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OSV Singapore 2007
24-25 September 2007 National University of Singapore
2nd International Conference on
Technology & Operation of Offshore Support Vessels
Proceedings
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Jointly Organised by
The Joint Branch of the Royal Institution of Naval Architects and the Institute of Marine Engineering, Science and Technology (Singapore)
Main Sponsor
The Centre for Offshore Research & Engineering, National University of Singapore
Sponsors
Content Keynote - “2005 Happiness, 2007 Elation, what will be the market temperament in 2009” An industry insiders perspective of the current boom cycle and what it may hold for us all Captain Mike Meade MNI - M3 Marine Pte Ltd
Page 1
Offshore Support Vessels - Ripe for an Independent Regulatory Regime Francis Tang, A.K. Seah & W.F. Cheung - American Bureau of Shipping
Page 6
Environmentally Friendly Class Notations Eivind Haugen & Øyvind Pettersen - Det Norske Veritas
Page 14
Dynamics of Single Point Mooring Configurations Thomas E. Schellin - Germanischer Lloyd
Page 26
Experiences from Hardware-in-the-loop (HIL) Testing of Dynamic Positioning and Power Management Systems Tor. A. Johansen, Asgeir J. Sørensen, Olve Mo & Thor I. Fossen - Marine Cybernetics, Ole J. Nordahl - Staoil ASA
Page 41
A Survey of Concepts for Electric Propulsion in Conventional and Ice Breaking OSVs Alf Kåre Ådnanes - ABB AS
Page 51
A Large Multipurpose Offshore Vessel that Integrates Industry Demands and Ownersí Requirement Jan-Paul de Wilde, Rolls-Royce Marine & B.H. Wong - EZRA Holdings Limited
Page 70
Performance and Concepts of uprated versions of MAN Diesel L27/38 Propulsion and GenSet Engines Thomas Thomsen & Klaus Petersen - MAN Diesel
Page 84
Advances in High Volume Bulk Handling System Pankaj Thakker - MacGregor Bulk AB
Page 97
MACS®-Multi Application Cargo System Anders Eide - PG Marine
Page 108
Offshore Supply Vessels equipped with Voith Schneider Propellers Ivo Beu - Voith Turbo Schneider Propulsion
Page 118
Innovation in Large AHT(S) Winches Helge Bartels & Ralf Nicolaisen - HATLAPA Uetersener Maschinenfabrik
Page 127
Insurance Coverage Issues for Offshore Support Vessels Keith Jones - Benfield Corporate Risk
Page 134
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2nd International Conference on Technology and Operation of Offshore Support Vessels OSV Singapore 2007 24th-25th September 2007, national University of Singapore Keynote address by Captain Mike Meade MNI, Managing Director of M3 Marine Pte Ltd 2005 Happiness, 2007 Elation, what will be the market temperament in 2009? An industry insiders perspective of the current boom cycle and what it may hold for us all
Good morning, distinguished guests and participants, fellow speakers, Ladies and Gentlemen. I am pleased to be invited here today for the opening of the second “OSV Singapore”, I trust you will join me in my appreciation for the fantastic organization that has been achieved by the joint branch of RINA and IMarEST and the Centre of Offshore Research and Engineering here at these wonderful facilities hosted by the National University of Singapore. Having been present at the inaugural OSV 2005 I am pleased to see that despite all of those present operating in the most buoyant times we have ever seen in this Industry, you have time to attend a seminar such as this and gain further knowledge to develop even more technological innovation within this, the truly most dynamic of the shipping sectors. Singapore continues to gather steam in becoming the centre of excellence for technological development, manufacturing, supply and support of the Offshore Maritime Industry. In Asia, Singapore remains the only other one stop shop for the Offshore business, outside of the triangle of cities around the North Sea and the bayous of the Gulf of Mexico. Akin to its older sister in Norway, Singapore has become the city of Choice for R&D and development of all manner of Offshore equipment, with OSV’s and their design and equipment being no exception. Likewise, regionally, as Singapore has grown in importance in the design and build of OSVs it has managed successfully through ‘home grown’ Companies to develop significant expertise and shipbuilding in China utilizing skill sets and know how developed here in Singapore. KeppelSingmarine, Yantai Raffles, POET, Jaya, Chuan Hup, Swissco, Pacific Radiance and Otto Marine to name a few should be proud of this significant achievement. In 2005, in his opening address, Jim Leibertz gave us a very concise but precise overview of the history of OSV development in Singapore and the growth from a zero base in the 70’s to the significant achievements of today. The euphoric rise in the business that was
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well underway two years ago has not slowed down and indeed has ‘geared’ up to an even faster pace. Lloyds List reported recently that within the past 6 months the global order book for OSVs has increased by some 122 ships, giving a current total order book of 586 vessels (in a fleet population of 3,613) or in percentage terms 16% of existing fleet …..not too many when you see it expressed like that ? Overbuilding? I wonder……
