wave impacts excitation on ship-type offshore structures in

Proceedings of OMAE-FPSO 2004
OMAE Speciality Symposium on FPSO Integrity
2004, Houston, USA
OMAE-FPSO'04-0062
WAVE IMPACTS EXCITATION ON SHIP-TYPE OFFSHORE STRUCTURES IN STEEP
FRONTED WAVES
Arjan Voogt and Bas Buchner
MARIN, The Netherlands
Joaquín Lopez-Cortijo Garcia
IZAR FENE Shipyards, Spain
ABSTRACT
Wave impacts in steep fronted waves can induce extremely
large impact pressures on the hull of moored ship-type offshore
structures. As part of the SAFE-FLOW Joint Industry Project
these loads were investigated with a dedicated series of model
tests. Based on the results a prediction method is developed,
which can be applied in the early design of the bow structure.
The paper first summarizes the test philosophy and the
developed prediction method. Then the practical design
considerations related to bow slamming are discussed in a
study focusing on the new DP-FPSO concept for a single
environment.
INTRODUCTION
Several incidents in recent years have revealed that wave
impact loading is an important issue for moored floating
production systems. Wave impact damage has been
experienced by both the Foinaven and Schiehallion FPSOs.
The present paper was developed within the SAFE-FLOW
project (SAFE-Floating LOading by Waves), which is initiated
in response to encouragement by the UK Health and Safety
Executive (HSE) following earlier joint research, such as the
JIP ‘F(P)SO Green Water Loading’, completed in 1997
(Buchner, 2002). Its main objective was to investigate water
impact loading hazards in detail.
Based on model tests carried out for both a free floating
FPSO and a fixed schematic bow a prediction tool (BowLab) is
developed which will be discussed in this paper.
EXPERIMENTS
As part of the project MARIN performed 2 series of model
tests at scale 1:60 in deep water. First tests were carried out on
a free floating Schiehallion FPSO model (Figure 1).
Figure 1: Bow impact event on free floating model
In these tests in irregular seas the incident wave data,
vessel motions and resulting relative motions, bow pressures
and structural response are measured. The tests showed that
more detailed load measurements were necessary and that an
investigation was needed into the relation between the
incoming waves and these loads. This resulted in the second
model test series on a highly instrumented fixed simplified
bow.
The simplified bow was instrumented with a large array of
pressure transducers and 3 force panels. The test program, also
making use of extensive video recordings, was designed such
that it was possible to determine the correlation between
undisturbed wave shape and the impact pressure time traces.
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Copyright © 2004 by ASME
From these tests irregular sea incident wave data and bow
pressure results are available on a fixed schematic bow
structure with varying rake and plan angles.
It was found that the magnitude of the wave impacts at the
front of the bow is dominated by the wave characteristics
(namely the local wave steepness), rather than by the motions
of the ship relative to the waves (relative wave motions).
Further the maximum pressures are measured close to the crest
of the incoming waves. An example of a steep wave front
reaching the bow structure is shown in Figure 2.
motion
RAO
WF ship
motions
location
bow slam
properties
Sea
Stat
im pac
non lin
wave
empirical
relations
structural
properties
structural
response
Figure 3: Load response prediction method
To predict the probability with which structural responses
occur within one selected sea state, BowLab uses the different
building blocks described in the sections below.
Non Linear Wave
The local wave steepness (dζ/dx) can be determined from
measurements of the wave elevations in an array of probes. An
example is shown in Figure 4, which shows the spatial wave
profile for successive steps in time. The time step between the
different lines is 0.31 seconds and the distance between the
probes 6 metres allowing for an accurate derivation of the local
wave steepness.
Figure 2: Steep wave approaching the bow
PREDICTION METHOD
To calculate the probability of wave impacts on any given
vessel structure, at any given elevation, a prediction method for
the simulation of the water surface and the occurrence of wave
impacts was created. Evaluating the available results, it was
decided to develop a method that split up the problem in two
main problems, see Figure 3:
• The relation between the local wave characteristics
and the magnitude (and other characteristics) of the
wave impact, the lower line in the figure.
• The position of the impact based on the related ship
motions, the upper line in the figure.
The combination results in a localised impact with specific
properties, resulting in the structural response of a local
structure with its specific structural properties.
p
MWL
6m
R1
R2
R3
R4
R5
R6
Figure 4: Visualisation of local wave steepness
The combined spatial and temporal information of the sea
state needed to derive the local wave steepness is not generally
available (in full scale data and model tests). Therefore the
vertical free surface velocity is preferred as input to the model.
