Comparison of monopile, tripod, suction bucket and gravity base
design for a 6 MW turbine
Michiel B. Zaaijer
Delft University of Technology, Section Wind Energy
Stevinweg 1, 2628 CN Delft, The Netherlands
Tel. +31 15 27 86426, Fax +31 15 27 85347, M.B.Zaaijer@citg.tudelft.nl
Abstract
In the medium term, offshore wind turbines will increase to around 6 MW rated
power. This paper analyses support structures for these turbines in medium water
depths (around 20 m). After a discussion of the design drivers for several concepts,
typical dimensions are obtained, as well as sensitivities of dynamic behaviour.
1
Introduction
Particularly for the offshore market, wind turbines continue to be proposed and built
to ever-greater sizes, with 3 MW to 5 MW set as the next target. Furthermore, after an
increase of demonstration wind farms in sheltered waters, exploitation of exposed
sites in deeper water has set off with the wind farms at Blyth (UK), Horns Rev (DK)
and Samsø (DK). This tendency toward larger sized turbines and hostile locations will
continue, as large-scale implementation of offshore wind energy proceeds. The
availability of benign locations and public acceptance of wind turbines at visible near
shore sites with larger numbers of migrating birds is waning. Since many of the
expensive offshore activities for installation and maintenance are more or less
proportional to the number of turbines, increase of turbine scale will continue to
provide economic benefits for some time to come. The subject of this paper is the
comparison of support structure concepts for these very large turbines. The
investigation is carried out in the framework of DOWEC (Dutch Offshore Wind
Energy Converter), in which a 6 MW turbine is designed for offshore conditions [1].
Numerous recent and older studies have investigated support structures for offshore
wind turbines. The current study continues these investigations in the context of
medium term developments for offshore wind energy. Monopiles and gravity base
structures are used in current offshore wind farms, but monopiles tend to become
extremely wide as turbines grow and gravity base structures are known to experience
very significant heave forces at exposed sites with intermediate water depths. To
overcome scaling problems of monopiles, several planned wind farms have already
proposed to use tripod structures. Due to its expected installation benefits, a full scale
prototype suction bucket foundation has recently been tried in Frederikshavn.
Although not directly copied, the considered concepts have a fair base in the offshore
industry, thus building on existing experience and minimising technical and
operational risks. The current study enables a direct comparison of the properties of
these concepts for the same conditions, unlike many other studies that focus on a
single concept. Furthermore, the fundamental principles of the concepts and the
design drivers are brought together and compared. The physical models that are used
in this study are deliberately simple, but still capture the main features required to
obtain proper insight in the designs.
2
Definitions and scope of work
Support structure, tower and foundation
Throughout this paper ‘support structure’ is used to indicate the entire structure below
the yaw system, including possible sub-seabed constructions. The support structure
consists of a tower and a foundation. The most arguable definition is the boundary
between these two and this has lead to miscommunication on various occasions.
Contactors prefer to divide the two at a structural boundary, such as the flange at the
access platform several meters above sea level, calling the entire sub sea structure the
foundation. In the more abstract context of a concept study this division is
impracticable, as concepts for the submerged structure and the structure in or on the
seabed can often be selected separately. For example, a tripod can be combined with
both piles and suction buckets. Therefore, ‘foundation’ will be reserved to indicate
that part of the structure that is in direct contact with the seabed and for which
geotechnical considerations are the design driver. For piles the foundation ends at the
seabed, whereas gravity bases may include a (ballasted) slab just above the seabed.
Lumping of mass and stiffness properties at this boundary is shown in [2] to lead to
acceptable predictions of dynamic behaviour, whereas lumping to effective values at
the flange above sea level is expected to lead to large errors at higher frequencies. To
acknowledge the differences between the design drivers of the upper and lower parts
of the tower, a distinction is made into a ‘marine segment’ and a ‘top segment’, as
illustrated in Figure 1. Note that not all support structure concepts are conveniently
covered by these definitions. For instance, the early gravity bases in shallow waters
extend above sea level and miss the marine segment of the tower.
