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International Conference on Technology &
Operation of Offshore Support Vessels
20 - 21 September 2005
National University of Singapore
Republic of Singapore
Jointly Organised by
The Joint Branch of the Royal Institution of Naval Architects and the Institute of Marine Engineering, Science and
Technology (Singapore)
Centre for Offshore Research & Engineering National University of Singapore
Main Sponsor
Co-Sponsors
In Conjunction withCop
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ISBN: 981-05-4245-3
Final Programme
Keynote
Offshore Support Vessels - A New Horizon
by James B. Liebertz, President & COO, American Bureau Of Shipping, Pacific
Division, Singapore
Paper 1
Bollard Pull Guarantee in terms of Propeller and Hull Design
by Sverre Tonheim, Scana Volda, Norway
Paper 2
New Demands on Offshore Terminal Tugs
by Robert G. Allan, Robert Allan Ltd, Canada
Paper 3
Tug Behaviour in Waves as Important Factor in the Operability of Offshore
LNG Berthing and Offloading Operations
by Bas Buchner, Olaf Waals (MARIN) and Pieter Dierx (TU-Delft), The
Netherlands
Paper 4
Assessment of Intact Stability Requirements of Offshore Supply Vessels for
Safe Operation
by Zafer Ayaz and Dracos Vassalos, The Ship Stability Research Centre,
Universities of Glasgow and Strathclyde, UK
Paper 5
Bridge System Safety: NAUT-OSV
by Hans Ramsvik, Det Norske Veritas, Norway
Paper 6
Modern Offshore Support Vessels: Class and Statutory Perspectives
by Ahmad Sarthy and Ham Joon Lok, American Bureau of Shipping, Singapore
Paper 7
Trends in Design of Liquid Handling System in Supply Vessels
by Anders Eide, Ing.Per Gjerdrum a/s (PG Marine), Norway
Paper 8
Is the OSV a Chemical Bomb?
by BH Wong, Ezra Marine Pte Ltd, Singapore
Paper 9
Integration of Pumping Systems in Supply Vessels matching IMO
Resolution
A. 673 (16)
by Anders Eide, Ing. Per Gjerdrum a/s (PG Marine), Norway
Paper 10
Roles of Offshore Support Vessels in Construction Activities for Oil & Gas
Industry
by Ng Eng Bin, Ng Yang Nee Elsie and Donikon Fajar, WorleyParsons Pte Ltd,
Singapore
Paper 11
The Geometric Aspects of Ship Position Coordinates Determination
Accuracy
by Andrzej Banachowicz, Maritime Academy of Gdynia / Adam Wolski, Maritime
University of Szczecin, Poland
Paper 12
Recent Advances in Dynamic Positioning, Positioning Technology
by Colin Soanes, The Dynamic Positioning Centre, Singapore
Paper 13
Improve Operability and Safety of DP Vessels Using Hybrid Control
Concept
by Asgeir Sorensen, Department of Marine Technology, NTNU, Norway,
Ser Tong Quek and Trong Dong Nguyen, Centre for Offshore Research and
Engineering, NUS, Singapore
Paper 14
Using Doppler Logs for Safer DP
by R I Stephens, A.J. Meahan and J. Flint, ALSTOM Power Conversion Ltd, UK
Paper 15
Hardware-in-the-Loop Simulation for Testing of DP Vessels
by Olav Egeland, Centre for Ships and Ocean Structures, NTNU and Roger
Skjetne, Marine Cybernetics AS, Norway
Paper 16
Ship Controllability Criteria and their Application for Small Vessels
by Albert Nazarov, Hull Co Ltd, Thailand
Paper 17
Rudder System
by Mr Henning Kuhlmann, Becker Marine Systems, Germany
Paper 18
Electric Propulsion in Field Support Vessels
by Alf Kare Adnanes and Gunnar Hide, ABB Marine, Norway
Paper 19
Propulsion of Offshore Support Vessels
by Jens Ring Nielsen, Rasmus Mandrup Jeppesen and Ege Lundgren, MAN B&W
Diesel A/S, Denmark
Paper 20
ACERT ® Technology: How Caterpillar Engines meet Current and Future
Emission Limits
by Stephen Phillips, Caterpillar Marine Asia Pacific Pte Ltd, Singapore
Paper 21
Offshore Supply Vessels equipped with Voith Schneider ® Propellers
by Ivo Beu, Voith Turbo Marine GmbH & Co. KG, Germany
Paper 22
Electric Propulsion in Ice-going Vessels
by Arto Uuskallio, ABB Marine, Finland and Alf Kare Adnanes, ABB Marine,
Norway
Paper 23
Electrical Propulsion Systems for Offshore Support Vessels
by Ivar Andersen, Offshore & Marine ASA, Norway
Paper 24
New Diesel-Electric Propulsion Concepts for OSVs
by Joachim Muller, SCHOTTEL GmbH & Co. KG, Germany
Paper 25
Common Rail Diesel Fuel Injection Technology
OSV Singapore 2005
Jointly organized by Joint Branch of RINA-IMarEST Singapore and CORE. 20-21 September 2005
Offshore Support Vessels – A New Horizon
Keynote Address by Mr James B. Liebertz
President & Chief Operating Officer, ABS Pacific Division
Good morning distinguished participants,
Ladies and gentlemen,
I am pleased to join you today for the opening of the International Conference on
Technology and Operations of Offshore Support Vessels, organized by:
- The Joint Branch of RINA and IMarEST, and
- Center of Offshore Research and Engineering at NUS.
Looking at the program, it is clear that this conference has brought together owners,
operators, engineers, equipment specialists and researchers from around the world.
Significantly, it symbolizes the current interest in and high level of activity for
offshore support vessels. There could be no more appropriate venue for such a
gathering than here in Singapore, the center for offshore-related construction and
management in the Asian theater..
Singapore has a tradition of building OSVs beginning in the 1970s. At that time
Singapore’s shipbuilding and marine industry was in its infancy. But it took on the
challenge of serving the emerging offshore industry in the region with a particular
focus on building mobile offshore drilling units and offshore supply vessels. We all
know that these activities declined in the 1980s as oil prices dropped. Many well
known names did not survive the lean years. Those that did are reaping the fruits now,
as demand for both MODUs and OSVs is once again very strong.
