Conference Paper

NREL/CP-500-47534

April 2010

Offshore Code Comparison 

Collaboration within IEA Wind Task 23: 

Phase IV Results Regarding Floating 

Wind Turbine Modeling 

Preprint

J. Jonkman

National Renewable Energy Laboratory

T. Larsen and A. Hansen

Riso National Laboratory, Denmark

T. Nygaard 

Institute for Energy Technology, Norway

K. Maus

Norwegian University of Life Sciences (UMB), Norway

M. Karimirad, Z. Gao, and T. Moan

Norwegian University of Science and Technology (NTNU), Norway

I. Fylling

Marintek, Norway

J. Nichols

Garrad Hassan & Partners Limited, United Kingdom

M. Kohlmeier

ForWind, Leibniz University of Hannover, Germany

J. Pascual Vergara and D. Merino

Acciona Energia, Spain

W. Shi and H. Park

Pohand University of Science and Technology, Korea

To be presented at European Wind Energy Conference (EWEC)

Warsaw, Poland

April 20-23, 2010
NOTICE

The submitted manuscript has been offered by an employee of the Alliance for Sustainable Energy, LLC 

(ASE), a contractor of the US Government under Contract No. DE-AC36-08-GO28308. Accordingly, the US 

Government and ASE retain a nonexclusive royalty-free license to publish or reproduce the published form of 

this contribution, or allow others to do so, for US Government purposes.

This report was prepared as an account of work sponsored by an agency of the United States government. 

Neither the United States government nor any agency thereof, nor any of their employees, makes any 

warranty, express or implied, or assumes any legal liability or responsibility for the accuracy, completeness, or 

usefulness of any information, apparatus, product, or process disclosed, or represents that its use would not 

infringe privately owned rights. Reference herein to any specific commercial product, process, or service by 

trade name, trademark, manufacturer, or otherwise does not necessarily constitute or imply its endorsement, 

recommendation, or favoring by the United States government or any agency thereof. The views and 

opinions of authors expressed herein do not necessarily state or reflect those of the United States 

government or any agency thereof.

Available electronically at http://www.osti.gov/bridge

Available for a processing fee to U.S. Department of Energy

and its contractors, in paper, from:

U.S. Department of Energy

Office of Scientific and Technical Information

P.O. Box 62

Oak Ridge, TN 37831-0062

phone: 865.576.8401

fax: 865.576.5728

email: mailto:reports@adonis.osti.gov

Available for sale to the public, in paper, from:

U.S. Department of Commerce

National Technical Information Service

5285 Port Royal Road

Springfield, VA 22161

phone: 800.553.6847

fax: 703.605.6900

email: orders@ntis.fedworld.gov

online ordering: http://www.ntis.gov/ordering.htm

Printed on paper containing at least 50% wastepaper, including 20% postconsumer waste
1

Abstract

Offshore wind turbines are designed and 

analyzed using comprehensive simulation 

codes that account for the coupled 

dynamics of the wind inflow, 

aerodynamics, elasticity, and controls of 

the turbine, along with the incident waves, 

sea current, hydrodynamics, and 

foundation dynamics of the support 

structure. This paper describes the latest 

findings of the code-to-code verification 

activities of the Offshore Code 

Comparison Collaboration, which operates 

under Subtask 2 of the International 

Energy Agency Wind Task 23. In the 

latest phase of the project, participants 

used an assortment of codes to model the 

coupled dynamic response of a 5-MW 

wind turbine installed on a floating spar 

buoy in 320 m of water. Code predictions 

were compared from load-case simulations

selected to test different model features. 

The comparisons have resulted in a greater

understanding of offshore floating wind 

turbine dynamics and modeling 

techniques, and better knowledge of the 

validity of various approximations. The 

lessons learned from this exercise have 

improved the participants’ codes, thus

improving the standard of offshore wind 

turbine modeling.

Keywords: offshore wind turbine, 

floating, spar buoy, aero-hydro-servoelastic analysis, code verification

Introduction

The vast offshore wind resource represents 

a potential to use wind turbines installed 

offshore to power much of the world. 

Design standardization is difficult, 

however, because offshore sites vary 

significantly through differences in water 

depth, soil type, and wind and wave 

severity. To ensure that offshore wind 

turbine installations are cost effective, the 

use of a variety of support structure types

is required. These types include fixedbottom monopiles, gravity bases, and 

space-frames—such as tripods and lattice 

frames (“jackets”)—and floating 

structures. In this context, the offshore 

wind industry faces many new design 

challenges.

Wind turbines are designed and analyzed 

using simulation tools (i.e., design codes)

capable of predicting the coupled dynamic 

loads and responses of the system. Landbased wind turbine analysis relies on the 

use of aero-servo-elastic codes, which

incorporate wind-inflow, aerodynamic

(aero), control system (servo), and 

structural-dynamic (elastic) models in the 

time domain in a coupled simulation 

environment. In recent years, some of 

these codes have been expanded to include 

the additional dynamics pertinent to 

offshore installations, including the 

incident waves, sea current, 

hydrodynamics, and foundation dynamics 

of the support structure [1]. The 

sophistication of these aero-hydro-servoelastic codes, and the limited data 

available with which to validate them,

underscore the need to verify their

accuracy and correctness. The Offshore 

Code Comparison Collaboration (OC3), 

which operates under Subtask 2 of the 

International Energy Agency (IEA) Wind 

Task 23, was established to meet this need.

Overview of OC3

To test the newly developed codes, the 

main activities of OC3 are (1) discussing 

modeling strategies, (2) developing a suite 

of benchmark models and simulations, (3) 

running the simulations and processing the 
2

simulation results, and (4) comparing and 

discussing the results. These activities fall 

under the much broader objectives of

•Assessing the accuracy and reliability of 

simulations to establish confidence in 

their predictive capabilities

•Training new analysts how to run and 

apply the codes correctly

• Investigating the capabilities and 

limitations of implemented theories

•Refining applied analysis methodologies

• Identifying further research and 

development needs.

Such verification work, in the past, has led 

to dramatic improvements in model 

accuracy as the code-to-code comparisons 

and lessons learned have helped identify 

model deficiencies and needed 

improvements. These results are

important because the advancement of the 

offshore wind industry is closely tied to 

the development and accuracy of 

dynamics models.