Tower - top segment
Transfer of loads from nacelle to marine segment
Blade tip clearance from tower and water
Monopile
Gravity base
Tripod
Tower - marine segment
Transfer of loads through the marine environment
Provide facilities for boat access
Foundation
Transfer of loads to the seabed
Geotechnical design
Suction buckets (‘Nicknames’)
Figure 1 - Definitions of the support structures
Analysed concepts
This paper treats two concepts for the marine segment of the tower and three
foundation concepts. The combinations of these concepts that are explored are shown
in Table 1 and were also illustrated in Figure 1. All concepts are combined with the
same conical top segment that starts at 9 meters above MSL.
Table 1 - Support structure concepts that are analysed in this study
(driven) Pile
Suction bucket Gravity base
X
X
Single column
X
X
Tripod
As mentioned in the introduction, these concepts have a fair base in the offshore
industry. Some other concepts, such as guyed support structures and lattice towers,
have reached an advanced status on the drawing board and a prototype of a suction
bucket has already been built (see for instance [3]). However, after a preliminary
qualitative and intuitive assessment, these concepts were currently left out of
consideration due to time restrictions. Floating concepts are out of the scope of this
study, as their known economic disadvantage will hamper implementation in medium
deep waters in the medium term [4].
Complexity of the models and used tools
Many disciplines are involved in the design of support structures for offshore wind
turbines: mechanics, dynamics, aerodynamics, electro mechanics, hydrodynamics,
geo-mechanics, material science, production and installation technology and design
methodology. Each of these disciplines has developed sophisticated models, either for
(offshore) wind turbines or for other applications. Some of these models are
implemented in simulation tools for offshore wind turbines, but these are particularly
directed at monopile and conical tower support structures. Furthermore, simulation
tools are rigorous, time-consuming and require detailed user input, which doesn’t
match the exploratory nature of this investigation. Therefore, new tools are developed
that try to capture the essence of the sophisticated models, but at the same time are
easily implemented and provide insight in the issues from a designer’s perspective
[5]. Comprehension of important design drivers and feasibility of concepts are
considered more important than obtaining fully optimised and approved solutions.
The reader must keep this in mind when the actual design results are presented.
The effect of interaction between the support structure and the turbine on dynamics
and fatigue loading is ignored. For a final design this would necessitate undesirable
conservatism in the safety factors and would consequently be unacceptable from an
economic point of view. Even for this exploratory study it is arguable that this
approach may fail to reveal important differences between different concepts, as
fatigue due to combined wind and wave loading is known to be an important aspect.
However, incorporation of a proper model for dynamic, combined (lifetime) loading
increases complexity of the design tools with an order of magnitude. The dynamic
response is implemented with a simple gust response factor of 1.5 for aerodynamic
loading and no dynamic amplification is applied to hydrodynamic loading.
The loading conditions during the lifetime of an offshore wind turbine are diverse and
depend on the instantaneous conditions of wind, wave and operational status and on
dynamic response. Literature describes a large variety of possible and realistic
extreme conditions, from which the following three load cases were selected from [6]:
•
•
•
E.2.1: Stand-by condition in a 50-year extreme gust and reduced wave
E.2.2: Stand-by condition in a 50-year extreme wave and reduced gust
S.1.3: Production with failed pitch control in extreme gust at rated wind speed
This selection of load cases covers extreme wind conditions at stand still and during
operation and extreme wave conditions. Therefore, it is expected that the order of
magnitude of the various types of loading is correctly represented by the load cases.
3
Description of the design conditions
Turbine
This study focuses on support structures for multi-Megawatt wind turbines in medium
deep waters. To obtain typical dimensions and to establish typical numerical values
for design drivers, a design case is specified. The design case is formulated in the
framework of the DOWEC project and comprises a 6 MW turbine. The concept of the
wind turbine is similar to the popular concept for onshore turbines of the MW+ class.
The rotor consists of three blades that are equipped with pitch regulation. Variable
speed operation is obtained with a doubly fed generator with a 30% power inverter.
Using a gearbox with a ratio of approximately 90, rotor speed varies between 7.5 and
14 rpm, with a nominal value of 11.8 rpm at 110% of synchronous generator speed.
The mass of nacelle and rotor equals 272·103 kg.
The aerodynamic design of the blade is based on 5 DU-airfoils up to 42.5 m rotor
span and one NACA-airfoil for the tip section. With a rotor diameter of 129 m,
nominal power is achieved at a rated wind speed of 12 m/s at the hub height of 91.4 m
above Means Sea Level (MSL).
