Conference Paper

NREL/CP-500-44810

January 2009

Offshore Code Comparison 

Collaboration within IEA Wind Annex 

XXIII: Phase III Results Regarding 

Tripod Support Structure Modeling 

J. Nichols and T. Camp

Garrad Hassan & Partners Limited, United Kingdom

J. Jonkman and S. Butterfield

National Renewable Energy Laboratory, United States

T. Larsen and A. Hansen

Risø National Laboratory, Denmark

J. Azcona, A. Martinez, and X. Munduate

National Renewable Energies Center, Spain

F. Vorpahl and S. Kleinhansl

Fraunhofer Center for Wind Energy and Maritime Engineering, 

Germany

M. Kohlmeier, T. Kossel, and C. Böker

Leibniz University of Hannover, Germany

D. Kaufer

University of Stuttgart, Germany

Presented at the 47th Annual Aerospace Sciences Meeting

Orlando, Florida

January 5 – 8, 2009
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1

Offshore Code Comparison Collaboration within IEA Wind 

Annex XXIII: Phase III Results Regarding Tripod Support 

Structure Modeling

J. Nichols and T. Camp

Garrad Hassan & Partners Limited (GH)

United Kingdom, St Vincent’s Works, Silverthorne Lane, Bristol BS2 0QD 

Telephone: +44 (0) 117 972 9900 Fax: +44 (0) 117 972 9901 

E-mail: james.nichols@garradhassan.com

J. Jonkman and S. Butterfield1

 1 Employees of the Alliance for Sustainable Energy, LLC, under Contract No. DE-AC36-08GO28308 with the U.S. Dept. of 

Energy have authored this work. The United States Government retains and the publisher, by accepting the article for publication, 

acknowledges that the United States Government retains a non-exclusive, paid-up, irrevocable, worldwide license to publish or 

reproduce the published form of this work, or allow others to do so, for United States Government purposes.

National Renewable Energy Laboratory (NREL), United States of America

T. Larsen and A. Hansen

Risø National Laboratory for Sustainable Energy, Technical University of Denmark

J. Azcona, A. Martinez, and X. Munduate

National Renewable Energies Center (CENER), Spain

F. Vorpahl and S. Kleinhansl

Fraunhofer Center for Wind Energy and Maritime Engineering (CWMT), Germany

M. Kohlmeier, T. Kossel, and C. Böker 

Leibniz University of Hannover, Germany

and

D. Kaufer

Endowed Chair of Wind Energy, University of Stuttgart, Germany

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 wind turbine, along with the incident waves, sea current, hydrodynamics, and 

foundation dynamics of the support structure. This paper presents an overview and 

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 Annex XXIII. In the latest phase of the project, a variety of project 

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

wind turbine installed on a tripod substructure in 45 m of water. The code predictions from 

a set of load case simulations—each selected to test different features of the models—were 

compared side by side. The comparisons have resulted in a more thorough understanding of 

the modeling techniques and better knowledge of when various approximations are not 

valid. Importantly, the lessons learned from this exercise have been used to improve the 

codes of the participants, hence improving the standard of offshore wind turbine modeling. 
2

I. Introduction

he 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 in water depth, 

soil type, and wind and wave severity. To ensure that offshore wind turbine installations are cost effective, the 

application of a variety of support structure types is required. These types include fixed-bottom monopiles, gravity 

bases, and space-frames—such as tripods, tetrapods, and lattice frames (e.g., “jackets”)—and floating structures. In 

this context, the offshore wind industry faces many new design challenges.

Wind turbines are designed and analyzed using comprehensive simulation tools (i.e., design codes) capable of 

predicting the coupled dynamic response and the extreme and fatigue loads of the system. Land-based 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 fully coupled (integrated) 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 structure1

. The sophistication of these aero-hydro-servo-elastic codes, and the limited data available with 

which to validate them, underscore the need to verify the codes to ensure their accuracy and correctness. The 

Offshore Code Comparison Collaboration (OC3), which operates under Subtask 2 of the International Energy 

Agency (IEA) Wind Annex XXIII, was established to meet this need.

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 (U.S.), Germany, Denmark, 

the United Kingdom (UK), Spain, the Netherlands, Norway, Sweden, and Korea. In this paper, specifically, results 

are presented from participants who come from Risø National Laboratory of Denmark; Garrad Hassan & Partners 

Limited (GH) of the UK; the National Renewable Energies Center (CENER) of Spain; and the Fraunhofer Center 

for Wind Energy and Maritime Engineering (CWMT), the Leibniz University of Hannover (LUH) and the Endowed 

Chair of Wind Energy at the University of Stuttgart (SWE), all from Germany.

T
3

II. Overview of the OC3 Project

A. Phase III in Relation to Other 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.

To encompass the variety of support structures required for cost effectiveness at varying offshore sites, different 

types of support structures (for the same wind turbine) are investigated in separate phases of the OC3 project:

• 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 platform in deep water.