Participants and Codes

The OC3 project is performed through 

technical exchange among a group of 

international participants who come from 

universities, research institutions, and 

industry across the United States of 

America, Germany, Denmark, the United 

Kingdom, Spain, the Netherlands, 

Norway, Sweden, and Korea. In this 

paper, specifically, results are presented 

from participants from the National 

Renewable Energy Laboratory (NREL), 

Risø National Laboratory of the Technical 

University of Denmark (Risø-DTU),

MARINTEK, the Norwegian University of 

Science and Technology (NTNU), the 

Institute for Energy Technology (IFE), the 

Norwegian University of Life Sciences 

(UMB), Garrad Hassan & Partners 

Limited (GH), the Leibniz University of 

Hannover (LUH), Acciona Energia, and 

the Pohang University of Science and 

Technology (POSTECH).

Most of the aero-hydro-servo-elastic codes 

that have been developed for modeling the 

dynamic response of offshore wind 

turbines are tested within OC3. The 

existing modeling capabilities of the 

simulation tools used by (and for some, 

developed by) each participant are 

summarized in Table 1. In the cases 

where Table 1 shows the same code being

used by multiple OC3 participants, the 

model development, simulation runs, and 

data processing were done independently. 

Further enhancements of the modeling 

capabilities are planned in the future.

Project Approach and Phases

The simulation of offshore wind turbines 

under combined stochastic aerodynamic 

and hydrodynamic loading is very 

complex. The benchmarking task, 

therefore, requires a sophisticated 

approach that facilitates the identification 

of sources of modeling discrepancies 

introduced by differing theories and model 

implementations in the various codes. 

This is possible only by (1) meticulously 

controlling all of the inputs to the codes 

and (2) carefully applying a stepwise 

verification procedure where model 

complexity is increased in each step.

The fundamental set of inputs to the codes 

controlled within OC3 relates to the 

specifications of the wind turbine. The 

OC3 project uses the publicly available

specifications of the 5-MW baseline wind 

turbine developed by NREL, which is a 

representative utility-scale multi-megawatt 
3

turbine. This conventional three-bladed 

upwind variable-speed variable-bladepitch-to-feather controlled turbine, is 

specified with detailed rotor aerodynamic 

properties; blade, drivetrain, nacelle, and 

tower structural properties; and generatortorque and blade-pitch control system

properties, the latter of which was 

provided to all OC3 participants in the 

form of a dynamic link library (DLL). 

The specifications of the NREL offshore 

5-MW baseline wind turbine are available 

in detail [2]. The hydrodynamic and 

elastic properties of the varying offshore 

support structures used in the project are

also controlled. Furthermore, the turbulent 

full-field wind inflow and regular and 

irregular wave kinematics are model 

inputs controlled within OC3. Risø-DTU

generated the turbulent wind velocity 

datasets and GH and NREL (for different 

phases of the project) generated the wave 

kinematics datasets; these datasets were 

then provided to all other participants. 

This approach eliminates any possible 

differences brought about by dissimilar 

turbulence models, wave theories, or 

stochastic realizations.

An important part of the comparison is a 

stepwise process that allows the origin of 

differences between code predictions to be 

discovered. Various combinations of 

wave and wind input were introduced with 

the rotor and tower being rigid or flexible, 

disentangling the contributions from wind 

and wave applied loads and dynamic 

response. Finally, the turbine was made 

operational so that the effect of the control 

system could be evaluated.

Emphasis within OC3 is given to the 

verification of the offshore support 

structure dynamics as part of the dynamics 

of the complete system. This emphasis is 

a feature that distinguishes OC3 from 

previous wind turbine code-to-code

verification exercises. To encompass the 

variety of support structures required for 

FAST Bladed ADAMS HAWC2 3Dfloat Simo SESAM / DeepC

Code Developer

NREL GH MSC + NREL

+ LUH

Risø-DTU IFE-UMB MARINTEK DNV

OC3 Participant

NREL + POSTECH GH NREL + LUH Risø-DTU IFE-UMB MARINTEK Acciona + NTNU

Aerodynamics

( BEM or GDW )

+ DS

( BEM or GDW )

+ DS

( BEM or GDW )

+ DS

( BEM or GDW )

+ DS

( BEM or GDW ) BEM None

Hydrodynamics

Airy+ + ME,

Airy + PF + ME

( Airy+ or Stream )

+ ME

Airy+ + ME,

Airy + PF + ME

Airy + ME Airy + ME Airy + PF + ME Airy+ + ME,

Airy + PF + ME

Control System (Servo)

DLL, UD, SM DLL DLL, UD DLL, UD, SM UD DLL None

Structural Dynamics (Elastic)

Turbine: FEMP +

( Modal / MBS ),

Moorings: QSCE

Turbine: FEMP +

( Modal / MBS ),

Moorings: UDFD

Turbine: MBS,

Moorings: QSCE,

UDFD

Turbine: MBS / FEM,

Moorings: UDFD

Turbine: FEM,

Moorings: FEM, UDFD

Turbine: MBS,

Moorings: QSCE,

MBS

Turbine: MBS,

Moorings: QSCE,

FEM

Airy+ – Airy wave theory

+) with free surface corrections 

BEM – blade-element / momentum

DLL – external dynamic link library

DNV – Det Norsk Veritas

DS – dynamic stall

GDW – generalized dynamic wake

FEMP – finite-element method

P) for mode preprocessing only

MBS – multibody-dynamics formulation

ME – Morison’s equation

MSC – MSC Software Corporation

PF – linear potential flow with radiation & 

diffraction

QSCE – quasi-static catenary equations

SM – interface to Simulink® with MATLAB®

UD – implementation through user-defined 

subroutine available

UDFD – implementation through user-defined forcedisplacement relationships

Table 1: Overview of aero-hydro-servo-elastic modeling capabilities
4

cost effectiveness at varying offshore sites, 

different support structures (for the same 

wind turbine) are investigated in separate 

phases of OC3:

• In Phase I, the NREL offshore 5-MW 

wind turbine is installed on a monopile 

with a rigid foundation in 20 m of water.

• In Phase II, the foundation of the 

monopile from Phase I is made flexible 

by applying different models to represent

the soil-pile interactions.

• In Phase III, the water depth is changed 

to 45 m and the monopile is swapped 

with a tripod substructure, which is one 

of the common space-frame concepts 

proposed for offshore installations in 

water of intermediate depth.

• In Phase IV, the wind turbine is installed 

on a floating spar buoy in deep water

(320 m).

The OC3 project started in fall of 2004

and ended in the winter of 2009. A follow 

up project to OC3 is being initiated in 

2010 through the IEA so that two more 

phases can be considered, including the 

analysis of a fixed-bottom jacket and a 

floating semi-submersible. A description 

of Phases I, II, and III and their results are 

presented in detail in [3], [4], and [5], 

respectively. This paper describes Phase 

IV and discusses its results.