One of the main changes we have seen in this current boom in Asia is the continued sustained growth of our home grown shipyards and owners with the interesting financial dynamic of some of our key yards and suppliers being acquired by major international companies and financial institutions. As examples here I point to the acquisition of Plimsoll by Macgregor, (who’s in turn acquisition of Hyrdramarine in Norway) has given one of our local leading suppliers a fantastic opportunity to produce cutting edge subsea lifting equipment to meet an ever increasing demand. Likewise, Pan United’s acquisition by Dubai World has brought a significant amount of cash to Singapore, with a leading ‘Company’ who’s Ownership is similar to a Singapore GLC’s. This can only be of benefit to Singapore and its bilateral ties to the Middle East. We should also mention here the Ownership changes in Jaya with two mega (in local shipyard terms) deals completed with ownership looking at one time being in the main Malaysian and now flipped to Private equity without losing its public presence. When talking boom, we cannot leave out our friends at Ezra an accolade. From humble beginnings in 2002 to today being one of Singapore’s shining lights in the public Maritime sector and having evolved into not only an OSV Owner but also a major player in the Offshore Construction, Pipelay and FPSO market. It should be noted that Ezra is the first of our peer group to have successfully floated in Norway (sorry Brian (Chang), you are now in China). So what is driving this phenomenal growth and should we concerned that the bubble is going to burst ?
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If we look at the Oil Production Forecast for just South East Asia we can see significant growth in Production going forward to meet what are known (and ever increasing) demands in the region more notably in China and India. Offshore Oil Production has grown on average about 1.8% per annum between 2000 and 2007 and is forecast to grow by 2.7% per annum from 2005 to 2012 whilst Gas Production, which I am sure we sill shortly see ‘indexed’ has experienced consistent growth since the 1980’s of around 6% with a forecast growth of nearly 6% continuing through to 2015. Worldwide Onshore and Offshore Oil Production (1960-2010 – forecasted) 100 Offshore
Onshore
Millions of barrels per day
80
58.7
60 54.8 51.5 40
47.1
46.2
15.8
16.7
1980
1990
46.3 20 19.0 0
2.0
4.5
1960
1970
23.4
26.2
29.0
2000
2005
2010
Note: Exact breakdown for year 2010 not available- estimated to be 29.0
Figure 1 SEA Oil Production Forecast Volume
4
3500 3250
Thousands of bbls oil per day
3000
Historical Data
Projected Data
Offshore Oil Onshore Oil
2750 2500 2250 2000 1750 1500 1250 1000 750 500 250 0 1940 1945 1950 1955 1960 1965 1970 1975 1980 1985 1990 1995 2000 2005 2010 2015 2020 Year
We can see the Oil and Gas demand growing and the supply is there and we will obviously need more equipment and we are seeing that built. A quick drive around Jurong and Tuas and a worthwhile daytrip to Batam will show those visiting the tremendous amount of activity in all facets of our business from Equipment to Service and the Construction of all manner of craft and of course the mammoth activity of our big brothers in the rig business who are absorbing any capacity they can in Singapore and the Riau Islands. The ‘Bears’ amongst us will tell you there are too many vessels being built and that the business is cyclic and we are in for a repeat performance of the horrors of 1985. I suppose 2 years ago when I stood here and waxed lyrical about ‘when does a boat die’ I somewhat shared that view. However, maybe as I get older I get softer and possibly more naive but from a statistical standpoint as long as those bloody old boats do die, I feel that the market can sustain the influx of newbuilds we are seeing deliver and have on order. I must stress here that my view is based on my ‘focus’ which tends to have a South East Asia and Middle East bias and doe not translate to how I would say see the North Sea market which has totally different drivers and is very much more volatile coupled with high spec and high cost equipment. World Fleet Age Profile – AHTS & PSV’s
Size Large AHTS 12,000 + BHP Medium AHTS 7-11,999 BHP Small AHTS 4-6,999 BHP Large PSV
4,000+ DWT
Ships in Operation 212
Ships on Order 98
Fleet > 20 Yrs 23%
357 531
111 82
70% 73%
114
54
0%
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Medium PSV 1500-3999 DWT 396 Small PSV 0.055 m-rads up lever curve to 30° angle of heel, and > 0.09 m-rads up to 40° or downflood. Area under righting lever curve between 30° and 40° or flood Righting lever GZ Max righting arm Initial metacentric height GM
> 0.03 m-rads. > 0.20 m at heel angle equal to or greater than 30°. Occurs at heel angle > 30° but not less than 25° > 0.15 m
