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Numerical Analysis of Bearing Capacity
of Suction Bucket Foundation for
Offshore Wind Turbines
Yun-gang ZhanSchool of Naval Architecture and Ocean Engineering,
Jiangsu University of
Science and Technology, Zhenjiang, Jiangsu, China
Email: [email protected]
Fu-chen LiuShandong Vocational Polytechnic College of Water
Resources
Rizhao, Shandong, China
Email: [email protected]
ABSTRACTSuction bucket foundations have been widely used in oil
and gas offshore structures.
Recently, it is attempted to be used as foundation for offshore
wind turbines. Though theloadings transferred to suction bucket are
all in combined mode for these two types of
applications, the eccentric lateral load, other than vertical
forces, has a dominating role of the bearing capacity of
suction bucket for offshore wind turbines. In this paper the
results of
numerical study on the bearing capacity of monopod suction
bucket installed in homogeneousclayey soil to support wind turbine
structures are presented, considering the frictional contact
behavior of interface between skirt and subsoil. Here
bucket aspect ratios are taken from 0.25
to 1.0 and eccentricity ratios of horizontal loads vary in the
range of 0.0~2.0. The bearing
capacity behavior of the bucket under pure vertical, lateral,
torsional loads was investigated
through displacement controlled method first, followed by the
interaction of these loads witheach other by using
load-displacement controlled analysis method. The interaction of
various
combinations of loading is presented in the form of failure
locus. It is shown that, the vertical
capacity from FE analysis agrees well with that of modified
conventional method, andhorizontal capacity factor decreases with
aspect ratio increasing. The torsional load has
significant effect on the vertical capacity but has trivial
effect on the horizontal capacity. The
normalized failure loci for suction bucket with different
eccentricity ratio can be fitted byelliptic curve and the fitting
curve is presented.
KEYWORDS: suction bucket; finite element; bearing capacity;
combined loading
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INTRODUCTION
In current major offshore wind turbine projects, two types of
foundations have been dominating,
the gravity base foundation and the monopile, which have
commonly been used as foundation foroil and gas offshore platform.
In contrast to typical oil and gas structures used offshore, for
a
wind turbine the foundation may account for up to 35% of the
installed cost (Byrne, Houlsby,
2003). Recent research and development projects (Byrne 2000;
Feld 2001; Houlsby et al ., 2005)
have shown that, suction bucket may be used as offshore wind
turbine foundations for itseconomic advantage and repetitive
usability. In addition, in-situ installation of the bucket
foundation is relatively convenient by active suction
installation method. Suction bucket
foundation is a hollow and large diameter cylinder, open at
bottom and closed at top as illustrated
in Fig.1. The suction bucket is installed initially to the
seabed by its self-weight to provide a seal
between the skirt tip and the soil. Then further
penetration is achieved by pumping out the water
within bucket through an opening in the top lid of the bucket,
thus developing differential
pressure between inside and outside of bucket forces it
downwards into desired depth of soil
(Ehlers et al ., 2004). Suction bucket can easily be
removed by pumping water back into the bucket cavity, forcing
it out of the seabed. Consequently, the suction bucket foundation
can be
used repetitively.
Figure 1: Suction bucket foundation
Suction bucket have previously been used as anchors to moor
large floating systems in deep
waters and foundations to support fixed offshore platforms in
coastal areas. For using as anchors,
the suction bucket is mainly subject to vertical (V ) and
horizontal ( H ) loads. However, by
comparison with offshore oil and gas platform, the vertical load
applied to wind turbines
foundations is relatively small, the horizontal load and
overturning moment (M) are substantial
compared with the vertical load (Byrne; Houlsby, 2003). A
typical assumption is that the
combined loads (V - H - M ) applied to
the structure are in-plane. Actually, in the case of wind
turbines the wind and wave directions may not be collinear.
Therefore, torsion moments (T ) about
the vertical axis of the structure might be applied to the
foundations along with overturning
moment, vertical and horizontal loads (Bienen, et al .,
2005). The combined loads
(V - H - M -T ),
which are not in co-planar, must be economically and safely
transferred to the sea ground.