Review of Results from Earlier 

Phases

Before discussing Phase IV, it is prudent

to summarize the key findings from 

Phases I through III because many of the 

results have followed through into 

subsequent phases. The code-to-code 

comparisons in Phases I through III have 

agreed very well, in general. The key 

reasons for the differences that have 

remained and the other findings from 

Phases I through III are discussed below

[3, 4, 5].

Phase I – Monopile with Rigid 

Foundation

•The modal-based codes predict slightly 

different second and higher coupled 

eigenmodes than what are predicted by 

the higher fidelity MBS- and FEM-based 

codes. This is because the higher 

eigenmodes in higher-fidelity codes 

include system couplings that are not 

accounted for in the lower fidelity 

modal-based codes. Differences in the 

dynamic response and energy content, 

therefore, appear in the higher frequency 

range.

•The codes that rely on full-field wind 

available in polar coordinates predict 

smoother aerodynamic loads (and thus 

smaller load deviations and smaller 

damage equivalent loads (DELs)) than 

codes that rely on rectangular 

coordinates. This results from the 

method in which the wind datasets were 

generated. To ensure that all participants 

used the same wind inflow, the full-field 

wind datasets were generated in 

rectangular coordinates and subsequently 

interpolated to polar coordinates for the 

codes that needed it. These differences 

were mitigated as much as possible by 

using a fine spatial resolution (32 × 32 

points across the rotor disk).

•The differences among the codes relating 

to the implementation of aerodynamic 

induction, tower interference, hub and tip 

loss, and dynamic stall models—and 

whether or not the aerodynamic loads are 

applied in the deflected or undeflected 

blade state—attribute to variations in the 

mean values of several key wind turbine 
5

loads (e.g., blade-root bending moments, 

rotor torque, and rotor thrust).

•The blade-pitch controller compensates 

somewhat for variations that might have 

been caused between codes that do and 

do not have blade-twist degrees of 

freedom (DOFs).

•Differing model discretizations for the 

aerodynamic and hydrodynamic loads 

lead to differences among the code 

predictions. This is most apparent in the 

substructure loads that depend highly on 

the discretization of hydrodynamic loads 

near the free surface, such as the bending 

moments in the pile at the mudline.

•Though every effort has been made to 

standardize model inputs, user error still 

occurs. It often takes several revisions 

before the model is developed and run as 

intended. It is also possible, in some 

instances, that errors still remain and 

account for otherwise inexplicable

modeling differences.

Phase II – Monopile with Flexible 

Foundation

•All of the results of Phase I also apply to 

the results of Phase II.

•All three of the simplified foundation 

models—apparent fixity, coupled 

springs, and distributed springs—can be 

derived and implemented to ensure that 

the overall response of the system above 

the mudline is identical under a given set 

of loading conditions (at least for the 

lowest system eigenmodes).

•The descretization problems described 

for the results of Phase I result in higher 

excitation in the second eigenmodes of 

the support structure in Phase II. 

However, this higher excitation is only 

visible when the turbine is not operating, 

because aerodynamic loading tends to 

limit the amount of additional excitation

(i.e., aerodynamic loading tends to damp 

out the excitation).

•The differing implementations of the 

aerodynamic models among the codes 

have more effect on the mean values of 

the wind turbine loads than on the power 

spectra.

Phase III – Tripod

•All of the results of Phase I also apply to 

the results of Phase III. (The results of 

Phase II have less bearing on Phase III 

because Phase III does not use 

foundation models).

•The most straightforward way to account 

for buoyancy loads in non-flooded multimember structures is through direct 

integration of the hydrostatic pressure 

that is dependent on the time-varying 

wave elevation (as opposed to 

accounting for buoyancy as a displaced 

volume with corrections for end effects). 

This is important for non-flooded 

members that are inclined, tapered, 

and/or embedded into the seabed (i.e., 

non-flooded piles). The OC3 code 

comparisons differed until everyone 

agreed on this approach.

•The most straightforward way to model a 

rigid multi-member structure is to 

increase the Modulus of Elasticity by 

several orders of magnitude consistently 

across all members. This permits 

calculation of how the loads are 

transmitted through what is a statically

indeterminate structure (because of the 

geometry of the tripod) in the limit of 

zero deflection. The OC3 code 

comparisons differed for those load cases 

that considered a rigid tripod until 

everyone agreed on this approach.
6

•Differing discretizations for the 

hydrodynamic and buoyancy loads along 

tapered members lead to differences 

among the code predictions. To 

eliminate the discrepancy, the 

hydrodynamic and buoyancy loads along 

tapered members must be finely 

discretized. Because the hydrodynamic 

inertia and buoyancy loads depend on the 

square of the member diameter, having 

too long a length between nodes can 

cause a large error in the total load.

•Because of the large diameter members 

of the tripod, significant surface areas 

and volumes are duplicated at the joints, 

which distorts the overall level of 

loading if the intersection is not 

accounted for in the mass, stiffness, 

hydrodynamic loading, and buoyancy 

loading. This was found to have a large 

effect on the overall loading and 

response of the tripod.

•The local shear deflection of the 

members in a multi-member support 

structure, which can be modeled with 

Timoshenko beam elements, was found 

to have a large effect on the distribution 

of loads through multi-member 

structures. This was a surprising finding 

because all of the beam members of the 

tripod analyzed in OC3 were thin and 

slender, such that the Bernoulli-Euler 

approach, which neglects shear 

deflection, was thought to be sufficient 

(and is sufficient in blades, towers, and 

monopiles). The shear effect was found 

to be much more important than 

originally assumed because the beam 

members are attached rigidly to other 

members and the relative displacement 

of each member influences the load 

distribution. The results of the code 

comparisons could clearly be grouped 

between the codes that do and do not 

have beam elements with shear 

deflection models.

•The initial transient solution takes a long 

time to dissipate due to the small amount 

of damping in the tripod and the method 

by which the hydrodynamic loads are 

initialized at the start of a simulation. 

The initial transient is longer when the 

turbine is not operating, because 

aerodynamic loading tends to damp out 

the initial solution quickly. Each model 

initializes its solution differently, so, the 

code comparisons differed during the 

start-up transient period.

