioning phase since offshore operations,
including mobilization of people and vessels offshore, are very expensive. The following points
are important to keep the cost down.
• The mooring system needs to be prelaid and ready to be connected.
• The anchor-handling vessel recovers the messenger lines from the platform and pulls in the
chain section of the mooring line. The connection to the wire section is done above the water.
• Tensioning of the mooring lines should be done from the platform with chain jacks. Space
limitations on the column supporting the tower and turbine should be considered carefully.
• Since the turbine will be already installed, the procedure involved to start up the turbine
should be simplified as much as possible.
• Installation and connection of the power cable are complex. The need to protect the subsea
cable for stability and to prevent damage should be assessed early on. Cable burying or
protective shells may be considered.


TABLE VII. Summary of stability characteristics.

Heeling angle in calm sea Down flooding angle Metacentric height
͑deg͒ ͑deg͒ ͑ft͒

Intact case at 0° wind heading 0 20.5 53
Intact case at 30° wind heading 0 22.5 53
Damaged case at 0° wind heading 4.5 18 38

VIII. TECHNICAL QUALIFICATION
Details on the methodology used to design the WindFloat, i.e., to predict its motion, size, and
structure, are discussed next. The work that has been performed to date includes the following:
• global sizing, including rules check and hydrostatics;
• stability;
• hydrodynamics, including model tests and hydroaerocoupling of the turbine and the hull;
• structural design, scantling, strength, fatigue of the trusses, and the mast.

IX. STABILITY
To assess the stability characteristics of the platform, the restoring moment is computed in
intact and damaged conditions at different wind headings. The downflooding angle—heeling angle
for which the vents above the top of columns are underwater—is also calculated and is shown in
Table VII.
The restoring moment curves obtained are compared to the curves of wind overturning mo-
ment to determine the heeling angle at equilibrium. Combined with a factor of safety, the com-
parison provides an estimation of the stability of the platform. A rough assessment of the wind
overturning moment under steady wind was carried out in this analysis, based on a range of thrust
coefficients for a 10 MW wind turbine. A worst case scenario ͑failure mode͒ is considered with a
combination of wind overturning moment and a faulty active ballast system. Wind headings every
30° are considered for this analysis.
Damage cases are also taken into account by assuming that a section of one column is flooded.
The damage remains limited due to compartmentation of the columns. In all considered configu-
rations, the angle of static equilibrium is smaller than the downflooding angle with a comfortable
safety margin and the platform remains stable in damaged conditions.

X. HYDRODYNAMIC MODEL
The time-domain software TIMEFLOAT was developed by the authors for coupled analysis of
floating structures. It uses WAMIT as a preprocessor to compute wave interaction effects and
computes the time-domain response of one or more floaters subjected to waves, wind, current, and
connected with moorings, tendons, hawsers, fenders, or any other mechanical connections. It takes
into account the viscous forces due to shedding around the hull and wave drift forces. The solution
is fully coupled, as the influence of vessel motion on tether forces is taken into account at each
time step, and conversely, the influence of tethers on vessel motion is also included at each time
step. A summary of the algorithm is presented next.
In the frequency domain, the equation of motion of a floater is

͓m + a͑␻͔͒x + b͑␻͒x + cx = F͑␻͒,
¨ ˙ ͑1͒
where a͑␻͒ and b͑␻͒ are frequency-dependent added mass and radiation damping coefficients, and
F͑␻͒ is the sum of forces applied to the floater including the wave-exciting force.
In the time domain, one can show that the equation of motion has the following general form:


FIG. 7. Wetted hull of the WindFloat for the WAMIT model.

͑m + aЈ͒x͑t͒ +
¨ ͵ −ϱ
t
K͑t − ␶͒x͑␶͒d␶ + cx͑t͒ = F͑t͒,
˙ ͑2͒

where aЈ is frequency independent and K is the retardation function,

͵
Ά ·
ϱ
1
aЈ = a͑␻͒ + K͑␶͒sin͑␻␶͒d␶
␻ 0
͑3͒
K͑␶͒ =
2
␲
͵
0
ϱ
b͑␻͒cos͑␻␶͒d␻ .

These integrals are calculated numerically.
TIMEFLOAT uses an explicit scheme to solve up to 12 degree of freedom ͑DOF͒ equations of
motion for a two-body system. The WindFloat is the only vessel considered in this analysis and
the software only solves 6-DOF equations. The general equation of motion is discretized in time
and the following linear vectorial equation is solved at each time step,

͓͑M͔ + ͓AЈ͔͒ak + ͓BЈ͔vk + ͓C͔xk = Fmem + Fdiff + Fvisc + Fdrift + Fmoor + Fwind . ͑4͒
The left-hand side of Newton’s equation of motion ͑4͒ contains terms proportional to the 6-DOF
acceleration ͑ak͒, velocity ͑vk͒, and motion of the floater ͑xk͒, with the following notations: ͓M͔ is
the mass matrix, ͓AЈ͔ is the 6 ϫ 6 infinite-frequency added-mass matrix, and ͓BЈ͔ is the 6 ϫ 6
matrix of retardation coefficients for t = 0, which are integrals of the frequency-dependent radiation
damping coefficients due to outgoing waves generated by the moving floater. The damping coef-
ficients are computed by WAMIT and integrated at the beginning of the time-domain simulation to
generate the retardation function matrix. ͓C͔ is the 6 ϫ 6 hydrostatic stiffness matrix computed by
WAMIT. Only the terms C͑3,3͒, C͑4,4͒, C͑5,5͒, C͑3,4͒, C͑3,5͒, and C͑4,5͒ are nonzero. Refer to
25
WAMIT manual for details. Figure 7 shows the hull geometry used in the WAMIT computations.
The right-hand side includes the various external forces. A brief description of the terms in
this equation is given below. Fmem represents the memory effect, i.e., the effects of wave compo-
nents generated by past motion of the floater, described by the convolution of the retardation
function with body velocity, as shown in Eq. ͑3͒ above.
Fdiff is the 6-DOF wave-exciting force determined by a Fourier series using the WAMIT


