I
MPROVING THE SEAKEEPING BEHAVIOUR IN HEAD WAVES
I
MPROVING THE SEAKEEPING BEHAVIOUR IN HEAD WAVES
Proefschrift
ter verkrijging van de graad van doctor aan de Technische Universiteit Delft,
op gezag van de Rector Magnificus prof. ir. K. C. A. M. Luyben, voorzitter van het College voor Promoties,
in het openbaar te verdedigen op donderdag 4 februari 2016 om 10:00 uur
door
Albert Aart Klaas RIJKENS
scheepsbouwkundig ingenieur Technische Universiteit Delft, Nederland
promotor: Prof. dr. ir. R. H. M. Huijsmans copromotor: Dr. ir. J. A. Keuning
Samenstelling promotiecommissie:
Rector Magnificus voorzitter
Prof. dr. ir. R. H. M. Huijsmans Technische Universiteit Delft Dr. ir. J. A. Keuning Technische Universiteit Delft
Onafhankelijke leden:
Prof. dr. M. R. Renilson Higher Colleges of Technology Prof. dr. G. Thomas University College London
Prof. dr. O. M. Faltinsen Norwegian University of Science and Technology Prof. ir. D. Stapersma Technische Universiteit Delft
Dr. ir. F. van Walree Maritime Research Institute Netherlands Prof. ir. J. J. Hopman Technische Universiteit Delft, reservelid
This research is supported by the Maritime Innovation Program in combination with the consortium:
Delft University of Technology
Maritime Research Institute Netherlands Damen Shipyards
Netherlands Ministry of Defence Bureau Veritas
Lloyd’s Register Imtech Marine Quantum Controls
Keywords: fast ships, transom flap, interceptor, thrust control, wave impacts
Printed by: Gildeprint - The Netherlands
Cover illustration: Daan Sijbring -www.daansijbring.nl
Copyright © 2016 by Albert A. K. Rijkens All rights reserved.
ISBN 978-94-6233-210-2
An electronic version of this dissertation is available at: repository.tudelft.nl
C
Nomenclature ix
1 Introduction 1
1.1 Research outline . . . 3
2 Operational Control 5 2.1 Non-linearities in the behaviour of fast ships . . . 6
2.1.1 Performance assessment based on vertical peak accelerations. . . . 8
2.2 Operational control variables . . . 13
2.2.1 Speed control . . . 13
2.2.2 Trim control . . . 22
2.3 Research questions . . . 32
3 Force Measurements on Active Trim Control Devices 33 3.1 Model Experiments . . . 33
3.1.1 Model and experimental set-up . . . 34
3.1.2 Test program and procedure . . . 37
3.2 Results . . . 39
3.2.1 Stationary condition . . . 40
3.2.2 Harmonic oscillations . . . 51
3.2.3 Superimposed harmonic oscillations . . . 60
3.3 Conclusions. . . 62
4 Design of the Proactive Control System 63 4.1 Proactive control strategy . . . 63
4.2 Simulation model . . . 65
4.2.1 Mathematical formulation . . . 66
4.2.2 Implementation of speed control . . . 72
4.2.3 Implementation of trim control . . . 76
4.3 Working principle of the proactive control system . . . 77
4.3.1 Decision making process. . . 78
4.4 Software architecture . . . 81
5 Numerical Computations 85 5.1 Numerical input and calibration . . . 86
5.1.1 Selected hull form . . . 86
5.1.2 Calm water sinkage and trim. . . 86
5.1.3 Regular waves . . . 89 vii
5.2 Convergence study for seakeeping computations . . . 96
5.2.1 Discretisation of the hull . . . 97
5.2.2 Time step . . . 97
5.2.3 Wave realisation . . . 100
5.3 Configurations for the proactive control system. . . 103
5.3.1 Proactive thrust control . . . 103
5.3.2 Proactive control of the interceptors . . . 110
5.3.3 Proactive control using thrust and interceptors . . . 114
5.4 Operational and environmental conditions . . . 117
5.4.1 Sea state . . . 117
5.4.2 Desired forward speed . . . 120
5.4.3 Criteria. . . 124
5.5 Performance degradation due to disrupted wave registrations . . . 126
5.5.1 Wave disturbance implementation. . . 126
5.5.2 Performance degradation . . . 128
5.6 Conclusions. . . 130
6 Proof of Concept 133 6.1 Wave prediction method . . . 134
6.1.1 Theoretical predictable zone . . . 136
6.1.2 Wave measurements . . . 138
6.1.3 Discussion and positioning of the wave probes . . . 144
6.2 Model experiments . . . 146
6.2.1 Experimental set up . . . 146
6.2.2 Data handling system . . . 152
6.2.3 Test program . . . 156
6.3 Results and comparison with numerical computations . . . 157
6.3.1 Calm water. . . 157
6.3.2 Regular waves . . . 158
6.3.3 Irregular waves. . . 164
6.4 Conclusions. . . 175
7 Conclusions and Recommendations 177 7.1 Conclusions. . . 177
7.2 Recommendations for future work . . . 179 A Additional results of the harmonic oscillation tests 183 B Speed tests with the towing carriage 191
References 200
Summary / Samenvatting 201
Dankwoord 209
Curriculum Vitæ 211
N
L
ATIN
L
ET TERS
A Submerged cross sectional area m2
a Reduction length of the near-transom flow correction m
a0 Near-transom flow correction coefficient
ab f Buoyancy correction coefficient
Ap Projected area of chines and transom m2
az bow Vertical accelerations at the bow m/s
az CG Vertical accelerations at the CG m/s
b Beam of planing surface m
Bc max Maximum beam over the chines m
Bt Waterline width of the transom m
Btr Beam at the transom measured over the chines m
BW L Beam on waterline m
CD Drag coefficient 1/2ρVFD2
mb2
Cd Damping coefficient
CD,c Cross flow drag coefficient
CF Skin friction coefficient (logRe−2)0.075 2
cg Wave group velocity m/s
CL Lift coefficient 1/2ρVFL2
mb2
CM Pitch moment coefficient 1/2ρVMθ2
mb3
Cm Added mass coefficient
CN Normal force coefficient 1/2ρVFN2
mb2
CT Tangential force coefficient 1/2ρVFT2
mb2 CV Speed coefficient pVmg b d t Time step s Er r Error indicator f Sectional force N FD Drag force N FL Lift force N ix
FN Normal force N
Fr Froude number pVs
g L
fr ed Reduction in sectional force due to near-transom flow correction N
Fr∇ Froude displacement number qVs
g ∇1/3
FT Tangential force N
g Gravitational acceleration m/s2
h Water depth m
H1/3 Significant wave height m
hi Deflection of the interceptor mm
hi nput Input reference position of the interceptor mm
b
h Normalized deflection of the interceptor hi
hi , max
Hw Wave height m
I Mass moment of inertia in pitch kgm2
i Discrete control cycle
k Wave number rad/m
KG Vertical position of CG m
K˙θ Pitch velocity control gain deg/sdeg
b
K˙θ Normalized pitch velocity control gain KK˙θ
˙θ max
ky y Pitch radius of gyration m
L Ship length m
Lc Length over the chines m
LCG Longitudinal position of CG m
LOA Length overall m
LW L Length on waterline m
M Mass of the ship kg
ma Sectional added mass kg
Mθ Pitching moment Nm
N Total number of wave components
Re Reynolds number VsL
ν
RT Total calm water resistance N
Rw Total mean resistance in waves N
S Number of strips
SW Wetted area m2
Sζ(ω) Spectral wave density m2/rad
T Draft m
t Time s
Tdes Thrust at the desired forward speed N
to f f set Time difference between the synchronisation pulse and the start
time of the carriage
s
Tp Wave spectrum peak period s
Tp,e Wave spectrum peak encounter period s
tph Prediction horizon s
tr un Run length simulation s
ttl Time lag s
U Water velocity tangential to the local planing surface m/s
u Thrust force fraction
V Water velocity normal to the local planing surface m/s
Vav Average forward speed of the ship m/s
Vc Speed of the carriage m/s
Vdes Desired forward speed of the ship m/s
Vend Achievable end speed within a prediction horizon m/s
Vi nput Input reference speed m/s
Vm Forward speed of the model m/s
Vr ed Maximum speed reduction within a prediction horizon m/s
Vs Forward speed of the ship m/s
W Weight of the ship N
wz Vertical orbital wave velocity m/s
x Surge motion m
xL Distance from the centre of gravity to centre of action for lift force m
xN Distance from the centre of gravity to centre of pressure for normal