Tower, top segment
The top segment of the tower, extending from 9 m above MSL to the yaw system, is a
tapered cylindrical tower, similar to its land-based counterparts. Although above sea
level, hydrodynamic loading may also cause stresses in this part of the tower, due to
dynamic response of the structure. For soft structures fatigue damage of the top
segment due to hydrodynamic loading may even be in the same order of magnitude as
below sea level [7]. However, since the quasi-static hydrodynamic loading considered
in this study doesn’t affect the structure above sea level, the design of the top segment
is the same for all support structures. Note that in a more detailed design study the top
segment might vary between the concepts due to consideration of dynamic response,
dimensions of the connection between top segment and marine segment, etcetera.
Assuming a diameter to wall thickness ratio of 200, the dimensions of the top segment
are found as given in Table 2. These dimensions provide sufficient strength and safety
margin for the three considered load cases. The diameter and wall thickness vary with
steps at small height intervals, which evidently reflects that manufacturing constraints
are not yet considered. The dimensions near the top reduce more than realistic,
because moments on the yaw system are not modelled.
Table 2 - Dimensions of the top segment of the tower for all concepts
Height above seabed
Diameter (m) Wall thickness (m)
30 (flange 9 m above MSL)
6.1
0.031
40
5.9
0.029
50
5.6
0.028
60
5.2
0.026
70
4.9
0.024
80
4.5
0.022
90
3.9
0.020
100
3.3
0.017
110 (tower top)
2.1
0.010
Tower mass above flange (kg) 226·103
Site conditions
The reference location is in the North Sea, 50 kilometres off the South coast of The
Netherlands. Within the area selected for the wind farm the surface elevation of the
seabed varies due to sand waves. The minimum water depth in the area of 21 m
(MSL) is selected for the design case. Soil conditions are based on geological maps
and borehole descriptions of TNO-NITG. The sand waves overlay 5 m firm clay,
which starts at 36 m below MSL. The next 35 m consist of fine to medium dense
sand, followed by clay with sand intercalations.
Wind and wave conditions are established from the NESS/NEXT database, which
contains hindcast data for over thirty winter seasons and nine summer seasons. The
hourly mean wind speed and extreme significant wave height are translated to
extreme and reduced gusts and wave heights, respectively, using the guidelines in [6]
and assuming a power law atmospheric boundary layer profile with exponent 0.082.
The relevant parameters that were determined are given in Table 3.
Table 3 - Wind and wave parameters for the design case
Parameter
Value
(wind speeds at 10 m above MSL)
Extremes with 1 year return period:
Extreme hourly mean wind speed (m/s)
18.8
Extreme/reduced gust (m/s)
24.8 / 22.6
Extreme significant wave height (m)
4.30
Extreme/reduced wave height (m)
8.0 / 5.7
Extremes with 50 year return period:
Extreme hourly mean wind speed (m/s)
27.3
Extreme/reduced gust (m/s)
36.0 / 32.8
Extreme significant wave height (m)
6.35
Extreme/reduced wave height (m)
11.8 / 8.4
Operating conditions:
Wave height at rated wind speed of 12 m/s at hub height (m)
2.1
4
Design principles and design drivers
Hydrodynamic loading
The calculation of hydrodynamic loading is usually performed in two stages: first, the
water particle kinematics is determined and second the pressure and drag forces are
calculated. A general indication of the applicability of models for water particle
kinematics and load calculations for wind turbines is provided by [8]. Small waves
can usually be treated with linear wave theory, whereas the higher waves in the
relatively shallow waters where wind turbines will be installed need to be treated with
non-linear theories, such as Stokes’ model or Dean’s stream function. Particularly at
spits and sand banks, a popular place for wind turbines, breaking waves with very
non-linear kinematics can occur. However, most common extreme stresses due to
wave loading of wind turbine towers will occur with highly non-linear, but nonbreaking waves [9]. In the current study linear wave theory is used nevertheless, as
this is much easier to implement. In general, this will lead to an underestimation of
the extreme forces for the given wave height.