The behavior of suction bucket foundations under vertical and
horizontal loads has been the
subject of various studies. Ei-Gharbawy (1998) conducted a
series of laboratory tests to study the
behavior and pullout capacity of suction bucket
foundations under vertical and inclined loading
H
V
T
L1
L
DSuction bucket
Mudline
Waves
H
V
T
L1
L
DSuction bucket
Mudline
Waves
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conditions. Sukumaran and McCarron (1999) documented an
application of the finite element
method to estimate the capacity of suction bucket foundations
installed in soft clays and subjected
to axial and lateral loads under undrained conditions. Aubeny et
al . (2003a 2003b) presented
upper bound solutions to estimate lateral load capacity of
suction bucket anchors as a function of
the load attachment point location and load inclination angle.
The effects of overturning moment
on bearing capacity of suction bucket have been studied using
finite element method by Wang
and Jin(2008), Bransby and Yun (2009). The results shown that
overturning moment induced byeccentric horizontal force reduce the
axial capacity and lateral resistance of suction bucket
foundation. Taiebat and Carter (2005) employed finite element
method to investigate the behaviorof suction bucket foundation
under the combinations of axial, lateral and torsional forces,
assuming the bucket fully bounded to subsoil, where it was shown
that torsional forces reduce the
axial and lateral capacity obviously. In their study, a typical
bucket with aspect ratio ( L/ D) 2 is
used.
In this paper the response of monopod suction bucket installed
in uniform clayey soil to support
wind turbine structures is studied by finite element method,
considering the combinations of
vertical, lateral, overturning and torsional forces
(V - H - M -T ), and the interaction
of these forces is
presented in the form of failure locus. Here bucket aspect
ratios are taken from 0.25 to 1.0, which
is within the range of bucket foundation for fixed platform.
Accordingly, a series of finite elementanalyses have been
conducted, which have been verified with published results and
theoretical
solutions. Based on numerical result, the dependencies of
bearing behavior of bucket foundation
on aspect ratio, loading combination are discussed.
NUMERICAL MODELING
Geometry of FE Models
ABAQUS FE package (ABAQUS, 2005) was used to investigate the
behavior of suction bucketfoundations with aspect ratios,
L/ D = 0.25, 0.5, 0.75 and 1.0. Due to
three-dimensional loading
conditions, a full-cylinder representing the sub-soil and the
bucket was considered. Thediscretized model area had a radius of
five times the bucket diameter. The bottom boundary of the
model was extended five times the bucket diameter below the toe
of the bucket. With these model
dimensions, the calculated behavior of the bucket is not
significantly influenced by the
boundaries. An example of the three-dimensional finite
element model for bucket foundation is
shown in Fig.2. The bucket was modeled with rigid body element
and the soil was modeled with
linear brick elements with reduced integration and hourglass
control (C3D8R). Relatively fine
meshes were employed at the edge of the bucket and immediately
below the bucket toe to capture
localized failure, while coarser meshes were used away from the
bucket to reduce computational
effort. No vertical and horizontal displacements were adopted as
boundary conditions for the base
of soil, as well as horizontal moments prevented on the side
boundary of soil.
Material propertyIn order to calculate undrained failure
conditions, the clayey soil response was taken as elasto-
perfectly plastic using a Tresca yield criterion with
uniformed undrained strength su. Young’s
modulus of the soil was assumed as 500 su and
Poisson’s ratio was taken as 0.49. The submerged
soil unit weight was taken as 36 kN/mγ = , and
the unit weight of bucket was taken as the same as
that of soil to establish the balance of initial stress.