In the Pacific region, my own organization, ABS, is very active in providing
classification services for OSVs that have been built or are under construction at a
large number of yards including established builders like Keppel Singmarine, Jaya,
Nam Cheong and Pan United, among others, as well as newer entrants such as ABG,
Bharati, Fujian Marwei, Piasau Slipway, and Yantai Raffles. We also work with
almost all the major OSV owners, both here in Asia and worldwide: Tidewater of the
US; Great Eastern of India; Chuan Hup, Jaya and Pacific Richfield of Singapore; Hadi
al Hammam and Halul Offshore of the Middle East, again to name just a few. This
exposure to so many different designs, yards and operators gives us a very
comprehensive understanding of the evolution, function and future of these
remarkable crafts.
The origin of today’s OSVs can be found in the Gulf of Mexico – when oil
exploration moved offshore in the 1950s. Then, surplus World War II vessels,
wooden fishing boats, and shrimp trawlers were used to supply offshore rigs with
cement, mud, spare parts, crews, fuel and food. In 1955, Alden and John Laborde
developed the first purpose-built vessel to supply offshore rigs and platforms. It
featured a bow wheelhouse and a long flat afterdeck that became the standard for
offshore supply vessels. If I can hazard a guess, the pioneers may have modeled OSVs
on a pickup truck – rugged, versatile, and capable of delivering goods and people to
the frontiers. The Labordes must be intensely proud today as their innovation has
done more than survive. It has become an industry standard that is an integral element
of the continuing search for energy resources under the world’s oceans.
Today we are once again seeing an almost fevered level of offshore activity as the
price of oil, and concerns over the adequacy of future supplies increase. Its impact on
the OSV sector has been to create a second wave, similar to the crest of the 1970s
after the long trough of the intervening years.
This wave is driven by two key factors: the first is the relatively low level of new
buildings in the OSV sector over the last two decades or so, which has precipitated a
dire need for fleet renewal. The second factor is obviously the unprecedented hike in
oil prices I mentioned a moment ago, which in turn, has stimulated the current high
level of offshore exploration.
Insiders and keen market observers would hasten to add that the first factor was felt
way before the second was even imminent. Indeed, the second factor has been a
godsend for the visionaries who ordered their vessels in the relatively lean time. They
can now more than reap the benefits of their vision.
According to Douglas Westwood – a leading energy consultant – 34% of oil
production in 2004 is from offshore; this will rise to 39% by 2015. Of this, product in
shallow waters has been declining, which naturally drives exploration and production
into deeper waters. It is projected that by 2010, most, if not all, of the growth offshore
will likely be from deepwater. And by 2015 deepwater will account for 25% of all
offshore production, a dramatic increase from the current 9% share.
Against this backdrop, the new generation of OSVs contains many features that cater
to the needs of deepwater support operations.
Consider the following:
- Deepwater drilling rigs will need more drilling supplies than shallow water
rigs.
- Deepwater locations are further out from shore bases.
- Deepwater anchor handling means heavier anchors and longer hawsers.
- Deepwater environments demand greater maneuverability and
position-keeping ability.
- Increasing safety and environmental standards for offshore exploration and
production will demand implementation of emergency preparedness such as
standby and rescue, fire-fighting and anti-pollution capabilities.
Next Paper >> Return to Main Content
The result is that the new generation of OSVs, built to support deepwater exploration
and production, are larger, more powerful, more maneuverable and are outfitted to
respond to a wide range of potential emergencies.
The engineering of all of these capabilities onto a relatively small platform is a
challenge that has been met through the adoption of modern technologies, such as
electric propulsion and integrated control systems. This new generation of OSVs is a
far cry from the one it replaces.
If anything, the OSV of today is more than a pickup truck; it’s like something
outfitted for James Bond.
Many of the papers that will be presented, including one that my own colleagues will
present on classification services for OSVs , will cover these developments in detail.
Given the current demand for this new generation of innovative, and often highly
specialized OSVs, this conference could not be more timely. I believe it provides a
valuable opportunity for an open exchange of views among the various participants in
this exciting field. I am confident that the discussions of the next two days will prove
useful and will lead to enhancements in the art and the science of OSV design and
operation. On this note, I wish all participants a most successful and rewarding
conference.
OSV Singapore 2005
Jointly organized by Joint Branch of RINA-IMarEST Singapore and CORE. 20-21 September 2005
Bollard Pull Guarantee in terms of Propeller and Hull Design
Sverre Kristoffer Tonheim
Navel Architekt 1969 Sales Manager
Scana Volda as, POB 105, N-6101 Volda Norway
ABSTRACTS
The paper’s objective is to emphasize that the final measured bollard pull of a tug depend on the SUM of many single factors. This factors are of different influence, and can be put in different groups, related to: 1) the hull design, 2) the propeller design, 3) the bollard pull definition, and 4) the bollard pull test conditions.
The paper do not present new scientific research, or knowledge on any of these factors. Just for illuminating the totality, the paper gives a brief description of the different influencing factors.
By this, the aim is to show that a bollard pull guarantee is something that involves many contributors.
Also the general level of bollard pull guarantees is briefly discussed in the paper.
An accurate prediction of the bollard pull test figure, is principally impossible. Therefore the guarantee figure should be set at a level which correspond with the given premises, and be based on well established documentation. A guarantee should also include suitable margins for the
uncertainties in calculations and in the bollard pull test result.
Besides the physical elements, a bollard pull guarantee also involve a commercial element. The propeller supplier have traditionally been the guarantor, even if the main premises for the potential bollard pull are beyond his influence. This paradox could perhaps be a subject of further discussion among the players in the propulsion business.
INTRODUCTION
Through experience in the propulsion business since 1969, the author have seen a growing interest and focus on the Bollard Pull topic.
Even if the term Bollard pull have been globally used for decades, and definitions have been submitted by several institutions for many years, there still exist some confusion in the market. This circumstance may represent a problem for the Guarantor, who is often confronting unspecified and rather diffuse guarantee demands.
Earlier it was quite common to give Bollard pull Predictions based on the actual constellation. Also Bollard pull Guarantees on specified figures were given. However, the requested figures were normally found quite sensible, and it was seldom a big problem to give the guarantee asked for.
Since the 60’s, propeller suppliers, have used the published results for Wageningen Ka / 19a combination, as basis for thrust predictions and design of nozzle propellers.
After some experience, the Kaplan blade outline was modified, while the nozzle section was used more or less unchanged.
Because of this uniform basis of calculation and design, there was established a universal level of bollard pull, accepted by the different actors in the market as an established “Normal”. By experience during so many years, this particular nozzle propeller have also been very well correlated with full scale measurements.
During the latest decades, more and more often a Bollard pull Guarantee is required, and normally there is a penalty clause involved in case that the guarantee figure is failed.
This fact must be accepted as a consequence of the strict operation conditions and competition in the Offshore / tug market to day.