For slender structures, such as monopiles and members of the tripod, the presence of
the structure can be ignored when calculating water particle kinematics. The most
commonly applied model to calculate the loads, Morison’s equation, compensates for
the influence of the structure through its inertia coefficient. More compact structures,
such as gravity bases, influence the movement of the water more significantly, which
would require diffraction theory and integration of the local surface forces. According
to [8], an adaptation of the inertia coefficient in Morison’s equation gives acceptable
results without diffraction theory. In this study Morison’s equation is used for lateral
loading of all submerged structural elements, with a size-dependent inertia coefficient
for gravity base structures.
For gravity base structures heave force on the horizontal surface is very significant. In
this study heave force is calculated with Bernoulli’s equation for the undisturbed
water kinematics, rather than with more accurate diffraction models. Hydrodynamic
loading complicates the design of a gravity base severely, as it requires a
simultaneous hydrodynamic and geotechnical analysis. The design of a foundation
pile (below the mudline) is much easier separated from the hydrodynamic analysis of
the marine segment of the tower. Heave force on the (smaller) topside of suction
buckets is ignored all together in this investigation, as this short term loading is
expected to be absorbed by the dynamic suction effect that is discussed later.
Soil mechanics
The dominating geotechnical principles are very different for the four foundation
concepts: the laterally loaded monopile, the axially loaded piles of the tripod, the
sealed suction bucket and the gravity base foundation under combined loading.
The lateral loads on the monopile are counteracted by a pressure difference between
both sites of the pile that is initiated by a displacement of the pile. For small
displacements the pressure increases linearly, but at larger displacements plastic
deformations cause the pressure to level off. The design of the laterally loaded pile is
based on Blum’s assumption, which considers a pile in perfectly plastic material with
an effective clamping depth. When a linearly increasing effective vertical soil
pressure is assumed the clamping depth can be expressed analytically in soil and pile
parameters [10]. The plastic behaviour of the soil gives a good engineering
representation of the limit state. The model corresponds reasonably well to finite
element analysis of pile deflections for actual pile penetrations exceeding the
clamping depth by 30% [11]. Thus, a pile design length of 1.3 times Blum’s clamping
depth is assumed to give negligible toe-kick.
The bearing capacity of an axially loaded pile comprises shaft friction and pile point
resistance. In case of hollow piles the soil inside the pile contributes to the bearing
capacity by friction with the inner wall, or by point resistance of the soil plug at the
pile tip, whichever is less. When a linearly increasing effective soil pressure is
assumed the shaft friction according to Coulomb’s relation for frictional material, the
pile point resistance following the theory of Prandtl, Terzaghi and Brinch Hansen and
the required bearing capacity give a quadratic expression for the pile length.
The type of suction bucket considered here is a cylinder with a cap that is sealed after
installation. Since no active suction is applied after installation of the bucket, the
geotechnical principles of axially loaded piles also apply here: skin friction and point
resistance. The point resistance of the pile tip is typically negligible, but the cap of the
bucket causes very significant bearing capacity by pressing on the soil plug inside the
bucket. Under tensile loads, the vertical displacement of the cap will result in a
pressure reduction below the cap. During longer tensile loading the suction area
below the cap will be drained and only skin friction will remain. Typical time-scales
for this process differ very much for different bucket dimensions and different soil
types [12]. In this study it has been assumed that the time scales for waves and wind
gusts are smaller than the drainage times. Furthermore, it has been assumed that the
suction force is always sufficient to withstand these dynamic loads and that the
relatively large static wind loading on the rotor and tower dominate the geotechnical
design. As a consequence, only a static resistance model is implemented. The skin
friction and end resistance are integrated numerically.
The gravity base must provide sufficient resistance against sliding and sufficient
vertical bearing capacity. The required sliding resistance determines the minimum
weight of the system, based on Coulomb’s relation for frictional material. Vertical
bearing capacity is calculated according to the theory developed by Prandtl, Terzaghi
and Brinch Hansen. Only the contribution of the soil weight is taken into account,
because this is commonly the largest contribution to bearing capacity for an offshore
GBS. Correction coefficients for inclined loading and overturning moment are
included, since these affect the design to a large extend. The bearing capacity of the
gravity base is checked for many phases of the incoming wave, due to its sensitivity to
the ratio between vertical and horizontal loading [13].