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 multimegawatt turbine that has also been adopted as the reference 

model for the integrated European Union UpWind research program. This wind turbine is a conventional threebladed upwind variable-speed variable blade-pitch-to-feather-controlled turbine. The specifications consist of 

detailed definitions of the rotor aerodynamic properties; blade, drive train, nacelle, and tower structural properties; 

and generator-torque and blade-pitch control system properties, the latter of which was provided to all OC3 

participants in the form of a dynamic link library. Reference 2 lists the specifications of the NREL offshore 5-MW 

baseline wind turbine in detail. The hydrodynamic and elastic properties of the varying offshore support structures 

used in the project are also controlled, and are discussed later. Furthermore, the turbulent full-field wind inflow and 

regular and irregular wave kinematics are model inputs controlled within the OC3 project. Risø generated the 

turbulent wind velocity datasets and GH 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. The first phase introduced the turbine with a rigid foundation. 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.

The second phase introduced the effects of a flexible foundation. Phase III examines the effects of wave loading 

on a more complicated support structure in deeper water. Emphasis within the OC3 project is placed on the 

verification of the offshore support structure dynamics as part of the dynamics of the complete system. This 

emphasis is a feature that distinguishes the OC3 projects from other wind turbine code-to-code verification exercises 

that have been performed in the past. Nevertheless, it was important to test the aerodynamic models separately so 

that modeling differences resulting from the aerodynamics could be identified. This identification is important 

because the aerodynamic models are known to be a routine source of differences in wind turbine code-to-code 

comparisons.
4

The OC3 project started in January 2005 and is scheduled to be completed in spring 2009. Since the start of the 

project, the reference 5-MW wind turbine, including the control system, has been developed; the wind and wave 

datasets have been generated; the simulations and code-to-code comparisons of Phases I, II, and III have been 

completed; and Phase IV has been initiated. A discussion of the wind and wave dataset generation and a description 

of Phases I and II and their results are presented in detail in Ref. 3 and Ref. 4, respectively. This paper describes 

Phase III and discusses its results. Phase IV will be presented in future papers.

B. Simulations To Be Performed for Phase III

In Phase III a set of three load case simulations has been defined for the NREL offshore 5-MW wind turbine2

installed on a tripod substructure with rigid foundations in 45 m of water (Figure 1). The specifications of each load 

case simulation are summarized in Table 1. Additionally, an Eigenanalysis is used to verify the full-system 

structural dynamics. 

The load case identifiers in Table 1 correspond to the identifiers used by the equivalent simulations from Phase I 

(see Ref. 3), which employed a monopile with a rigid foundation model. In Phase III, though, it was not necessary to 

independently test the aerodynamic, hydrodynamic, and aero-servo-elastic models—as was done in Phase I—

because these models were identical between Phases I and III. Fewer combinations of wind and wave conditions 

were also needed to test the tripod substructure models in Phase III. Consequently, the set of simulations from Phase 

III is much smaller than the set used in Phase I; as a result, the load case identifiers are not sequential.

Figure 1. NREL 5MW wind turbine on tripod support structure
5

Load

Case

Flexible Subsystems Wind Conditions Wave Conditions

2.6 None None: air density = 0 NWH: Stream Function (Dean), 

H = 8m, T = 10s

2.7 None None: air density = 0 None: water density = 0

4.3 Substructure, Tower None: air density = 0 NWH: Stream Function (Dean), 

H = 8m, T = 10s

5.1 Substructure, Tower,

Drivetrain, Blades

Steady, uniform, no shear: 

Vhub = 8m/s

NWH: Stream Function (Dean), 

H = 8m, T = 10s

Vhub – hub-height wind speed 

averaged over 10 minutes

H – individual wave height

T – individual wave period

The tripod support structure is shown in Figure 1. The tripod is a good test for the offshore structure modeling 

capabilities of codes because it incorporates a number of features not present in conventional monopile support 

structures:

• There is no overall axial symmetry; especially there is asymmetry between forward and backward and 

between fore-aft motion and side-side motion.

• Different numbers of members connect at various nodes.

• The central member is significantly tapered.

• Members are at varying angles to the vertical.

For each load case simulation, 57 model output sensors were initially analyzed. In addition to the 47 sensors 

analyzed in Phase I for the rotor, drive train, nacelle tower, and environment (again, see Ref. 3), 35 more outputs 

were analyzed in Phase III. These included more tower loads, as well as water particle velocities and accelerations, 

to better understand the differences between the results of the different codes.

Table 1. Load case definition for Phase III
6

C. Review of Phase I and Phase II Results

Before discussing Phase III, it is important to summarize the key findings from Phases I and II because many of 

the modeling differences that led to code-to-code discrepancies in Phase I have followed through into subsequent 

phases. Though the code-to-code comparisons in Phase I agree very well in general, the key reasons for the 

differences that remained were as follows3,4:

• The modal-based codes (FAST, Bladed, and FLEX5) predict slightly different second and higher coupled 

eigenmodes than those predicted by the higher fidelity multibody- and FEM-based codes (ADAMS and

HAWC2). Differences in the dynamic response and energy content are therefore to be expected in the 

higher frequency range.