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Figure 2: Finite element model used in analyses
Unlike FE analyses carried out by Monajemi and Abdul Razak
(2009), in which bucket was fully
bonded to surrounding soil, here the contact behavior of
the interface between skirt and soil was
simulated by contact pair algorithm in the ABAQUS. Normal “hard”
contact model was used to
describe the detachment or contact of skirt and soil. When skirt
and soil were in contact, normal
pressure, as well as tangential frictional resistance was
transferred between interfaces,
accompanied by Coulomb’s friction activated. Otherwise, no
forces were transferred between
skirt and soil. The ultimate frictional resistance was set as
soil-undrained strength su. When shear
stress on the contact surface was less than su, both are
stuck together and no slip happened,
whereas slip occurred along the contact surface when shear
stress was up to the ultimate frictional
resistance. The connection between top lid of bucket and soil
was considered as fully bonded to
account for suction developed within bucket to some degree.
Analysis method
The finite element calculations were executed stepwise. At
first, the initial stress state was
generated by applying gravity to the whole model. Subsequently,
the contact conditions between
bucket and soil was activated. Then displacement analyses
were carried out as recommended by
previous researchers (e.g. Bransby, Randolph, 1997; Wang,
Jin, 2008) to determine the pure
ultimate vertical (V ult ), horizontal
( H ult ), overturning ( M ult ) and
torsional (T ult ) loads. For
determining combined loading loci, load-displacement controlled
analyses were performed
beginning with initial stress state, where, for example, a
constant vertical load lower than V ult was
applied to the bucket foundation following by a horizontal
displacement analysis step. This
method has been found to provide more accurate and effective
calculation of combined failure
loads. For the combination of horizontal load and overturning
moment, it was determined by
imposing horizontal displacement at a fixed point on the
foundation with arm lever L1 above the
mud line. The failure loci
of H - M would be depicted by several
analyses with different lever-arm
heights, that is to vary the eccentricity
ratio L1/ D of lateral load.
In contrast to offshore floating system, wind turbines are
relatively sensitive to the displacement
of foundation. The ultimate capacity was defined as a ‘cut-off’
value where either of the
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following criterions was satisfied: the vertical settlement of
reference loading point approaching
0.05 D; or the horizontal tilt 0.02( L+ L1); or
the rotation about vertical axis reaching 0.01 radians.
With these displacement limitations, trouble free operation
could be secured for wind turbines.
BEARING CAPACITY UNDER PURE LOADING
CONDITIONS
Vertical capacity
The conventional vertical bearing capacity for embedded
foundation can be estimated as:
u c s d uV N A sζ ζ = ⋅ ⋅ ⋅ ⋅ (1)
where A is the plane area of the bucket,
N c is the bearing capacity factor of 5.14,
sζ is the shape
factor and 1.2 sζ = is suggested for circular
footing, d ζ is the embedment factor and
1 0.4 arctan( / )d L Dζ = + ⋅ suggested
for circular footing. Deng and Carter (1999) modified the
formula to take into account the adhesion developed on the
bucket skirt. Based on the result of aseries of finite element
analyses, they recommended bearing capacity factor
9.0c N ≈ and
embedment factor 1 0.4( / )d L Dζ
= + . If the skirt adhesion u DLsπ is added to
the capacity
predicted by conventional method, Eq.1 should be modified
by appending an add-in 4 ( / ) u L D As⋅ .
The results of the FE analyses for suction buckets under
vertical loads are presented in Fig.3. It
can be seen that the major part of ultimate resistances has
mobilized at the prescribed vertical
displacement limit. A summary of the failure loads obtained from
FE analyses and that from Eq.
1 and the method suggested by Deng and Carter are given in Tab.
1. The results of FE analyses
are close to the value calculated using modified Eq.1, and less
than that of Deng and Carter’s
method. This is because Deng and Carter’s method was based on
the results of FE analysis with
fully bounded conditions between skirt and subsoil, and the
vertical resistances were determinedat vertical displacement
reached 0.5 D. The cut-off value of this paper is more
practical than that
of Deng and Carter’s method.
