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INSTALLATION OF LARGE SUBSEA
PACKAGE
François Pétrié (Océanide), Benjamin Rousse (Océanide),
Christophe Ricbourg (Stat Marine), Bernard Molin (Ecole Centrale Marseille),
Guillaume de Hauteclocque (Bureau Veritas)
ABSTRACT
Deepwater field developments are usually associated with large submarine package to be
installed such as suction anchors, manifolds and other subsea equipment.
Key data for the lowering operation engineering are the slamming, added mass and drag
coefficients of the package. These hydrodynamic values are not well known. They are usually
considered as a constant whereas they are depending on many parameters such as crane
motion amplitude and period, package porosity...
This lack of data may lead to over-design the lifting equipment and/or to over-restrain the
operating environmental conditions. Consequently, this could have a significant impact on the
cost & planning of the operation.
This paper presents the main results of the recent work that has been performed for the joint
industry project “Offshore installation of heavy package”, in 2008 and early 2009. The
objective of this program is to establish more refined hydrodynamic coefficients and, if
needed, to propose a detailed methodology for the engineering of such operation. In
particular, model tests results are compared to DNV recommended practice.
1. INTRODUCTION
The aim of this document is to present the results from the French joint industry project
“Installation of Large Subsea Package”. This research project is leaded by Océanide, in
partnership with Ecole Centrale de Marseille, Stat Marine and Bureau Veritas. It is sponsored
by Total, Technip, Doris Engineering and Saipem.
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After a brief analysis of the large subsea package installation challenges, some of the project
results are presented. First, the selection of the studied cases is briefly introduced. Then,
devices and results from the model tests performed in the offshore tank basin BGO FIRST
located at La Seyne sur mer, France are detailed. Finally, theoretical model for hydrodynamic
coefficients calculation and design recommendations for heavy lifting are proposed.
2. Subsea package installation challenges
During the installation of offshore fields, very large subsea packages are laid down on the
seabed. Some of these packages are over 100 ton. Those operations are completed from
specific vessels equipped with heavy lifting devices.
With the growth of the packages and the very deep water depth of new fields, the engineering
offices are often at the limit for lay down design. This comes from the lack of data on
hydrodynamic coefficients used for the design. Those hydrodynamics data are the added
mass, the radiation and quadratic damping and the slamming load in the splash zone.
This lack of data leads to over-conservatism and so to increase the lifting device capacities or
to reduce drastically the operating environmental conditions. The consequence is an important
operating cost impact and a planning drift.
Packages shapes
Large packages can generally be classed into two main shapes:
• Cylindrical shape: mainly suction anchors. It is a cylinder shape opened at bottom and
partly opened at top with events. The size and the number of the events impact the
hydrodynamic characteristics of the package.
• Parallelepiped shape: number of subsea equipment can be mentioned (manifold, flet,
sled…). They are usually supported by a mudmat for geotechnical stability. Due to its
shape and its size, the mudmat is often the preponderant component of the overall
hydrodynamic loads. It is usually perforated to let the water escape during the seabed
laying.
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Installation phases
Water entry
The water entry is a critical phase for object laying. The package is suspended to the lifting
vessel crane. During the lowering through the free surface and the wave zone, the package
buoyancy and the slamming loads can destabilize the equilibrium. During this phase, cable
slack events have to be avoided.
When the package is immerged, hydrodynamic damping is influenced by the free surface
proximity.
Lowering object in infinite water
The lowering phase is continuous. There is no brutal equilibrium change like for water entry
or seabed laying. In very deep water, the design of this phase shall take into account the
dynamic answer of the system cable+package and check that the natural periods of the system
are away from the wave periods on site.
Lowering close to seabed
The problematic is the same as for the infinite water lowering. The only change is that the
seabed proximity influences the added mass and the hydrodynamic damping of the system.
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3. Studied cases
Context
Installation design of large structures requires knowing the hydrodynamic parameters like the
added mass, the added damping but also the slamming loads. Depending on the installation
phase (see above), these parameters may vary with the motion of the structure, with the waves
parameters, with the distance from the free surface or from the seabed, but also with other
parameters like the “porosity” of the structure and its inclination.