• The codes that rely on full-field wind that is supplied in polar coordinates (FLEX5) predict smoother 

aerodynamic loads (and thus smaller load deviations and smaller damage equivalent loads) than codes that 

rely on rectangular coordinates (FAST, Bladed, ADAMS, and HAWC2). This follows from the method in 

which the wind datasets were generated. To ensure that all participants used the same wind inflow, the fullfield wind datasets were generated in rectangular coordinates and subsequently interpolated to polar 

coordinates for the codes that needed it. This cause for differences was mitigated as much as possible by 

using a fine spatial resolution (32 × 32 points across the rotor disk).

• The differences between 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 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 (ADAMS and HAWC2) and do not (FAST, Bladed, and FLEX5) have blade-twist degrees of 

freedom (DOFs).

• Differing model discretizations for the aerodynamic and hydrodynamic loads lead to differences between 

the code predictions. This is most apparent in the substructure loads that depend strongly on the 

discretization of hydrodynamic loads near the free surface.

• Even though every effort has been made to standardize model inputs, user error still happens. 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 unexplainable modeling differences.

Phase II was based on Phase I with the addition of three different foundation models. Phase III does not have any 

foundation DOF modeled; therefore, the results from Phase II, when differing from those from Phase I, have less 

bearing on this report. However, there was further evidence of some codes producing more excitation in the second 

tower frequencies when elements pass through the wave surface. The agreement between different codes was better 

when all the DOFs of the turbine were enabled and it was operating. This is because the substructure loads are 

heavily influenced by the aerodynamics and the higher frequency discrepancies are damped out.
7

D. Wider Objectives of the Project

To test the newly developed codes, the main activities of the OC3 project are (1) discussing modeling strategies, 

(2) developing a suite of benchmark models and simulations, (3) running the simulations and processing the 

simulation results, and (4) comparing the results in a side-to-side fashion. But these activities fall under the much 

broader objectives of

• Assessing the accuracy and reliability of results obtained by simulations to establish confidence in the 

predictive capabilities of the codes

• Training new analysts how to run and apply the codes correctly

• Identifying and verifying the capabilities and limitations of implemented theories

• Investigating and 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 deficiencies and needed improvements in existing codes. 

These results are important because the advancement of the offshore wind industry is closely tied to the 

development and accuracy of dynamics models.

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 have been discussed in papers on previous phases of the project3,4.
8

III. Phase III Modeling

Phase III has stretched the abilities of the codes that have been mainly developed to model simple monopile 

towers. HAWC2 models the wind turbine in the multibody representation; ADCoS models the wind turbine as a 

nonlinear finite element system with beam elements used for the support structure; FAST and GH Bladed both use a 

modal representation. In the past couple of years GH Bladed has released a multimember support structure module 

that allows the modes of complicated structures to be modeled. CENER has used the NASTRAN code to model the 

wave interaction with a tripod model and a Guyan technique5 to use these results as input to FAST. Risø has 

modeled the tripod structure with its multibody wind turbine code, HAWC2. CWMT has developed a model of the 

NREL turbine using the nonlinear, finite element code ADCoS, with hydrodynamic forces applied as nodal forces 

derived from ASAS. The University of Hannover has modeled the tripod structure in Ansys and used WaveLoads to 

apply the forces from the sea state. The University of Stuttgart has used a modal based wind turbine code, Flex, 

coupled with Poseidon, a finite element tool.

Overall, the Phase III results have been encouraging. Modal-based and multibody codes have produced similar 

results. There have been many stages of comparison, as modeling complex offshore support structures is a relatively 

new development in wind turbine design. With each comparison, the results have converged.

A. Full-System Eigenanalysis

Figure 2 gives the lowest 13 natural frequencies calculated for the stationary—but fully flexible—offshore wind 

turbine atop a tripod support structure. 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.

0.0000

0.5000

1.0000

1.5000

2.0000

2.5000

3.0000

1st Tower Fore-Aft

1st Tower Side-to-Side

1st Drivetrain Torsion

1st Blade Collective Flap

1st Blade Asymmetric Flapwise Pitch

1st Blade Asymmetric Flapwise Yaw

1st Blade Asymmetric Edgewise Pitch

1st Blade Asymmetric Edgewise Yaw

2nd Tower Fore-Aft

2nd Tower Side-to-Side

2nd Blade Collective Flap

2nd Blade Asymmetric Flapwise Pitch

2nd Blade Asymmetric Flapwise Yaw

CENER FASTNASTRAN Natural Frequency (Hz)

CENER Bladed Natural Frequency (Hz)

CWMT ADCoS Natural Frequency (Hz)

GH Bladed Natural Frequency (Hz)

GH Bladed (Timoshenko) Natural Frequency (Hz)

LUH WaveLoadsANSYS Natural Frequency (Hz)

Risoe HAWC2 Natural Frequency (Hz)

Risoe HAWC2_BE Natural Frequency (Hz)

SWE FLEX5Poseidon Natural Frequency (Hz)

The tripod support structure is stiffer than the monopile used in Phase I, increasing the tower natural frequencies 

by 15%–20%. The increased stiffness of the tower has little effect on the natural frequencies of the blades; however, 

the drive train frequency is increased slightly. The discrepancy in second tower fore-aft modes is partially due to the 

Figure 2. Eigenanalysis of the NREL 5-MW turbine with tripod support structure
9

coupling between rotor and tower modes. In particular, the Bladed version used by CENER is a previous version in 

which there was no coupling term between the tower nodding motion and the rotor out-of-plane motion. The lower 

values of the second blade asymmetric flapwise yaw mode found by Risø with HAWC2 and CWMT with ADCoS 

are due to the inclusion of the tower torsional mode. For this mode, the vertical blade remains stationary, whereas 

the other two blades move out of phase with each other. If the support structure can twist, the effective stiffness of 

this mode is reduced and therefore the frequency decreases.