force
m
xst ar t Initial distance between the model and the wavemaker m
y Sectional wetted half beam m
z Heave motion m
G
REEK
L
ET TERS
αf Angle of the transom flap deg
αi Geometrical interceptor angle of an equivalent transom flap deg
β Deadrise angle deg
δf Deflection-beam ratio of the trailing edge of the transom flap
δi Deflection-beam ratio of the interceptor
ε Relative error
εφn Disturbance of the wave phase angle
εζφ Wave disturbance factor
ηx Time trace of signal x
κ Wave steepness ratio Hw
λw
λ Chord-beam ratio of the transom flap
λw Wave length m
µx Average value of signal x
∇ Volume of displacement m3
ν Kinematic viscosity m2/s
ω Wave frequency rad/s
ωf Oscillation frequency transom flap rad/s
ωi Oscillation frequency interceptor rad/s
ωp Peak frequency of wave spectrum rad/s
ωp,e Peak encounter frequency of wave spectrum rad/s
φ Phase angle rad
ρ Mass density of water kg/m3
σ Span-beam ratio of control device
σx Standard deviation of signal x
σ2x Variance of signal x
τ Trim angle deg
θ Pitch motion deg
ζ Wave elevation m
bζ Fourier transform of the wave elevation
ζi nput Input reference wave elevation m
ζs ync Wave synchronisation pulse
eζ Complex wave amplitude
S
UBSCRIPTS
a Amplitude of the signal
CG Centre of gravity
f Flap
i Interceptor
max Maximum value of the signal
1
I
NTRODUCTION
A high speed performance at sea can give an operator a tactical or economical advan-tage. The success rate of some missions, like Search And Rescue (SAR) operations or patrol duties, is even largely dependent on a quick arrival time of the ship and thus de-mands a high operational speed during daily operations, irrespective of the particular ambient wave conditions. However, a high speed performance in a seaway is often asso-ciated with rather violent motion behaviour in which the resulting large wave impacts, or vertical peak accelerations, pose a rather important limiting factor for the vessel’s op-erability. These wave impacts not only endanger the safety of the people on board but may also compromise the structural integrity of the ship.
Over the past decades numerous studies have been carried out to improve the sea-keeping behaviour of fast ships. Most of these studies focus on hull form developments which resulted in a wide variety of advanced marine vehicles for high speed transporta-tion at sea. Alongside hull form modificatransporta-tions to the more conventransporta-tional fast monohull designs all sorts of alternative concepts have been developed in an effort to increase forward speed and to improve seakeeping behaviour. Well-known examples of these ad-vanced marine vehicles are amongst others: hydrofoils, catamarans, Small Waterplane Area Twin Hulls (SWATH), hovercraft and Surface Effect Ships (SES). All of these designs have their own merits and shortcomings and they have matured to a certain stage of per-fection over the last decades. However, a common denominator for each of these con-cepts is that they are considerably more complicated compared to the monohull design, which makes them rather expensive both in construction as in operation.
Moreover, most of these alternative marine vehicles need to be relatively large in size with respect to the surrounding waves, as the wave height poses a rather strict limit-ing factor for their operability. For instance, multihull vessels need to maintain suffi-cient clearance between their main deck and the water surface, as the greater waves can cause severe wet deck slamming which needs to be prevented at all times. Air cushion supported vessels, on the other hand, have difficulties in retaining their air pocket if the waves become comparatively large. Hence, these concepts need to be of considerable size if they are required to operate in all weather conditions which further adds to their
overall costs. For these reasons, many operators still tend to favour the relatively simple and low-cost construction of the fast monohull design.
However, these fast monohulls can exhibit rather severe motions and accelerations in higher sea states, which also limits their operability and so a continuous demand exists for improved seakeeping behaviour. Traditionally, most design efforts were focussed on optimising calm water performance. This resulted in rather efficient hull forms, from a resistance point of view, but these ships were also notorious for their violent motions and high acceleration levels in waves. After the Second World War research focus steadily shifted towards improving the motion behaviour of fast ships in waves and this gained real interest as from the late sixties (Savitsky, 1968).
From that time, a considerable amount of work has been carried out in order to im-prove the seakeeping performance of fast ships. New innovative hull forms were devel-oped which significantly reduced the wave exciting forces at speed. One of the design strategies developed at the Delft University of Technology is called the Enlarged Ship Concept (ESC) in which the length of the hull is increased to enable an optimisation of the bow sections from a strictly hydrodynamic point of view (Keuning and Pinkster, 1995). This design procedure later evolved into the Axe Bow Concept (ABC) which is characterised by an even more radical bow shape to further improve the behaviour of the ship in waves (Keuning, 2006).
Both of these design strategies have been successfully applied to a wide range of fast ships. However, most of these vessels are still relatively large in size with overall lengths ranging between 35 – 55 m. This is partly because the design philosophy entails an in-crease of the ship’s length to accommodate the strong modifications in geometry of the forward part of the hull. Although certain aspects of the design philosophy can also be applied to smaller ships, as was demonstrated more recently during the development of the new SAR boat for the Royal Netherlands Sea Rescue Institution with an overall length of 19.3 m (Keuning, 2012).
Especially these smaller ships suffer from adverse operational effects while sailing in a seaway at high forward speeds. This is partly due to their limited overall dimensions which means that the wave heights in prevailing or “normal” circumstances may already be relatively large. In addition, the typical mission purposes of these smaller type of ships often call for high speed performances, in which they need to operate at Froude numbers that are generally not economically or technologically viable for their bigger counterparts. These typical conditions can result in rather severe wave impacts that not only take place in the bow region of the ship but which can also occur further aft during sequences in which the hull is (almost) completely emerged from the water surface.
The crews of this smaller class of fast ships often use an operational control method called “throttle control” to evade the larger wave impact events. Full scale trials have indicated that well-trained helmsmen are able to estimate the response behaviour of the ship based on their visual observations of the incident wave in front of the bow (van Deyzen, 2014). They consequently react to the more extreme waves by reducing the thrust several seconds before an expected wave impact takes place. These timely con-trol actions allow the vessel to decelerate, for a short period of time, resulting in a lower forward speed at the moment of impact which, in general, results in a more comfortable ride.