frequency-dependent wave-exciting force components and wave amplitude components represent-
ing the specified wave spectrum. A random phase and random frequency algorithms are used to
generate irregular wave trains efficiently and accurately.
Fvisc is the 6-DOF viscous force resulting from drag effects on the vessel columns and water-
entrapment plates. These are computed using a modified Morison equation model based on the
relative velocity of the wave/current kinematics and of special line members. Results of multiple
model test campaigns have been used to calibrate the empirical viscous force model. The effect of
ocean currents is captured with this viscous force model.
Fdrift is the 6-DOF drift force on the vessel computed based on the WAMIT mean drift
frequency-dependent coefficients obtained with the pressure integration or momentum approach
and the wave amplitude components. Newman’s approximation is used. Alternatively, a full
second-order diffraction model can be used if the WAMIT second order module is run. Previous
work has shown that the second-order potential solution was not required for the WindFloat.
Fmoor is the 6-DOF force on the vessel resulting from all mooring lines. Mooring lines are
modeled either with cable elements or nonlinear springs. For cable elements, a finite-difference
scheme is used to yield the dynamic mooring line configuration and mooring tensions at each time
step. The mooring dynamics and hydrodynamic loads are included using a Morison type formu-
lation. The nonlinear finite-difference equations are solved using a Newton–Raphson algorithm, as
described by Chatjigeorgiou and Mavrakos.26
Fwind is the 6-DOF wind turbine force on the vessel superstructure. The wind force model was
modified to capture some of the aerodynamic coupling between the turbine and the WindFloat
platform. It was assumed that the wind force applied on the rotor was proportional to the square
of the relative velocity between the wind and the hub. It was determined that an equivalent disk in
the rotor plane with 72.7 m diameter would provide the maximum rated thrust of a 10 MW
turbine, assuming a 1.2 drag coefficient on the disk. The wind force is perpendicular to the disk
and its direction varies in time with the platform rotations. The gyroscopic moment was estimated
from

Mgyro = I⍀ ϫ p, ͑5͒

where I is the moment of inertia of the spinning rotor, p is the rotational velocity vector of the
rotor around its axis, and ⍀ is the rotational velocity vector of the platform around the pitch and
yaw axes. The gyroscopic moment Mgyro is added to the moment contribution of Fwind.
Newton’s equation is applied in an inertial frame of reference which coincides with the vessel
frame of reference at t = 0. The origin of the vessel frame of reference is located at the mean water
level directly under the center of gravity. The X-axis points toward the bow, i.e., the column
supporting the wind turbine tower, the Y-axis toward port side, and the Z-axis upward.
TIMEFLOAT is written in FORTRAN. Information is provided to the software through an input
file in text format, with all vessel, mooring, and numerical parameters. Additional input consist of
the WAMIT files and the wind and current coefficients files. After reading the input, TIMEFLOAT
solves an initial static phase, in which mean wind and current loads are applied as well as the
mooring line pretension. This phase serves to reduce the transient phases and quickly provides
static information if needed. Then, the solution is advanced in time using a Runge–Kutta algorithm
for the 6-DOF rigid-body motion and velocities. At each of the four fractional steps used in this
process, external forces are updated.
WAMIT6.3 software was used to compute added-mass and damping coefficients as well as
wave-exciting forces and mean drift coefficients. Only the underwater part of the hull is modeled.
The model includes the columns, water-entrapment plates, and main tubulars connecting columns.
The bracings are only modeled as line members using the Morison equation. Dipole elements are
used to discretize the water-entrapment plate since they are thin structural elements.


FIG. 8. Picture of the WindFloat model.

XI. DESIGN CASES
In the preliminary design phase, a selected number of design cases were deﬁned based on a
combination of offshore mooring design codes and offshore wind turbine design codes, i.e.,
28
API-RP2SK ͑Ref. 27͒ and Germanischer Lloyd. The design cases that were thought to be the most
onerous for the platform motions were checked. These included the extreme coherent gust ͑ECG͒
and the 100 year storm ͑13.5 m Hs͒ shown in Tables IV and V. In addition, a number of operating
cases were run corresponding to the turbine maximum thrust wind speed ͑ϳ12 m / s͒ with asso-
ciated waves ͑ϳ2 m Hs͒, and the maximum wind speed with turbine spinning ͑ϳ25 m / s͒ with
associated waves ͑ϳ4 m Hs͒.
For detail design and certiﬁcation, a much larger number of design cases will have to be
considered; however, the return period of the maximum events will likely be 50 years in accor-
dance with wind turbine design codes, rather than the 100 year return period selected for this
preliminary study. Space does not allow for an extensive presentation of the hydrodynamic simu-
lations; however, some results of numerical predictions are provided later and compared to model
test results for key parameters.