B. Specific Problems in Phase III that Have Been Resolved

The participants noted that there are no models of wave run-up and run-down or slap and slam loading 

implemented in codes. As far as we are aware, no model is implemented in a code for offshore structures.

One difference that could not be overcome between the multibody codes and modal codes has been the issue of 

modal damping. For the HAWC2, Rayleigh damping has been used and the damping for a mode calculated from a 

sum over the members. It was not possible to reduce the Rayleigh damping coefficient sufficiently to produce the 

same modal damping that was used in Bladed and FAST without causing numerical difficulties.

1. Discretization of wave loads near the sea surface

One of the first problems that was observed in Phase III was the appearance of stepwise jumps in the member 

loads near the free surface of the sea. These were reduced to negligible levels by discretizing the hydrodynamic 

loads more finely near the sea surface. Figure 3 shows the axial force in the upwind leg of the tripod with both a 

coarse and a fine discretization.

-4000.0000

-3500.0000

-3000.0000

-2500.0000

-2000.0000

-1500.0000

-1000.0000

-500.0000

0.0000

0 5 10 15 20 25 30 35

Time [s]

Shear Force [kNm]

Upwind leg axial shear

force (coarse

discretisation)

Upwind leg axial shear

force (fine

discretisation)

Figure 3. Effect of increasing resolution of hydrodynamic loads near the free surface
10

2. Discretization of tapered members

The discretization of the tapered, central element of the tripod had a large effect on the loads. Buoyancy and 

inertial Morison’s forces per unit length both depend on the square of the diameter of the member, so having too 

long a length between members can cause a large error in the total force. Figure 4 shows the fore-aft shear force in 

the central member when the tapered member is divided into 6 sections and the difference when the member is 

divided into 17 sections.

-900.0000

-850.0000

-800.0000

-750.0000

-700.0000

-650.0000

-600.0000

0.0000 5.0000 10.0000 15.0000 20.0000 25.0000 30.0000 35.0000

Time [s]

Shear Force [kN]

Longitudinal

Shear Force

Central Member

Coarsely Divided

Longitudinal

Shear Force

Central Member

Finely Divided

3. Overlap of tripod members

Another large effect is caused by the overlapping of members at joints. Figure 5 shows an example of the 

overlapping region close to the mean sea level in the tripod. For the large diameter members of the tripod 

configuration, significant surface areas and volumes are duplicated, distorting the overall level of wave and 

buoyancy loading. The intersecting members will also have an influence on the mass of the tripod. This later 

influence does not show up in the code to code comparison, however, as all the participants made the assumption 

that the overlapping sections of mass are negligible. 

Figure 4. Effect of increasing number of members making up the central, tapered member

Figure 5. Schematic of overlapping region between two members
11

In all, 18 sections overlap, neglecting the sections in which the overlap is between three members in total. The 

difference between assuming the members intersect at the defined nodes and removing these overlapping sections is 

159 m3

, or 8% of the total volume below sea level. Therefore, it has a significant effect on the buoyancy calculation 

as well as the applied Morison’s equation forces.

4. Shear deflection in tripod members

For a long time it was not possible to get a good match in results between the codes. To find the reason, we 

performed a simplified study of the turbine mounted on the tripod (Section IV-A). The simplest case of this was a 

static analysis with gravity as the only external load. Even for this simple case there were major discrepancies. 

Actually the HAWC2 results were deviating from the rest of the results. The reason for the discrepancies was the 

difference in beam model, which was demonstrated by CWMT using different kinds of beam elements in ANSYS. 

The importance can be seen in Figure 6, where only the static loads from gravity are shown. In HAWC2 a 

Timoshenko beam is used as default, where all other codes use Bernoulli-Euler beams. Because all beams are thin 

and slender, the Bernoulli-Euler approach should be sufficient, but because the beams are attached rigidly in all ends 

of the tripod structure, the shear effect is much more important than originally assumed. To verify these findings, we 

delivered HAWC2 results delivered using both Timoshenko beams and the same beams neglecting the shear effects 

by increasing the form factor by 104

. Also, the Bladed results were delivered using both Timoshenko beams and 

Bernoulli-Euler beams, also showing the same results. In Figure 6 the axial compression force in the vertical center 

member of the tripod is shown. The results are different with a factor of 1.9. However, this is the part of the 

structure where the difference is largest. This difference in solution will also be present for external loads 

transmitted from the turbine to tripod as, for example, aerodynamic loads.