During the latest years, a trend of increased guarantee level have been observed. This increased level is often found very optimistic, not to say unrealistic. At least, they are considered too marginal as a real sound guarantee figure.
This trend must also be understandable, because every tugboat owner will like to have the highest possible bollard pull for the lowest possible investment.
propellers. Especially the nozzles have been subject to investigation during last years, and several new nozzles have been promoted. They may be optimized for either free running, or for bollard conditions. The common designation for those new nozzles is “High efficient nozzles”
Most of these new nozzles proclaim improvements in the magnitude of 3-4 %, compared to the 19a. Nozzle.
One of this new nozzles proclaims improvements up to 8-10 %, both in bollard and free running, with one and the same nozzle profile.
Even much higher improvements are published from comparative full scale tests, where old 19a nozzles with the old propeller blades, have been replaced by the new nozzles and new blades.
It may be that such impressive reports have promoted a general rise of the bollard pull guarantee level in the market.
However, for none of those new nozzles exist the officially published, and very comprehensive model test documentation, as for the Wageningen series. Because the performance of the new nozzles is not generally available, they have not obtained the same status as a universal standard like the Wageningen Ka/19a and Ka/37 series.
None of these two classic nozzles pretend to be the ultimate nozzle sections. Originally they were developed with genuine airfoils, but was thereafter simplified from practical reasons, with a strait conical outer surface to adapt for easy production and low price.
Therefore it is no big surprise if the all-curved nozzle sections shows improvements. The open question is; how large is this improvement. Is it documented good enough, and is it enough to justify the increased cost ?
It is no secret that identical ships, and even the same ship in different tests, may show great
differences in bollard pull test figures. This fact demonstrate generally a large uncertainty and lack of repeatability as far as bollard pull tests are concerned.
The guarantee demand should be based on reliable documentation, and should also define the type of nozzle section desired.
DEFINITIONS OF BOLLARD PULL.
The general meaning of the term Bollard Pull is the force in the towing line, measured during a full scale bollard pull test. The term is widely used in common language, without further precision.
In a guarantee context there is a need of a more precise definition, and some of the most actual
definitions are listed below:
1) Sustained (or continuous) Bollard Pull
This is the mean towing force, recorded every 30 sec., during a period of 5 minutes, at 100 % MCR power, all applied for propulsion.
2) Maximum static Bollard Pull
This is the mean of the 2 highest subsequent towing force recorded every 30 sec. at 100% MCR power.
3) Maximum Bollard Pull
This is the highest single figure recorded during the test, also with 100 % MCR power.
Even several definitions exists in the market, but this is of less importance and will not be mentioned here.
During a Bollard pull test, the highest peak value is normally obtained shortly after the full power is applied on the propeller. The readings may fluctuate, but the mean value will decrease
asymptotically toward the sustained figure.
When using the general term “Bollard Pull”, the Sustained- or Continuous pull is proposed as the specific meaning.
This is the figure offering the best repeatability and comparability, and also the figure indicating the tug’s real service capability in the best way.
BOLLARD PULL IN GENERAL.
The real physical quantity of Bollard Pull, BPo, is created by the thrust force To, supplied by the propeller system, corrected for the interaction effects, caused by the presence of the hull.
The relation is: BPo = To (1 – t),
To, in this equation is the so-called “open water” thrust, which is the thrust that the actual design propeller would perform working in homogenous, parallel, flow.
To must be predicted, either by an open water model test, or by theoretical- or empirical based calculations.
t, in the equation is the thrust deduction factor for bollard condition. Considering the bollard condition, this factor incorporate all the effects caused by the presence of the hull, struts, bushings, rudders. etc. (In bollard condition the ship speed is zero, and the classic propulsion factors used in free sailing constellations, like wake, thrust reduction, relative rotation efficiency, are not defined. The term t for bollard condition will therefore have a somewhat different content than in free sailing considerations.)
The equation above look very simple. However, each of the quantities is a complex subject, influenced by many different factors.
This different factors will be briefly described in the following.
As most tugs are equipped with CPP Nozzle propellers, and the view will be more or less the same for all types, the following review is adapted to this propeller type only.
FACTORS INFLUENCING ON THE OPEN WATER THRUST.
The factors influencing on the open water thrust can be classified in two main groups, the Primary and the Secondary factors. The Primary factors are the initial basic factors, given in the ship specification, or implicit in the ship’s design, and in the ship’s service profile.
The secondary group is conditional factors, which are derived based on the primary ones, in such a way to optimize the propulsion system.
1) THE PRIMARY FACTORS,
a) Installed Engine power.
The engine power is the most decisive of the thrust-influencing factors. For bollard test considerations, the engine characteristic is not so important as the loading point will be 100% power / 100% rpm.
b) The engine speed.
For a direct coupled propeller, the engine rpm will be of outmost importance for the obtainable thrust, because it will be also the propeller rpm. With a geared solution, the
engine rpm is of minor importance, only influencing on the gear / transmission losses to a modest extent.
c) The ship speed.
The ship resistance curve, implicit in the hull geometry, will give the
potential ship speed. Normally the contract ship speed is given with indication of the corresponding service power, ship draft and eventual sea margins.
d) The effective wake fraction, and the nominal wake flow distribution in the propeller plan. This may be found by model tests, or else it must be estimated based on the hull parameters and lines, and by experience from similar hulls .
e) The contract bollard pull.
The required bollard pull will influence on the propeller design in order to obtain the corresponding necessary bollard thrust.
f) Propeller main type (open or nozzle propeller)
This is also a very important factor as far as the bollard thrust is concerned. The nozzle propeller will perform better than an open propeller of the same diameter.
With the same power and diameter the hydrodynamic load on the propeller itself will be reduced, and this is also beneficial for the cavitations and vibration conditions etc.
g) Type of nozzle section.
Traditionally the Wageningen 19 a, have been applied with different L/D ratios (also the Dn/Dp will vary correspondingly), adapted for different purposes. For tugs and trawlers, long sections have been used, and for free sailing ships shorter sections.
For typical push-pull tugs, the Wageningen 37 section have been used. This nozzle perform less than 19a in forward, but more in astern.
Besides this two classical types, different new types are available.
h) The Propeller Aperture
The propeller diameter is of the same importance for the bollard thrust as the power.
Within certain limits, the bollard thrust will be approx equal as long as the product of power and diameter is constant.
For open water bollard thrust, the larger propeller will always be the superior. Therefore, if the propeller rpm is not already fixed with the engine, but can be freely adjusted, the practice
should be to select the largest possible diameter. (under careful consideration of the actual constrains given by the propeller aperture, to secure the sufficient clearances.)
i) The number of blades.