Table 1: Vertical failure loads V ult
/ A su
L / D 0.25 0.5 0.75 1.0
Modified Eq.1 7.77 9.31 10.76 12.11
FE analysis 8.51 9.79 10.82 11.86
Deng & Carter method 11.88 12.96 14.04 15.12
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Figure 3: Vertical load-displacement relationship for suction
bucket
Horizontal capacity
The pure horizontal capacity of suction bucket foundations were
obtained from FE analyses by
applying lateral displacement on the central point of bucket at
the mudline level. The capacity
commonly expressed as:
ult h u H N LDs= (2)
Where H ult is the ultimate lateral capacity of
bucket foundation and N h is the lateral
capacity
factor. Fig. 4(a) shows the FE result of N h. It can
be seen that the value of N h decreases
with L/ D
increasing. The trend of this varying of N h is
depicted in Fig. 4b and the formula of the fitting
curve is
212.23 -14.06 / 6.12 ( / )h
N L D L D= + (3)
(a) (b)
Figure 4: Horizontal capacity of suction bucket foundation
0 2 4 6 8 100
2
4
6
8
10
12
14
L / D = 0.25
L / D = 0.5
L / D = 0.75
V e r t i c a l L o a d V u l t / ( A
. s u )
Vertical Displacement / D %
L / D = 1.0
0 1 2 3 40
2
4
6
8
10 L / D = 0.25
L / D = 0.5
L / D = 0.75
H o r i z o n t a l L o a d H u l t / ( L . D . s u )
Horizontal Displacement / L %
L / D = 1.0
0.00 0.25 0.50 0.75 1.00
4
5
6
7
8
9
N h
L / D
FE
Fitting Curve
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The trend of N h varying
with L/ D is different from that of Deng and
Carter’s study, in which the
aspect ratio of bucket, L/ D, has no significant
effect on the lateral capacity factor N h. The
reason
is as above mentioned that they assumed skirt was fully bounded
to subsoil.
Torsional capacity
Based on the assuming that the shear strength of the interface
between skirt and soil is the same
as that of subsoil, the theoretical value for the ultimate
torsional capacity of bucket foundations,
T ult, is expressed as:
2
ult t uT N LD s= (4)
where N t is the torsional capacity factor
of (0.5 /12 / )t N D Lπ = + . Fig.5 shows the
load-deflection
curves predicted by the FE analyses for buckets with different
aspect ratios, as well as the
theoretical solutions. It can be seen that the torsional
resistances were fully mobilized at the
prescribed rotational limitation and are in excellent
agreement with the theoretical value
represented by dash-dot-lines from Eq. 4.
Figure 5: Torsional capacity of suction bucket foundation
COMBINED LOADING CAPACITY
For suction bucket foundations, the failure loci of combined
loading are usually described in V - H ,
V - M , H - M , H -T ,
V - H - M load spaces. While for
wind turbines foundation, the overturning moment
mainly induced by horizontal loads applied eccentrically. The
effect of overturning moment on
the bearing capacity of bucket foundation can be represented by
ultimate horizontal loads applied
with different level above seafloor. On the other hand, the
vertical loads always exist. So the
failure loci of combined loading are depicted in
V - H and
V -T - H load spaces with different
value
of L1/ D, based on large numbers of FE analyses.
Vertical-horizontal capacity
Tab. 2 gives the horizontal failure loads with
L1/ D=0, 0.2, 1.0, and 2.0 under no vertical load
conditions. The horizontal resistances under
V - H and V -T - H
loading conditions can be
normalized using these corresponding capacities.