Figure 6. Axial force in the vertical center beam of the tripod. A clear difference can be seen between codes 

using Bernoulli-Euler or Timoshenko beams. The HAWC2, Bladed, and ANSYNS, results are submitted in 

two versions with either Timoshenko beams or Bernoulli-Euler beams.
12

IV. Loading Simulations

Six locations were chosen for output to best capture the influence of the waves on the structure and are shown in 

Figure 7. The first station, 2 m above the mean sea level, illustrates the effects of wave kinematics near the top of the 

wave; the second station is at the intersection of three slanted legs and one central member; the third is a tapered 

member; the fourth and fifth are in the center of the two upwind slanted members; and the final is the load at the 

start of the foundation. 

Six participants contributed time domain simulations to this phase of the project with six different code 

combinations being used. The legend that will be used for the results of these simulations is shown in Figure 8.

Figure 7. Locations for load output

Figure 8. Legend for loads
13

A. Load Case 2.7: Loads in the Structure out of the Water

The tripod structure presents new challenges in wind turbine modeling. For instance, it is not a tree-like structure 

and has to support dynamic loads through axial forces rather than bending moments. Therefore, it was decided to 

compare the loads on the structure caused solely by gravity, without complications caused by buoyancy and wave 

loads, to iron out differences in the structural modeling. 

Because of the difference in results between the two beam definitions (Section 3B4), the results are presented in 

two groups; one for the Euler-Bernoulli formulation and one for the Timoshenko formulation. GH Bladed, SWE 

Flex Poseidon, CWMT ADCoS, LUH Waveloads Ansys, Risø HAWC2_BE and CWMT ANSYS Beam all used the 

Euler-Bernoulli element. Risø HAWC2, CWMT ANSYS Pipe, and GH Bladed (Timoshenko) used the Timoshenko 

element. The overall spread is also shown to illustrate which loads are most affected by the beam formulation. As 

shown in Tables 2 and 3, the shear force and bending moment at the mud line split significantly into two groups, as 

does the axial force in the central member and the lower brace (see Table 4).

CWMT CWMT CWMT GH GH LUH Risoe Risoe SWE

ADCoS ANSYS 

Beam ANSYS Pipe Bladed Bladed (Timoshenko)

WaveLoads 

ANSYS HAWC2 HAWC2_BE FLEX5 Poseidon

1 -7087.10 -7100.70 -7100.70 -7101.00 -7101.00 -7100.69 -7101.87 -7101.87 -7054.62 0.67% 0.02% 0.67%

2 -7223.80 -7237.40 -7237.40 -7237.00 -7237.00 -7237.38 -7238.64 -7238.64 -7191.31 0.65% 0.02% 0.65%

3 432.70 392.00 731.35 422.05 755.80 391.97 736.25 397.38 391.54 10.17% 3.30% 70.49%

4 -3902.70 -3889.00 -4020.00 -3877.73 -4013.30 -3880.68 -3861.51 1.06% 0.17% 4.04%

5 -1107.80 -1127.00 -792.40 -1149.59 -813.63 -1150.95 -1149.84 3.79% 2.64% 34.42%

6 -5028.80 -5027.70 -5027.80 -5028.00 -5028.00 -5027.73 -5031.63 -5031.53 -5013.91 0.35% 0.08% 0.35%

Overall 

Spread

Spread EulerBernoulliSpread 

Timoshenko

Axial Force

Location

CWMT CWMT CWMT GH GH LUH Risoe Risoe SWE

ADCoS ANSYS 

Beam ANSYS Pipe Bladed Bladed (Timoshenko)

WaveLoads 

ANSYS HAWC2 HAWC2_BE FLEX5 Poseidon

1

2

3 43.33 38.24 37.83 38.03 37.76 38.24 37.72 38.13 38.96 13.53% 0.28% 14.49%

4 37.66 35.97 37.00 35.73 36.93 35.77 35.88 5.34% 0.17% 5.31%

5 58.05 59.28 60.02 54.97 56.98 55.04 54.68 8.15% 5.19% 9.36%

6 -2824.30 -2879.60 -2228.90 -2865.00 -2219.00 -2879.58 -2232.41 -2882.17 -2872.23 2.02% 0.60% 24.99%

Overall 

Spread Location

Shear Force

Spread EulerBernoulliSpread 

Timoshenko

CWMT CWMT CWMT GH GH LUH Risoe Risoe SWE

ADCoS ANSYS 

Beam ANSYS Pipe Bladed Bladed (Timoshenko)

WaveLoads 

ANSYS HAWC2 HAWC2_BE FLEX5 Poseidon

1 -1382.00 -1386.60 -1386.60 -1382.00 -1382.00 -1386.65 -1383.07 -1383.07 -1435.49 3.84% 0.33% 3.85%

2 -1382.00 -1386.60 -1386.60 -1382.00 -1382.00 -1386.65 -1383.07 -1383.07 -1435.48 3.84% 0.33% 3.85%

3 -746.29 -814.74 -810.80 -811.86 -811.86 -814.74 -808.41 -812.39 -826.61 9.98% 0.42% 9.96%

4 855.37 870.37 893.03 871.29 894.53 871.67 875.40 2.31% 0.17% 4.47%

5 207.15 232.16 239.50 213.32 226.09 213.39 213.36 11.59% 5.76% 14.66%

6 -9469.80 -8972.90 -7318.20 -8975.00 -7320.00 -8972.86 -7330.73 -8983.10 -8930.95 5.95% 0.17% 25.39%

Overall 

Spread Location

Bending moment

Spread EulerBernoulliSpread 

Timoshenko

Overall, the agreement is good, once the beam element formulation is taken into account. Results are typically 

within a few percentage points of each other, but more when the load is small compared to the loads in surrounding 

members; for example, the axial force at location 3, the central, tapered member.