Manual throttle control requires an experienced operator who is able to apply the necessary throttle corrections. However, even the most skilful operators may still make occasional mistakes in their assessments. These misjudgements can be aggravated due to various causes like fatigue, distraction or loss of concentration. Moreover, the oper-ator’s view can be hampered by excessive spray, severe weather conditions or darkness. These type of conditions make it rather difficult (or even impossible) to apply the well-timed throttle adjustments.
In an attempt to fully exploit the potential benefits of this operational control princi-ple and alleviate the human responsibility, a new research initiative was set up to inves-tigate the feasibility of an automated thrust control system. A initial study to the theo-retical performance increase of this proactive thrust control system was already carried out by Van Deyzen (2014). Van Deyzen used a motion prediction program to compare the results of simulations with proactive thrust control to the outcome of a benchmark simulation at constant forward speed. During the controlled simulations continuous short-term predictions were made at regular time intervals to approximate the vertical acceleration in the near future. These numerical predictions were accordingly used by the system to determine whether it needed to reduce the forward speed of the ship in order to evade a high peak value in the vertical accelerations. These simulations demon-strated that the maximum acceleration level could be significantly reduced with this op-erational control method, while the average attainable forward speed was similar to the particular benchmark simulation.
In this study, the proactive system is expanded in which it can not only change the speed of the ship but, in which it can also modify the trim during a control action. The trim (and consequently the sinkage) is affected by the application of active trim control devices at the transom of the ship. This instantaneous trim can have a strong influence on the magnitude of the vertical peak accelerations as was already indicated in the exten-sive systematic model experiments of Fridsma (1969). Hence, these mechanisms could possibly be used to change the orientation of the ship to “a more favourable position” at the time of the wave impact event in order to lower its magnitude. The combined control actions of both the thrust and the trim devices could potentially lead to a further reduction of the vertical peak acceleration levels, while average attainable forward speed is increased in comparison to the configuration with only proactive control of the thrust.
1.1.
R
ESEARCH OUTLINE
Chapter 2 provides the background information to formulate the main research ques-tion of this study. It first discusses the origin of non-linear behaviour of fast ships in head waves and it identifies the limiting factors in operability of these small and fast ship types using existing full scale measurement data. The current day practices of manual throt-tle control actions applied by experienced helmsmen is visualised and results with and without active throttle control are compared. In addition, the theoretical improvements of an automated proactive thrust control system are presented using numerical results of the preliminary study carried out by Van Deyzen (2014). Hereafter, the potential ben-efits are discussed when this proactive system is expanded with the additional control parameter to actuate the trim control devices. At the end of this chapter the main re-search question is posted which is succeeded by a number of related sub-questions that
will all be addressed in the present study.
The objective of Chapter 3 is to obtain a detailed description of the hydrodynamic characteristics of active trim control devices. For this purpose a series of systematic model tests are carried out in order to quantify and compare the hydrodynamic per-formance of a transom flap and an interceptor design. Measurements are conducted in different stationary modes, where the devices are kept in a fixed predetermined position, but also in various dynamic modes in which forced oscillatory motions are imposed on the control surfaces. The generated hydrodynamic forces are analysed and several em-pirical relations are formulated based on these experimental results. These emem-pirical relations can be implemented into a motion prediction program in order to quantify the effects of these active devices on the motion behaviour of fast ships.
Chapter 4 describes the design of the proactive control system. First, the general nu-merical implementation of the control strategy is explained. Hereafter, the mathemati-cal formulations of the simulation program and the numerimathemati-cal implementations of the two control variables are discussed. Various control options are computed whenever the predicted acceleration level exceeds a specified threshold value and the selection proce-dure for obtaining the ultimate desired solution is explained. At the end of this chapter some words are devoted to the architecture of the program to enable the required real-time performance of the proactive control method.
In Chapter 5 the theoretical performance of various configurations of proactive con-trol system are examined. The predicted motions and accelerations are validated with experimental results in regular waves, at the beginning of this chapter, and a sensitiv-ity study is made to investigate the influence of different numerical input settings of the program. Subsequently, the theoretical performance of three different configurations are evaluated. The performance of two configurations using a only a single control vari-able and a configuration in which both varivari-ables can be modified simultaneously, are compared to benchmark simulations without the proactive control system.
Chapter 6 presents a proof of concept study in the towing tank of the Delft University of Technology to demonstrate the capabilities of the proactive system in a physical test environment. The experiments are conducted in irregular head wave conditions and an active interceptor is installed on the transom of the model to be able to modify its trim position during the tests. Temporary speed reductions of the model are effectuated by regulating the forward speed of the towing carriage. Both the instantaneous position of the interceptor and the change in forward speed of the model are governed by the real-time response predictions of the proactive control system.
Finally, the main conclusions are presented in Chapter 7 which also provides a num-ber of recommendations for future work.
2
O
PERATIONAL
C
ONTROL
The current research aims to develop an automated proactive control system that is specifically designed to reduce the vertical peak accelerations experienced by small and fast ship types operating in head wave conditions. In this context, it is considered im-portant to first briefly explain the physical nature of these undesired peak accelerations, as well as discussing existing ways to reduce them. Hereafter, the rationale is given for the potential performance increase of fast ships using the proposed operational control strategy. At the end of this chapter, the relevant research questions will be posted using the discussed literature and available methods as a framework. These research questions form the guiding principle for the remaining chapters presented in this thesis.
Section 2.1 elaborates on the non-linearities in the behaviour of fast ships in head waves. The origin of the non-linear behaviour is addressed and it discusses how this be-haviour differs from that of low speed displacement ships. Moreover, it identifies several known design methods that can improve the seakeeping behaviour of fast ships and it describes how this knowledge was put into practise during the development of several novel fast hull form designs. In addition, a visualisation tool is introduced in order to quantify and assess seakeeping performance of these fast ship types. Section 2.2 dis-cusses different operational control methods that can contribute to a faster and safer ride at sea. It first summarises the results obtained from full scale trials that were set up to gain insights in the important aspects of the limiting behaviour of fast ships for the crew. In addition, it visualises the throttle control actions applied by experienced helms-men during high speed operation in (large) waves. Next, the results of a numerical study carried out by van Deyzen (2014) are discussed. He designed a so-called proactive thrust control system which is inspired by this anticipatory manual throttle control principle. At the end of this section an explanation is given about how this proactive method could be further employed for an additional control variable, where not only the thrust is con-trolled, but in which the system also actuates active trim control devices to improve the operability of the ship. Finally, the various research questions are formulated which are presented in Section 2.3.
2.1.