XII. MODEL TESTS SETUP
A model test campaign was conducted at the UC Berkeley 200 ft long ͑61 m͒ ship-model
testing facility to test the validity of the numerical analysis tools. A 1/105 scale model of the
platform was fabricated out of the acrylic. Lead weights were placed inside the columns and on
the water-entrapment plates to adjust the center of gravity to its target position; item ͑1͒ in Fig. 8.


The platform motion was measured using a digital video camera tracking the motion of light
emitting diodes placed on model ͑2͒. The system provides 3-DOF measurements of the motion in
the plane of the camera.
Tower ͑3͒ was made of a thin ͑not-to-scale͒ 1 in. outside diameter acrylic pipe because the
device used to model the wind turbine was relatively heavy and it was not possible to obtain the
correct center of gravity with the lead weights if the tower was modeled with a 3 in. diameter
acrylic pipe, as originally planned. Stays made of thin string were connected to the tower to
increase its stiffness.
The turbine model device was connected to the top of the tower onto a load cell ͑4͒, which
measured the axial force perpendicular to the tower. A large disk ͑5͒ made of foam board was
placed on the model to attract wind loads corresponding to the design wind force. No attempts
were made to match the atmospheric turbulence. The wind maker naturally produces turbulence
and the turbulent wind fluctuations are somewhat averaged by the large disk. In the end, the wind
force was measured and the turbulence level will be compared to variations in the aerodynamic
forces generated by a prototype wind turbine. The disk diameter is a third of the total area covered
by the rotor. The drag coefficient on the disk is estimated to be 1.2.
An electrical motor ͑6͒ was placed at the top of the tower to model the gyroscopic effect. This
well-known mechanical force arises when a rotor spinning around a certain axis undergoes a
rotation around a different axis. For instance, platform pitch and yaw would lead to gyroscopic
forces applied on the tower. These forces are a significant design issue for the blades and the
shaft/bearings, but they may also have a contribution to the global response of the floater. The
motor was adjusted to spin at the Froude-scaled turbine speed of 2 Hz ͑approximately 12 rpm in
prototype scale͒, and the inertia of the blades was approximately modeled with two weights ͑7͒
positioned on an aluminum rod ͑8͒.
The model was kept in position in the tank using four soft springs—two of them connected to
column 1 which holds the turbine and one on each of the other columns. The mooring lines were
connected at the edges of a 7 ϫ 7 ft2 square frame placed on the tank floor. This provided a top
angle for the mooring lines of approximately 45°. This equivalent mooring model provided hori-
zontal stiffness similar to that of the prototype six line catenary mooring system, yielding a 65 s
resonant period in surge. However, the prototype mooring design has not been finalized and the
focus of these tests was placed on platform motion. No attempts were made to measure mooring
tension or validate mooring dynamics.
A plunger type wave maker is located at one end of the tank and a parabolic wave absorption
beach at the other end. A set of five large wind fans was assembled to generate the required wind
loading on the turbine model, as shown in Fig. 9. The effect of the active ballast system was
modeled by shifting lead ballast on the model to compensate for the mean wind overturning
moment.
A 3 h long realization of the 100 year waves was generated. The associated wind is 25 m/s,
which is the maximum wind speed at which the wind turbine is allowed to rotate. Such wave
events may occur at the site with wind speed under the cutoff speed due to swells. Most likely, the
rotor would be parked if such wave conditions arise; however, this conservative design case was
generated to establish upper bounds of platform motion. The 100 year wave run was repeated
without wind. Additionally, regular waves were run with and without wind to determine response
amplitude operators ͑RAOs͒.

XIII. RESULTS
Results of the 100 year storm simulation are summarized in Table VIII. Time series of
platform surge, heave, and pitch were processed to yield rms, maximum, and minimum values.
These show a satisfactory agreement between the model test results and numerical simulations
performed with TIMEFLOAT. The pitch rms is slightly underpredicted by the software ͑1.15° versus
1.27° measured͒, and the minimum and maximum pitch angles are off by 1° due to some differ-
ences in the predicted versus measured wind overturning moment; the platform response is,
however, deemed extremely well behaved, with maximum pitch angle of 5° in a 13.5 m Hs sea


FIG. 9. WindFloat model in the 100 year storm.

state. The maximum crest to trough pitch is 7° with a 21.3 m maximum wave height ͑crest to
trough͒. Similar responses and trends were observed for all tested platform headings ͑0° and 90°͒
and for runs with and without wind. The maximum yaw angle measured in the 90° runs was under
10°.
RAOs were computed for wave periods between 6 and 18 s. Figure 10 shows the RAOs in
surge, heave, and pitch for 0° wave heading. The presence of wind does not affect surge or sway
significantly, but its effects are slightly more pronounced on the pitch RAOs. Although wind speed
is constant in all the regular wave runs, it does impact the regular wave response because the
wave-induced motions generate a sinusoidal variation in the relative speed between the wind and
the disk, which results in an additional periodic force component on the disk leading to a corre-
sponding periodic pitch moment.
Regular wave tests were repeated with 90° wave heading to investigate the platform yaw
response; i.e., the model orientation was changed by rotating the anchoring frame to 90°. There is
no wave-induced yaw for 0° heading since the platform is port/starboard symmetric; the yaw RAO
at 90° is shown in Fig. 11. Additional tests were carried out by adding two large triangular vertical
plates on each column ͑named yaw plates͒ with the bottom edge extending outward to the edge of
the heave plate and the side extending from the heave plate to 20 ft below the mean water level in
prototype scale. The effects of “yaw plates” in reducing first-order yaw were minimal. The irregu-

TABLE VIII. Numerical and model test results in the 100 year storm with 0° wave heading and 25 m/s steady wind.