The blade number is of smaller importance for the bollard thrust, but the influence depend on the circumstance. If the propeller rpm is fixed, (direct coupled) the optimum diameter will be larger the lower the blade number. In this case there will be a slightly thrust gain by using the lower blade number with its larger diameter. On the contrary, if the diameter is the fixed parameter ( the max, possible in the aperture), than the propeller speed will be adjusted to the lower optimum as the blade number increases. In such cases the bollard thrust will be only marginally influenced, or nothing at all, by the blade number.
j) Classification rules / Ice classes.
The classification rules will influence the thrust marginally by defining the blade thickness. Especially in cases with high ice class this influence can be of some significance. Also the propeller hub diameter will depend on the ice class, and this will also influence on the thrust.
k) Other hydrodynamic design criterions.
Normally there are given several types of propeller design criterions with different priority. Such criterions may be related to bollard pull, efficiency in different service conditions, cavitations, noise and vibration levels, etc. etc.
Highest priority for tugs may be the bollard pull, but even for tugs the bollard
condition is not the only service condition. Typical multi purpose “working vessels” must first of all carry out their different purposes efficiently, but also undertake sailings. Some ships have a service profile representing a very wide power range. Specially in combination with constant propeller speed, the low power conditions must be considered carefully to avoid pressure side cavitations. The final propeller became normally more or less of a compromise, not really optimized for one of the particular conditions, but hopefully the best possible over all solution for the whole range of conditions.
l) The arrangement of the propulsion system in the ship.
The propulsion arrangement may differ very much from a single screw tug with a short shaft length, to a twin screw anchor handling vessel with the engines located in the fore ship. The gear and shaft losses will therefore be different.
This is the main primary factors, constituting the criterions for the propeller design.
It should be noted that these Primary factors, are totally decisive for the obtainable bollard thrust level.
2) THE SECONDARY FACTORS.
Based on the above given primary factors, the detailed propeller design can go on, to decide the secondary factors, which means the geometrical main data of the propeller.
a) The propeller design speed (rpm)
Based on the factors given in the primary group, the first and foremost is to decide the optimum propeller speed in such a way that the design criterions and priorities are taken into proper consideration.
This is an iterative process which involves more or less all propeller parameters in the same process.
b) Blade area ratio EAR, / radial area distribution.
The blade area ratio is decided based on cavitations analysis for the different service conditions.
The efficiency decrease with increasing area, so the area should not be taken larger than necessary to avoid cavitations with a certain margin. The radial distribution of area should be co-ordinated with the desired radial distribution of circulation.
c) Skew back / Rake angle
Skew and rake are partly linked together, and they are to be decided based on the actual wake pattern, the radial load distribution, and the vibration criterion. For several reasons a very high skew is not used on nozzle propellers. Skew and rake are not very influential on the bollard thrust.
d) The design pitch / radial pitch distribution / “off design” pitch distribution.. The design pitch is decided to fit in the best way with the service profile. For nozzle
propellers with high bollard pull, the radial pitch distribution should be constant or even increased toward the tip, combined with long tip cord length. A high tip load, combined with small clearances in the nozzle, increases the nozzle circulation and thereby the efficiency. However a too high tip load may result in tip vortex cavitations, noise, vibration and erosion inside the nozzle. For CPP the “off design” conditions must be specially considered and analysed, because when the pitch setting is changed, also the radial pitch distribution, the
blade sections, the tip clearance and the tip longitudinal position in the nozzle also changes.
e) Blade section camber. / Camber distribution
The radial section camber must be fixed according to the desired radial propeller load to secure ideal working condition for each individual radial section. Because of inhomogeneous wake flow, the section camber may be reduced below the ideal camber, and a part of the section lift is obtained by angle of attack.
The cord-wise camber distribution is selected to obtain a uniform pressure distribution along the section cord. At the leading edge, modifications may be necessary to counteract cavitations in a variable wake flow, and specially to avoid pressure side cavitations in low power conditions.
f) Blade thickness / cord wise thickness distribution
The blade section thickness is calculated for the maximum load condition according to the Classification rules. The cord wise thickness distribution is selected to obtain uniform pressure distribution. High ice class will result in thicker blades. It may be necessary to increase also the section cord length to maintain a sub critical t/c ratio and to avoid cavitations. Large blade thickness and large area will reduce the efficiency and the bollard thrust slightly.
g) Propeller hub diameter.
The propeller hub is selected based on a strength analyses of the different hub parts exposed to the internal and the external forces and torques acting on the hub. Also here the actual ice class is of great influence. The aim is to incorporate in all propellers so called
progressive strength, i.e. the strength of the propeller increases all the way from the blade tip to the propeller shaft.
The ratio Hub diameter / Propeller diameter will therefore depend on the actual nominal propeller load, and of the ice class.
A high hub/diam. ratio, will reduce slightly the efficiency and the bollard thrust.
h) Propeller tip clearances in the nozzle.
In order to maintain a high hydrodynamic tip load and increase the nozzle’s thrust
contribution, both beneficial for high efficiency and bollard thrust, the clearance between the propeller tip and the nozzle inner surface must be kept on a minimum. For a CPP the
minimum effective clearance is constrained because the blade tip must pass zero pitch during manoeuvrings , and also because the blades shall be demountable inside the nozzle.
i) The Nozzle section profile / Lengths – Diameter ratio.
The Wageningen sections no. 19a, or no. 37, may be used with different L/D ratios dependant on the actual average thrust coefficient, and type of tug. The higher the thrust coefficients, the longer nozzle is favourable.
Other nozzle sections are also available for different purposes.
THE FINAL PROPELLER / NOZZLE DESIGN.
With the secondary factors the propeller main geometry is developed. The detailed geometrical description and production tolerances for propeller and nozzle are decided, and the propeller design is thereby completed.
The influence from the secondary factors on the obtainable bollard thrust level, is very limited compared to the major influence of the primary factors.
.
PREDICTION OF THE OPEN WATER BOLLARD THRUST
The open water performance of the final design propeller may be verified by an open water model test. However, model test of design propellers are not often carried out.
Bollard model test is normally undertaken at an early stage with a model of the actual hull equipped with a fixed pitch stock propeller / nozzle.
It is important that the used stock propeller/nozzle at least have the same main parameters as the actual propeller, so that the test is run at the correct model conditions.
If the stock propeller differ significant from the actual design propeller, the test is less valuable, but one can at least obtain an approximate thrust deduction factor.