N
ON
-
LINEARITIES IN THE BEHAVIOUR OF FAST SHIPS
Forward speed has an important effect on both the calm water and the seakeeping be-haviour of fast ships. At low speeds, i.e. Fr < 0.4, buoyancy forces still play a major role, also for fast ship types, and their behaviour is much like that of “normal” displacement ships. Some hydrodynamic lift may already be developed on the hull at this speed, how-ever, its effect on the calm water and the seakeeping behaviour is still relatively mild. With a further increase in speed a larger hydrodynamic pressure is developed and a big-ger fraction of the weight of the craft is carried by this hydrodynamic lift component. This results in a clearly observable effect in the still water orientation of the ship as the sinkage and trim may change considerably with respect to its reference position at zero forward speed. A large hydrodynamic lift is usually favourable for resistance considera-tions as the change in position reduces both the wetted surface and the submerged vol-ume of the ship, which consequently lowers the frictional and wave making resistance. Particularly the earlier fast hard chine hull forms, designed around 1930s and 40s, were specifically optimized for their calm water characteristics. The main design philosophy, at that time, was to develop a hull with minimal resistance resulting in a maximal for-ward speed, which was stimulated by the military applications of these fast ship types during World War II (Keuning, 1994).Hard chine ships built at the end of World War II could reach very high speeds in still water, not only because of their hydrodynamic efficient hull forms with low resistance, but also due to the large installed propulsive power as a result of the advances in diesel engine and gas turbine technology. However, they were also notorious for their rather severe seakeeping behaviour. Specific issues were; large heave and pitch motions and more importantly, the occurrence of very large impact accelerations when these ships were operated in a seaway. After World War II more emphasis was put on the rough-water performance of these fast craft as they were increasingly used for civilian appli-cations like Coast Guard patrol, Search and Rescue missions and passenger transport. Comfort, operability and safety became more important factors for these “new” mar-itime applications, which required a different type of hull than one that was designed for pure speed alone.
In the late sixties more research effort was put into the seakeeping behaviour of fast ships. Savitsky (1968) made a thorough analysis of the hydrodynamic problems associ-ated with high speed operations in waves. Later, Fridsma (1969; 1971) conducted note-worthy work by quantifying the importance of various design parameters using an ex-tensive series of systematic model tests. Three prismatic models with different deadrise angles were tested in various operating conditions in both regular and irregular waves. Fridsma noted that the response motions and the vertical peak accelerations in particu-lar were strongly non-linear related to the wave height. Van den Bosch (1970) carried out experiments with two more realistically shaped high speed models. The first model had a low deadrise angle of 12° and its geometry was quite similar to the parent hull form of the Clement and Blount (1963) series. The deadrise angle of the second model was dou-bled to 24° while all other dimensions were kept equal as far as possible. A considerable gain in seakeeping ability was achieved with the high deadrise hull at the cost of some additional resistance.
The realisation that almost all high speed operations at sea were limited by the mo-tions and particularly the vertical acceleration levels on board, resulted in a shifted de-sign focus towards improving the seakeeping characteristics of the ship instead of only optimising its calm water performance. This had a clear impact on the hull geometry. Keuning (1994) indicated that the actual shape of the bow has an important effect on the extent of the non-linearity of both the wave exciting forces and the hydrodynamic lift. Fast ships are usually limited in size meaning that the waves are comparatively large and thus induce large wave exciting forces on the hull. This results in large relative motions, the more so because the peak of the motion response occurs at longer (and thus larger) waves due to the high encounter frequency at high forward speed (Keuning, 1994). These large relative motions, in their turn, induce large variations of the submerged geometry, which is especially true for fast ships as they usually have distinct V-shaped sections in the forward part of the hull. The strong change in submerged volume yields rapid vari-ations in the hydrodynamic lift which, in itself, is the main cause of the strong peaks in the vertical accelerations.
These better insights into the physical cause of the wave impact phenomena on high speed craft were also utilised to improve hull form designs and this was done during the development of the Enlarged Ship Concept (ESC) introduced by Keuning and Pinkster (1995). The design philosophy of the ESC entails an increase of the ship length while all other dimensions and functionalities are kept as equal as possible. The “additional length” was added to the bow region which could subsequently be shaped from a strictly hydrodynamic point of view without having to compromise on other (interior) demands. The bow sections were designed with large deadrise angles, or little flare, which reduces the change in submerged volume in waves and consequently the severity of the wave impact events. Moreover, the lengthening of the forward part of the hull also resulted in the fact that the accommodation could be located further aft, in relative terms, where motion and accelerations levels are usually lower.
This design strategy later evolved into the development of the Axe Bow Concept (ABC) which is featured by an even stronger modification of the bow region (Keuning et al., 2001). This concept has almost vertical bow sections together with a downward sloping centre line resulting in a maximal draft at the forward perpendicular. The deep fore foot postpones bow emergence significantly while the change in submerged volume in waves was considerably lower due to the distinct shape of the bow sections. The deck sheer was greatly increased towards the bow to maintain the same reserve buoyancy and pitch restoring moment as the original benchmark design.
The ESC and especially the ABC showed significant improvements in the seakeeping characteristics compared to non-optimized conventional fast ship designs. A detailed performance comparison is given by Keuning (2006) which also presents a full descrip-tion of the design philosophies of both concepts. The ESC was first built in full scale at the end of the nineties and several years later the maiden voyage with an Axe Bow was made. Ever since their market introductions numerous ships have been sold and they are deployed for various high speed operations. Figure 2.1 displays two ships designed according to the ESC and ABC respectively.
(a) Damen Stan Patrol 4207 (b) Damen Fast Crew Supplier 5009
Figure 2.1 Photographs of an Enlarged Ship Concept (Figure 2.1a) and an Axe Bow Concept (Figure 2.1b). (Photographs courtesy of Damen Shipyards)
2.1.1.
P
ERFORMANCE ASSESSMENT BASED ON VERTICAL PEAK ACCELERA-TIONS
The improved insights into the non-linear behaviour of fast ships around the seventies, which differed considerably from the “known” linear behaviour of low speed displace-ment ships, also started a discussion about making meaningful measures to define and assess the ride quality of fast ships in waves. Payne (1976) indicated that conventional methods based on the Root Mean Square (RMS) values, used for displacement hulls, may not be applied to define comfort levels in cases where the motions of a vehicle suffer from “impacts” or impulsive velocity changes. The reason for this was that RMS values have no relation to the actual peak values in the acceleration signal of fast marine vehi-cles. Payne even warned that the sole use of RMS values in the ride assessment of high speed ships may cause lethal accelerations values to appear as being perfectly safe.
A rather extensive study to quantify and compare the peak acceleration responses of fast ships was conducted recently by Riley et al. (2013). This study presents a generalised approach to obtain the peaks in a time record of the rigid body accelerations measured on small and fast ships. The relevant vertical peak accelerations are identified by a search algorithm, which are subsequently sorted from the largest to smallest value and this tab-ulated data is plotted in a (cumulative) distribution plot. The work of Riley et al. (2013) includes both recommendations for low-pass filtering boundaries and detailed settings for peak finding algorithms.
At the Ship Hydrodynamic Laboratory of the Delft University of Technology extensive use is made of a so-called Rayleigh distribution plot in order to visualise and compare the non-linear response signals of fast ships in irregular waves. The computational search algorithm for the peaks and troughs is quite similar to the procedure described by Ri-ley et al. (2013), however a somewhat different method has been chosen to present the distributions themselves. The reasoning behind the particular visualisation method was initially given by Keuning (2006), and later de Jong (2011) gave a detailed description of the computational procedure and the theoretical assumptions behind the construction of the Rayleigh distribution plots. This type of plot will also be frequently used in the
current thesis to indicate the performance of various settings of the proactive control system. Hence, it seems to be of interest to first introduce the mathematical background and construction method of this particular visualisation tool.