Equivalent Intact Stability Criteria Area under righting > 0.070 m-rads up to 15° lever curve when max. righting lever occurs at 15° and > 0.055 m-rads up to 30° when max. righting lever occurs at 30° or above.* Area under righting > 0.03 m-rads. lever curve between 30° and 40° or flood Righting lever GZ > 0.20 m at heel angle equal to or greater than 30°. Max righting arm Occurs at heel angle > 15° Initial metacentric height GM
> 0.15 m
*When max righting lever occurs at between 15° and 30°, the corresponding area under the righting lever curve should be: 0.055 + 0.001(30° - θmax) m-rads. Where θmax is angle of heel at which the righting lever curve reaches its max. Table 1. Intact Stability Requirements for OSVs [IMO Res. A.469(XII)]
REFERENCES International Maritime Organisation (1989). “Guidelines for the transport and handling of limited amounts of hazardous and noxious liquid substances in bulk on offshore support vessels”, IMO Resolution A.673(16). International Maritime Organisation (1997). “Code of Safe Practice for the carriage of cargoes and persons by offshore supply vessels”, IMO Resolution A.863(20). International Maritime Organisation (2003). “Convention on the International Regulations for Preventing Collisions at Sea, 1972”, Consolidated Edition. International Maritime Organisation (2004). “Consolidated text of the International Convention for the Safety of Life at Sea, 1974, and its Protocol of 1988”, Consolidated Edition. International Maritime Organisation (2005). “International Convention on Load Lines, 1966, and Protocol of 1988, as amended in 2003”, Consolidated Edition. International Maritime Organisation (2006). “Amendments to the Guidelines for the transport and handling of limited amounts of hazardous and noxious liquid substances in bulk on offshore support vessels”, IMO Resolution MEPC.158(55). International Maritime Organisation (2006). “Adoption of amendments to the Guidelines for the transport and handling of limited amounts of hazardous and noxious liquid substances in bulk on offshore support vessels”, IMO Resolution MSC.236(82). International Maritime Organisation (2006). “Adoption of the Guidelines for the design and construction of offshore supply vessels, 2006”, IMO Resolution MSC.235(82).
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International Marine Contractor Association Guidelines for Design and Operation of Dynamically Positioned Vessels (199), IMCA M 103 International Marine Contractor Association Guidelines for the Safe Operation of Dynamically Positioned Offshore Support Vessels, IMCA M 182 International Marine Contractor Association Guidelines for the Training and Experience of Key DP Personnel, IMCA 117
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OSV Singapore 2007 Jointly organized by Joint Branch of RINA-IMarEST Singapore and CORE. 24-25 September 2007
Environmentally Friendly Class Notations Eivind Haugen*, Øyvind Pettersen** *Maritime Technical Consultancy, Maritime Production and Technology Centre, Det Norske Veritas, Høvik, Norway **Offshore Support, Special Ships & Diving Systems, Maritime Production and Technology Centre, Det Norske Veritas, Høvik, Norway
ABSTRACT This paper presents key elements in Det Norske Veritas’ voluntary scheme for owners with ships and ship operations implementing improved environmental measures through an environmental class notation. There are two different notations named CLEAN and CLEAN DESIGN, and the paper describes some of the key areas covered by the notations, including individual differences. INTRODUCTION Since 2000 Det Norske Veritas has given Owners an option for applying a voluntary scheme combining built in technical features in the ship with more detailed operational procedures, all with an aim to reduce the environmental impact from emissions and discharges. The scheme has been identified as a class notation where maintenance of standards is verified through regular class surveys. Market interest in the class notations has been high, although varying when considering the various shipping segments. When analyzing the ships with such class notation we find that relatively new offshore support vessels make one of the larger groups. Among recent contracts for offshore support vessels intended for the new oil exploration areas in harsh and environmentally vulnerable areas the major part of the vessels are contracted with environmental class notation CLEAN DESIGN. The aim of the notations is to assist the owner making a systematic approach to achieve an improved environmental performance, which is confirmed through the class surveys. Basis for a number of the requirements in the class notations are future requirements known e.g. from discussions in IMO’s Marine Environment Protection Committee (MEPC) identifying future regulations, others may be chosen based on technical development giving improved environmental performance. As a consequence ships with such notation will fulfil requirements beyond those mandatory today. An example of requirements introduced for ships with such class notation is a requirement for Green Passport as defined by IMO in their Guidelines on Ship Recycling. This is a requirement that will become mandatory for all ships in due course, even if IMO has not yet decided when.