In most cases there is no model test carried out at all, and the prediction of bollard thrust must be based entirely on calculations. If such calculations are based on any of the standard model series corrections must be made for the geometrical differences, and for scale effect.
Such open water calculations may also be done by special hydrodynamic programs based on the actual design geometry of the propeller and nozzle.
CONDITIONS INFLUENCING ON THE THRUST DEDUCTION FACTOR.
As mentioned above, the thrust deduction factor for bollard condition is best obtained from a bollard model test.
lines and the propeller /nozzle/shafting/rudder arrangement. For the propeller designer it is important to receive such documentation, enabling him to evaluate the construction before a bollard pull guarantee shall be confirmed.
The thrust deduction fraction in bollard condition is less than in free running conditions, but the same factors are generally influencing, although to different extent.
a) Blunt waterline angels or steeply rising buttocks in the aft body will increase the t.
The low pressure zone induced by the propeller and nozzle will act on the hull surface plane with a large angle to the direction of thrust, resulting in a sucking effect from the hull.
To reduce this factor, all hull surfaces in front, above, and aft of the propeller should have the least possible angle with the thrust direction.
b) The distance between the propeller and the hull, in forward and upward direction should be as large as possible to reduce t.
c) The rudder type and thickness will influence on the t. Specially blunt leading edges and fish tail formed trailing edges, may increase t markedly.
d) The clearance between the hull surface and the nozzle should be sufficient to secure a free circulation around the nozzle section, all around the circumference.
e) The attachment for the nozzle should be arranged with slender aerodynamic profiles, oriented parallel to the induced flow. Wide closed boxes and heel pieces, especially with sharp edges and corners, will increase t and may also induce cavitations and erosion problems.
f) Side thruster openings close in front of the propeller will increase t.
g) All types of flow separation and turbulence will also increase t significantly. Too sharp curvature and changes of flow direction must be avoided.
h) Inclination of the propeller shaft and nozzle will reduce the horizontal force component and normally increase the hull surface angle with the force direction. The increased angle relative to the inflow direction, will reduce the thrust compared to ideal open water condition.
i) Diverging shafts in twin screw ships will have the same effect as inclined shaft. In principal also converging shafts will have such an effect, but as the shafts here are more parallel to the
aft body flow direction, moderate converging shafts may be favourable.
j) Rounded Transom stern, exposed to the induced propeller flow may induce drag by coanda effects, which will also increase t.
For normal aft body lines and propeller arrangements, the thrust deduction factor for bollard condition, may vary between 0,04 and 0,08.
Because the viscous fraction of t is considered to be small compared to the pressure component, there is no scale effect on t.
THE PREDICTED BOLLARD PULL.
With the estimated open water thrust To for the actual design propeller, and the estimated bollard thrust deduction coefficient t, the Bollard Pull can be estimated.
The result is (hopefully) identical with the real physical force, supposed to exist in the towing line under ideal bollard test conditions.
THE VERIFIED BOLLARD PULL.
The verification of the Bollard Pull figure is done by a full scale bollard pull test.
The final recorded figure will be a result of the real force in the towing line, and the accuracy of the measurement. The real force in the towing line is the ideal force, influenced by the actual environmental disturbances, and deviations caused by the test set up.
Provided that the theoretical prediction is correct, the propeller and nozzle are fabricated within the tolerances , the arrangement and mountings in the ship are proper, the test conditions are
ideal, the engine power is correct, and the force is measured by an accurate scale, the verified figure would be identical with the predicted figure.
On the contrary, if one or several of the a.m. premises are not fulfilled, the verified figure will deviate from the predicted figure.
BOLLARD TEST CONDITIONS
The Bollard test procedure and environmental conditions are not topic of this paper, but for the sake of completion, it should be shortly mentioned.
Generally one can say that any restriction and disturbance of the free propeller stream, caused by the surrounding test basin, will influence on the real pull figure. Also forces created by current, wind and waves will influence, in most cases with a negative effect on the pull.
Generally, the real engine power may deviate substantially from the nominal power. High air temperature and humidity may reduce the power, and this will not be reflected by reading the fuel rack, the rpm indicator and the exhaust temperature alone.
The real power delivered to the propeller should be measured by a correct torque / rpm measurement directly on the tail shaft.
The test set up should be arranged with one single free horizontal towing line of sufficient length and the scale should be calibrated to an accurate level. The rudder should be in mid ship position and should not be moved during the test period. Course corrections obtained by use of side thrusters will also disturb the bollard pull, and must be avoided.
Another important matter for the power stability during the test, is to disengage the overload protection system, normally applied on CPP systems.
The submersion of the propeller must be sufficient to avoid air suction, which may occur at high propeller load and low advance coefficient as in the bollard condition. Air suction will be
manifested by pull fluctuations. If necessary, a moderate trim by stern should be allowed, even if this will reduce the horizontal force component marginally.
(A forward tilt of the nozzle may reduce the potential risk for air suction.)
The different Bollard Pull test Codes are giving detailed guidance for the test procedure. Also advice for calculation of eventual corrections are given in those test codes.
BOLLARD TEST ACCURACY.
It is considered that, even with very good environmental conditions, proper test procedure, and good measuring accuracy, the final recorded bollard pull figure will have a margin of error of approx. 5 %. In most test cases, the ideal conditions are missing, and the margin of error is correspondingly higher.
A SOUND BOLLARD PULL GUARANTEE FIGURE.
A guarantee should generally be a thrust worthy term.
The guarantor normally know nothing about the test basin or the conditions of the future bollard
pull test. Accordingly, a sound bollard pull guarantee figure, given on a general basis, should be fixed say 7- 8 % below the estimated figure, for compensating the uncertainties in the test.
CONCLUDING REMARKS
The propeller supplier is traditionally the honoured guarantor of bollard pull.
Often a very high bollard pull guarantee may be requested at an early stage, long before the necessary hull information is available. Not seldom, the guaranty figure required from the propeller supplier may even be higher than the official guarantee figure in the ship contract Such guarantee demands is naturally found very problematic, and will normally be refused.
Obviously the bollard thrust, delivered by the propeller, is the definite main component of the bollard pull. Only a small fraction, the thrust deduction, is considered associated directly with the presence of the hull. This may lead to the conclusion that the hull design is of minor importance compared to the propeller design.
This conclusion is false because not only the thrust deduction factor, but also the potential propeller thrust, is given with the ship concept.
The propeller supplier is normally involved in the project at a later stage, when the main premises are already fixed, and his possibility of influence is thereby very limited.
The propeller supplier is also without influence on the bollard test conditions.
After the above arguments, it should be reason to ask why the guarantee burden should be placed on the propeller supplier.