RAYLEIGH DISTRIBUTION PLOTS
For this analysis it is necessary to make the assumption that the frequency spectrum of the sea waves under consideration is narrow banded and Gaussian distributed. Longuet-Higgins (1952) demonstrated that, when the above assumption is satisfied, the crest-to-trough wave height in the wave record follows a Rayleigh distribution. In addition, Longuet-Higgins (1952) considered that when the envelope of the wave elevation is sym-metrical with respect to the zero mean line, the probability-distribution of the wave am-plitudes is the same as the probability-distribution of the crest-to-troughs wave heights. This Rayleigh distribution is formulated according to,
f (x) = x σ2x exp µ − x σxp2 ¶ (2.1) in which x is the variable under consideration and σx is its standard deviation. The probability that an amplitude xa of the signal, with Rayleigh distributed minima and maxima, exceeds the threshold value a is given by,
P¡xa> a¢= Z∞ a x σ2x exp µ − x σxp2 ¶ dx = exp µ − a 2 2σ2x ¶ (2.2)
Equation (2.2) is subsequently used to modify the scaling of the horizontal axis in the Rayleigh distribution plot. This is done in such a way that the probability of exceedance of Rayleigh distributed crest and troughs are indicated with a straight line in the distribu-tion plot. This deformadistribu-tion of the horizontal scale can be obtained when Equadistribu-tion (2.2) is rewritten to,
a = σx q
−2ln¡P¡xa> a¢¢ (2.3) Finally, the horizontal scale is inverted which means that a probability of exceedance of 100 %, i.e. a = 0, intersects with the vertical axis.
For maritime engineering purposes it is often desirable to make a statistically-based approximation of the maximum wave height at, for instance, once every 1000 wave cy-cles. This maximum value is mostly used as a design condition which can be directly related to the associated significant wave amplitude, for Rayleigh distributed peaks and troughs, by means of Equation (2.2). This significant amplitude xa1/3of a signal is related
to the standard deviation σxaccording to,
xa1/3= 2σx (2.4)
The deformed horizontal axis of the Rayleigh distribution plot ensures that the signif-icant value of the wave amplitude and the peak value occurring once every 1000 wave
cycles are represented on a single straight (inclined) line. The same type of plot can sub-sequently be used to visualise the response behaviour of a ship in the particular wave conditions. These response motions will also be Rayleigh distributed as long as the ship behaves in a linear fashion. The maximum amplitude values, for this linear case, are thus automatically embedded in the significant value of the response signal and the dis-tribution of the crests and troughs will again appear as a straight line in the Rayleigh distribution plot. However, this straightforward relation does not exist whenever the ship displays a (strong) non-linear behaviour. The distribution of the response minima and maxima, for this non-linear case, will not correspond to the theoretical Rayleigh dis-tributed line and these deviating values can consequently not be linked to the significant value of the signal. Hence, the full probability distribution is required for non-linear re-sponse signals. The level of deviation from the straight Rayleigh line can also be used as a visual indicator for the extent of the non-linearity of the particular response signal. PEAK IDENTIFICATION TECHNIQUE FOR ACCELERATION SIGNALS
The measured or computed time traces need to be analysed to identify the relevant peaks and troughs used for the distribution plot. This is a delicate process, especially in the case of measured acceleration signals, as these signals are highly sensitive for high fre-quency noise which distorts the rigid body peak accelerations that are typically of in-terest. A proper data analysis starts with a reliable measurement set up and Zseleczky (2012) gives a rather useful overview of the important steps for acceleration measure-ments, including accelerometer selection and mounting, data acquisition equipment and the use of anti-aliasing filtering. After data collection a post-processing step is re-quired to remove the high-frequency content from the signal using low-pass filtering techniques. The cut-off frequency of the low-pass filter has to be carefully selected in order to eliminate electrical noise or structural vibrations as much as possible while the shape of the rigid body peak accelerations should be retained, as too aggressive filtering affects the magnitude of these wave impact events. McCue et al. (2012) presented a study of the effects of different filter types, filter order and cut-off frequencies on the charac-terisation of vertical peak accelerations of fast ships. Riley et al. (2013) recommended a 10 Hz cut-off frequency, as a starting point, for full scale acceleration measurements of high speed ships, although he also suggest that a Fourier spectrum analysis should be performed to confirm whether this value is suited for the individual data set.
Not all high frequency vibrations are removed from the measured signal after low-pass filtering and there will be an abundance of lower peak values when the local max-ima of these artificial oscillations are counted by the peak search algorithm. This will skew the statistics towards zero, giving a distorted distribution of the response signal of the ship. Grimsley et al. (2010) presented four different numerical peak identification methods that can help to extract the relevant peaks from a filtered acceleration record of a planing craft. Unfortunately, no performance ranking was given between the de-scribed methods. However, the most common peak identification method in today’s use is to apply a horizontal (time) threshold which is sometimes preceded by a minimum vertical peak (acceleration) distance (de Jong, 2011; Riley et al., 2013; Zseleczky, 2012).
The procedure applied here adopts the recommended 10 Hz cut-off frequency at full scale from Riley et al. (2013). This frequency boundary is also much applied in past pub-lished work and it may be desirable to use the same cut-off frequency as it enables one
to make consistent comparisons with historical data (McCue et al., 2012). The horizon-tal threshold in the peak search algorithm was set equal to half the encounter period, following the suggested value of McCue et al. (2012). This combination resulted in satis-fying results in which the significant amplitude of the filtered time trace agreed well with the significant value of the counted peaks.
APPLICATION OF THE VISUALISATION TOOL
An example of the application of these Rayleigh distributions plots is presented in Fig-ure 2.2. These figFig-ures are obtained from de Jong (2011), who re-evaluated the model data of an ESC hull form that was originally tested by Keuning and van Walree (2006). The main full scale particulars of this ship are listed in Table 2.1. Figure 2.2 shows the distribution of the measured wave realisation together with the response motions and accelerations of the model in a wave spectrum with a peak period of Tp = 7.8 s and a significant wave height of H1/3= 3.5 m. The forward speed of the model during these
tests was set to Fr∇= 2.03, corresponding to 35 knots in full scale. In addition to the
dis-tributions of the signals, the straight inclined “Rayleigh lines” are also indicated in every figure, which are based on the significant values of the measured signals. The significant amplitude of each signal is presented with the horizontal dashed line which intersects with the Rayleigh line at exp(−2) ≈ 13.5 %, by definition. The Rayleigh distributions are provided with 95 % confidence bounds which were determined by de Jong (2011).
Table 2.1 Main particulars ESC hull shape
Designation Symbol Full scale Unit
Length waterline LW L 55.00 m
Beam on waterline BW L 8.46 m
Draft T 2.66 m
Volume of displacement ∇ 516.00 m3
Longitudinal position CG LCG 22.49 m Pitch radius of gyration ky y 13.75 m
The distribution of the wave elevation is presented in Figure 2.2a. From this figure it can be observed that the crest and troughs are Rayleigh distributed, as their distribu-tions agree (for the most part) with the straight Rayleigh line. The results of the heave and pitch motions are indicated in Figures 2.2b and 2.2c, in which a certain degree of non-linearity can already be distinguished. The crests in heave motions are consider-able larger than the troughs for the lower probability of exceedance values. The opposite can be noticed from the pitch motions in which the troughs, representing bow up am-plitudes, are significantly larger in magnitude relatively to the crests in the signal. How-ever, the most profound non-linear behaviour can be seen in the vertical accelerations at the bow displayed in Figure 2.2d. The troughs in the acceleration signal are limited by the gravitational acceleration augmented by a small acceleration component due to the pitching motion of the ship. The crests, on the other hand, contain values as large as five times the gravitational acceleration and the large deviation from the Rayleigh line clearly indicates the great extent of non-linearity in this acceleration signal.