Table 2: Horizontal failure loads ( H ult
/ L Dsu) with different L1 / D
0 1 2 3 4 50.0
0.5
1.0
1.5
2.0
2.5
3.0
L / D = 0.25
L / D = 0.5 L / D =
0.75
T o r s i o n a l L o a d T u l t / ( L . D
2 . s u )
Rotation Radian %
L / D = 1.0
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L/ D L1/ D
0.0 0.2 1.0 2.0
0.25 9.11 7.59 2.86 1.55
0.5
6.32
4.71
1.89
1.07
0.75 5.12 3.73 1.77 1.07
1.0 4.34 3.36 1.72 1.06
The failure loci obtained from FE analyses are shown in Fig. 6
for bucket foundation under
vertical and horizontal loads. It can be seen that, though the
eccentricity ratios, L1/ D, are different
for bucket foundation with given aspect ratio L/ D,
the normalized loci in V - H loading spaces
are
almost identical. The locus of L1/ D=0 can be chosen
as typical one to describe the interaction of
vertical and horizontal loading, and are shown on Fig. 7. The
overall size of normalized failure
envelope enlarges as aspect ratio L/ D, increases. For
a bucket foundation of given aspect ratio, the
vertical ultimate capacity, V ult and horizontal
ultimate capacity, H ult of different eccentricity
ratio
L1/ D can be obtained from Tab. 1 and Tab. 2,
then the horizontal capacity varying with vertical
load could be determined from Fig.7.
Figure 6: Normalized failure loci for
V - H
0.0 0.2 0.4 0.6 0.8 1.0
0.0
0.2
0.4
0.6
0.8
1.0
H / H u l t
V / V ult
L1 / D = 0.0
L1 / D = 0.2
L1 / D = 1.0
L1 / D = 2.0
L / D = 0.25
0.0 0.2 0.4 0.6 0.8 1.00.0
0.2
0.4
0.6
0.8
1.0
H / H u l t
V / V ult
L1 / D = 0.0
L1 / D = 0.2
L1 / D = 1.0
L1 / D = 2.0
L / D = 0.5
0.0 0.2 0.4 0.6 0.8 1.00.0
0.2
0.4
0.6
0.8
1.0
H / H u l t
V / V ult
L1 / D = 0.0
L1 / D = 0.2
L1 / D = 1.0
L1 / D = 2.0
L / D = 0.75
0.0 0.2 0.4 0.6 0.8 1.00.0
0.2
0.4
0.6
0.8
1.0
L / D = 1.0
H / H u l t
V / V ult
L1 / D = 0.0
L1 / D = 0.2
L1 / D = 1.0
L1 / D = 2.0
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Figure 7: Normalized typical failure loci for
V - H
Vertical-torsional-horizontal capacity
Fig. 8 presents the failure loci for bucket caissons with aspect
ratio ranging from 0.25 to 1.0
subjected to combinations of vertical and torsional loads. It
can be seen that torsional loads havesignificant effects on
vertical capacity when it reaches 0.8T ult. A failure locus
for the bucket of
L/ D = 1.0 subjected to combination of horizontal and
torsional loads is presented in Fig. 9, which
shows that the horizontal resistances are not affected by the
torsional load until torsional loading
limitation state reached. The ultimate horizontal resistances,
H ult for bucket of different
eccentricity ratio L1/ D can be made as
reference values to normalize that for bucket foundation
under V -T - H loading conditions.
Figure 8: Normalized failure loci for V -T
Figure 9: Normalized failure loci for
T - H
An additional set of FE analyses, in which only suction bucket
foundation of aspect ratio 1.0 wasconsidered, were carried out to
investigate the effects of torsional loads on the
V - H capacity.
Representation of the normalized failure loci in
V -T - H loading space is shown in Fig.