The aim is not to change this tradition, but rather to show that the responsibility for the final bollard pull result is a mutual concern.
If this fact is clearly realized by the involved parties, the result could be secured by a good co-operation between the ship designer and the propeller supplier at an earlier project stage.
Finally, the most important factors in order to secure the bollard pull guarantee, is to install sufficient engine power, and to arrange for a sufficient propeller size, corresponding reasonably with the bollard pull demand. Suitable margins for the unavoidable uncertainties should also be considered.
The bollard pull test should be carried out in accordance with acceptable test codes, and under controlled conditions in a good test basin.
As a basis for issuance of bollard pull guarantees, all available model tests and hull data, including the intended type of nozzle, should be submitted to the guarantor.
REFERENCES:
M.W.C. Oosterveld, Wake adapted ducted propellers, NSMB publ. No. 345. 1970
Bollard Pull Trial code for Tugs with Steeprop propulsors.
Steerprop Ltd, March
2001
G. Kuiper, The Wageningen propeller series, MARIN publication 92-001
OSV Singapore 2005
Jointly organized by Joint Branch of RINA-IMarEST Singapore and CORE 20-21 September 2005
New Demands on Offshore Terminal Tugs
Robert G. Allan, P. Eng., FRINA, FSNAME
President, Robert Allan Ltd., Canada
Key Words: Offshore, Terminal Operations, Tugboats, Seakeeping, Research, Towing
ABSTRACT
This paper describes the international trend to the construction of oil and gas terminals
in ever more exposed locations, with the resultant increase of demand on tugs, mooring vessels,
and mooring systems to accommodate the berthing of ships under higher wind and sea-state
con-ditions than had previously ever been considered for ship-handling operations. The lack of
pub-lished information on how tugs can perform in such conditions has led to demands for more
ab-solute information on the performance degradation of tugs in a seaway, and simultaneously
de-mands for larger and more powerful tugboats to perform these functions. The author's company
has conducted independent research into the subject of tug performance in a seaway, as well as
evaluating tug-ship interaction forces in a seaway, leading to a new generation of powerful tug
designs capable of performing reliably under more extreme sea-states. Two new classes of
off-shore tug resulting from this research are described in detail, as well as the results of some of the
research.
Tugboats have been providing ship-handling services at offshore installations for many
years. The majority of these are essentially harbour tugs, pressed into service in locations where
sea-states may exceed the tugs capacity to operate effectively. The extreme motions result in
reduced reliability of the ship-handling operations, and cause fatigue and stress on the crews,
with resultant risk of injury to personnel, or damage to the tugs, ships or terminals. In many
cases, to overcome the shortfalls of typical tugs in these exposed locations, operators have
re-sorted to the use of larger, powerful AHTS or OSV class vessels, which in most cases represent
too much equipment for the job, and an inappropriate use of the resources.
As a new generation of LNG tankers looms on the horizon, the proposed terminals for
these ships (with their perceived risks) will in most instances be positioned in locations remote
from conventional sheltered harbours with their associated population concentrations. Almost by
default these terminals will be located in positions more exposed to high winds and sea-states
than has ever been the case before. These high-sided, large tankers demand higher-powered and
capable tugs to escort them into the terminal areas, and to dock them safely. Typical
perform-ance specifications currently in circulation describe providing reliable continuous tug services in
wind conditions at least up to 30 knots, with associated sea-states at least up to significant wave
height (H
s) of two metres.
The scope of functions required of these tugs frequently include at least the following:
•
Escort towing capability in approaches to the terminals
•Ship handling/ship-assist functions at terminal
•
Tail boat functions at loading buoys or FSOs
•Fire-fighting
•
Oil spill response
•Oil spill recovery
•Pilot boarding
In addition, the tug crew must be able to connect to LNG tanker shipside fittings at
rela-tively low freeboards, restricting the bow height of the tug.
There is, to the author's knowledge, very little published data available on the
inter-action forces between a tug and its tow in close-coupled berthing operations in a seaway. There
is a growing body of operational experience however, which clearly indicates that the forces
in-volved are extreme, and exceed the capability of many deck fittings and winches etc., on both the
tug and the ship. There is clearly a need to develop a better understanding of these interaction
forces in both pulling and pushing modes, and indeed there has been a recent Joint Industry
Pro-ject (JIP), designated as the SAFETUG proPro-ject, initiated by MARIN of the Netherlands to
per-form an extensive research program aimed at filling this void. The results of a pilot study into
this topic were published by Buchner et al [1] in 2005. The JIP is sponsored by a combination of
oil companies, towing companies, some shipyards and design consultants, including the author's
company. The MARIN tests, illustrated in Figure No. 1, unfortunately used a style of tug which
is not typical of a modern ship-handling tug, and which was not self-propelled, hence the results
are of questionable value in application to understanding modern tug behaviour.
Nevertheless the initial tests performed at MARIN as a precursor to the SAFETUG project
high-lighted some of the very important issues of the problem, including the range of tug-ship contact
areas, the need for elasticity in the towline/connection system, and the need to provide maximum
possible motion-damping characteristics in the tug itself.
Figure No. 1 Pilot Program Model Tests, Performed at MARIN
Prior to initiation of the SAFETUG Project, Robert Allan Ltd. determined that there was
a need to provide a higher standard of tug design for this type of application. A contract award
in 2004 for a major terminal/escort tug design provided the opportunity to test ideas for a new
generation of hull form for operation in exposed conditions. In previous model and full-scale
tests on very high-performance escort tug designs, namely the AVT tugs Ajax (Figure No. 2) and
Velox (Figure No. 3) designed and built for Ostensjo Rederi AS of Norway, it was observed that
the hull characteristics developed to provide maximum possible indirect steering forces resulted
in exceptional sea-keeping and motion-damping capabilities. The large skegs obviously
contrib-ute to reduced roll motions, however the sponsoned hull form acts dramatically to reduce roll
amplitudes and roll accelerations.
Figure No. 2 AVT 40/90 Class Escort Tug Ajax
Figure No. 3 AVT 37/60 Class Escort Tug Velox
These results led to the conclusion that a tug designed for offshore operations, even if
NOT
required to perform escort duties, would benefit significantly from the same type of hull
form.