100 50 20 10 5 2 1 0.1 0 1 2 3 4 Probability of exceedance [%] ζC G [m]
(a) Wave elevation
100 50 20 10 5 2 1 0.1 0 1 2 3 4 Probability of exceedance [%] z [m] (b) Heave motion 100 50 20 10 5 2 1 0.1 0 2 4 6 8 Probability of exceedance [%] θ [deg] (c) Pitch motion 100 50 20 10 5 2 1 0.1 0 2 4 6 Probability of exceedance [%] az b o w /g [-] Crests Troughs Rayleigh line Significant value
(d) Vertical acceleration at the bow
Figure 2.2 Rayleigh distribution plots of peaks and troughs and significant amplitudes of the waves, ship motions and vertical accelerations. The waves have a peak period of Tp= 7.8 s and a significant
wave height of H1/3= 3.5 m. The forward speed of the ship is equal to Fr∇= 2.03. From de Jong
2.2.
O
PERATIONAL CONTROL VARIABLES
The major modifications to the geometry of the hull, especially to the bow region, signif-icantly improved the seakeeping performance of fast ships over the last decades. These performance improvements are usually assessed by a comparative method in which a conventional and an optimized model-sized hull form are both tested in a research en-vironment at (nearly) identical operating conditions. The difference between the be-haviour of both designs is often quantified and used as a measure to express their mutual ranking.
A question that remains, after this scientific model-scale comparison, is how the new and improved full-scale design is utilized in reality by the crews out on open seas. On the one hand, they could choose to increase safety and comfort levels on board and sail at the same average forward speeds as the previous “non-optimized” design. The expected performance increase in seakeeping is in this case quite similar to indicated research results. On the other hand, they could also increase the average forward speed and “ex-change” the better seakeeping behaviour for this additional gain in forward speed. This will enhance the operability of the ship as a higher forward speed can be achieved, how-ever it may also partly or even completely undo the intended improvements in comfort and safety levels for the crew.
Although it is not the aim of this research to give advice on which method should be followed by the crew, as it is probably highly dependent on the particular type of oper-ation and no single comprehensive advice can be formulated, it does signify the impor-tance of the operational element irrespective of the actual design and performance of the ship itself.
Two operational control variables are discussed in the upcoming sections. In sec-tion 2.2.1 a more detailed descripsec-tion is given about the incentives for the helmsmen to change the forward speed of the ship and it indicates the possible effects it may have on the seakeeping performance. Section 2.2.2 discusses various motives for trim control on fast ships and it describes common methods and mechanisms to be able to effectuate a change in trim.
2.2.1.
S
PEED CONTROLExtensive series of full-scale trials have been carried out by the Ship Hydromechanics Laboratory of the Delft University of Technology to determine the most important limit-ing criteria for the operation of fast ships in irregular waves. These trials were conducted with different ship types to include the effect of the ship length into the developed crite-ria. From the various trials the following data was available:
• Comparative full scale side-by-side tests of two SAR ships with an overall length of 14 m and 15 m respectively and a maximum forward speed of 31 knots. These ships are part of the fleet of the Royal Netherlands Sea Rescue Institution and trials were conducted in the North Sea area (Ooms and Keuning, 1997).
• Full scale measurements onboard the Valiant; a 42 m fast patrol boat designed according to the ESC and capable of reaching 27 knots. The ship is owned by UK Customs and it was tested off the West Coast of Scotland (Keuning and van Walree, 2006).
All ships were instrumented and monitored to measure the motions along with the acceleration levels in the wheelhouse and on the bow. In addition, throttle settings were recorded to detect the voluntary speed reductions applied by the crew. Environmental conditions such as wind and waves were obtained and video footage was made from the standpoint of the bridge. The ships were operated by a professional crew during the trials and they were instructed to perform their usual tasks under real circumstances. A team of experts from the Delft University of Technology was present at all tests which was extended with specialists from the Maritime Research Institute Netherlands and the US Coast Guard during the trials with the Valiant. Apart from the measured signals the observations from the expert teams and interviews with the crews were also included in the data analysis.
The observations and conclusions of these tests are presented in Ooms and Keuning (1997) and Keuning and van Walree (2006), and the main findings can be summarized as:
• All crews imposed voluntary speed reduction at roughly the same conditions, in terms of ships motions and vertical peak accelerations.
• The incentive for the voluntary speed reduction was not based on the prevailing magnitude of the significant amplitude of the motions or vertical accelerations; but the occurrence of a single big peak, or a small group of larger peaks, in the vertical accelerations signal proved to be the most important reason to voluntarily reduce speed.
• The main intention of the speed reductions was to prevent any potential consecu-tive impacts from happening.
Also a number of differences were observed between the SAR ships and the larger ESC type:
• The occurrence of large peaks in the vertical accelerations at the bow and the re-sulting structural vibrations due to an wave impact proved to be the main deter-mining factor to reduce speed in the case of the larger Valiant design. The typical maximum accepted vertical acceleration values were 0.8 g in the wheelhouse and 2.0 g at the bow.
• A higher level of acceleration was accepted by the crew on the smaller SAR ships and the level in the wheelhouse was far more important for the crew. The accepted value in the wheelhouse was typically 1.3 g and at the bow a value of 2.5 g was considered acceptable.
• Manual throttle control was only applied by the helmsmen on the smaller SAR boats, in order to evade the higher peaks in the vertical accelerations.
The presented acceleration values are only used for indicative purposes and should not be applied in a generalised way. The ultimate acceptable levels are probably de-pendent on a highly complex interaction of numerous variabilities, including motion characteristics like frequency, amplitude and duration of the oscillation, but it may also be compounded by different psychological and physiological aspects of the persons on-board (Townsend et al., 2012). These personal aspects do not only differ among people
but they may also vary on a day-to-day basis for a certain individual. In that sense, it may be convenient for the crew to have a more direct way of controlling the magnitudes of these vertical peak accelerations for each separate mission.
INFLUENCE OF THE SPEED ON THE VERTICAL PEAK ACCELERATIONS
A (temporary) reduction in speed lowers the total hydrodynamic lift on the hull, however this does not necessary mean that the change in hydrodynamic lift in waves is also re-duced under all circumstances. For instance, a different speed is likely to cause a change in the running sinkage and trim and this new position may result in worse ship motion characteristics. Moreover, a lower speed will also decrease the wave encounter frequency which could be moved closer to the resonance motion region of the ship. This could in-crease the vertical peak accelerations as the larger motions may result in larger variations in the hydrodynamic lift.