10, where
the torsional load varies from 0.2T ult to
0.9T ult and eccentricity ratio varies from 0.0 to 2.0
to
account for the overturning moment. The vertical resistance was
normalized using the vertical
load from Fig. 8 and the horizontal resistance was normalized by
H ult from Fig. 9. As can be seen,
the numerical data for combination of vertical and horizontal
load are well in coincidence for a
given torsional load, though the eccentricity ratio varying. The
failure loci can be approximated
with elliptical curve
0.0 0.2 0.4 0.6 0.8 1.00.0
0.2
0.4
0.6
0.8
1.0
H / H u l t
V / V ult
L / D = 0.25
L / D = 0.50
L / D = 0.75
L / D = 1.00
0.0 0.2 0.4 0.6 0.8 1.00.0
0.2
0.4
0.6
0.8
1.0
L / D = 0.25 L / D =
0.50
L / D = 0.75
L / D =
1.00 V / V u l t
T / T ult
0.0 0.2 0.4 0.6 0.8 1.00
1
2
3
4
5
L1 / D = 2.0
L1 / D = 1.0
L1 / D = 0.2
L1 / D = 0
H / L . D . s u
T / T ult
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1
a b
ult ult
H V
H V
+ =
(5)
where a and b are fitting coefficients and are
presented in Tab. 3. The fitting curves for eachtorsional loading
case are shown in the last frame of Fig. 10, which shows that the
effect of
vertical loading on lateral resistance decreases with the
increasing of torsional load. Namely, the
range of horizontal resistance affected by vertical load is
larger under small torsional load than
that under great torsional load.
Table 3: Fitting coefficients of normalized failure loci
for V -T - H
L/ D = 1T / T ult
0.2 0.4 0.6 0.8 0.9
a 1.79 1.80 1.85 1.92
2.22
b 0.28 0.25 0.19 0.15
0.13
0.0 0.2 0.4 0.6 0.8 1.00.0
0.2
0.4
0.6
0.8
1.0
H / H u l t
V / V ult
L1/ D = 0.0
L1/ D = 0.2
L1/ D = 1.0
L1/ D = 2.0
Fitting Curve
T = 0.2T ult
0.0 0.2 0.4 0.6 0.8 1.00.0
0.2
0.4
0.6
0.8
1.0 T = 0.4T ult
H / H u l t
V / V ult
L1 / D = 0.0
L1 / D = 0.2
L1 / D = 1.0
L1 / D = 2.0
Fitting Curve
0.0 0.2 0.4 0.6 0.8 1.00.0
0.2
0.4
0.6
0.8
1.0T = 0.6T ult
H / H u l t
V / V ult
L1 / D = 0.0
L1 / D = 0.2
L1 / D = 1.0
L1 / D = 2.0
Fitting Curve
0.0 0.2 0.4 0.6 0.8 1.00.0
0.2
0.4
0.6
0.8
1.0T = 0.8T ult
H / H u l t
V / V ult
L1 / D = 0.0
L1 / D = 0.2
L1 / D = 1.0
L1 / D = 2.0
Fitting Curve
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Figure 10: Normalized failure loci for
V -T - H
CONCLUSION
Suction bucket foundation for offshore wind turbines under the
combination of vertical,horizontal, torsional and overturning loads
are investigate by the FE method, considering the
frictional contact of the interface between skirt and subsoil.
The bucket was assumed to beembedded in a homogeneous clayey soil
which deforms under undrained conditions. A
displacement controlled method was used to predict the capacity
under pure loading condition
and a load-displacement controlled method was used to derive the
failure loci under combined
loading conditions. Based on a series of numerical analyses, the
following conclusion can be
drawn:
(1) The vertical capacity from FE analysis agrees well with that
of modified conventional method
considering the adhesion between skirt and subsoil. Horizontal
capacity factor, N h is not a
constant, which decreases
with L/ D increasing.
(2) The torsional load has significant effect on the vertical
capacity, but has trivial effect on the
horizontal capacity until it nearly reaching its ultimate
value.
(3) The normalized failure loci for suction bucket with
different eccentricity ratio are almost the
same in vertical-horizontal load space, if torsional load fixed.
The locus expands with torsional
load increasing and can be represented by elliptic curve.
(4) An individual component of combined loads can be determined
using the failure loci
presented in the paper with other components are given in
advance.
Supported by: Natural Science Foundation of the Jiangsu Higher
Education Institutions of China
(No. 09KJD570002)
REFERENCES1.
Abaqus, (2005). User’s Manual. Version 6.5.
2.
Aubeny, C. P., Han, S. W., and Murff, J. D., (2003a) “Inclined
Load Capacity of
Suction Caissons”, International Journal for Numerical and
Analytical Methods in
Geomechanics, 27, 1235-1254.