The vertical free surface velocity (dζ/dt) is related to the local
wave steepness (dζ/dx) by the wave celerity.
In steep waves that cause the bow impact, linear theory
clearly under predicts the wave steepness. The most suitable
method of simulating the water surface to give a reasonable
probability of vertical free surface velocity was found to be
second order wave theory, as described by Sharma and Dean
(1981) for instance. Applying second order wave theory results
in an improved prediction of dζ/dt, as shown in Figures 5 and 6
for the basin waves applied.
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Copyright © 2004 by ASME
measured
2nd order
linear
linear
2nd order
measured
dζ/dt
6
10
12
Vertical free surface velocity
Figure 5: Measured, first and second order waves
Figure 7: Empirical relation for local impulse
The relation is independent of the sea state (within the
range tested) and holds for a schematic flat plate bow. Within
the design method the mean fit is used as a maximum that can
occur. For more realistic curved bow shapes the loads are
reduced as described in the next sections. The spreading around
this mean can be used as input to the derivation of the load
factors in a first principles reliability approach.
Figure 6: Probability vertical free surface velocity
It is clear that the second order theory is not capable of
describing the asymmetry in the measured non-linear wave.
However within the accuracy of the present design tool this is
not considered a critical aspect and the distribution of the
vertical free surface velocities do match the measured nonlinear distribution reasonably well.
Empirical Relations
To estimate the probability of an impact at a certain
vertical free surface velocity, the percentage of impacts
occurring were counted for bins of vertical free surface
velocities (Guedes Soares, et al. 2004). It is observed that up to
a certain vertical free surface velocity no impacts occur. Above
this threshold level, the probability increases linearly with the
vertical free surface velocity up to the probability of 100%.
Beside the slam probability, the slam magnitude is of vital
importance. After analysis of all data, it was decided to relate
the slam impulse (I), the area under the load time trace, to
vertical free surface velocity (dζ/dt).
Figure 7 shows the measured impulses versus the
corresponding vertical free surface velocities. For different
velocity bins the mean and standard deviation of the occurring
impulses is added to the figure, resulting in straight lines.
The comparison with the wave heights of the incoming
undisturbed waves show that most slam loads occur relatively
close to the wave crest. This is consistent with the observation
that the maximum velocity occurs close to the crest front of the
incoming wave.
WF Ship Motions
The test results showed that the slam centre occur close to
the crest of the undisturbed non-linear wave. However for the
design of a bow panel in a floating structure, also the ship
motions should be taken into account to determine the location
of the impact. To check the assumption that these motions in
steep waves can still be described with linear diffraction theory,
the model tests reported with a floating model (Voogt, 2001)
were analysed further.
An iterative procedure has been developed that fits a linear
wave and its second-order contributions to a measured wave
train. The procedure is described in Voogt and Buchner (2004)
where a good comparison between measurements and linear
calculations is found, which is confirmed in the time traces in
Figure 8.
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Copyright © 2004 by ASME
and stiffeners in the bow area this is not always sufficient to
determine an accurate response with resulting stresses. The
monitoring on the Schiehallion bow (Hodgson, 2003) allows
for this comparison between pressures and stresses. Based on
these results the pressure duration, immersion velocity and
spatial extent are considered critical (Figure 10).
Data already described
Impulse
Magnitude
Structur
Geometry
Bow
Shape
Figure
motions
8:
Measured
and
calculated
vessel
Impact Location
Even for these large differences between the linear wave
and the measured wave, the linear wave is still capable of
describing the major part of the ship response. Therefore the
following procedure is implemented to determine the location
of the impact:
• For a time trace of the linear wave the motions of the
vessel are calculated.
• With this linear wave non-linear wave components
can be calculated and added to the time traces.
• Combining the time trace of the non-linear wave and
the linear ship motions gives a relative wave motion in
front of the vessel.
• From these relative motions the slam centre can be
determined if a slam occurs.
Figure 9 visualizes the procedure. The yellow ship is
moving on the blue linear seas. While the red line indicates the
second order wave which determines the impact size and
location.