As several studies were conducted in the past to characterize the hydrodynamic behaviour of
suction piles, it was decided at the beginning of the present research project, to focus on the
large parallelepiped shaped structures and especially on the perforated mudmats with a skirt,
as it is usually one of the main components of the hydrodynamic loading.
The parameters of the tests were defined as described hereafter.
Wave conditions:
Three sea-states were considered:
• Hs = 1m, Tp = 4s
• Hs = 1.5m, Tp = 6s
• Hs = 1.5m, Tp = 8s
where Hs is the significant wave height and Tp the peak period.
On installation site, the peak period may be larger, in particular in West Africa. But, for long
wave lengths, motions of the installation vessel are more or less in phase with the swell.
These long wave conditions are generally not critical regarding the crane tip motions and thus
were disregarded.
L = 1.5 x l
l h = l/8
Skirt
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Lowering velocity:
Typical lowering velocity recommended by DNV is 0.5 m/s (crane hook velocity).
Practically, the crane operator adjusts this velocity depending on the observed situation for
each phase. At the splash zone crossing, the lowering may be stopped when the mudmat is
just above a wave crest, and then the crossing is started with the proper velocity, avoiding to
have the mudmat crossing again the free surface but also to have snap loads. Hence, the
lowering velocity may be different from 0.5 m/s, and this value is deemed to be an upper
bound velocity.
During the tests of splash zone crossing, two lowering velocities were used: 0.5m/s and a
lower velocity.
Mudmat motions:
The motions of the mudmats are derived from the wave conditions and from RAOs of typical
installation vessels (see figure below).
Crane tip RAO of vertical displacement
0.00
0.20
0.40
0.60
0.80
1.00
1.20
1.40
1.60
1.80
2.00
4 6 8 10 12 14 16 18 20
Wave period (s)
RA
O a
mpl
itude
(m/m
)
-150
-100
-50
0
50
100
150
RA
O p
hase
(deg
)
Vessel 1 amplitude Vessel 2 amplitude Envelop (210deg heading)
Vessel 1 phase Vessel 2 phase
The motions of the mudmat may depend on the depth, since the stiffness of the system is
linked to the paid-out length of cable. Consequently, the period of the motion may be in the
vicinity of the natural period of the system (cable, mudmat and hydrodynamics) or not.
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Impact velocity:
For the slamming, the impact velocity was estimated through the DNV formula. Two formula
were used and the results analyzed (oldest one provided in the “Rules for Planning and
Executions of Marine operations” dated 1996, and the recent Recommended Practice DNV-
RP-H103).
The highest value obtained is 3 m/s which accounts for several conservatisms and is
consequently considered as an upper bound. Finally, four impact velocities where selected:
0.5, 1, 2 and 3 m/s.
Porosity:
Most of the mudmats made of perforated plates (and not truss) have a rate holes/plate (called
porosity) designed for soil stability purpose. Nevertheless the porosity hydrodynamic impact
is investigated. Three porosity rates are studied: plain mudmat (no porosity), low porosity
(named porosity 1), high porosity (named porosity 2).
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4. Model tests
Model tests setup
Model tests have been conducted in Froude’s similitude. The model scale is 1/16th.
The model is composed of:
• a mudmat
• a 1D motion generator
Each mudmat tested is composed of:
• a 5mm thickness steel plate perforated or not (model scale)
• 4 vertical skirts
Figure 1: mudmat model
The 1D motion to be imposed to the model is heave translation. The motion generator is then
composed of:
• a vertical electric jack
• a vertical beam mounted on a trolley and guide rail equipped with one load sensor
• 4 load sensors between the mudmat model and the vertical beam
• an inclination assembly to allow the mudmat model inclination
Mudmat 1 Mudmat 2
Mudmat 3
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Figure 2: overall assembly
15° 5°
Figure 3: inclination assembly
Measurement
The wave elevation is measured from a wave probe located at the same distance from the
wave maker than the model.
The model displacement is measured from the optical system without contact Krypton and its
acceleration by an accelerometer.
The load is measured in two areas: directly on the mudmat model and between the trolley and
the jack.