Wind surge 85 ft/s heave Steady pitch
Heading 0 ͑ft͒ ͑ft͒ ͑ft͒

rms Model tests 10.56 6.88 1.27
Time float 9.18 6.40 1.15
Maximum Model tests 48.46 18.97 4.87
Time float 43.51 16.18 5.77
Minimum Model tests Ϫ22.16 Ϫ22.05 Ϫ3.87
Time float Ϫ17.28 Ϫ22.61 Ϫ2.67


FIG. 10. RAO in surge, heave, and pitch at 0° with and without wind.

lar wave test showed that the second-order yaw was also not significantly reduced. Overall, the
experiment did not point to serious limitations of the numerical modeling ability.

XIV. COUPLED AEROHYDRODYNAMIC MODEL
The forces generated by the wind turbine are reasonably well computed by the modified
TIMEFLOAT software and are correspondingly well modeled experimentally for a steady wind
speed. However in reality, the wind speed is constantly changing due to naturally occurring
turbulence in the atmosphere. Large wind turbines are equipped with sophisticated control systems
generally designed to keep the rotor speed constant at all times using a variable torque generator
and a blade pitching mechanism ͑changing the angle of attack of the blades by rotating them
around their local axis͒. This technique, known as “blade pitching,” can have significant effects on
floating platforms, as observed by Nielsen et al.9 and by Jonkman.29 The control system may

FIG. 11. RAO in yaw at 90° without wind.


induce negative damping, which results in resonant oscillations of the platform at its pitch natural
period.
In order to assess the effects of blade pitching on the floater, as well as to provide accurate
computation of all loads induced by the wind turbine on a moving foundation, a software dedi-
cated to wind turbine design, FAST, was interfaced with TIMEFLOAT to provide a fully coupled
aeroservoelastic/hydrodynamic time-domain numerical model of the WindFloat platform with a 5
MW wind turbine.
FAST, which stands for “fatigue, aerodynamics, structures, and turbulence” is an aeroser-
voelastic modal code for horizontal axis wind turbines developed by the National Renewable
Energy Laboratory ͑NREL͒. FAST models the wind turbine as a combination of rigid and flexible
bodies. The rigid bodies are the earth, nacelle, hub, and optional tip brakes. The flexible bodies
include blades, tower, and drive shaft. The model connects these bodies with several DOFs,
including tower bending, blade bending, nacelle yaw, rotor teeter, rotor speed, and drive shaft
torsional flexibility. FAST uses Kane’s method to set up equations of motion, which are solved by
numerical integration. The AERODYN subroutine package developed by Windward Engineering is
used to generate aerodynamic forces along the blades.
The FAST and TIMEFLOAT FORTRAN source codes were modified to change TIMEFLOAT into a
subroutine called by FAST. Hydrodynamic forces, including wave-exciting forces, viscous forces,
and mooring forces are computed by TIMEFLOAT and passed to FAST, which solves the coupled
turbine tower problem and passes platform motion back to TIMEFLOAT.
The FAST model of a utility-scale multimegawatt turbine known as the “NREL offshore 5 MW
baseline wind turbine” was developed by Jonkman et al.21 using publicly available information
from turbine manufacturers. This wind turbine is a conventional three-bladed upwind variable-
speed variable blade-pitch-to-feather-controlled turbine. A conventional control system was used
with a generator-torque controller whose goal is to maximize power capture below the rated
operation point and a blade-pitch controller designed to regulate rotor speed above the rated
operation point.
The coupled FAST-TIMEFLOAT model was run using the validated WindFloat hydrodynamic
model described in Sec. X. Sample results are provided for a 4 m significant sea state with 12 s
peak period and a 12 m/s steady wind. Waves and wind are at 0° heading, along the symmetry axis
of the WindFloat. A Jonswap wave spectrum is assumed with peakness factor ␥ = 2.4. No atmo-
spheric turbulence is assumed in this simulation.
Figure 12 shows sample time series of the platform roll, pitch, and yaw over a 5 min duration
after the initial transients generated at the beginning of the numerical simulation have disappeared.
A slight asymmetry is present due to the rotation of the rotor in one direction, generating a small
mean roll ͑ϳ1°͒ and yaw ͑ϳ2°͒ component. A background platform pitch oscillation of approxi-
mately Ϯ2° is caused by the blade-pitch controller, which excites the platform at its pitch resonant
period around 30 s. This was later tuned out by modifying the controller coefficients and adding
an additional filter. Superposed to the resonant pitch cycles are wave-induced pitch oscillations,
which result in slight changes between resonant cycles, but are overall a small contribution to the
platform pitch in this sea state.
In Fig. 13, time series of the base of the tower are shown. Wave-induced surge is clearly
visible in this 4 m irregular wave sea state. Mean surge is primarily driven by mean aerodynamic
loads on the turbine. The platform pitch oscillation results in vertical movement of the tower base
at the same period as the pitch cycles.
Figure 14 presents the blade-pitch angle time series ͑at the bottom͒ and power out-take ͑at the
top͒. The blade-pitch controller locks into the platform pitch resonance with 30 s cycling of the
blades. A drop in produced power occurs for approximately 2 s at each cycle when the relative
speed between the nacelle and incoming wind drops below the threshold for maximum power
output. This does not have a large impact on mean produced power, which is 4.95 MW on average,