Table 2. Static bending moment at locations 1–6

Table 3. Static shear force at locations 1–6

Table 4. Static axial force at locations 1–6
14

B. Load Case 2.6: Regular Waves Acting on a Support Structure Modeled as a Rigid Structure

Overall, the results for bending moment (Figure 9) agree well. Above the waterline (locations 1 and 2) there are 

offsets, but the amplitude of the change caused by the passing of the wave is similar. There is good agreement in the 

tapered member (location 3) and in the upwind leg (location 4). By the lower brace and pile (locations 5 and 6), 

there are variations between the codes caused by different methods of accounting for the overlap and for 

implementing the beam model.

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0

5000

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5 10 15

Simulation Time (s)

Bending Moment (kNm)

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Bending Moment (kNm)

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Bending Moment (kNm)

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Bending Moment (kNm)

1 2

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5 6

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Simulation Time (s)

Bending Moment (kNm)

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Bending Moment (kNm)

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Bending Moment (kNm)

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Simulation Time (s)

Bending Moment (kNm)

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5 10 15

Simulation Time (s)

Bending Moment (kNm)

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0

5 10 15

Simulation Time (s)

Bending Moment (kNm)

1 2

3 4

5 6

Figure 9. Fore-aft bending moment at locations 1–6
15

The shear forces agree less well than the bending moments (Figure 10). There are some differences because of 

the implementation of the wave kinematics. Although a standard set was created, some codes were unable to use this 

as input. Therefore, there are some differences caused by the wave models used.

The widest variation is seen in the central member (location 3). There are several complications for the 

calculations of this force: the member is tapered; it is not inside a tree-like structure so there are many paths that the 

loads can be distributed along; and it is at the height at which the wave loads are strongest. Better agreement is seen 

in the two braces and at the mudline.

-100

-50

0

50

100

150

5 10 15

Simulation Time (s)

Shear Force (kN)

-200

-150

-100

-50

0

50

100

150

200

250

300

5 10 15

Simulation Time (s)

Shear Force (kN)

-20

0

20

40

60

80

100

5 10 15

Simulation Time (s)

Shear Force (kN)

-20

-10

0

10

20

30

40

50

5 10 15

Simulation Time (s)

Shear Force (kN)

-60

-40

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0

20

40

60

80

5 10 15

Simulation Time (s)

Shear Force (kN)

-15000

-10000

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0

5000

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5 10 15

Simulation Time (s)

Shear Force (kN)

1 2

3 4

5 6

-100

-50

0

50

100

150

5 10 15

Simulation Time (s)

Shear Force (kN)

-200

-150

-100

-50

0

50

100

150

200

250

300

5 10 15

Simulation Time (s)

Shear Force (kN)

-20

0

20

40

60

80

100

5 10 15

Simulation Time (s)

Shear Force (kN)

-20

-10

0

10

20

30

40

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5 10 15

Simulation Time (s)

Shear Force (kN)

-60

-40

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0

20

40

60

80

5 10 15

Simulation Time (s)

Shear Force (kN)

-15000

-10000

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0

5000

10000

5 10 15

Simulation Time (s)

Shear Force (kN)

1 2

3 4

5 6

Figure 10. Fore-aft shear force at locations 1–6
16

Overall, the agreement is good for the axial forces (Figure 11). One difference that can be seen is between the 

three Bladed simulations. One (in red) has the model strictly defined as in the specification; the others have some 

parts of the members made invisible to waves and buoyancy to estimate the effect that the members do not overlap. 

During the course of the project we found that the area of overlap is a relatively large effect for tripod-like structures 

and should not be ignored.

-7103

-7102

-7102

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-7101

-7100

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-7099

5 10 15

Simulation Time (s)

Axial Force (kN)

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0

5 10 15

Simulation Time (s)

Axial Force (kN)

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0

200

400

600

5 10 15

Simulation Time (s)

Axial Force (kN)

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0

500

1000

5 10 15

Simulation Time (s)

Axial Force (kN)

-6000

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0

5 10 15

Simulation Time (s)

Axial Force (kN)

-7250

-7240

-7230

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-7190

-7180

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5 10 15

Simulation Time (s)

Axial Force (kN)

1 2

3 4

5

6

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5 10 15

Simulation Time (s)

Axial Force (kN)

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0

5 10 15

Simulation Time (s)

Axial Force (kN)

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5 10 15

Simulation Time (s)

Axial Force (kN)

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5 10 15

Simulation Time (s)

Axial Force (kN)

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5 10 15

Simulation Time (s)

Axial Force (kN)

-7250

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-7170

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5 10 15

Simulation Time (s)

Axial Force (kN)

1 2

3 4

5

6

Figure 11. Axial force at locations 1–6
17

In general the design of the tripod will be determined by the maximum stresses for extreme loads and the cycles 

of loads for fatigue loads. A comparison of the absolute maximum and range of load for the different codes is shown 

in Tables 5 and 6. Results are not presented for all codes, as not all of the codes were able to produce full results for 

each simulation. 