Overview of Phase IV

Numerous floating platform concepts are 

possible for offshore wind turbines, 

including spar-buoys, tension leg 

platforms (TLPs), barges, and hybrid 

concepts. In OC3 Phase IV, the spar-buoy 

concept called “Hywind,” developed by 

Statoil of Norway, was imitated. This 

concept was chosen for its simplicity in 

design, suitability to modeling, and the 

existence of a full-scale prototype. Statoil 

graciously supplied detailed platform and 

mooring system data for the conceptual 

version of the Hywind platform that was

developed to support a 5-MW wind 

turbine, as analyzed in [6, 7, 8]. Per the 

request of Statoil, the original data was

condensed and sanitized to make it 

suitable for public dissemination. Aspects 

of the original data were also adapted 

slightly so that the platform design 

appropriately supports the NREL 5-MW 

baseline turbine, which has slightly 

different properties than the turbine Statoil 

used to develop their system. The rotornacelle assembly of the NREL 5-MW 

turbine—including the aerodynamic and 

structural properties—remains the same as 

in Ref. [2], but the support structure (tower 
7

and substructure) and control system 

properties have been changed. The new 

system is referred to as the “OC3-

Hywind” system, to distinguish it from 

Statoil’s original Hywind concept. The 

OC3-Hywind system features a deeply 

drafted, slender spar buoy with catenary 

mooring lines as illustrated in Figure 1. 

Some of the system data is listed in 

Table 2; the concept is documented in 

much greater detail in [9].

Floating 

offshore wind 

turbines are a 

good test for 

aero-hydroservo-elastic 

codes because 

they incorporate

a number of 

features not 

present in 

conventional 

fixed-bottom 

offshore support 

structures:

•The introduction of very low frequency 

modes that can impact the aerodynamic 

damping and stability of the system;

•The possibility of significant 

translational and rotational motions of 

the support structure, which can couple 

with the motions of the rotor-nacelle 

assembly;

•The mooring system and anchoring is a 

new component that must be considered

in the overall analysis; and

•The support structure need not be slender 

and cylindrical (although this is nearly 

the case in the OC3-Hywind system), 

such that hydrodynamic radiation,

diffraction and other wave effects can 

become important.

In Phase IV, a set of load-case simulations 

has been specified for the OC3-Hywind 

system, as summarized in Table 3. Most 

of the load cases in Table 3 correspond to

those of equivalent simulations from 

earlier phases (see Refs. [3, 4, 5]). In 

Phase IV, though, it was not necessary to 

run all of the earlier simulations, and as a 

result, the load-case identifiers are not 

sequential.

The load cases new to Phase IV are cases 

1.3, 1.4, and 5.4. Case 1.3 is used to 

compare the static equilibrium condition 

of the full system void of wind or wave 

excitation. This case is useful for 

checking that the combined weight of the 

rotor-nacelle assembly, tower, and 

platform balances with the static buoyancy 

(i.e., weight of the displaced fluid) and 

mooring reactions identically between the 

models. Case 1.4 is used to test the OC3-

Figure 1: Illustration 

of the NREL 5-MW 

wind turbine on the 

OC3-Hywind spar

Depth to Platform Base Below SWL (Total Draft) 120 m

Elevation to Platform Top (Tower Base) Above SWL 10 m

Depth to Top of Taper Below SWL 4 m

Depth to Bottom of Taper Below SWL 12 m

Platform Diameter Above Taper 6.5 m

Platform Diameter Below Taper 9.4 m

Platform Mass, Including Ballast 7,466,330 kg

CM Location Below SWL Along Platform Centerline 89.9155 m

Platform Roll Inertia about CM 4,229,230,000 kg•m2

Platform Pitch Inertia about CM 4,229,230,000 kg•m2

Platform Yaw Inertia about Platform Centerline 164,230,000 kg•m2

Number of Mooring Lines 3

Angle Between Adjacent Lines 120o

Depth to Anchors Below SWL (Water Depth) 320 m

Depth to Fairleads Below SWL 70.0 m

Radius to Anchors from Platform Centerline 853.87 m

Radius to Fairleads from Platform Centerline 5.2 m

Unstretched Mooring Line Length 902.2 m

Mooring Line Diameter 0.09 m

Equivalent Mooring Line Mass Density 77.7066 kg/m

Equivalent Mooring Line Weight in Water 698.094 N/m

Equivalent Mooring Line Extensional Stiffness 384,243,000 N

Additional Yaw Spring Stiffness 98,340,000 Nm/rad

Table 2: Summary of OC3-Hywind spar properties
8

Hywind system’s hydrodynamic damping 

through free-decay tests. This case 

considers only the six rigid-body DOFs of 

the platform in still water; each DOF is 

perturbed separately, and the time-series 

responses are compared. Case 5.4 is used 

to test the OC3-Hywind’s frequency 

response. This case considers steady 

uniform winds without shear at 8 m/s, 

regular Airy 2-m high waves with 

different frequencies in each simulation, 

the turbine control system and all DOFs 

enabled, and simulates until a periodic 

steady-state condition is reached. The 

outputs of case 5.4 are the time-seriesgenerated Response Amplitude Operators 

(“effective RAOs”). An “effective RAO” 

is defined here to mean the difference in 

response amplitudes between nonlinear 

time-domain simulations run with and 

without wave excitation. That is, the 

outputs are the effective amplitudes of the 

platform motions, turbine displacements, 

loads, etc. normalized by the wave 

amplitude (1 m) versus wave frequency. 

The word “effective” is used to distinguish 

these values from RAOs that are typically 

found using linear frequency-domain 

techniques.

For each load-case simulation, a total of 

57 model outputs were analyzed. In 

addition to the 39 outputs analyzed in 

Phases I through III for the rotor, 

drivetrain, nacelle, tower, and environment 

(again, see Refs. [3, 4, 5]), 18 outputs were 

used in Phase IV to analyze the loads and 

deflections of the floating platform and 

mooring system.

Load

Case

Enabled DOFs Wind Conditions Wave Conditions Analysis Type

1.2 Platform, tower,

drivetrain, blades

None: air density = 0 Still water Eigenanalysis

1.3 Platform, tower,

drivetrain, blades

None: air density = 0 Still water Static equilibrium solution

1.4 Platform None: air density = 0 Still Water Free-decay test time series

4.1 Platform, tower None: air density = 0 Regular Airy: H = 6 m, T = 10 s Periodic time-series solution

4.2 Platform, tower None: air density = 0 Irregular Airy: Hs = 6 m, Tp = 10 s,

JONSWAP wave spectrum

Time-series statistics, DELs,

power spectra

5.1 Platform, tower,

drivetrain, blades

Steady, uniform, no shear:

Vhub = 8 m/s

Regular Airy: H = 6 m, T = 10 s Periodic time-series solution

5.2 Platform, tower,

drivetrain, blades

Turbulent: Vhub = Vr (11.4 m/s),

σ1 = 1.981 m/s, Mann model

Irregular Airy: Hs = 6 m, Tp = 10 s,

JONSWAP wave spectrum

Time-series statistics, DELs,

power spectra

5.3 Platform, tower,

drivetrain, blades

Turbulent: Vhub = 18 m/s,

σ1 = 2.674 m/s, Mann model

Irregular Airy: Hs = 6 m, Tp = 10 s,

JONSWAP wave spectrum

Time-series statistics, DELs,

power spectra

5.4 Platform, tower,

drivetrain, blades

Steady, uniform, no shear:

Vhub = 8 m/s

Regular Airy: H = 2 m,

ω = 0.1, 0.2,..., 3.5 rad/s

Time-series-generated

“effective RAOs”

H – individual wave height

Hs – significant wave height

T – individual wave period

Tp – peak spectral period

Vhub – hub-height wind speed averaged 

over 10 minutes

Vr – rated wind speed

σ1 – longitudinal wind speed standard 

deviation

ω – individual wave frequency

Table 3: Summary specifications for the Phase IV load-case simulations
9

Phase IV Results

Each load-case simulation of 

Phase IV was run by the OC3 

participants. The legend in 

Figure 2 delineates how the 

results are presented in the

figures that follow. The color 

shade and line type distinguish 

the results from separate 

participants and codes. Some

results were not processed by all 

of the OC3 participants, which is 

why some of the predictions are 

blank or have a zero value. Only 

a small subset of the results is 

presented.

Full-System Eigenanalysis

Figure 3 gives the lowest 19 natural 

frequencies calculated for the stationary

OC3-Hywind system in still water from 

load case 1.2. Results were 

obtained from all codes except 

FAST by POSTECH and Bladed. 

While case 1.2 calls for enabling 

all system DOFs, the rotornacelle assembly is modeled 

rigidly in ADAMS by LUH and 

3Dfloat, and the rotor-nacelle 

assembly and tower are modeled 

rigidly in Simo, SESAM, and 

DeepC. The designation of 

“pitch” and “yaw” in the 

asymmetric flapwise and 

edgewise blade modes identifies 

coupling of the blade motions 

with the nacelle-pitching and 

nacelle-yawing motions, respectively.

Most of the codes agree on their 

predictions of the lowest six natural 

frequencies, which correspond to the six 

(primarily) rigid-body modes of the 

floating support platform (with only minor 

Figure 2: Results Legend

Figure 3: Full-system hydro-elastic natural frequencies from load case 1.2
10

couplings to the turbine flexibilities). The 

platform roll and pitch natural frequencies 

are predicted higher by Risø-DTU with the 

HAWC2 code. While these differences

are still being investigated, one thought is 

that they are caused by the treatment of 

gravity in the model linearization process, 

as gravity has a strong influence on the 

roll and pitch restoring of a deep-drafted 

spar.

Not all of the codes predict the natural 

frequencies of the flexible body modes. 

The codes that model the tower and blade 

bending and drivetrain torsion modes

predict similar natural frequencies, with 

only slight discrepancies. The discrepancy

in the second blade asymmetric flapwise 

yaw frequency has been seen in earlier 

phases of OC3, particularly Phase II [4]. 

In this mode, the vertically positioned 

blade remains stationary, while the two 

other blades flap out of phase with each 

other. These blade motions couple with 

the torsion of the tower and—in Phase 

IV—the yaw of the platform (both being 

rotations about the tower centerline). The 

codes which do not account for tower 

torsion (FAST) predict higher natural 

frequencies—corresponding to the stiffer 

compliance about the tower centerline—

than the codes that do (ADAMS, 

HAWC2).

Free Decay

Figure 4 through Figure 6 give the 

platform surge, heave, and pitch time 

series from perturbations in platform 

surge, heave, and pitch, respectively, from 

the free-decay tests of load case 1.4. 

Results were obtained from all codes 

except Bladed and DeepC.

For the surge free-decay test of Figure 4, 

all codes except FAST by POSTECH 

agree on the surge displacement. Both 

NREL and POSTECH used FAST and 

their results compare well, except that 

there is less hydrodynamic damping in 

POSTECH’s results. This is caused by 

POSTECH’s FAST model missing one 

hydrodynamic damping term (the socalled “additional linear damping”

specified for the OC3-Hywind spar). 

When comparing the heave-surge coupling

between the codes, 3Dfloat predicts less

coupling, HAWC2 predicts more 

damping, and FAST by POSTECH predict 

less damping. The pitch-surge coupling 

can be placed into two groups—FAST by 

POSTECH and ADAMS by LUH in one, 

and FAST and ADAMS by NREL, 

HAWC2, 3Dfloat, and SESAM in another. 

The pitch offset results from the 

overhanging mass of the rotor-nacelle 

assembly from the centerline of the tower 

and platform.

For the heave and pitch-heave coupling 

responses in the heave free-decay test of

Figure 5, all codes agree well except 

SESAM, which predicts no coupling to 

pitch, and HAWC2, which has too much 

heave and pitch damping.

All codes agree on the platform-pitch 

response in the pitch free-decay test of

Figure 6, except ADAMS by LUH, which

predicts too little damping. HAWC2 

predicts different surge-pitch and heavepitch couplings than the other codes.
11

Figure 7 gives the platformyaw response from the yaw 

free-decay test of load case 1.4. 

All codes agree very well.

Hydro-Elastic 

Response with 

Regular Waves

Figure 8 shows time histories 

of platform surge, heave, and 

pitch displacements; tower-top fore-aft 

deflection and shear force; and 

downstream fairlead tension (from 

mooring line #1) from load 

case 4.1. This case considers 

the response of the platform 

and flexible tower (with rigid 

rotor-nacelle-assembly) excited 

by regular (i.e., periodic) 

waves. Results were obtained 

from NREL with FAST and

ADAMS, Risø-DTU with 

HAWC2, IFE-UMB with 

3Dfloat, MARINTEK with 

Simo, and Acciona with SESAM. The 

responses for two wave passages (i.e., 

Figure 4: Free decay in platform surge from load case 1.4

Figure 5: Free decay in platform heave from load case 1.4

Figure 6: Free decay in platform pitch from load case 1.4

Figure 7: Free decay in 

platform yaw from case 1.4
12

periods) are shown. All codes removed 

start-up transients from their results. The 

instantaneous wave elevation at the 

platform centerline is highest at 0 and 10 s 

and lowest at 5 and 15 s, except in 3Dfloat 

and Simo, which are phase-shifted relative 

to the other codes.

In these results, all codes agree on the 

platform-surge amplitude except HAWC2, 

which may simply have output the wrong 

parameter. For pitch and heave, all codes 

agree. For tower fore-aft deflection—

deflection induced by platform motion—

most codes agree quite well. Simo and 

SESAM predict no deflection because the 

tower was modeled rigidly in those codes. 

The tower-top fore-aft shear force

(induced by the oscillating mass of the 

rotor-nacelle assembly) and downstream 

fairlead tension agree well in all codes 

except SESAM.