FIG. 12. WindFloat rotations in 4 m seas with 12 m/s wind.

but would require ﬁltering. Further investigations of the control system have been performed
following the recommendations of Jonkman29 and Nielsen9 to eliminate this resonant response in
order to maximize power production and minimize fatigue loading of all components and systems.
Results will be published shortly.

FIG. 13. Tower base motion in 4 m seas with 12 m/s wind.


FIG. 14. Power outtake and blade pitch in 4 m seas with 12 m/s wind.

XV. DESIGN STANDARDS AND ENVIRONMENTAL CONDITIONS
The WindFloat is a novel offshore structure, which combines a wind turbine and a floater. No
formal design code has been developed yet for the design of structural reinforcement and scant-
ling. Existing standards for offshore wind turbines were developed in the past decade from knowl-
edge of onshore wind turbines and growing experience in near-shore operations of wind energy
devices. However, their scope remains limited to wind turbines in shallow waters with fixed
foundations.
The WindFloat is a moored platform with a complex dynamic behavior, which cannot be
overlooked in the structural design of critical elements, such as the tower. Although offshore wind
energy codes, such as the Germanischer Lloyd Guidelines for the Certification of Offshore Wind
Turbines,28 provide critical information about the extent of wind loading on the structure, the
design criteria may not be sufficiently conservative for a floater.
To ensure a high reliability of the design, the structural analysis of the WindFloat is largely
based on standards from the oil and gas industry, including the ABS rules for Mobile Offshore
Drilling Units30 and the API Recommended Practice for Fixed Offshore Platforms.27 The DNV
Recommended Practice C202 ͑Ref. 31͒ is used to assess shell buckling of the tower. These design
criteria need to be combined with a realistic model of the wind loading effects and conservative
estimation of environmental loadings on the hull.
The environmental loadings in both cases are obtained for sea states in the wave scatter
diagram encountered at the intended location of the WindFloat, off the coast of Northern Califor-
nia. For each peak period ͑Tp͒ in the wave scatter diagram, the sea state with highest significant
wave height ͑Hs͒ is identified. The 12 resulting sea states with characteristics listed in Table IX
represent the steepest wave conditions for each peak period. The strength analysis may be based
on these sea states. All peak periods are included in the strength analysis since wave loading
depends on wavelength. The largest wave height does not necessarily result in largest loading on
the platform. The fatigue analysis requires the generation of extensive numerical data. The fatigue
damage must be calculated for all sea states in the wave scatter diagram, based on a time series of
nominal stress. To avoid the production of large amounts of data and to save CPU time, the stress
range is computed only for those 12 identified sea states. For a given peak period, the level of


TABLE IX. Sea states for structural strength analysis.

Tp Hs Hs
Case No. ͑s͒ ͑m͒ ͑ft͒

01 20 12.5 41.0
02 16.7 11.5 37.7
03 14.3 9.5 31.2
04 12.5 8.5 27.9
05 11.1 7.5 24.6
06 10 7.5 24.6
07 9.1 7.5 24.6
08 8.3 6.5 21.3
09 7.1 5.5 18.0
10 6.3 4.5 14.8
11 5.3 3.5 11.5
12 4.2 1.5 4.9

stress is assumed to be linear with significant wave height. Thus, the level of stress is scaled with
significant wave height to complete the wave scatter diagram and determine the fatigue life of all
structural elements.
For the truss and the tower of WindFloat, strength and fatigue analyses are carried out. The
computation of local forces and moments is achieved with finite-element software SAP by Com-
puter & Structures, Inc., Berkeley, CA, using beam theory. The structural calculations are linear. A
static analysis is sufficient on the truss since the natural period of its elements are too low to be
excited by environmental loading. However, a dynamic analysis is necessary to account for the
excitation of the natural period of the tower. The applied loads are obtained from TIMEFLOAT time
series for each sea state. External forces and moments are applied at the extremities of the tubular
elements in the finite-element model or as distributed loads. For the dynamic analysis of the tower,
the acceleration load calculated in TIMEFLOAT is directly applied at the base of the tower.
The purpose of this study is to identify the weakest points on the elements and to run a
preliminary structural analysis to ensure the reliability of the elements. For the strength analysis,
the most extreme stresses are used to compute recommended strength ratios. When necessary, the
thickness of the tubular elements was adjusted to meet the appropriate safety factors in strength.
On tubular elements, fatigue assessment is especially critical at the joints. A hot-spot stress ap-
proach as recommended in API is used to estimate the fatigue at the joints between bracing
elements. This method entails the calculation of stress concentration factors ͑SCFs͒ at the joints.
The fatigue life is computed based on the nominal stress as provided by a beam-column finite-
element model multiplied by the SCF. The damage and fatigue life are computed with a formu-
lation from DNV Recommended Practice RP-C203 for a short term Rayleigh distribution of stress
levels.
The annual damage for all sea states and in three directions is combined with Miner’s rule,

D=
Td
A
ͩ ͚ͪ
ϫ⌫ 1+
m
2 seastates
pi␯i͑2ͱ2␴͒m , ͑6͒

where ␴ is the range of the nominal stress, pi the probability of occurrence of a sea state in any
given year, and ␯i is the frequency of cycles, which may be taken to the zero-up crossing fre-
quency. Recall that Td is the design life and A and m are parameters of the API X S-N curve.