Maximum Absolute Bending Moment [kNm]

CENER CENER GH Risoe CWMT LUH

FAST Bladed Bladed HAWC2 ADCoS

WaveLoads

ANSYS

1 1452.57 1465.27 1426.00 1537.84 1423.00 1437.86 7.88%

2 1638.13 1659.97 1621.00 1739.21 1639.00 1638.33 7.14%

3 7975.49 8143.26 7913.00 7888.92 7805.70 7942.55 4.25%

4 2413.70 2349.06 2354.00 2354.08 2367.00 2490.04 5.90%

5 604.25 534.93 545.96 548.24 545.86 553.61 12.48%

6 9829.42 8393.65 8201.00 9362.83 9210.20 8414.37 18.29%

Location Spread

Bending Moment Range [kNm]

CENER CENER GH Risoe CWMT LUH

FAST Bladed Bladed HAWC2 ADCoS

WaveLoads

ANSYS

1 174.18 171.90 173.00 181.20 158.30 180.98 13.22%

2 710.74 712.26 716.04 737.70 684.36 706.22 7.50%

3 14252.62 14549.55 14149.00 13978.36 14081.80 14277.51 4.02%

4 1925.56 1889.33 1877.89 1860.78 1872.07 2042.09 9.49%

5 100.44 78.15 55.86 71.95 76.88 93.28 56.13%

6 16647.57 16746.48 15583.00 15064.66 15297.70 15229.90 10.67%

Location Spread

Table 5. Maximum bending moment at locations 1–6

Table 6. Bending moment range at locations 1–6
18

C. Load Case 4.3: Regular Waves Acting on a Flexible Support Structure (Rigid Rotor and No 

Aerodynamics)

Figure 12 shows that the tower top displacement is relatively similar in character, but that there is a fairly wide 

variation in the phase of the motion between the codes. This could be due to the initial conditions that are used for 

the simulation.

The mean sea level displacement (see Figures 13 and 14) shows much better agreement. In this case, most of the 

motion is dominated by the action of the waves.

-0.035

-0.030

-0.025

-0.020

-0.015

-0.010

-0.005

0.000

0.005

0.010

5 10 15

Simulation Time (s)

Tower Top Displacement (m)

-0.005

-0.004

-0.003

-0.002

-0.001

0.000

0.001

0.002

0.003

0.004

0.005

5 10 15

Simulation Time (s)

Mean Sea Level Displacement (m)

Figure 12. Tower top fore-aft displacement

Figure 13. Mean sea level fore-aft displacement
19

-0.010

-0.008

-0.006

-0.004

-0.002

0.000

0.002

0.004

0.006

0.008

5 10 15

Simulation Time (s)

Mean sea level pitch angle (deg)

To show how important the consideration of dynamics is to the loading on the tower, the differences in

maximum bending moment and bending moment range between load cases 2.6 and 4.3 are presented in Tables 7 and 

8.

Bending Moment Range [kNm]

CENER CENER CWMT GH GH LUH Risoe Risoe

FAST 

NASTRAN Bladed ADCoS Bladed

Bladed 

(Timoshenko)

WaveLoads

ANSYS HAWC2 HAWC2_BE

1 115.12% 87.01% 114.50% 118.02% 173.28% 106.53% 149.46% 123.18%

2 93.36% 78.46% 102.18% 105.86% 151.14% 88.87% 129.51% 98.00%

3 -31.36% 9.17% 12.31% 12.55% 19.04% 11.83% 15.97% -4.33%

4 -1.07% 0.63% 0.96% 0.93% 1.23% 0.85% 1.29% 0.44%

5 0.79% -0.02% 0.05% 0.32% 0.20% -0.26% 0.28% 0.74%

6 0.61% -0.24% 0.23% 0.09% 0.55% 0.19% -1.44% -0.90%

Location

Bending moment range [kNm]

CENER CENER CWMT GH GH LUH Risoe Risoe

FAST 

NASTRAN Bladed ADCoS Bladed

Bladed 

(Timoshenko)

WaveLoads

ANSYS HAWC2 HAWC2_BE

1 1658.89% 1499.32% 2100.49% 1968.61% 2919.08% 1709.17% 3310.23% 1701.81%

2 343.06% 370.64% 498.80% 488.67% 712.80% 435.99% 788.80% 349.34%

3 -50.27% 9.39% 13.63% 13.61% 19.53% 12.28% 5.40% 0.45%

4 -41.46% 1.75% 2.48% 2.45% 3.56% 2.17% 1.18% 0.02%

5 -30.18% -1.70% -2.29% 9.55% 3.89% 2.73% 10.96% 17.41%

6 -40.14% -0.17% -0.26% -0.24% -0.10% -0.10% -1.10% -0.04%

Location

Figure 14. Mean sea level fore-aft pitch angle

Table 7. Increase in maximum bending moment taking dynamics into account

Table 8. Increase in bending moment range taking dynamics into account
20

V. Conclusion

Engineers from six institutions, including universities and private companies, participated in Phase III of the 

OC3 by submitting results. Many more institutions, including those with a background in traditionally offshore 

industries such as oil and gas, as well as those previously involved in onshore wind energy, have been involved in 

discussing and analyzing these results. 