Installation
Tripod piles are of similar type and size of conventional offshore structures and the
installation of these is currently routine. Monopiles for 6 MW wind turbines have
much larger diameters than current piles, which may cause practical problems, such as
the lack of sufficiently heavy hammers. However, no physical restrictions are
expected for 5-6 m piles. The ratio of diameter to wall thickness of the tripod piles
and monopile are fixed at 60 and 100, respectively, as a preliminary criterion to avoid
buckling during pile driving. This is slightly more optimistic than the guidelines in
[14] but follows the same general increasing trend with increasing pile diameter.
For suction buckets the installation process is a significant factor for the structural
design. The driving force during installation of the suction bucket is the hydrostatic
pressure difference over the cap and the deadweight of suction bucket, ballast and preassembled parts of the tower. It has been assumed that the pressure inside the suction
bucket can be reduced to zero, although in reality this may cause liquefaction of the
soil at the point of critical suction [15]. The assessment of installation is based on the
procedure outlined for skirted foundations in [16]. A “highest expected” skin friction
and end resistance are used to determine the resistance during installation, while
“most probable” values are used in the calculation of bearing capacity. Effects of pore
pressures on skin friction and tip resistance are not considered, although an indication
of this simplification is given in Chapter 5.
For gravity bases the installation process may result in requirements for the structural
design for practical reasons, such as the capacity of the installation vessel or the size
of the workspace. As these are not specified, no requirements are used in this study.
Dynamics
The dynamic behaviour of the support structure is an important design driver for
offshore wind turbines [17]. The foremost criterion is avoidance of resonance at wave
excitation frequencies, the rotor frequency and blade passing frequencies. In this
study the natural frequency of some design results are determined with Rayleigh’s
method or a finite element model, to get insight in its value and variability. As
turbines get higher, the natural frequency of monopiles comes down into the highenergy part of the wave spectrum. It is expected that the stiffer tripod suffers less from
wave-resonance and provides more opportunities to tune the natural frequency. Only
the stiffness of pile foundations is modelled, using springs that represent forcedisplacement relations at regular intervals along the pile.
Scour
Due to changed currents around the structures, erosion of the seabed will occur. Due
to the scour hole that originates from this process the soil supporting the foundation
starts at a lower level and the overburden pressure on deeper layers reduces. As a
consequence, bearing capacity and resistance of the foundation reduces and the
natural frequency drops. Initially, protection of the seabed against scour is assumed,
for instance by rock dumping. For the monopile and piled tripod the effect of
omission of scour protection is looked at. For the gravity base and suction buckets
scour is expected to be unacceptable, due to their high reliance on near-surface soil.
On larger scale natural sediment displacements may result in rise or drop of the entire
seabed around the structure. The magnitude is independent of size of the construction
and can be several meters at some North Sea sites. These sites are particularly
unsuitable for gravity bases, suction buckets and to some extent piled tripods.
5
Design space and typical results
Monopile
The design freedom of the monopile is very limited. The main parameter that can be
influenced is the ratio between diameter and wall thickness. In this study a fixed ratio
of 200 is taken at all heights above the seabed and a fixed ratio of 100 for the
foundation pile. Increase of this ratio results in a lighter construction, but buckling
risk imposes a limit. Land-based towers for wind turbines are built with a ratio up to
300, but local impacts of waves, ships and flotsam and jetsam increase buckling risks.
Table 4 gives the dimensions of the monopile for these more or less optimum ratios.
Table 4 Dimensions of the monopile
Mass (kg)
158·103
Marine segment
Height above seabed
Diameter (m) Wall thickness (m)
30 (= 9 m above MSL) 6.1
0.030
20
6.3
0.032
10
6.6
0.033
0
6.9
0.035
Mass (kg)
199·103
Foundation pile
Penetration depth (m) Diameter (m) Wall thickness (m)
26
5.6
0.056
Mass (tonnes)
500
0.40
Foundation+marine segment
400
Natural frequency
300
Marine segment
Foundation
200
100
100
120
0.35
140
160
60
180
80
Natural frequency (Hz)
Variation of the ratio between diameter and wall thickness is also a means to adapt the
natural frequency of the monopile support structure. The effectiveness of this means
is shown in Figure 2, together with the effect on structural mass. In Figure 2 the ratio
between diameter and wall thickness of foundation and marine segment are coupled.