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5. Slamming Loads
Slamming on still water
For slamming on still water tests, the mudmat impacts the still water surface with a constant
velocity. The results for these tests are summarized on the following figures for three
inclinations (0°, 5°, 15°):
Porosity 1
0
500
1000
1500
2000
2500
3000
3500
4000
4500
0 0.5 1 1.5 2 2.5 3 3.5
impact velocity (m/s)
Cs*
S (m
2 )
0deg
5deg
15deg
Figure 4: slamming on still water – porosity 1
Porosity 2
0
500
1000
1500
2000
2500
0 0.5 1 1.5 2 2.5 3 3.5
impact velocity (m/s)
Cs*
S (m
2 )
0deg
5deg
15deg
Figure 5: slamming on still water – porosity 2
Those results lead to the following conclusions:
• Cs decreases with the porosity growth
• Cs decreases for inclined mudmat
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Slamming on irregular waves
With the aim to be more realistic, some tests have been performed on regular and irregular
waves. Only the irregular wave tests are presented here.
For irregular wave tests, the approach is different than for still water. The slamming load is
treated as a statistical value. The equivalent of 1 hour tests (scale 1) has been performed.
Around 400 slamming impacts have been measured for each test.
The mudmat imposed motion is a harmonic oscillation at wave Tp, around the free surface
mean elevation.
For each water entry, the slamming load is measured.
The measures of Cs for each impact during one test are presented on Figure 6 below:
Porosity1 - Irregular Waves Tests
0
500
1000
1500
2000
2500
3000
3500
0.0 0.5 1.0 1.5 2.0 2.5 3.0 3.5 4.0 4.5
impact velocity (m/s)
Cs*
S (m
2 )
Figure 6: Cs repartition – 0deg inclination
The Figure 6 gives the repartition of the Cs with the impact velocity. With these results, the
Cs is characterized and its variation with the impact velocity can be taken into account for the
design.
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Design impact velocity
In order to compute the design loads, a design velocity has to be associated with the slamming
coefficient. Figure 7 sums up the velocities to take into account in the impact velocity
calculation.
Figure 7: Velocities
Lv : Lowering velocity
)(tvct : Crane tip velocity ctv : significant value
)(tvw : Wave particle vertical velocity wv : significant value
)(tvrel : Relative motion between the crane tip and the surface relv : significant value
The temporal impact velocity is: LrelLcrwimpact vtvvtvtvtv +=+−= )()()()(
In a design approach, an impact design velocity should be available from the significant
value. An impact design value is thus defined as:
relLimpact vCvv ⋅+= )(α
Where )(αC is function of the risk α to be taken. )(αC is defined by the distribution of relv (a
Rayleigh law has been shown to be a conservative estimate).
The significant relative velocity relv is often taken as the quadratic sum of ctv and wv :
²²² wctrel vvv +=
This implies that the wave velocity and the crane velocity are independent; moreover, this
does not take into account the diffraction/radiation of the waves by the vessel. The link
between )(tvrel and )(tvct leads to overestimate the velocity for high period (the free surface
and the crane tip are moving together). The lack of diffraction significantly overestimates the
velocity in shielded area. (See Figure 8)
Crane tip vertical velocity
Lowering velocity
Wave velocity
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Figure 8: relv around a supply for different heading and Tp (Pierson-Moskowitz spectra, long crested)
As the impact forces are proportional to the square of the velocity, a better estimation of relv
can lead to significant differences in the hydrodynamic loadings. Figure 9 shows the evolution
of relv in the lowering area shielded by the vessel (30° relative to the wave).
0
0.1
0.2
0.3
0.4
0.5
0.6
0.7
0.8
0.9
1
5 5.5 6 6.5 7 7.5 8Tp (s)
Vre
(m
/s)
Exact
Quadratic sum
Figure 9: relv
13/19
6. Added mass and damping
Oscillation tests
To measure the mudmat added mass and damping, oscillation tests have been performed with
the following variable parameters:
• Amplitude and period of the harmonic motion
• Sea surface or seabed distance
• Porosity of the mudmat
A total of 72 tests have been performed.