Hydro-Elastic Response with 

Irregular Waves

Figure 9 shows power spectra computed in 

load case 4.2 for the same parameters 

shown in Figure 8. Case 4.2 tests the same 

model used in case 4.1, but with excitation 

from irregular (stochastic) waves. Results 

were obtained from all codes except FAST 

by POSTECH, ADAMS by LUH, and 

3Dfloat. Because several codes contain 

start-up transients, the statistical results 

from this case are difficult to compare and

are not presented.

The wave spectrum has the highest 

amount of energy at 0.1 Hz, corresponding 

to the peak-spectral period of 10 s. All 

codes agree on the energy content of the 

system responses at this frequency, except 

SESAM in platform pitch. The platform 

natural frequencies (about 0.008 Hz in 

surge, 0.032 Hz in heave, and 0.034 Hz in 

pitch) are easily visible in the platformdisplacement power spectra of all codes; 

Figure 8: Hydro-elastic time series with regular waves from load case 4.1
13

however, the codes differ on their 

predictions of energy content at and 

around these frequencies (with the farthest 

outliers being Bladed in surge, Simo in 

heave, and SESAM in pitch). Spectral 

shape estimates in this frequency range are

sensitive to time-series sampling. There is

considerable statistical uncertainty of 

response spectra because of the relatively 

short duration of the prescribed time series 

compared to the wave and platform 

periods.

In the tower-top fore-aft deflection and 

shear force power spectra, the platformpitch and first tower bending (about 0.46 

Hz) natural frequencies are clearly visible. 

Bladed, however, shows more energy just 

below the first tower bending natural 

frequency and less energy at the platformpitch natural frequency. SESAM and 

DeepC predict higher energy than the 

other codes in fairlead tension above 0.1 

Hz, likely the result of undamped highfrequency motions in their FEM-based 

mooring line solution.

Aero-Hydro-Servo-Elastic 

Response with Regular Waves

Figure 10 gives time histories of out-ofplane blade-tip deflection; generator 

power; rotor speed; platform surge, heave, 

pitch, and yaw displacements; tower-top 

fore-aft deflection, shear force, and rotor 

torque; and downstream and upstream 

fairlead tensions (from mooring line #1 

and #2) from load case 5.1. This case is 

like 4.1 except that the rotor and drivetrain 

are flexible, the controller is enabled, and 

the turbine is operating in uniform wind 

below rated wind speed. In addition to the 

wind excitation, which leads to nonzero 

mean loads and displacements, the system 

Figure 9: Hydro-elastic power spectra with irregular waves from load case 4.2
14

is again excited by regular waves that

induce oscillations about the means. The 

responses are shown for two wave 

passages (i.e., periods)—which equates to 

about three rotor revolutions—after all 

start-up transients have died out. Results 

were obtained from all codes but Bladed, 

ADAMS by LUH, and DeepC. The 

instantaneous wave elevation at the

platform centerline is highest at 0 and 10 s 

and lowest at 5 and 15 s, except in 3Dfloat

and Simo, which are phase-shifted relative 

to the other codes.

The generator power agrees quite well 

between all codes (at least as well as they

have in earlier OC3 phases), except FAST 

by POSTECH, which predicts a lower 

Figure 10: Aero-hydro-servo-elastic time series with regular waves from load case 5.1
15

mean and amplitude. Likewise, all codes 

agree well in rotor speed, except FAST by 

POSTECH and 3Dfloat, which predict 

lower means and amplitudes.

The out-of-plane blade-tip deflection 

agrees reasonably well between all codes 

except FAST by POSTECH, which has 

less overall oscillation, and HAWC2, 

which may simply have output the wrong 

parameter. 3Dfloat, Simo, and SESAM

predict no deflection because their rotors 

were modeled rigidly.

Except for FAST by POSTECH, all codes 

predict a similar mean value of the towertop rotor torque (i.e., roll moment), but the 

oscillations about this mean vary. FAST 

by POSTECH predicts a lower mean and 

oscillation amplitude. All codes agree on 

the tower-top fore-aft shear (thrust) force, 

but 3Dfloat results in a larger tower-top 

deflection. Simo and SESAM predict no 

deflection because the tower was modeled 

rigidly in those codes.

For the platform-surge displacement, all 

codes but HAWC2 agree on the amplitude 

of oscillation, but the mean values vary

despite the agreement on thrust—Simo,

3Dfloat, and SESAM are a bit higher and 

HAWC2 is very low (HAWC2 also has 

less amplitude). The results are similar for 

platform-pitch displacement, except that 

HAWC2 matches the majority of codes 

much better. The nonzero means in these 

displacements are the result of the nonzero 

mean rotor thrust. The higher amount of 

platform-surge and -pitch displacements in 

3Dfloat results from the slightly higher 

mean thrust.

The rotor thrust is counteracted by 

tensions in the mooring lines, with the 

downstream mooring line less loaded than 

the upstream lines. The variations in the 

fairlead tensions are fairly consistent with 

the variations in platform displacement. 

That is, the downstream fairlead has a 

lower tension and the upstream fairleads 

have higher tension in 3Dfloat and 

SESAM than in the other codes.

The differing platform-heave 

displacements imply that the codes differ 

in how they couple heave with surge and 

pitch. The slightly negative mean 

platform-yaw displacement—predicted 

similarly by most codes—results from the 

rotor-shaft tilt, which causes some of the 

rotor torque to act about the yaw axis. The

oscillation about this mean—resulting 

from a gyroscopic moment produced by 

platform pitching (at the wave period) in 

combination with spinning rotor inertia—

varies a bit between the codes.

Aero-Hydro-Servo-Elastic 

Response with Irregular 

Waves

Figure 11 shows power spectra computed 

in load case 5.3 for the same parameters 

shown in Figure 10. Case 5.3 tests the 

same model used in case 5.1, but with 

excitation from stochastic wind above 

rated wind speed and irregular (stochastic) 

waves. Results were obtained from NREL 

and POSTECH with FAST, GH with 

Bladed, NREL with ADAMS, 

MARINTEK with Simo, and Acciona with 

SESAM.
16

Figure 11: Aero-hydro-servo-elastic power spectra with irregular waves from load case 5.3
17

The power spectra agree very well in most 

parameters. As in case 4.2, the wave

spectrum has the highest amount of energy

at 0.1 Hz—corresponding to the peakspectral period of 10 s—and all codes 

agree on the energy content of the system 

response at this frequency, except SESAM 

in platform-pitch displacement. Unlike 

case 4.2, there is a high amount of energy 

at low frequencies (below 0.1 Hz)—

corresponding to the highest energy in the 

wind—and all codes agree on the energy 

content of the system response at these 

frequencies, except SESAM again in 

platform-pitch displacement and Simo in 

generator power. Simo also has more 

energy in the generator power above 0.1 

Hz. As in the previous SESAM results, 

the higher energy content in the fairlead 

tensions above 0.1 Hz is clearly visible. 