0.0 0.2 0.4 0.6 0.8 1.00.0
0.2
0.4
0.6
0.8
1.0T = 0.9T ult
H
/ H u l t
V / V ult
L1 / D = 0.0
L1 / D = 0.2
L1 / D = 1.0
L1 / D = 2.0 Fitting Curve
0.0 0.2 0.4 0.6 0.8 1.00.0
0.2
0.4
0.6
0.8
1.0
H
/ H u l t
V / V ult
T / T ult = 0.2
T / T ult = 0.4
T / T ult = 0.6
T / T ult = 0.8
T / T ult = 0.9
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Vol. 15 [2010], Bund. F 644
3.
Aubeny, C. P., Han, S. W., and Murff, J. D. (2003b) “Refined
Model for Inclined
Load Capacity of Suction Caissons”, Proc. 22nd International
Conference on
Offshore Mechanics and Arctic Engineering, Cancun, Mexico, OMAE,
2003, 37502.
4.
Bienen, B., Byrne, B. W. and Houlsby, G. T. (2005) “Six
degree-of-freedom loading
of a circular flat footing on loose sand”, Experimental data,
Report No OUEL2289/05. Oxford: Department of Engineering Science,
University of Oxford.
5. Bransby, M. F. and Randolph, M. F. (1997) “Shallow
Foundations Subject to
Combined Loadings”, Proc. 9th Int. Conf. on Computer Methods and
Advances in
Geomechanics, Wuhan, 1947-1952.
6. Bransby, M. F., Yun, G. (2009) “the undrained capacity
of skirted strip foundations
under combined loading”, Geotechnique, 59(2):115-125.
7. Byrne, B.W. (2000) “Investigations of Suction Caissons
in Dense Sand”, Ph.D.
Thesis, University of Oxford, Oxford.
8.
Byrne, B. W. and Houlsby, G. T. (2003) “Foundation for Offshore
Wind Turbines”,
Phil. Trans. of the Royal Society of London, Series A 361,
2909-2300.
9. Deng, W. and Carter, J. P. (1999) “Analysis of Suction
Caissons in Uniform Soils
Subjected to Inclined Uplift Loading”, Report No. R798,
Department of CivilEngineering, The University of Sydney,
Australia.
10.
Ehlers, C. J. Young, A. G. and Chen, W. (2004) “Technology
Assessment of
Deepwater Anchors”, Proc. Offshore Technology Conference,
Houston, Texas, OTC,
16840.
11.
Ei-Gharbawy, S. L. (1998) “The Pullout Capacity of Suction
Caisson Foundations”,
PhD Thesis, University of Texas at Austin.
12.
Feld, T. (2001) “Suction Buckets, a New Innovative Foundation
Concept, Applied to
Offshore Wind Turbines” Ph.D. Thesis, Aalborg University
Geotechnical
Engineering Group, Feb.
13.
Houlsby, G. T., Kelly, R. B., Huxtable, J. and Byrne, B. W.
(2005) “Field Trials of
Suction Caissons in Sand for Offshore Wind Turbine Foundations”,
Geotechnique.
55(4): 287–296.14. Monajemi, H. and Abdul Razak, H (2009)
“Finite element modeling of suction
anchors under combined loading” Marine Structures,
22(4):660-669.
15.
Sukumaran, B., and McCarron, W. O., (1999) “Total and Effective
Stress Analysis of
Suction Caissons for Gulf of Mexico Conditions”, Proc. OTRC'99
Conference,
Geotechnical special publication No. 88, 247-260.
16. Taiebat, H. A., Carter, J. P. (2005) “Interaction of
Forces on Caisson in Undrained
Soil”, Proc. 15th Int. Offshore and Polar Engineering
Conference, Seoul, Korea, 625-
323.
17. Wang, D., Jin, X. (2008) “Failure Loci of Suction
Caisson Foundations Under
Combined Loading Conditions”, China Ocean Engineering,
22(3):455-464.
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