200
Tower - marine segment
0.30
100
Foundation pile
Ratio diameter/wall thickness (-)
Figure 2 - Variation of monopile mass and natural frequency with diameter to wall
thickness ratio (natural frequency determined with Rayleigh’s method)
When no scour protection is applied, a scour hole of approximately 1.5 times the pile
diameter is expected. The foundation pile has to be adapted to this hole by increasing
diameter and wall thickness with 7% each and increasing pile penetration with 4.8 m,
resulting in an increase of mass by approximately 35%. Additionally, a means to cross
the scour hole with the electricity cable has to be provided. The marine segment can
remain more or less the same, lest manufacture and installation allows. The effect of
scour on the natural frequency is shown in Figure 6 below.
Gravity base structure
The marine segment of this concept is nearly equivalent to that of the monopile, being
cut-off at the GBS top surface. Considerations of manufacturing, installation and
dynamic behaviour may result in differences, but the basic design principles are the
same. Therefore, this section focuses on the foundation of this concept. The weight of
a gravity base has to be sufficient to avoid uplift, tilting and sliding, while at the same
time avoiding failure of the subsoil. The main parameters to achieve this balance are
the diameter and height of the gravity base. Figure 3 shows which gravity bases
provide a stable foundation and where boundaries of failure mechanisms occur. The
correction factors on bearing capacity reduce to zero at severely inclined loads or
tilting, making the boundary for bearing an envelop for all failure mechanisms.
GBS height (m)
10
Stability boundaries
Instable
Bearing
Stable
Sliding
5
Tilting
Bearing (no corrections)
0
0
10
20
30
40
50
Lifting
GBS diameter (m)
Figure 3 - Stability boundaries for gravity base structure
Overturning moment
1E+8
Horizontal
0E+0
Vertical
-1E+8
0.0
0.5
1
3E+8
Capacity
0E+0
0.0
1.0
0
0.5 -Vertical load 1.0
Wave phase (-)
Correction and utilisation (-)
Effective area correction
Inclination correction
6E+8
Utilisation
2E+8
Force and capacity (N)
External forces (N), (Nm)
The minimum (dry) mass of a stable gravity base of 4100·103 kg is obtained for a
GBS with 22 m diameter and 2.7 m height. Curious enough, this is nearly the same
mass of a gravity base for a 3 MW turbine, studied in [13] in a slightly more benign
environment. Apparently, heave force on the gravity base itself is more dominant than
the large overturning moment due to aerodynamic loads. The forces on this gravity
base, including wind and wave loading on turbine and tower, are shown in Figure 4,
along with its utilisation of the vertical bearing capacity. The correction factors for
inclination and effective area indicate the effect of inclined and off-centre loading,
respectively. The bearing capacity for a purely vertically loaded situation is multiplied
with these correction factors to obtain the displayed capacity. The narrow peak of the
utilisation demonstrates that bearing capacity needs to be checked at small intervals.
Wave phase (-)
Figure 4 - Forces and utilisations of gravity base (ø = 22 m, height = 2.7 m)
Tripod
The main design parameters for the tripod are the height of the joint and the base
radius. Only main member-forces are determined, using a statically determinate beam
model. The resulting mass for the marine segment and piles are shown in Figure 5.
3 Piles
10
15
20
25
30
35
45
40
40
40
35
35
35
30
30
30
25
25
25
20
20
20
15
15
15
10
5
40
10
15
60-100
100-140
20
25
30
35
10
5
40
10
15
140-180
200-250
250-300
300-350
20
25
30
35
10
40
Base radius
Base radius
Base radius
20-60
Piles + marine segment
45
Node height
5
Marine segment
45
350-400
250-300
300-350
350-400
400-450
Figure 5 - Variation of tripod mass with joint height and base radius (in 103 kg)
Variation of the ratio between diameter and wall thickness (equal to 100 for the braces
and 50 for the central column in Figure 5) did not have a significant effect. At the
reference site the splash zone is estimated to be between 14 and 29 m above the
mudline. Considering corrosion, maintenance and wave impacts, the splash zone is an
unfavourable location for the tripod joint. Therefore, a joint height of 30 m above
mudline is selected, giving a minimum mass of tripod and piles at a 20 m base radius.
The corresponding dimensions are given in Table 5.