XVI. STRENGTH AND FATIGUE DESIGN OF THE TRUSS
The primary function of the truss is to provide the WindFloat hull with sufficient global
structural stiffness to withstand environmental loads. A three-dimensional ͑3D͒ model of the
WindFloat is created ͑Fig. 15͒. The columns and bracing are modeled with tubular grade 50 steel


FIG. 15. Truss finite-element model.

beam-column elements. Main horizontal bracing members are 150 ft ͑45.7 m͒ long cylinders that
support the horizontal loads between columns. Light bracing members provide reinforcement at
1/3 of their length. These bracing members are diagonal between the main bracings and columns
for vertical stiffness and horizontal between main bracing elements to provide horizontal stiffness.
The joints between the column and the bracing are modeled with an element of stiffness, ten times
that of the bracing element, and consistent with API recommendations. The water-entrapment
plates are not included in this model but the applied forces on the plates are calculated externally
and transferred to the base of the columns.
External and inertia forces applied to each structural member are computed using dedicated
software, based on the TIMEFLOAT program, which computes hydrodynamic loads by integration of
the diffraction and radiation pressures on each part of the structure. The software also matches the
hydrodynamic panels with corresponding structural elements. The time-domain force components
passed to the finite-element model include weight of all elements, radiation, and diffraction pres-
sures, as well as mass inertia and hydrostatic stiffness effects. Wave exciting forces, including
Froude–Krylov effects, are passed via the diffraction pressure. The viscous forces, reflecting


viscous loads on heave plates, columns, and truss members, are applied to the corresponding parts
of the structure. The mooring forces are applied vertically to the chain stoppers at the top of
column since they set the column in compression. The horizontal component of the mooring is
applied to the fairlead at the keel, with a 45° top angle. The wind-induced forces ͑thrust and
torque͒ are applied horizontally at the top of the tower. Drift forces are neglected since they are
relatively small on individual elements. It is verified that the sum of external forces and inertia
forces on all parts of the structure is approximately null.
The truss consists of unstiffened tubular elements. For the analysis of tubular members, API
RP2A-WSD defines allowable axial, bending shear, and hoop stresses. Maximum predicted
stresses on the elements in design environmental conditions are computed with finite-element
analysis. The overall structural reliability of a member is estimated by combination of the maxi-
mum to allowable stress ratios with appropriate safety factors. All computed ratios must be less
than 1 to comply with API.
The stress on the truss is determined using a static finite-element algorithm on the model
subject to all environmental loads including rigid-body dynamics contributions. To capture the
highest stress level, the forces are calculated for a 1 min snapshot of the most extreme wave of a
1 h simulation on all relevant sea states for three headings.
The maximum API stress ratios increase with larger sea states. Thus, sea state 1, with the
largest significant wave height, is associated with the maximum stress ratio at 90°, heading for
most frame elements. Figure 16 represents the maximum API ratios calculated in the worst case,
at 90° heading for sea state 1, plotted directly on the structure. The shell thickness and diameter of
the truss elements were adjusted to ensure compliance with API criteria.
It was determined through further analysis that the wind loads were driving the design of the
truss in strength analysis. Figure 17 illustrates the effect of wave and wind loading on the shape of
the truss: The main horizontal bracing elements undergo significant bending.
Next, the fatigue analysis is performed on the truss. The target design life of the WindFloat is
20 years. In this design cycle, a safety factor of 10 is applied and a calculated fatigue life of 200
years is required. This is very conservative and can be reduced as the engineering is refined.
The fatigue analysis is critical at the joints between bracing elements and the fatigue life of the
connection is determined based on the stress ranges calculated by beam theory. To apply the
hot-spot stress curve, the SCF needs to be determined. For a nominal stress away from the welding
toe on a beam model of the tubular element, it is reasonable to expect the SCF to be between four
and six for a well-designed connection. The Von Mises stress at the connection obtained from
beam-column finite-element modeling is used as nominal stress in this case. A sensitivity analysis
is carried out on the value of the SCF. The exact stress ratio between the maximum stress at the
weld toe and Von Mises stress in beam theory will be determined precisely by finite-element
analysis with a 3D model of the connection in follow on studies. It should also be noted that weld
profile control is assumed at the joints of truss elements so that the API X-curve may be used to
define the relationship between hot-spot stress range and number of cycles to failure.
The maximum levels of Von Mises stress in the truss are observed for peak periods between
6 and 10 s depending on the heading. This is consistent with wave loads on the columns when the
wavelength is half the distance between columns.
The stress ranges are determined for all sea states in the scatter diagram and combined to
obtain the fatigue life. Results are summarized in Table X. Assuming that the stress at the weld is
accurately computed by the beam model and that no increase in wall thickness is implemented at
the connection, the minimum fatigue life of the nodes is 670 years. This optimistic assumption
will be verified with a detailed finite element of the connection. It is likely that increase in wall
thickness over a short section near the node will be required to achieve fatigue life targets.
Estimates of fatigue life based on a can with wall thickness equal to twice the nominal wall
thickness of the tubular members are provided in Table XI.