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General Introduction This document describes the main issues covered by DNV Rules for Classification of Ships Part 6 Chapter 12, Environmental Class.
Interpretation of the Rules When writing the Rules and requirements for the classification of ships it is impossible to foresee each and every vessel configuration, problem and interpretation that might occur. It is therefore important to keep the intention of the Rules in mind when interpreting the Rules, which in general is to limit the operational emissions and discharges from a vessel, as well as to limit the likelihood and consequences of accidents. It is also important to understand that the Rules for Environmental Class only cover the environmental aspects of design and operation of vessels, and that the safety aspects for vessel and crew are covered elsewhere in the Rules. Where a conflict between safety and environmental considerations occur, the safety of the passengers, crew and vessel must have first priority.
Environmental Class Notations The Environmental Class Notations CLEAN and CLEAN DESIGN are voluntary Class Notations, limiting the emissions of harmful pollutants, and limiting the probability and consequences of accidents. CLEAN: MARPOL compliance with additional requirements CLEAN DESIGN: As for CLEAN, but with more stringent requirements, and in addition provisions for accident prevention and limitation. The Rules for Environmental Class are under constant development as legislation comes into force and new legislation is proposed. Vessels holding the Class Notation CLEAN or CLEAN DESIGN are in the forefront of the international legislative regime on environmental issues. This also means
that as some requirements in the Rules for CLEAN and CLEAN DESIGN are becoming mandatory, the Rules must be developed by adopting new legislation not yet ratified.
Environmental considerations The main environmental considerations addressed by DNV Rules for Environmental Class CLEAN and CLEAN DESIGN are discussed and explained below.
Emissions to Air NOx – Nitrogen oxides Nitrogen oxides are created in internal combustion engines such as diesel engine, as a function of pressure and temperature. NOx is measured in g/kWh.
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On a local level, NOx contributes to the formation of low-level ozone, thereby contributing to respiratory problems, damaging forests, and other problems. On a regional level it will, when present in the atmosphere together with sulphur dioxide, undergo chemical transformations giving acid rain when absorbed by water droplets in clouds. In addition NOx will also cause over-fertilization of inland and coastal waters. In the latter years there has been a strong emphasis by national, regional and global legislative bodies on the reduction of NOx-emissions. The emission of NOx can be limited by: “Upstream” technologies such as Direct Water Injection, Humid Air Motor or fuel/water emulsion, all aiming to reduce the peak temperature in the cylinder during combustion. “Downstream” technologies such as Selective Catalytic Reduction, introducing a reduction agent such as urea into a catalytic converter unit in the funnel, “cleaning” the exhaust gas. The use of NOx abatement techniques is mostly relevant to vessels with the Class Notation CLEAN DESIGN. Vessels with CLEAN carry IMO NOx-certificates or equivalent for the relevant engines, thereby fulfilling the requirements of the Rules. For CLEAN DESIGN the engines must emit about 20 % less NOx than specified by the IMO NOx-curve. This is difficult to achieve by engine tuning alone without compromising engine efficiency. In cases where one or more of the engines do not fulfil the requirements for NOx emission for CLEAN DESIGN, a separate study for the complete power/propulsion plant may be carried out to show overall vessel compliance with the Rules. In order to prove adherence to the Rules, the Operational Procedures and logs must show 1. That the engine manufacturers’ instructions with regards to spare parts and maintenance are followed, that the Engine Technical File is kept in order 2. That any NOx-reduction units (DWI, SCR etc.) are operated according to instructions, and that logs are kept to prove the use (e.g. urea consumption, water/fuel consumption etc.) 3. This must be verified at Annual survey by checking the technical file and NOx-equipment log if installed.
SOx – Sulphur oxides Sulphur oxides are formed in internal combustion engines, boilers and other systems using fuels containing sulphur. SOx concentration is measured as g/kWh. On a local level, SOx contributes to respiratory problems, acid attack on vegetation and limestonebased structures, and other problems. On a regional level it creates acid rain (ref. also comments on NOx above) and over-fertilization of inland and coastal waters. There has been a strong focus on the reduction of SOx-emission on a national and regional level, especially within the legislative frameworks of the EU and USA. At an international level (IMO) has introduced restrictions on SOxemissions in MARPOL Annex VI. Discussions are continuing in IMO on whether their targets for limiting SOx-emissions in Annex VI shall be adjusted. The emission of SOx in exhaust gases is wholly dependent on the sulphur content of the fuel utilised, and can be controlled by: “Upstream” measures such as limitations on the sulphur content used onboard “Downstream” technologies such as exhaust gas scrubbers A great proportion of the operating cost of cargo ships in particular comes from bunker expenses. Many ship-owners therefore react to the requirement of limitations of the sulphur content of fuels used onboard vessels with CLEAN. It is important to emphasise this requirement early in the design phase, so that properly sized low-sulphur fuel tanks and an appropriate system arrangement can be provided onboard.