The second tower and blade bending 

natural frequencies are visible in the 

FAST, Bladed, and ADAMS predictions 

of the tower-top shear force and rotor 

torque. However, ADAMS predicts less 

energy at these frequencies, which may 

result from the increase in numerical 

damping with frequency that is typical of 

ADAMS simulations.

“Effective RAOs”

Figure 12 shows “effective RAOs” 

computed in load case 5.4 for the same 

parameters shown in Figure 10 and Figure 

11. Case 5.4 tests the same model used in 

case 5.1, but with a set of regular waves at 

varying frequency. Results were obtained 

from NREL and POSTECH with FAST, 

NREL with ADAMS, IFE-UMB with 

3Dfloat, and Acciona with SESAM.

Typical of conventional RAOs, the 

“effective RAO” responses are largest at 

the system’s natural frequencies. For 

platform pitch and heave, the “effective 

RAOs” show excitation at and around the 

pitch and heave natural frequencies (both 

about 0.2 rad/s). For platform surge, the 

most excitation occurs at the very lowest 

frequencies and at the platform-pitch 

natural frequency. The latter indicates 

coupling between surge and pitch. The 

behavior of the fairlead tension “effective 

RAOs” is similar to surge, indicating that 

the surge is what most influences the 

tensions. For platform yaw, the “effective 

RAO” shows excitation at and around the 

yaw natural frequency (0.75 rad/s), which 

is caused by gyroscopic loading from the 

spinning rotor combined with platform 

pitching. The “effective RAOs” for outof-plane blade-tip deflection and tower-top 

fore-aft deflection, shear force, and rotor 

torque show broadband excitation across 

all frequencies, with extra excitation at the 

platform-pitch and first tower-bending (3.0 

rad/s) natural frequencies. The generator 

power and rotor speed “effective RAO” 

behavior are similar, but contain less 

energy at the first tower-bending natural 

frequency.

In the “effective RAO” code-to-code 

comparisons, ADAMS had trouble 

converging at 1.5, 2.0, and 2.2 rad/s

(hence the “spikes”); otherwise, the 

ADAMS results agree with the FAST 

results by NREL.

3Dfloat also agrees well with NREL’s 

FAST results, except that 3Dfloat predicts 

lower excitation in platform yaw and higher 

excitation in the fairlead tensions and at and 

around the first tower-bending natural 

frequency in all parameters. This latter issue 

is still under investigation; possible 

explanations are (1) differences in the 

aerodynamic damping due to the rigid rotor 

or differing aerodynamic theories or (2) the 
18

modeling of the rigid spar with an artificially 

high stiffness in 3Dfloat.

SESAM predicts higher excitation in the 

platform-surge and -pitch displacements 

and fairlead tensions. The latter are 

Figure 12: Aero-hydro-servo-elastic “effective RAOs” with regular waves from case 5.4
19

attributed to a lack of convergence in 

SESAM’s FEM-based mooring line 

solution. SESAM also predicts lower 

excitation in platform yaw than does 

FAST and ADAMS.

The “effective RAOs” from FAST by 

POSTECH differ with those from FAST 

by NREL over several parameters, 

including out-of-plane blade-tip 

deflection; generator power; rotor speed;

platform-surge, -heave, and –pitch 

displacements; and fairlead tensions. The 

differences in platform displacements and 

the resulting differences in the fairlead 

tensions are caused by the lesser amount 

of hydrodynamic damping in POSTECH’s 

FAST model, as described earlier. 

Interestingly, this drop in damping leads to 

surge displacement and resulting fairlead 

tensions with negative “effective RAOs,” 

which means that there is less system 

motion with waves than in still water.

The reason for negative “effective RAOs”

is as follows. In simulations with wave 

excitation, the wave-induced motions 

occur at the wave-excitation frequency, 

which brings about (potential-flow-based) 

wave-radiation damping. The waveradiation damping in surge is not 

negligible above 0.2 rad/s (see [9]). In 

still water, the wave frequency is 

eliminated from the excitation/response, 

thereby eliminating the wave-radiation 

damping, which leads to a minor 

controller-induced instability of the 

platform-surge mode. Because of the 

system’s nonlinearities, this instability 

leads to a limit-cycle oscillation of the 

platform-surge mode at the surge natural 

frequency (where there is negligible 

hydrodynamic damping) when operating 

in still water. The wave-radiation 

damping resulting from wave excitation 

reduces or eliminates the platform-surge 

instability (depending on the waveexcitation frequency) and resulting limitcycle oscillations, thereby reducing the 

overall motions considerably.

The controller-induced instability and 

resulting limit-cycle oscillations in still 

water are largest at rated wind speed and 

are present in the system even with the 

proper amount of hydrodynamic damping. 

Due to the strong influence of the 

(potential-flow-based) wave-radiation 

damping in these results, it is noted that 

neglecting this damping will likely lead to 

vastly different “effective RAOs” at and 

around rated wind speed.

Conclusions

Offshore wind turbines are designed and 

analyzed using comprehensive simulation 

codes that model the systems’ coupled 

dynamic aero-hydro-servo-elastic 

response. The OC3 project, which

operates under Subtask 2 of the IEA Wind 

Task 23, has performed work to verify the 

codes to assess their accuracy and 

correctness.

In Phase IV of OC3, participants used an 

assortment of codes to model the coupled 

dynamic response of the NREL 5-MW 

wind turbine installed on a deeply drafted, 

slender spar buoy with catenary mooring 

lines in 320 m of water. Code predictions 

were compared from load-case simulations

selected to test different model features. 

The comparisons have resulted in a greater

understanding of offshore floating wind 

turbine dynamics and modeling 

techniques, and better knowledge of the 

validity of various approximations.
20

The verification activities performed in 

OC3 are important because the 

advancement of the offshore wind industry 

is closely tied to the development and 

accuracy of dynamic models. Not only 

have vital experiences and knowledge 

been exchanged among the project 

participants, but the lessons learned have 

helped identify deficiencies in existing 

codes and needed improvements, which 

will be used to improve the accuracy of 

future predictions.

References

[1].Passon P and Kühn M. “State-of-theart and Development Needs of Simulation 

Codes for Offshore Wind Turbines.” 