FIG. 16. Maximum API design ratios on WindFloat platform in 90° heading sea state 1.

XVII. STRUCTURAL ANALYSIS OF THE TOWER
The design of the tower must take into account wind and wave-induced motions. A dynamic
analysis of the tower is required since the first lateral mode of resonance is near 3 s. At such
periods, some wave energy may be transmitted to the tower through the platform rigid-body
motions.
The tower is a tapered unstiffened 220 ft ͑67 m͒ high tube, with increasing wall thickness
from top to bottom. It supports a 300 ton nacelle and rotor at the top. It is connected to the column
at the bottom with a bolted or welded flange joint. The buckling strength of the tubular element is
determined for extreme environmental conditions and the fatigue life of the joint at the base of the
column is calculated.
The numerical model is composed of a number of beam elements with decreasing diameter
and thickness from bottom to top. Beam elements are sufficient for this study since there is no
external pressure distribution on the tower. A convergence analysis is carried out to determine the
minimum number of elements necessary to correctly represent the dynamic characteristics of the
tower. With eight elements, the mass and stiffness of the structure have converged.
1 h time series of accelerations at the base of the tower are generated for all twelve relevant
sea states in the wave scatter diagram. Additionally, the largest wind force ͑the maximum of the


FIG. 17. Deformed ͑50ϫ͒ shape of WindFloat structure in worst loading conditions ͑sea state 1 at 90° heading͒.

thrust versus wind speed curve͒ is applied horizontally at the top and the tower supports its own
weight as well as the weight of the turbine. The deﬂections of the tower are computed using linear
beam theory with a time-domain ﬁnite-element algorithm.
In Fig. 18, the bending moment ͑top͒ and the sway motion at the base of the tower ͑bottom͒
are plotted during the largest wave event of the 1 h time series for sea state 1, which corresponds
to the 100 year storm. The maximum horizontal excursion at the base of the turbine tower is 60 ft
͑18 m͒ crest to trough during a single wave cycle, corresponding to a 70 ft ͑21 m͒ wave crest to

TABLE X. Fatigue life on connection between bracings based on nominal wall thickness.

Damage Fatigue life
SCF ͑per year͒ ͑year͒

1 1.7ϫ 10−03 670
1.5 8.2ϫ 10−03 121
2 2.8ϫ 10−02 36


TABLE XI. Fatigue life with double wall thickness at the connection.

Damage Fatiguelife
SCF ͑per year͒ ͑year͒

1 6.7ϫ 10−05 14 934
1.5 3.8ϫ 10−04 2623
2 1.5ϫ 10−03 670

trough. The bending moment time series clearly shows the dynamic response of the tower, which
includes oscillations with a period below 3 s superimposed to the wave-induced component with
a period around 20 s.
A 2% ratio of critical damping is applied to the numerical model. This is the level of damping
expected on the tower when the turbine is parked. In most scenarios, when the turbine rotates, the
damping ratio increases on the tower due to aerodynamic drag. A sensitivity analysis is performed
to evaluate the effect of damping on tower fatigue.
In Fig. 19, the bending stress at the base of the tower is plotted for 2% and 5% critical
damping ratios, highlighting the variations in the dynamic response of the tower. Yet, the energy
at the natural period of the tower is small compared to wave-induced variations in bending stress.
The structural damping does not affect the fatigue results signiﬁcantly: The rms of bending
moment varies by only 1% when damping is increased from 2% to 5% of critical damping in this
high sea state.
The natural period of the tower is low enough to not interfere with wave-induced motion of
the platform. The unsupported section of the WindFloat tower is much shorter than onshore towers
because the hub is slightly lower than onshore, and the platform truss provides lateral stiffness to
the tower up to 33 ft ͑10 m͒ elevation above the mean water line.
Bending moment at the base of the tower is also plotted in Fig. 20 for sea state 12 ͑Tp
= 4.2 s͒. The bottom of the ﬁgure shows a time series of the sway motion, which is a combination
of linear wave dynamics with period of 4.2 s and slow-drift cycles with period of approximately

FIG. 18. Bending stress ͑kips/ft͒ and sway ͑ft͒ at the base of the tower at the largest wave of sea state 1.


FIG. 19. Sensitivity of bending stress with damping ratio in sea state 1 at 90° heading.

50 s. Only the energy from ﬁrst-order wave dynamics at low periods is transmitted to the tower.
Excitation of the tower natural periods is not apparent due to the small magnitude of tower base
motion.
For the strength analysis, the design recommendations from DNV-RP C202 are used. The
shell buckling assessment is based on formulas for unstiffened tubular elements. The column

FIG. 20. Bending stress ͑kips/ft͒ and sway ͑ft͒ at the base of the tower at the largest wave of sea state 12.