Impact &
location
Target
Location
of Slam
Load- Response
Pre diction
Pressur
Duration
Immersion
Velocit
Spatial
Extent
Structural Response
Equivalent Pressure
Slam
Figure 10: Bow impact characteristics
For some impacts the pressure front travels upward over
the plate, while others occur more instantaneously. If we
determine the time between the impacts over the height of the
flat plate differences of 0.0 to 0.3 seconds can be found for two
adjacent transducers. This corresponds to a velocity of 10 m/s
and higher. Up to immersion velocities of 100 m/s the pressure
time trace of the slam show the traditional shape (progressive
slams), above this velocity only short sharp (instantaneous)
slams occur.
Different immersion velocities tend to occur at comparable
values of vertical free surface velocity. Therefore the
probability distribution of the immersion velocities is
determined. This distribution is independent of the slam size.
The average distribution is shown in Figure 11. This figure
shows two types of slams:
The blue bars on the left result in slams that progress over
the bow. The duration of the resulting load time traces on a
bow segment are direct related to these velocities.
For velocities above 100 m/s (the red bar) an instantaneous
slam occurs. The duration of the resulting load time trace is
independent off the target size and thus independent of the
immersion velocity.
2nd order wave
Figure 9: Wave impact location
Figure 11: Probability of immersion velocities
Bow Slam Properties
So far the method focussed on the local impact pressures
on the bow and on its position. However, to design the plating
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Copyright © 2004 by ASME
For the schematic bow with the large number of pressure
transducers the spatial extent is derived by integrating the bow
pressures downward from the maximum slam over increasing
areas, with steps of 1.5 metre. This resulted in two typical types
of slams:
Instantaneous slam: The whole pressure time trace
decreases with the integral over the area. This type of impulse
is recognized as an instantaneous slam. The decrease of the
pressure depends on the size of the pressure front and drops
down fast during integration.
Progressive slam: The rise time increases while the
pressure impulse remains relatively constant. This second type
of impulse progresses along the structure. Due to the pressure
integration the rise time increases with the size of the area and
the impulse remains approximately the same.
In the prediction method, the time traces for instantaneous
and progressive slams are scaled to represent the correct
impulse and immersion velocity for each calculated impact.
With the resulting curves and the location of the slam centre the
equivalent static pressure on a specified panel can be
determined.
characteristic of bow impact loads, particularly instantaneous
slams, that pressure loads are applied very quickly. Dynamic
response of the structure is therefore likely, requiring that
dynamic amplification be included in any design calculations.
The structural evaluation of bow slam loading and
structural response is in principle a coupled hydro elastic
problem. These coupled effects are likely to be most important
when the natural period is longer than twice the duration of the
slam. The full coupled problem has not been investigated in the
present project. Instead a simplified approach has been used.
With the rise and decay time (and their characteristics) of
the slam and the natural frequency of the structure the dynamic
response can be calculated with a single degree of freedom
mass spring system. Excluding hydro-elastic effects, the
structural response can be described with the equation of
motion in which a constant damping, added mass and mass are
assumed.
For a typical large local slam the rise and decay times are
about 10 ms and for a panel natural period of about 1/10 sec the
DAF will be 0.96 (Figure 13). This DAF smaller than 1 occurs
if the structure has a long period and does not have time to
respond to the impact load before it starts to reduce.
Structural properties
The width and curvature of the panel can result in a
reduction of the equivalent static pressure. For very small panel
width the results can be considered two dimensional and
multiplication of the pressure with the width of the panel gives
the design load. For larger panels the curvature of the bow
becomes important. Even for a flat plate the average pressure
reduces with the width of the panel due to local wave effects.
An empirical formula is fitted through the measurements
performed in Glasgow and documented in Barltrop and Xu
(2004). Since one would expect the curved bow factor to be
always less than the flat plate value, the pressure for a curved
plate is limited to the flat plate value. The resulting factor
between the local pressure and the pressure average over the
width of the plate is shown in Figure 12.
DAF
0.96
Rise time 0.01 s
Decay half time 0.01 s
Structural freq. 10 Hz
Pressure
Pmax
Pmax/2
Tr
T½d
Time
Figure 12: Pressure reduction for curved panels
Figure 13: Example Dynamic Amplification Factor
Structural response
The previous sections focussed on the bow impact loading.