An example that a lower speed does not automatically have to lead to lower maxi-mum impact levels is illustrated in Figure 2.3. This figure compares the acceleration lev-els at the bow measured during a systematic series of model tests at different constant forward speeds and in various irregular wave conditions. The tests were conducted by Keuning and van Walree (2006) with the ESC model that was already introduced in Sec-tion 2.1.1. The hull shape of this scientific ESC model is quite similar to the geometry of the commercial Valiant design. The plots present the measured acceleration distribu-tions for three different wave condidistribu-tions with a constant spectral peak period Tp = 7.8 s and variable significant wave heights H1/3ranging from 2.0, 3.0 and 3.5 m. Figure 2.3a
shows the acceleration distributions for three different constant forward speeds at the lowest wave height. It can be seen that a higher speed gives higher acceleration levels in the particular circumstances. However, looking at the two larger wave heights shown in Figures 2.3b and 2.3c; a different type of trend can be noticed. Higher speeds still give larger significant values of the vertical accelerations (at the 13.5 % mark), but in the maximum region a clear reduction can be seen in the accelerations while going from the intermediate to the highest forward speed. Thus, in these operating conditions it can actually be beneficial to increase speed to obtain lower maximum acceleration val-ues. However, it must be noted that the highest speed of Fr∇= 2.90 corresponds to a
rather “challenging” value of 50 knots for the full scale 55 m prototype ship which lim-its the practical use of this method. In addition, the maximum values in these model tests are already 50 % to 150 % higher than the “acceptable” values obtained from the full scale trials, so it may be doubtful if the actual ship would ever be exposed to these conditions. It does, however, underline fact that these highly non-linear systems can ex-hibit rather complex motion behaviours in which it can be rather difficult to foresee the consequences of a change in speed under all circumstances.
MANUAL THROT TLE CONTROL
To increase insights into the manual throttle control actions of the operators on the smaller SAR ships, additional full scale trials were conducted by van Deyzen et al. (2012a) and van Deyzen (2014). These tests were specifically set up to quantify the effects of the manual throttle control actions on the vertical accelerations and to study the different throttle control strategies among different operators. The trials were carried out with
100 50 20 10 5 2 1 0.1 0 2 4 6 Probability of exceedance [%] az b o w /g [-]
(a) Wave condition 1: H1/3= 2.0 m and Tp= 7.8 s
100 50 20 10 5 2 1 0.1 0 2 4 6 Probability of exceedance [%] az b o w /g [-]
(b) Wave condition 2: H1/3= 3.0 m and Tp= 7.8 s
100 50 20 10 5 2 1 0.1 0 2 4 6 Probability of exceedance [%] az b o w /g [-] Fr∇= 1.45 Fr∇= 2.03 Fr∇= 2.90
(c) Wave condition 3: H1/3= 3.5 m and Tp= 7.8 s
Figure 2.3 Rayleigh distribution plots of the vertical accelerations at the bow of the ESC hull form for various forward speeds and in different wave conditions
Figure 2.4 Search and Rescue boat from the “Arie Visser” class (Photograph courtesy of Flying Focus)
Table 2.2 Main particulars “Arie Visser” class
Designation Symbol Full scale Unit
Length overall LAO 18.80 m Length waterline LW L 14.50 m Beam on waterline BW L 4.14 m Draft T 1.04 m Volume of displacement ∇ 27.32 m3 Vertical position CG KG 1.25 m Longitudinal position CG LCG 6.10 m Pitch radius of gyration ky y 4.50 m
two similar SAR ships of the “Arie Visser” class of the Royal Netherlands Sea Rescue Insti-tution and the tests were divided over six days including calm, moderate and rough sea states. Four different professional helmsmen were asked to operate the ships during the separate test runs. All trials were performed on the North Sea off the coast of the Hoek van Holland and IJmuiden rescue stations. Figure 2.4 presents a photograph of the Arie
Visser which is part of the eponymous series. The main particulars of the “Arie Visser”
class are indicated in Table 2.2.
The layout and the type of instrumentation, on these ships, was quite similar to the previous trials with the two other SAR ships presented in Ooms and Keuning (1997). Dur-ing this new series special attention was devoted to quantify the effects of the throttle control actions. This was done by comparing two different run types. For the first type, a desired constant speed was selected and the operator was not allowed to make any changes to the throttle handle. During the second run type, throttle control actions were allowed according to their usual operating procedure.
15 20 25 30 Vs [kn] 0 5 10 15 20 25 30 35 −1 0 1 2 Time [s] az b o w /g [-] 0 25 50 75 100 Th rot tle [%]
Figure 2.5 Throttle control action registered on the Jeanine Parqui of the “Arie Visser” class. The waves have a significant wave height of H1/3= 1.0 m and a peak period of Tp= 3.9 s. From van Deyzen (2014)
Figure 2.5 shows a registered sample of one of these normal operating trials and it includes the time traces of the throttle position, the forward speed and the vertical ac-celerations at the bow. The operator seemingly expected a wave impact and reduced the throttle setting from about 95 % to 50 %, which resulted in an almost instant response in ship speed with an ultimate speed drop of about 10 knots over a 5 s period. Unfortu-nately, the effect of the control action on the vertical accelerations cannot be determined based on this short time fragment as no comparison is available to the situation without throttle control.
In an attempt to visualise the performance increase due to active throttle control, the distributions of three different runs were compared by van Deyzen (2014). Two of these trials were carried out with “active” throttle and the other was performed with a fixed throttle setting selected by the helmsman. The results are displayed in Figure 2.6 and each measurement lasted about 14 minutes. The average speed of the constant throt-tle test was 23 knots, while the average speed during the tests with throtthrot-tle control was 27 knots. A higher average forward speed could be attained during the trials with throttle control, but at the same time also higher vertical accelerations were measured compared to the fixed throttle reference run, see Figure 2.6. This makes it rather difficult to assess the performance increase of the evasive control actions of the operator. Van Deyzen sug-gested that these larger acceleration values could be attributed to misjudgements or loss of concentration of the helmsman. Another explanation could be that the crew has a dif-ferent perception of the acceleration level due to a certain “feeling of being in control” when the operator is allowed to make throttle adjustments (van Deyzen et al., 2012a).
100 50 20 10 5 2 1 0.5 0.1 0 1 2 3 4 5 6 Probability of exceedance [%] az b o w /g [-] Fixed throttle, Vs= 23 kn Throttle control #1, Vav= 27 kn Throttle control #2, Vav= 27 kn
Figure 2.6 Rayleigh distribution plot of the vertical acceleration; comparing a trial at fixed throttle with two runs using throttle control on the Jeanine Parqui of the “Arie Visser” class. The waves had a signifi-cant wave height of H1/3= 1.0 m and peak period of Tp= 3.9 s. From van Deyzen (2014).
Also the motivation of the person in command could play an important role during the measurements “to show his capabilities in terms of throttle control”, while on the other hand a conservative fixed throttle setting could be selected “just to be on the safe side” during the trials in which the operator was not allowed to intervene. In any case, it seems that the human factor has a rather big influence on the outcome of the results, which dis-turbs the comparison, thereby making it difficult to formulate any conclusive statements about the performance increase of the manual throttle control method.