FIG. 21. ͑Left͒ Axial force in compression. ͑right͒ Bending moment at largest event of sea state 1 at 90° heading.

buckling does not need to be computed since ͑kL / i͒2 Ͻ 2.5 E / fY, where kL is the effective length,
i is the radius of gyration of the cross section, E is Young’s modulus, and fY is the yield strength
of steel.
The largest events are identified over a 1 h time series. The shell buckling ratio is calculated
at the lower end of each of the eight elements using the local wall thickness and diameter for this
element. Stress is largest at these lowest ends since the axial force and bending moment increase
toward the base of the tower, as illustrated in one time step in Fig. 21.
It may be noted that even for extreme events of the largest sea states, wind force on the turbine
contributes up to 70% of the axial stress on the tower. The wind force is critical to the design of
the tower in strength.
Shell buckling ratios are computed for these extreme events according to DNV recommenda-
tions. At the base of the tower, the largest design equivalent to the Von Mises stress to design shell
buckling strength ratio is 0.4, which is 40% of the maximum allowed. Thus, the tower will not be
affected by buckling from dynamic wave loads and wind thrust.
The fatigue analysis is assessed at the joint between floater and turbine at the base of the
tower. The column and the tower meet in a flange connection, which is bolted or welded. The
standard deviation of the Von Mises stress is determined over a 1 h simulation of the structural
response to the 12 relevant sea states. The bending moment is computed at a point at the base of
the tower for a number of wave directions between 0° and 180°, to account for the directionality
of waves at the Northern California location. Each heading is given identical probability of
occurrence for this analysis.
The hot-spot stress S-N curve with a Rayleigh approximation is used to determine the damage
per year on the connection. The SCF should be computed from a 3D finite-element analysis of the
connection. However, this work will be performed in a later phase of the project once structural
details of the connection are established. In a preliminary analysis, a sensitivity study is carried out
on the SCF at the base of the tower.
Results are summarized in Table XII. The calculated fatigue life is 37 280 years based on
nominal wave-induced stress. Damping level is conservatively assumed to be 2% of critical for all
sea states, although it will likely be higher when the turbine is spinning. The design of the
connection between the tower base and top of column will have to be carefully designed to reduce
SCFs to acceptable levels based on fatigue life targets. Fatigue due to cycling of the wind loads
and tower vibrations due to the spinning rotor have not been included in this model. Detailed
aerodynamic calculations will be performed to account for these additional fatigue sources.


TABLE XII. Summary of sensitivity of fatigue damage on SCF.

Damping Damage Fatigue life
͑%͒ SCF ͑per year͒ ͑year͒

2 1 2.68ϫ 10−05 37 280
2 2 5.59ϫ 10−04 1790
2 4 1.16ϫ 10−02 86
2 6 6.87ϫ 10−02 15

XVIII. CONCLUSION

The work presented herein was aimed at providing sufficient technical information about the
system to highlight challenging areas for any offshore floating wind turbine foundations. The most
prominent areas are as follows.
• The turbines in their “as-is” configurations may not be able to withstand some of the floater
induced motions. It is therefore critical to involve the turbine manufacturers, to verify that the
new motion envelopes are within their design criteria. It is further important to minimize the
floater motions, most critically the pitch motion, to eliminate any potential for the blade
interference with the mast due to the gyroscopic force, which maintains the blades in their
turning plane.
• Fabrication and installation: The foundation should be fabricated and integrated near the
installation site. However, the infrastructure required for the construction of such a large
system may not exist near some of the potential wind farm areas and might have significant
cost implication on the project.
• Steel cost has been rising significantly recently, but so has the welding and fabrication costs.
Optimizing the structure for steel weight may not yield the most inexpensive hull. Under-
standing the fabricator constraints during the design phase is very important to reduce fab-
rication complexities and associated cost run-ups.

This paper also discusses the hydrodynamic analysis of the WindFloat. Numerical analysis
was first carried out with simplified models of the wind turbine forces. This work was done with
a fully coupled time-domain algorithm, which accounts for diffraction-radiation effects, as well as
viscous forces and the influence of the mooring. Model tests were performed to validate the
predictive ability of the numerical hydrodynamic algorithm. This experimental work consisted of
generating wave loads in a wave tank facility, as well as wind loads using fans and a drag disk
placed on the model, and a rotor to model gyroscopic effects.
A coupled aeroelastic-hydrodynamic model was then implemented to provide better resolution
of wind turbine loads and take into account the effects of the turbine control system. For this work,
the validated hydrodynamic model discussed above was interfaced with FAST software developed
by NREL for design of wind turbines. It was shown that interactions between the wind turbine
control system and the platform generate small rotational oscillations with long periods ͑
ϳ30 s͒, which, in some cases, could result in slightly reduced power output. Further work will be
carried out to improve the turbine control system, and assess the effects of coupled aeroelastic-
hydrodynamic loads on the WindFloat components.
Lastly, this paper discusses the preliminary structural assessment of the WindFloat. It focuses
on the methodology designed to estimate the strength and fatigue of WindFloat’s novel structural
components. It is assumed that structural loading on the underwater elements of the platform, such
as the columns and the water-entrapment plates, is mostly dependent on wave loading. Their
preliminary design can be conservatively established using design guidelines developed for the
offshore industry. Novel elements, such as the truss or the interface between the wind turbine and
the columns, i.e., the tower, must be analyzed thoroughly due to the importance of aerodynamic

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