Copenhagen Offshore Wind 2005 

Conference and Expedition, 26–28 

October, Copenhagen, 2005. (CD-ROM)

[2].Jonkman J, Butterfield S, Musial W, 

and Scott G. Definition of a 5-MW 

Reference Wind Turbine for Offshore 

System Development. NREL/TP-500-

38060. NREL: Golden, CO, 2009.

[3].Passon P, Kühn M, Butterfield S, 

Jonkman J, Camp T, and Larsen TJ.

“OC3—Benchmark Exercise of AeroElastic Offshore Wind Turbine Codes.” 

The Second Conference on The Science of 

Making Torque From Wind, 28–31 

August, Copenhagen, 2007. [Online].

Available: http://iopscience.iop.org/1742-

6596/75/1/012071/pdf?ejredirect=.iopscie

nce.

[4].Jonkman J, et al. “Offshore Code 

Comparison Collaboration within IEA 

Wind Annex XXIII: Phase II Results 

Regarding Monopile Foundation 

Modeling.” 2007 European Offshore Wind 

Conference & Exhibition, 4–6 December, 

Berlin, 2007. [Online]. Available: 

http://www.eow2007

proceedings.info/allfiles2/206_Eow2007

fullpaper.pdf.

[5].Nichols J, et al. “Offshore Code 

Comparison Collaboration within IEA 

Wind Annex XXIII: Phase III Results 

Regarding Tripod Support Structure 

Modeling.” 47th AIAA Aerospace Sciences 

Meeting Including The New Horizons 

Forum and Aerospace Exhibition, 5–8 

January, Orlando, FL, 2008. (DVD-ROM)

[6].Nielsen FG, Hanson TD, and Skaare B.

“Integrated Dynamic Analysis of Floating 

Offshore Wind Turbines.” 25th

International Conference on Offshore 

Mechanics and Arctic Engineering, 4–9 

June, Hamburg, 2006. (CD-ROM)

[7].Skaare B, Hanson TD, and Nielsen FG.

“Importance of Control Strategies on 

Fatigue Life of Floating Wind Turbines.” 

26th International Conference on Offshore 

Mechanics and Arctic Engineering, 10–15 

June, San Diego, CA, 2007. (CD-ROM)

[8].Larsen TJ, and Hanson TD. “A Method 

to Avoid Negative Damped Low Frequent 

Tower Vibrations for a Floating, Pitch 

Controlled Wind Turbine.” The Second 

Conference on The Science of Making 

Torque From Wind, 28–31 August,

Copenhagen, 2007. [Online]. Available:

http://iopscience.iop.org/1742-

6596/75/1/012073/pdf?ejredirect=.iopscie

nce.

[9].Jonkman J. Definition of the Floating 

System for Phase IV of OC3. NREL: 

Golden, CO, 2009 (to be published).
F1147-E(10/2008)

REPORT DOCUMENTATION PAGE Form Approved

OMB No. 0704-0188

The public reporting burden for this collection of information is estimated to average 1 hour per response, including the time for reviewing instructions, searching existing data sources, 

gathering and maintaining the data needed, and completing and reviewing the collection of information. Send comments regarding this burden estimate or any other aspect of this 

collection of information, including suggestions for reducing the burden, to Department of Defense, Executive Services and Communications Directorate (0704-0188). Respondents 

should be aware that notwithstanding any other provision of law, no person shall be subject to any penalty for failing to comply with a collection of information if it does not display a 

currently valid OMB control number.

PLEASE DO NOT RETURN YOUR FORM TO THE ABOVE ORGANIZATION.

1. REPORT DATE (DD-MM-YYYY)

April 2010

2. REPORT TYPE

Conference Paper

3. DATES COVERED (From - To)

4. TITLE AND SUBTITLE

Offshore Code Comparison Collaboration within IEA Wind Task 23: 

Phase IV Results Regarding Floating Wind Turbine Modeling; 

Preprint

5a. CONTRACT NUMBER

DE-AC36-08-GO28308

5b. GRANT NUMBER

5c. PROGRAM ELEMENT NUMBER

6. AUTHOR(S)

J. Jonkman, T. Larsen, A. Hansen, T. Nygaard, K. Maus, 

M. Karimirad, Z. Gao, T. Moan, I. Fylling, J. Nichols, M. Kohlmeier, 

J. Pascual Vergara, D. Merino, W. Shi, and H. Park

5d. PROJECT NUMBER

NREL/CP-500-47534

5e. TASK NUMBER

WE101211

5f. WORK UNIT NUMBER

7. PERFORMING ORGANIZATION NAME(S) AND ADDRESS(ES)

National Renewable Energy Laboratory

1617 Cole Blvd.

Golden, CO 80401-3393

8. PERFORMING ORGANIZATION

REPORT NUMBER

NREL/CP-500-47534

9. SPONSORING/MONITORING AGENCY NAME(S) AND ADDRESS(ES) 10. SPONSOR/MONITOR'S ACRONYM(S)

NREL

11. SPONSORING/MONITORING

AGENCY REPORT NUMBER

12. DISTRIBUTION AVAILABILITY STATEMENT

National Technical Information Service

U.S. Department of Commerce

5285 Port Royal Road

Springfield, VA 22161

13. SUPPLEMENTARY NOTES

14. ABSTRACT (Maximum 200 Words)

Offshore wind turbines are designed and analyzed using comprehensive simulation codes that account for the 

coupled dynamics of the wind inflow, aerodynamics, elasticity, and controls of the turbine, along with the incident 

waves, sea current, hydrodynamics, and foundation dynamics of the support structure. This paper describes the 

latest findings of the code-to-code verification activities of the Offshore Code Comparison Collaboration, which 

operates under Subtask 2 of the International Energy Agency Wind Task 23. In the latest phase of the project, 

participants used an assortment of codes to model the coupled dynamic response of a 5-MW wind turbine installed 

on a floating spar buoy in 320 m of water. Code predictions were compared from load-case simulations selected to 

test different model features. The comparisons have resulted in a greater understanding of offshore floating wind 

turbine dynamics and modeling techniques, and better knowledge of the validity of various approximations. The 

lessons learned from this exercise have improved the participants’ codes, thus improving the standard of offshore 

wind turbine modeling.

15. SUBJECT TERMS

offshore wind turbine, floating, spar buoy, aero-hydro-servo-elastic analysis, code verification

16. SECURITY CLASSIFICATION OF: 17. LIMITATION 

OF ABSTRACT

UL

18. NUMBER 

OF PAGES

19a. NAME OF RESPONSIBLE PERSON

a. REPORT

Unclassified

b. ABSTRACT

Unclassified

c. THIS PAGE

Unclassified 19b. TELEPHONE NUMBER (Include area code)

Standard Form 298 (Rev. 8/98)

Prescribed by ANSI Std. Z39.18