Robert Allan Ltd. embarked upon a series of model tests for the new ASD 36/70-E
Class terminal escort tug, under contract to IRSHAD of the UAE to prove the merit of the
appli-cation of advanced hull forms to more conventional tug appliappli-cations. The mandate for the new
generation of IRSHAD tugs was to perform full escort and terminal operations throughout their
operating region, including terminals in which sea-states up to H
s= 2.0 metres are common. The
test program included measurements of conventional performance parameters such as speed in
calm water and in waves, seakeeping in waves, and the indirect steering and braking
perform-ance. Although the results of these tests provided absolute numbers which were considered to be
very encouraging, the Client had no frame of reference by which to judge whether or not the
model test results were better than experienced on their existing tugs. Therefore a second series
of seakeeping tests were initiated using the model of the proposed new tug side-by-side against a
model of one of the existing vessels for which full-scale performance, albeit generally empirical
evidence, was available.
The two models were ballasted to simulate drafts and GMs corresponding to actual
op-erating conditions for the existing tug, which had a typical load GM of 2.0 metres, and to
pre-dicted drafts and associated GM values of 2.5 metres prepre-dicted for the new tug. However in
or-der to measure the benefit of the hull form characteristics independent of GM value the new
de-sign was also ballasted to a GM of 2.0 metres. Tests were performed in head, quarter and beam
seas up to 3.0 m Hs.
Figure No. 4 and No. 5 illustrate the results of these tests at H
s= 2.0 metres, indicating
roll amplitudes and wheelhouse height roll accelerations respectively. The data shows very
clearly the potential benefits that can be achieved through the use of advanced hull forms. The
amplitude of roll motions were reduced by 48%, and roll accelerations were reduced by 68% in
the new design. Interestingly, the motions of the advanced hull do not appear to be particularly
sensitive to GM, and the lower value of GM, not surprisingly shows lower accelerations, but
more notable were the lower roll amplitudes also associated with the lower GM values.
Advanced vs. Conventional Hull Forms
Figure No. 4
Figure No. 5
Comparison of Roll Amplitudes:
Comparison of Roll Amplitudes:
Advanced vs. Conventional Hull Form
Advanced vs. Conventional Hull Form
The
new
ASD 36/70-E Class tug (Figure No. 6) is currently being tendered for
con-struction internationally, and a design contract award is expected in the 3rd quarter 2005.
Par-ticulars and performance predictions of the new tug are as follows:
•
Length, overall
-
35.80 metres
•
Beam, maximum
-
13.50 metres
•
Depth, moulded
- 7.06 metres
•
Hull draft, moulded
- 4.05 metres
•
Maximum draft, over-drives - 5.75 metres
•
Power -
4,400
kW
•
Bollard pull
-
85 tonnes, maximum
- 70 tonnes minimum, at 90% MCR, in design
ambient conditions
•
Indirect steering force (F
s)
-
120 tonnes at 10 knots
•
Indirect braking force (F
b)
-
148 tonnes at 10 knots
It is believed that this class of tugs will be the first true purpose-designed escort tugs
built in an ASD configuration, as opposed to a number of large offshore ASD tugs claiming to be
escort capable but not really able to perform indirect towage effectively. This new design
em-bodies the results of an extended program of research and development into improving escort tug
performance undertaken by the Author's company, in collaboration with the Institute of Marine
Dynamics in St. John's, Newfoundland, as described by Allan and Molyneux [2]. That research
program has ultimately led to testing vessel designs with all of the conventional VSP, Z-tractor
and ASD propulsion arrangements has demonstrated clearly that with proper hull design, an
ef-fective escort tug can be designed regardless of propulsion system configuration. Figure No. 7,
extracted from [3] indicates the relative indirect steering performance of the three configurations
at 10 knots.
Figure No. 7 Indirect Steering Performance of Three Alternate Escort Tug Configurations
Tug-Ship Interactions
Seizing the opportunity to use the model for the ASD 36/70-E Class tugs, a series of
additional tests were performed, simulating the interaction between a tug and a tanker in a
sea-way, measuring the connecting line forces for varying sea-states as a function of tug power,
posi-tion, orientaposi-tion, line length and angle, and towline elasticity. Figure No. 8 and No. 9 illustrate
the test setup and program elements.
Figure No. 8 Test Arrangement: ASD 36/70-E in Stern Sea
The most salient results of this quick test program were as follows:
•
Highest towline forces occur when the tug is at 0° or 45° to the wave direction, and
the tanker is at 90°, 45°, or 135° to wave direction
•
Towline force increases with tug proximity to the tanker, and with the steepness of
the towline
•
Maximum RMS values of towline force were measured at 3.85 times Bollard Pull.
It is estimated that the associated maximum single amplitude towline force could be
as much as 12 times Bollard Pull
•
Reducing the height of the connection point significantly reduces the forces in the
towline
Figure No. 10 illustrates the maximum line forces in calm, 2 metre and 3 metre seas, with no
spring in the system, tug at 90° to waves, tanker at 90° degrees.
Figure No. 11 illustrates the RMS values of roll angle in 3 metre seas, with the tug at 0° degrees
(stern to) the seas and tanker at 90° (beam to) to the seas.
The results of this test series will be fed directly into the SAFETUG Project.
Figure No. 10 Comparison of Towline Forces for Calm, 2 Metre, and
3 Metre Seas for Varying Powers and Connection Positions
Figure No. 11 Comparison of Roll Angles in 3 Metre Seas for Varying
Varying Powers and Connection Positions
In 2003, in response to the demands of clients for major offshore support tugs, Robert
Allan Ltd., commenced the development of a new generation of powerful offshore support tugs,
which were designated as the RAmpage Class. There was a distinct demand from industry for
tugs of 90 to 100 tonne BP, capable of operating in severe offshore conditions, but which did not
necessarily incorporate an escort capability or large cargo-handling capabilities in the design.
This would result in smaller, less expensive vessels than the class of AHTS and OSVs currently
used for this type of service. A series of powerful RAmpage Class tug designs were developed
throughout 2003/2004 incorporating a variety of propulsion and deck machinery options. The
distinguishing features of these tugs include:
•
High power/size ratios
•
Incorporation of fore deck handling winches akin to more conventional
ship-handling tugs
•
Large anchor-handling/towing winches and line control gear aft
•Some liquid (fuel and potable water) cargo capacity
The first of the new RAmpage 5000-ZM class of tug, the Seabulk Angola (Figure No. 12
and No. 13) was delivered in May 2005 from Labroy Shipbuilders of Singapore to the
Owners Seabulk Angola Inc. Vessel particulars are as follows:
•
Length, overall
-
49.5 metres
•
Beam, maximum
-
49.5 metres
•
Depth, moulded
- 6.75 metres
•
Maximum draft, over-drives
- 6.40 metres
•
Power
-
5,940
kW
•
Bollard pull
-
100 tonnes, minimum
•
Free running speed
- 14 knots
•
Maximum deadweight
- 900 tonnes
Figure No. 12 Seabulk Angola - First of Class RAmpage 5000 Tug
Figure No. 13 Seabulk Angola - First of Class RAmpage 5000 Tug
The use of Z-drive propulsion, in conjunction with a powerful bow-thruster provides the
RAmpage Class tug with the maximum possible manoeuvring capability, and the thrust
neces-sary to provide effective berthing and hold-back capability for the largest tankers.