The mudmat trajectory can be written as
Z = A.sin(ω.t) (1)
The hydrodynamic load is
)cos()sin()( 332332 tb
Atm
AtF ωρω
ωρωρ
ωρ −= (2)
with
• ρ = water density
• A = harmonic motion amplitude
• ω = motion pulsation
• m33 = added mass term
• b33 = damping term. Drag, radiation and pressure loss through perforations
From the measured load during the harmonic oscillation tests, the added mass and damping
terms are identified thank to a sliding Fourier analyse. This sliding Fourier analyse allows
identifying the part of the load in phase with the velocity and the part of the load in phase
with the acceleration (cosines and sinus term in equation Erreur ! Source du renvoi
introuvable.)).
Those results (m33 and b33) are compared to theoretical ones in the next section.
14/19
Theoretical calculation
Model description
Figure 10 below gives the reference geometry: a perforated horizontal plate with vertical
skirts. The problem is solved for axisymmetric (3D) or 2D geometries.
The axisymmetric geometry is applicable for suction anchors or nearly square mudmats
whereas the 2D geometry is applicable for elongated mudmats.
Figure 10: geometry
• The structure has a harmonic imposed motion with ω the motion pulsation and A the
amplitude. The problem is solved within the scope of linearized potential flow theory. The
velocity potential ( )x z tΦ , , or ( )R z tΦ , , is written:
i i( ) ( ) e ou ( ) ( ) et tx z t x z R z t R zω ωϕ ϕ− −� � � �� � � �� � � �
Φ , , = ℜ , Φ , , = ℜ , (3)
The fluid domain is split into three subdomains:
• Subdomain 1: outside, from R (or x) = a to infinity
• Subdomain 2: between the seabed and the perforated plate
• Subdomain 3: above the structure
In each subdomain the velocity potential is written as an Eigen-function expansion.
The pressure loss condition at the porous plate is written as a quadratic expression of the
relative vertical velocity through the perforated plate:
2 3 2
1( cos ) cos
2 z zp p A t A tτρ ω ω ω ω
µτ−− = Φ − | Φ − | . (4)
Here τ is the porosity, or open-area ratio, and µ a discharge coefficient, close to 1.
The velocity potentials and normal velocities are then matched as the common boundaries.
For details see Molin & Nielsen (2004) [1] or Molin et al. (2007) [2]. A no-flow condition is
15/19
written on the skirt. The pressure loss equation (4) is satisfied through an iterative procedure
initiated from the solid case solution.
Two porosity cases are studied at several motion amplitudes. We shall note that the
calculations with the two porosities are redundant since the hydrodynamic coefficients only
depend on the « porous Keulegan Carpenter » number defined as
aA
cK 221~µτ
τ−= (5)
Calculations are first performed at mid-water, far from the sea surface. The hydrodynamic
coefficients are then nearly insensitive to the oscillation frequency.
In the figures, the added mass is divided by the fluid density and the damping by the fluid
density times the frequency to be comparable to the added mass.
If the forced motion is tAZ ωsin= , the hydrodynamic load is:
tb
Atm
AtF ωρω
ωρωρ
ωρ cossin)( 332332 −= (6)
0
100
200
300
400
500
600
700
800
0 0.2 0.4 0.6 0.8 1 1.2 1.4 1.6 1.8 2
A (m)
M33
/ ρρ ρρ B
33/ ρ
ωρω ρωρω
M33 porosity 1M33 porosity 2B33 porosity 1B33 porosity 2
Figure 11: porous mudmat at mid-water
16/19
Axisymmetric (3D) versus 2D model
The Figure 12 below shows the comparison between the axisymmetric and the 2D model. For
the axisymmetric model, the diameter used is calculated to have the same area than the
rectangular model.
0
100
200
300
400
500
600
700
0 0.2 0.4 0.6 0.8 1 1.2 1.4 1.6 1.8 2
A (m)
M33
/ ρρ ρρ B
33/ ρ
ωρω ρωρω
M33 2DM33 3DB33 2DB33 3D
Figure 12: Axisymmetric (3D) versus 2D comparison for porosity 2 – mid-water
These models are used to evaluate the m33 and b33 damping for several configurations. The
main variable parameters are the porosity and the distance to the sea surface or to the seabed.