However, for the evaluation of ship-type offshore structures,
the structural response under this impact loading determines
whether a structure is able to survive a certain event. It is a
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Copyright © 2004 by ASME
DP FPSO
The proposed method was now applied, as a case study, to
the DP FPSO concept, developed by IZAR FENE Shipyards in
co-operation with FMC Sofec, DNV and MARIN, Cortijo et al
(2003) and see Buchner and Cortijo (2003). The hull of the DPFPSO has been designed taking the following key aspects into
consideration:
a) Environmental forces induced by waves, wind and current
should be minimised in order to keep the CAPEX/OPEX
of the Unit within reasonable limits, as well as the
emissions produced by the dual (diesel/gas) generators,
which must be continuously running in a fully DP vessel.
b) The hull configuration must allow for an adequate
arrangement of the thrusters at the fore and aft ends, to
avoid thruster-thruster and thruster hull/turret interactions.
Moreover, the layout of the machinery spaces and the ease
of thruster overhauling is significantly affected by the
selected vessel hull forms.
c) In heavy storms (hurricanes, squall events, etc.), the vessel
behaviour against green waters and slamming occurrences
must be acceptable, insuring adequate safety of the people
onboard and sensitive equipment (fire fighting, lifesaving
equipment, etc.).
d) Layout requirements, such as: Accommodation forward
(for navigation in sail away condition after disconnection),
turret amidships (to minimise riser system dynamics and
vessel motions for turret buoy (dis)connection operations),
and offloading equipment at the stern for tandem
offloading, have to be catered for in the design.
e) The hull design must also account for constructability
issues, i.e. a simpler structure is easier to built, inspect and
maintain.
In order to accomplish the above mentioned design
requirements of the DP-FPSO, the vessel hull has been
provided with the following features:
- Hull forms typical for new-built FPSOs, with prismatic
mid-body in way of the cargo tanks, sloped flat transom
and triangular bow. Rounded bilge of adequate radius to
accommodate bilge keels, as well as camber of suitable
height.
- The fore and aft ends have been designed to properly
accommodate the envisaged number and size of thrusters
(three forward and three aft). The triangular bow allows a
suitable arrangement for the three forward thrusters,
minimises the wind, current, wave drift forces and forward
resistance, and yields a simpler construction. In order to
provide room for installation of the offloading equipment
and flare, the hull shape aft at main deck level will be
square (full beam) in a cantilever structure of rounded
shape to prevent slamming.
- A forecastle deck of suitable height is provided to insure
acceptable green water loads in extreme design conditions.
A poop deck is provided as well for similar reasons.
Bulwark of reduced height (to avoid large induced loads
on the stanchions and supporting deck) is fitted forward
and aft as required. A bow flare angle of 30 degrees has
been provided above the maximum draft.
A moonpool of diameter as specified by turret designer is
located amidships.
-
In Figures 14 and 15 the artist impression and body plan of
the vessel are shown. The vessel characteristics are shown in
Table 1.
Figure 14: Body plan of the DP FPSO
Table 1: Main particulars of the DP FPSO
Parameter
Displacement (fully loaded)
Length between perpendiculars
Breadth
Depth
Draft (fully loaded)
Value
214,757
260
46
28
20.5
Unit
Metric tonnes
m
m
m
m
Figure 15: Artist impression of the DP FPSO
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Copyright © 2004 by ASME
WAVE IMPACT ANALYSIS
The prediction method can be used to estimate a suitable
long-term distribution of slam response pressures on any
relevant structure and this in turn can be used as input to a
reliability assessment of the design. From this assessment load
factors can be derived which needs to be applied to asses the
design pressure on the construction as described by Fyfe
(2004). Although in a real design the most critical weather
conditions should be determined for a specific vessel, the
present paper focuses on design methodology rather than the
design for a specific area. Therefore, the present paper makes
use of the 100-year Hurricane for the Gulf of Mexico, see Table
2.
5m
4m
Table 2: 100-year Hurricane for the Gulf of Mexico
Parameter
Significant wave height
Spectral peak period
JONSWAP gamma
100-year wave
12.5
13.0
3.3
Unit
Metre
Seconds
-
The structural configuration is checked both for plating
and stiffener design at different elevations on the bow and for
the supporting transverse frames.
The characteristic dimensions are derived from the
drawings in Figure 16. The right hand side drawing shows the
vertical distance of 5 metres between the transverse frames.
Four stiffeners are located between these frames, resulting in a
horizontal stiffener spacing of 1 metres. The figure also shows
the stiffener design from elevation 24000-29000 with 4 metre
web spacing.