Table 5 - Tripod dimensions
Node height, above seabed (m)
Base radius, from tower to pile (m)
Outer diameter (m)
Braces
Wall thickness (m)
Outer diameter (m)
Base
Wall thickness (m)
Central column Top diameter (m)
Top wall thickness (m)
Base diameter (m)
Base wall thickness (m)
Outer diameter (m)
Piles
Wall thickness (m)
Length below mudline (m)
Tripod mass (marine segment) (kg)
Mass
Pile mass, total of 3 piles (kg)
30
20
Overall
1.52
0.015
0.77
0.015
3.90
0.078
2.00
0.040
1.05
0.017
32.4
216·103
43·103
Figure 6 shows the first and second natural frequency of several design solutions of
Figure 5. These natural frequencies are obtained from FEM analysis and cannot be
compared directly with the monopile results of Figure 2, which were obtained with
Rayleigh’s method. Apparently, the influence of the main design parameters on the
first natural frequency is marginal, but some variation in the second natural frequency
can be achieved. A similar conclusion was found for variation of the tripod topology.
Local scour
45
2.89
0.81
40
35
0.249
1.32
25
0.254
1.39
15
30
20
5
10
15
20 25
30
35
10
40
Base radius
250-300
300-350
350-400
Node height .
0.269
1.05
2.67
0.90
Relative natural frequency (-)
Design variation
0.287
0.92
1.0
Monopile
0.9
Tripod
0.8
1st Mode
0.7
2nd Mode
0.6
0
0.5
0.0
0.4
0
2
2
4
6 *Diameter Tripod (-)
0.5
1.0
1.5 *Diameter Baseline (-)
4
6
Scour depth (m)
8
400-450
Figure 6 - 1st and 2nd natural frequency variation (in Hz)
When no scour protection is applied, the scour hole will be smaller than that of the
monopile, but current around the marine structure is also expected to contribute to
seabed erosion. Assuming a scour hole of 7 m, the pile diameter has to increase with
approximately 35% and its mass is increased with nearly a factor 2. The effect of
scour on the natural frequency of tripod and monopile is also shown in Figure 6.
Suction buckets
The tower of this support structure concept is nearly equivalent to that of the piled
tripod, although the connection of the larger suction buckets may require some
adaptation of the tripod base. In this study the previously discussed tripod is used and
this section focuses on the suction bucket foundation. There are two basic
requirements for the suction buckets: installation has to be possible with the
achievable hydrostatic force and resistance to operational loads has to be sufficient.
The main dimensions that can be varied are the bucket diameter and penetration
depth. Figure 7 shows which buckets can be installed at the reference site and which
buckets provide sufficient resistance to the loads, resulting in feasible buckets for this
design case. In this study the tripod is connected to the three suction buckets during
installation, providing some additional deadweight.
Compression
3
5
7
9
Diam eter
11
13
Installation
All
1
1
1
1
6
6
6
6
11
11
11
11
16
16
16
16
21
15
1
3
5
7
9
Diam eter
11
13
21
15
1
3
5
7
9
Diam eter
11
13
21
15
Penetration depth
1
Tension
21
1
3
5
7
9
11
13
15
Diam eter
Figure 7 - Feasibility of suction buckets (dark area meets criterion)
The lightest bucket in Figure 7 has a diameter of 8 m and a height of 8 m. With
average wall thicknesses of 20 mm and 40 mm for rim and cap, respectively, this
results in a mass of nearly 50·103 kg per bucket. (These wall thicknesses are not
thoroughly engineered). Note that critical suction is not analysed, resulting in a
possibly larger required diameter after more detailed study of the installation.
The results of some alternative studies are presented in Figure 8. Drainage currents
are known to reduce the resistance during installation. This was initially ignored, but
Figure A shows the results of a very optimistic assumption of reduced resistance.
Even though the resistance is reduced by nearly 50%, the effect on the lightest
feasible suction buckets on the left-hand side is very small, due to the dominance of
the operational requirements.
Previous studies indicated that suction buckets are more economic in clays than in
sandy soils [18]. In clays the margin between “most probable” and “highest expected”
resistance is smaller, resulting in a smaller margin between the expected undesired
installation resistance and desired operational resistance. Furthermore, the skin
friction in clay is already fully available at small penetration depths, whereas skin
friction in sand builds up with penetration. Figure B shows the feasibility in case the
5 m thick clay layer is shifted upward to the mudline.