PROACTIVE THRUST CONTROL
To demonstrate the basic idea and increase insights into the theoretical effect of the tem-porary speed reductions on the maximum vertical peak accelerations, a fundamental mathematical model was developed by van Deyzen et al. (2012b) based on a non-linear mass-damper-spring system. This mathematical model was set up to characterise the typical motion behaviour of a fast ship in head waves, including an assumed non-linear coupling between the horizontal translating speed and the motions in vertical direction. During time simulations with this mathematical model, continuous short-term predic-tions were made at regular time intervals in order to predict the near-future behaviour. These numerical forecasts represented the line of sight of the human operators, which could accordingly be used to determine the temporary reductions in speed whenever the (predicted) vertical accelerations exceeded a predefined threshold value. A control action was selected from an additional series of predictions at various discrete thrust
0 5 10 15 20 25 30 35 40 45 50 −1 0 1 2 3 4 5 Time [s] az b o w /g [-]
Figure 2.7 Vertical accelerations registered on the Koos van Messel of the “Arie Visser” class. The waves have a significant wave height of H1/3= 2.65 m and a peak period of Tp= 5.7 s. From van Deyzen (2014)
settings. The vertical acceleration level of the implemented condition had to comply the acceleration threshold at the expense of a minimal speed loss. With this numerical ap-proach the random effects of the human interventions were eliminated and replaced by a more consistent automated controller. The numerical results indicated that temporary speed reductions could significantly lower the maximum acceleration levels compared to benchmark simulations at an equivalent constant translating speed.
The mass-damper-spring system was subsequently exchanged for a more dedicated simulation model in van Deyzen (2014) that could be used to predict the “actual” be-haviour of fast ships operating in a particular sea state. Moreover, this more advanced system was also developed as an initial step towards a fully automated controller that can be installed aboard and which eventually could take over the manual throttle con-trol actions applied by the helmsman.
Reducing the peak acceleration levels with an automated system is however not an easy task to perform. One of the issues is the large difference in time scale of the fierce peak disturbances in relation to the relativity slow motion response of the ship on a (speed) control action. Figure 2.7 shows one of the larger wave impact events mea-sured during the rough sea state trials described by van Deyzen et al. (2012a). It can be seen that the impact only lasts for a few tenths of a second, while the deceleration of a ship typically takes an order of magnitude longer as is indicated in Figure 2.5. The helmsman uses his visual observations of the incident wave in front of the bow to an-ticipate the behaviour of the ship. An experienced operator may evade the foreseeable wave impacts by applying an appropriate reduction in speed before the actual impact takes place. However, this manual throttle control method requires a rather intense con-centration throughout the full mission duration. Fatigue issues due to long, complex missions or harsh sea states can significantly reduce the alertness of the operator. More-over, the view from the bridge may be partly or even complete obstructed by excessive spray, severe weather conditions or darkness, which can make it rather difficult (or even impossible) to apply the well-timed throttle adjustments.
An automated system can alleviate the human responsibility and it may exploit the full potential capabilities of this anticipatory control principle. Van Deyzen (2014) iden-tified three essential components for a so-called proactive thrust control system:
• A wave sensor to obtain the incident wave information in front of the bow
• A simulation program that computes the near-future response predictions in a very short period of time
• A control system that intervenes and implements the required control action Although all three components are required for a proactive system on a ship out on open seas, an initial selection was made between them at the start of this research. The development of the wave sensor is not taken into consideration just yet, as the first prior-ity is to study the feasibilprior-ity and potential performance of the proactive control system in a research environment and alternative solutions exists for measuring the wave profile in these laboratory conditions (see Chapter 6). When the performance of the proactive system in these research conditions is satisfactory, further investigations into the devel-opment of the wave sensor itself may be undertaken according to the specific design requirements that may result from the current study.
Figure 2.8 shows a comparison of the numerical results of the proactive thrust con-trol system presented by van Deyzen (2014). This Rayleigh distribution plot shows two individual simulations with a SAR ship of the “Arie Visser” class; one at a constant speed and the other with the proactive thrust control system. In both conditions a similar av-erage forward speed is attained while the differences in the maximum acceleration level are evident. 100 50 20 10 5 2 1 0.5 0.1 0 1 2 3 4 5 6 Probability of exceedance [%] az b o w /g [-] Benchmark, Vs= 20.0 kn Thrust, Vav= 19.8 kn
Figure 2.8 Rayleigh distribution plot of the vertical accelerations on a SAR ship of the “Arie Visser” class; com-paring a benchmark case at a constant speed of 20 knots with the proactive thrust control config-uration using a prediction horizon of 4 seconds. The waves in these simulations have a significant wave height of H1/3= 2.15 m and a peak period of Tp= 7.5 s. From van Deyzen (2014).
2.2.2.
T
RIM CONTROLThe promising results of the proactive thrust control system raises the question whether this anticipatory method can be further employed for additional control variables. Be-sides the forward speed, another well-known operational parameter that can have an important effect on the magnitude of the vertical peak accelerations is the instantaneous trim position of the ship.
A considerable amount of research has been carried out in order to quantify the ef-fects of the running trim at high forward speeds. The vast majority of these studies focus on the powering performance in the calm water, as they attempt to reduce the resis-tance due to an optimisation of the trim angle (Jensen and Latorre, 1992; Millward, 1976; Mossaad et al., 2005; Savitsky and Brown, 1976; Tsai et al., 2003). While other work may be more related to safety issues, for example, to avoid undesired oscillatory heave and pitch motions in calm water (porpoising) that can manifest due to dynamic instabilities in the hydrodynamic lift on hulls that operate at high trim angles (Blount and Codega, 1992; Savitsky, 1964). Low trim angles, on the other hand, may cause hazardous effects like bow diving and/or loss of directional stability (Dawson and Blount, 2002). Finally, the running trim can also have an important effect on the ride quality of fast ships in waves. Fridsma (1969) indicated that an increase in trim from 4° to 6° produced 50 % to 100 % higher values in the vertical acceleration signal on a hull with a constant deadrise angle of 20° operating at a speed coefficient of CV= 2.7 in regular head waves.
A change in the trim of a ship is usually applied with some sort of hydrodynamic lift-generating device. One of the earliest trim control mechanisms was based on the “Lürssen effect” and this system was applied on various classes of the German
Schnell-boote, or S-boats, which served during World War II. These boats were equipped with
two small auxiliary rudders, called lifting rudders, situated at either side of the main rudder and these rudders were angled outboard by 30° during the acceleration phase of the ship. This created a low-pressure separation zone, behind each of the lifting rudders, which was ventilated with air from the surface at the transom edge (Saunders, 1957). The air-filled cavities behind the rudders were now at atmospheric pressure while the region ahead of the rudders was characterised by a high hydrodynamic pressure. This effect re-sulted in a net upward force at the stern of the ship, reducing the trim and increasing its speed for the same amount of thrust. The angle of the lifting rudders could be decreased again to about 17° once the separation zones were filled with air, without losing the lift-ing effect, while at the same time the induced drag of the liftlift-ing rudders was reduced allowing even higher ship speeds.
The overall system was, however, quite complex both mechanically as in operation and it was not widely used after World War II as wedges, transom flaps and later intercep-tors became favoured for trim control applications (Blount, 2013). Wedges and transom flaps are unified in their hydrodynamic action, but these mechanisms differ in their ar-rangements. A wedge is usually located in a recess underneath the transom area where it is integrated into the hull bottom plating of the ship. Transom flaps, on the other hand, generally extend aft of the hull itself. Both mechanisms consist of a flat plate which is fitted at an angle relative to the ship buttock lines. The interceptor is a vertical mounted plate on the transom which extends below the hull, perpendicular to the incoming flow. An impression of both a transom flap and an interceptor are given in Figure 2.9.