The results from the calculation are compared to the model tests results.
Implementation of the edge drag
The model described above takes into account the flow separation through the perforations,
idealized as porosity, but not the flow separation at the outer edges or at the skirt base.
In Molin et al. (2007) [2], an empirical correction of added mass and damping calculated is
proposed, based on a drag term related to the averaged flow velocity through the plate.
This correction has been implemented in the axisymmetric and 2D models even if the
pertinence of this method for structure equipped with skirts can be questioned.
Based on works of Graham (1980) [3] on plates, followed by Sandvik et al. (2006) [4], the
drag coefficient is related to the Keulegan-Carpenter number by the relation: 1 3
D CC Kα − /= (7)
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Comparison with model tests
-Added Mass
0
100
200
300
400
500
600
700
800
0 5 10 15 20
KC porous
Ma/
rho
2D
3D
Model test Porosity 1
Model test Porosity 2
Figure 13 : Added mass in infinite water
The fact that cK~
is the main parameters of interest is thus confirmed by the experiment. The
calculation slightly over-estimates the results, the axisymetric model being closer to the
experiment than the 2D model.
Compare the added mass of the opaque mudmat without skirts; the added mass of a circular
disk is 8% higher (45% for the 2D plate). Those differences are however not big enough to
explain the difference between the model tests and the calculations. The difference is more
likely due to the drag on the edge of the mudmat, which is not taken into account in the
calculation on Figure 13. Figure 14 presents the results obtained taking into account an
empirical correction for the drag on the edge. The agreement is slightly closer to the
experiment (Note: cK~
is not the only parameter anymore).
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0
100
200
300
400
500
600
0 5 10 15 20
KC porous
Ma/
rho 3D porosity 1 alpha=4
3D porosity 2 alpha=4
3D without edge effect
Model test Porosity 1
Model test Porosity 2
Figure 14 : Edge effect on the added mass
-Damping
0
5
10
15
20
25
30
35
40
45
0 5 10 15 20 25
KC porous
b33/
rho
3D calculation
Model test, porosity 1
Model test, porosity 2
Figure 15: Damping in infinite water
Regarding the damping coefficient, the agreement with the experiment is not as good as for
the added mass. cK~
does not seem to be the only governing parameter. This is due to fact that
the damping is more sensitive to the drag on the edge of the object. The results with the
correction for the edge effects are presented on Figure 16.
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0
5
10
15
20
25
30
35
40
45
50
0 5 10 15 20
KC porous
b33/
rho
3D porosity 1 alpha=43D porosity 2 alpha=43D without edge effectModel test, porosity 1Model test, porosity 2
Figure 16 : Edge effect on the damping
Taking into account the edge effect thus allows a much more accurate calculation of the
damping coefficient. However, the parameter alpha used for the correction is expected to be
sensitive to the thickness of a plate (or skirt length).
7. Conclusions
This research project has lead to a significant progress in the knowledge of hydrodynamic
coefficients used for heavy lifting operations. This better knowledge will allow optimising
heavy lifting operations and then to reduce their cost.
We want to thank the sponsors of this project: Total, Technip, Saipem and Doris Engineering.
A second phase of this project is under discussion for studying more package shapes.
8. References
[1] B. MOLIN & F.G. NIELSEN 2004 Heave added mass and damping of a perforated disk
below the free surface, Proc. 19th Int. Workshop Water Waves & Floating Bodies,
Cortona (www.iwwwfb.org).
[2] B. MOLIN, F. REMY & T. RIPPOL 2007 Experimental study of the heave added mass and
damping of solid and perforated disks close to the free surface, Proc. IMAM Conf., Varna.
[3] GRAHAM J.M.R. 1980 The forces on sharp-edged cylinders in oscillatory flow at low
Keulegan-Carpenter numbers, J. Fluid Mech., 331–346.
[4] SANDVIK P.C., SOLAAS F. & NIELSEN F.G. 2006 Hydrodynamic forces on ventilated
structures, Proc. 16th International Offshore & Polar Eng. Conf., San Francisco.