Figure C shows the effect of reduction of the vertical (non-gravity) loads by 50%, due
to an increase in tripod base radius. The lightest suction bucket at ø 6 m and 7 m
height is 30% lighter than the initial design. Thus, the optimum tripod base radius for
a suction bucket foundation will be larger than for driven piles.
The extra deadweight of a preassembled turbine and top segment of the tower can
facilitate the installation of the suction buckets. The effect, as shown in Figure D, is
marginal, as the structural weights are marginal compared with the hydrostatic forces.
Installation benefit
3
5
7
9
Diam eter
11
13
Wider tripod
Extra deadweight
1
1
1
6
6
6
6
11
11
11
11
16
16
16
16
21
15
1
3
5
7
9
11
13
21
15
1
3
Diam eter
5
7
9
Diam eter
11
13
21
15
1
3
5
7
9
11
13
Penetration depth
1
Clay toplayer
1
21
15
Diam eter
Figure 8 - Feasibility of suction buckets for alternative situations
6
Conclusions and recommendations
As a result of the different models for the design of each of the support structures, the
uncertainties partly correspond and are partly independent. The last type of
uncertainty evidently hinders a comparison of the concepts. A thorough analysis of
the uncertainty would require more detailed design of each structure, with more
precise models, but that is outside the scope of this study. Therefore, the conclusions
drawn from the presented results must be put into perspective of the applied simple
models. Likewise, it is emphasised that the conclusions are based on a particular
design case and are not necessarily valid for all conditions.
Table 6 summarises the structural masses found in this design study. The piled tripod
is the lightest structure, due to the light foundation piles. This conclusion still holds
when no scour protection is applied. However, fabrication of the tripod is likely to be
more costly than the monopile and requires more space. Optimisation of the
manufacturing process is therefore a beneficial and essential task for tripods. The
costs and duration of the installation, as well as the availability of equipment will play
an important role in the selection between monopiles and piled tripods.
Table 6 - Structural masses for foundation and marine segment (in 103 kg)
(Top segment
226)
(Mono)
(Tripod)
Suction
Gravity
pile
piles (3)
buckets (3) base
199
43
150
4100
158
357
4258
Single column
216
259
366
Tripod
As known and shown, the installation requirements of the suction bucket tend to result
in wide, shallow foundations. Because the pile-soil friction for cohensionless soils
increases with penetration depth, suction buckets require a larger surface than driven
piles. This is only partly compensated by the smaller wall thicknesses that are
required for the benign installation procedures of suction buckets. Therefore, the
tripod with suction buckets is the heaviest all-steel construction. Although slightly
different conditions may render this structure lighter than the monopile, fabrication
will be even more costly and spacious than that of the piled tripod. The largest benefit
of this concept is expected from its suitability to install the support structure with a
preassembled turbine.
The gravity base structure is more difficult to compare with the other structures, due
to the very different type of material and manufacturing process. Considering that
manufacturing costs per kg of steel structures and reinforced concrete structures differ
in the order of 30, the gravity base may be competing with the other structures. This
reversal of the trend to steel structures seen in current practice would be caused by the
smaller increase of gravity bases with turbine scale. However, it must be noted that
gravity base designs are very case and site specific and this conclusion is not generic.
Furthermore, installation, seabed preparation, scour-protection and non-technical
issues also differ significantly from those of the steel structures.
Adaptation of the first natural frequency in the design phase has a smaller impact on
structural weight for the tripod than for the monopile, but the range of variation is
smaller than expected. The second natural frequency of the tripod can be more easily
adapted. The counter effect is a larger sensitivity of the tripod’s second natural
frequency to scour. A more detailed analysis of natural frequencies that is not
presented in this paper showed that the first natural frequency of the tripod was only
10% higher than that of the monopile, while the second natural frequency was 10%
lower. Therefore, the extra stiffness of the tripod has a relatively small effect on
resonance frequencies. However, it may contribute to a smaller crosstalk of
hydrodynamic (fatigue) loading to the top segment (see [19]).
Acknowledgement
The work presented in this paper is partly performed within the DOWEC project,
which is subsidised by the Dutch Ministry of Economic Affairs through the EET
programme. The contributions of the partners in the DOWEC project are gratefully
acknowledged.
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