In learning from the previous phases of the project, the modeling of an onshore turbine and an offshore turbine 

mounted on a monopile foundation including various models of the soil interaction, the modeling issues that 

specifically affect the modeling of tripod, and by extension, other space frame substructures, have been exposed. 

These include:

The sensitivity of wave loads and loads on tapered members to the discretization of the structure.

The proportion of the structure that is double counted if the overlapping sections are not analyzed fully.

The shear deflection of the members, which is not an important effect in a monopile, becomes important when 

the relative displacements of different sections of the support structure affect the distribution of load.

By examining these modeling differences, progress has been made in knowledge of the best way to use existing 

codes to produce accurate models as well as in modifying existing codes to include effects that were previously 

thought to be of only minor influence. The most important result of the analysis is that a Bernoulli-Euler beam 

element is not sufficient for modeling tripod structures as the shear effect included in the Timoshenko formulation is

of utmost importance. For some beam members, the difference in load can be up to a factor of 1.9.

In addition, a baseline set of calculations has been formulated against which new offshore wind turbine codes 

can be tested. These calculations can also be used to train engineers and promote a rigorous approach to modeling 

new structures.

A final, fourth phase to the project is underway, focusing on a floating “spar-buoy” structure supporting the same 

5-MW wind turbine as the other three phases. As well as attracting involvement from more parties, this stage of the 

project is expected to stretch the modeling capabilities of existing codes as nonlinear effects caused by the large 

motion of a floating structure become important.
21

References 1 Passon P and Kühn M, “State-of-the-art and Development Needs of Simulation Codes for Offshore Wind Turbines,” 

Copenhagen Offshore Wind 2005 Conference and Expedition Proceedings, 26–28 October 2005, Copenhagen, Denmark [CDROM], Copenhagen, Denmark: Danish Wind Energy Association, October 2005.

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, Golden, CO: National Renewable Energy Laboratory, February 2007 (to be published).

3 Passon P, Kühn M, Butterfield S, Jonkman J, Camp T, and Larsen TJ, “OC3—Benchmark Exercise of Aero-Elastic 

Offshore Wind Turbine Codes,” Journal of Physics: Conference Series, The Second Conference on The Science of Making 

Torque From Wind, Copenhagen, Denmark, 28–31 August 2007, [online journal], Vol. 75, 2007, 012071, URL: 

http://www.iop.org/ EJ/article/1742-6596/75/1/012071/jpconf7_75_012071.pdf? request-id=8kI1Ig5u3BGgUobT2wi7Kg, [cited 

28 August 2007].

4 Jonkman J, Butterfield, S Passon P, Larsen T. Camp T, Nichols J, Azcona J, and Martinez A, “Offshore Code 

Comparison Collaboration within IEA Wind Annex XXIII: Phase II Results Regarding Monopile Foundation Modeling,” 2007 

European Offshore Wind Conference.

5 Guyan RJ, Reduction of Stiffness and Mass Matrices, AIAA Journal, Vol. 3, No. 2, 1965.
F1147-E(10/2008)

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4. TITLE AND SUBTITLE

Offshore Code Comparison Collaboration within IEA Wind Annex 

XXIII: Phase III Results Regarding Tripod Support Structure 

Modeling

5a. CONTRACT NUMBER

DE-AC36-08-GO28308

5b. GRANT NUMBER

5c. PROGRAM ELEMENT NUMBER

6. AUTHOR(S)

Jonkman, J.; Butterfield, S.; Nichols, J.; Camp, T.; Larsen, T.; 

Hansen, A.; Azcona, J.; Martinez, A.; Munduate, X.; Vorpahl, F.; 

Kleinhansl, M,; Kohlmeier, M.; Klossel, T.; Boker, C.

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NREL/CP-500-44810

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WER9.4001

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National Renewable Energy Laboratory

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Golden, CO 80401-3393

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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 wind turbine, along with the 

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

an overview and 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 Annex XXIII. In 

the latest phase of the project, a variety of project participants used an assortment of codes to model the coupled 

dynamic response of a 5-MW wind turbine installed on a tripod substructure in 45 m of water. The code predictions 

from a set of load case simulations—each selected to test different features of the models—were compared side by 

side. The comparisons have resulted in a more thorough understanding of the modeling techniques and better 

knowledge of when various approximations are not valid. Importantly, the lessons learned from this exercise have 

been used to improve the codes of the participants, hence improving the standard of offshore wind turbine modeling.

15. SUBJECT TERMS

wind turbine design codes; offshore wind development; wind energy; wind turbine aerodynamics

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