As each plate is supported by a stiffener at the top and
bottom the typical panel dimensions for the calculation of the
loads are the equivalent. For the present design a panel of 4
metre wide and 1 metre height is used. For the transverse frame
design a panel height of 5 metre is used which is the distance
between the decks.
Figure 16: Structural Configuration
As mentioned before the width and curvature of the panel
can result in a reduction of the equivalent static pressure. For
very small panel width the results can be considered two
dimensional and multiplication of the pressure with the width
of the panel gives the design load. For larger panels the
curvature of the bow becomes important. This curvature is
described for the present design with the local panel radius
derived from a plan view of the deck as shown in Figure 17.
14 m
Figure 17: Panel diameter
For this typical panel the design loads and response are
calculated with the prediction method described above. The
results are shown in Table 3 for different elevations of the panel
above the mean surface level. The table shows both the static
and dynamic pressure. In the calculation of the dynamic
pressure the panel stiffness and added mass are used to
calculate the period of the response (30 Hz).
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Copyright © 2004 by ASME
Table 3: Equivalent pressures on plate and stiffeners
for different elevations on the bow
Elevation above
MSL [m]
7
8.5
10
12
Static Pressure
[kPa]
206
262
301
258
Dynamic Pressure
[kPa]
323
411
468
412
CONCLUSIONS
In the present paper the practical design considerations
related to the wave impact problem are discussed. The
following can be concluded:
•
•
For this specific structure the dynamic response results in a
significant increase of the equivalent pressures. The plate and
stiffeners are designed against these dynamic pressures. In the
chosen condition the maximum pressures are experienced at 10
metres above the mean surface level which is close to the
crests.
With the prediction method the sensitivity of these loads
against the radius of the panel can easily be checked. Table 4
shows the results for the typical panel located at 10 metres
above the MSL. The results show that the pressures increase
significantly when a blunt bow is used.
•
•
Table 4: Equivalent pressures on plate and stiffeners
for different curvatures of the bow
Local Panel
Radius [m]
7
15
30
40
Static Pressure
[kPa]
301
431
586
660
Dynamic Pressure
[kPa]
468
669
910
1025
For the transverse frame design a panel of 4 metres wide
and 5 metres height is used. For the increased are the response
periods are longer. The prediction method can easily be used to
perform a sensitivity study on this period and the elevation of
the transverse frame. The calculated pressures are shown in
table 5 and are lower than for the plate and stiffeners,
especially when accounted for the longer natural periods. For
the design loads these pressures needs to be multiplied by the
panel area resulting in larger loads on the transverse frames
than on the stiffeners as expected.
Table 5: Equivalent pressures on transverse frames
for different elevations and natural periods
Elevation Static Pressure
[kPa]
[m]
8.5
10
12
209
264
211
Dynamic Pressure [kPa]
30 Hz
20 Hz
10 Hz
328
414
331
321
405
324
249
305
250
•
Wave impact is a significant problem for floating
offshore structures in steep waves and needs to be
assessed in the early design of the structure.
The developed design evaluation method can assist in
this process as it predict the bow slam loading
problem from the input (scatter diagram) to the output
(predicted load and response levels) based on a clear
description of the bow slam physics.
The method is based upon second order wave theory
describing the wave steepness and an empirical
relation between the wave steepness and the local
impact. The position of this impact follows from a
time domain analysis of the linear ship motions. The
method assumes long crested waves which are
considered a worst case scenario.
Applying this tool shows the strong influence of the
local panel radius on the calculated loads. The loads
can be minimized by decreasing the bow radius as this
minimizes the wave entrapment.
The structural response determines the design
pressures on the constructions and need to be carefully
considered in the design process.
ACKNOWLEDGMENTS
The SAFE-FLOW project (SAFE-FLOating offshore
structures under impact loading of shipped green water and
Waves) is funded by the European Community under the
‘Competitive and Sustainable Growth’ Programme (EU Project
No.: GRD1-2000-25656) and a group of 26 industrial
participants (oil companies, shipyards, engineering companies,
regulating bodies). The participants are acknowledged for their
interesting discussion of the results during the project and the
permission to publish the present paper. The authors are solely
responsible for the present paper and it does not represent the
opinion of the European Community.
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Glasgow & Strathclyde Universities (NAME)
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Copyright © 2004 by ASME
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