Figure No. 14 illustrates the General Arrangement of the RAmpage 5000 Class tug.
Several variations on the RAmpage theme have been developed, including options up to 65
metres in length, and various powering arrangements, including diesel-electric power and
con-ventional propeller/nozzle arrangements. Figure No. 15 and No. 16 illustrate some recent
pro-posals for RAmpage 6000 and RAmpage 6500 Class tugs
Figure No. 15 RAmpage 6000 Class, 60 Metre Offshore Tug
Figure No. 16 RAmpage 6500 Class, 65 Metre Offshore Tug
CONCLUSIONS
The advent of offshore LNG and oil loading/offloading terminals in areas exposed to
more severe wind and sea-states than was previously the norm has created a demand for a new
generation of tugboats capable of providing sustained control over the attended ships in these
conditions. The docking and control of ships in sea-states of 2 metres significant wave height or
more results in line forces at least ten times the tug BP. This demands a new approach to hull
de-sign and motion control in the basic hull form, and tug-ship connection systems capable of
ac-commodating the extremely high shock loads imparted. Research by the author's company into
the motions of various hull forms has enabled the development of new high-performance tug
de-signs capable of providing a much higher standard of performance in offshore operations than
was previously possible. Motions and accelerations can at least be halved by intelligent and
in-novative design approaches as illustrated by the ASD-E Class hull designs. Similarly, powerful
tugs of more compact dimensions such as the RAmpage Class can replace large AHTS and
OSVs currently used as terminal tugs in offshore applications.
The current dynamic tug market internationally provides the opportunity to develop
new and innovative designs that will better serve the market demand for a better standard of
tow-ing vessel performance. New vessels are on the horizon that will truly set new benchmarks for
offshore towing performance.
* * *
REFERENCES
[1]
Buchner, Dierx, and Waals, The Behaviour of Tugs in Waves assisting LNG Carriers
dur-ing Berthdur-ing along Offshore LNG Terminals, OMAE Conference, June 2005.
[2]
Allan, R.G. and Molyneux, David, Escort Tug Design Alternatives and a Comparison of
Their Hydrodynamic Performance, National Research Council of Canada, The Society of
Naval Architects and Marine Engineers Maritime Technology Conference & Expo 2004,
Washington, DC, September 30-October 1, 2004.
[3] Molyneux,
David,
A Comparison of Hydrodynamic Forces Generated by Three Different
Escort Tug Configurations. National Research Council of Canada, Institute for Ocean
Technology, November 2003.
OSV Singapore 2005
Jointly organized by Joint Branch of RINA-IMarEST Singapore and CORE 20-21 September 2005
Tug Behaviour in Waves as Important Factor in the Operability of
Offshore LNG Berthing and Offloading Operations
Bas Buchner1*, Pieter Dierx2** and Olaf Waals2*
*
Maritime Research Institute Netherlands (MARIN)
**
Delft University of Technology, The Netherlands
ABSTRACT
For future offshore LNG terminals tugs are planned to assist LNG carriers during berthing and offloading operations. A model test study was carried out to better understand the tug behaviour in waves and to make a first step in the quantification of the related weather limits. Scale 1:35 model tests were performed in the two important ‘modes’ of a tug during this type of operation: the ‘push’ mode and the ‘pull’ mode. Realistic weather conditions were used and the tugs were working at the unshielded and shielded sides of the LNG carrier. Based on the results presented in this paper, it can be concluded that the motions of tugs in waves are significant, even in wave conditions that are considered to be mild for the berthing and offloading LNG carriers. The resulting push or pull loads may hamper these tug operations significantly. Special measures are necessary to take this behaviour into account in tug design, LNG carrier design and development of operational procedures and equipment. The paper gives insight in the typical tug behaviour in different weather conditions. The tug behaviour in waves was also simulated numerically. Considering the complex behaviour of the tug in the push mode, the comparison between the tests and simulations is remarkably good.
Figure 1. Assisting tugs in a more sheltered condition.
INTRODUCTION
Tugs play an important role in the availability of terminals in ports and offshore. Terminals and offloading operations in general, are more often located in exposed areas (wind, waves and currents) or require tug assistance in the exposed approaches (escorting). However, at the same time downtimes of the related operations are to be kept to a minimum. The more effective and enduring tug operation in open seas will enable these lower downtimes. This requirement puts a larger demand on tugs and their crews as they have to assist vessels in exposed areas. Operators, in particular for LNG, require low downtimes combined with a safe operation, which can be achieved by proper (escort) tugs during (port) approaches. Equally important is the persistent operation of tugs in offshore circumstances during side to side, tandem offloading operations and future offshore offloading concepts where tugs are needed. In and around ports often escorting regimes are put in place to cope with emergencies or assist large vessels where speed control is important.
The tug operator and designers wish to operate their existing tugs and to design their future tug as best for the job. Best is defined as the tug having the lowest downtime, giving the highest assist capability, as safe as needed and for the lowest costs. The focus on this type of application of tugs will lead to better equipped, operated and future designed tugs and help to increase the safety and overall efficiency in the matrix of terminal operation conditions.
In the current fleet of tugs lots of uncertainties exist around tug vessel type selection, equipment and operations about the ultimate capability of these tugs in exposed conditions. The optimal tug for the job requires a tug adapted to the type of vessel and circumstances under consideration. Important aspects are its assist capabilities, speed, agility, workability and safety in representative sea states and for a competitive price.
In addition to the fundamental hull design characteristics related to performance and seagoing capabilities the interface equipment like winches, tow points, line/fenders and made fast points onboard the ships have a considerable influence on the capabilities of tugs in waves. Knowing these influences could greatly help to determine a more effective and efficient operation of tugs and hence the ship operation in offloading conditions.
Last but not least tug crews have to be made familiar with these new areas of operating tugs and be convinced of the safety of the operation.
BACKGROUND
So far this type of assisting tugs has mainly been used in sheltered conditions in harbours or other more shielded conditions around terminals, see Figure 1.
Figure 2. Experience with tugs assisting crude carriers during lightering operations has shown that waves may hamper tug operations significantly (courtesy Capt. Mark Scholma).