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In order to prove that the vessel is operated in accordance with the Rules, operational procedures should: 1. Make sure only fuel with sulphur content less than the specified maximum limit is ordered. The Fuel Oil Management plan (or equivalent) should explicitly specify the maximum sulphur content to be used in general, and in ports and “SOx-controlled area” in particular. 2. Ensure that the sampling regime conforms with the Rules. In order for the Master/owner to prove adherence with the maximum sulphur requirements, a fuel sample should be kept onboard for at least one year, and the bunker receipt should be filed onboard for three years. Most vessels already have stringent sampling procedures in place in case they have received sub-standard fuel causing damages to machinery or poor engine performance. At Annual survey the bunker receipts and required logs will be checked to make sure the maximum sulphur limits have been complied with.
Refrigerants Refrigerants used onboard in cargo refrigeration plants, air-conditioning plants and provision refrigeration/freezing systems escapes to the atmosphere through system leaks and spillages during maintenance and not least during recycling/recovery. Refrigerants based upon halogenated carbon substances contribute strongly to the depletion of the Earth’s ozone layer, causing increased ultraviolet radiation and subsequently increased risk of skin cancer. They also contribute to global warming. International legislation such as the Montreal protocol and various EU regulations have been adopted to phase out the unwanted substances and replace them with more environmentally friendly alternatives. The Ozone Depleting Potential (ODP) of a substance is a measure of how potent it is with respect to destroying high altitude ozone. The Global Warming Potential (GWP) is a measure of how strongly the substance acts with respect to retaining and returning heat radiation from the Earth’s surface, the global warming effect. This is measured relatively to the effect of CO2, over a 100 year life span (as most of these substances are destroyed in the atmosphere within a matter of years). Refrigerants used onboard vessels with the Class Notation CLEAN must have an ODP of zero, and the GWP should be maximum 3500. In no case is chlorofluorocarbons (CFCs and HCFCs) allowed, but hydrofluorocarbons (HFCs) are allowed in addition to “natural” refrigerants such as ammonia (NH3) and CO2. For vessels with Class Notation CLEAN DESIGN the maximum GWP is 1890. The refrigeration systems onboard should be of such a design as to easily facilitate the removal of the whole of the system “charge”, the full volume of refrigerant, without spillage. A recharge valve/connection should be placed in a position where a receiver or tank can be connected and the system compressor can pump the refrigerant into the tank. A system of discovering system leakages at an early stage should be in place. This can either be in the form of detectors at suitable locations, or by checking the system for leaks and the liquid level at regular intervals. Leak detectors/sniffers are often unsuitable where the refrigeration machinery and compressors are located in engine rooms where the air volume is large and the air change ratio is high, but can be useful when systems are located in enclosed spaces. Please note that such a system might be required by other Rules and regulations regarding safety. The system and the inspection regime should be worked out in co-operation with the system manufacturer, and must include a separate log for recording refrigerant levels and maintenance events.
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The operational procedures for the refrigeration systems must describe the corrective actions to be taken when the refrigerant level sinks below a pre-defined limit. The corrective actions must include leakage reducing measures. The maximum annual leakage must be less than 10 % of the total system charge. This should be documented by a Refrigerant Recharge log. In cases where the refrigerant properties do not meet the requirements regarding the GWP, a caseto-case evaluation of the Total Equivalent Warming Impact (TEWI) can be undertaken by the supplier. A TEWI analysis considers the efficiency of the refrigerant and the system’s energy consumption to show that the lowered CO2-emission by systems providing energy to the refrigeration plant compensates for the higher impact of the refrigerant itself.
Fire fighting substances Fire fighting substances with an Ozone Depleting Potential are prohibited (see Refrigerants above). Maximum Global Warming Potential is to be 4000 for vessels with Class Notation CLEAN, 1650 for CLEAN DESIGN. This means that e.g. halon is prohibited. CO2, water fog, argon and other natural substances are acceptable.
Cargo evaporation Volatile Organic Compounds (VOC) are contributors to the formation of low-level ozone, responsible in turn for amongst other things respiratory problems. They are also carcinogenic, and pose a great safety risk as they can ignite at the right concentrations. VOC are light fractions of oil cargoes evaporating especially during cargo loading and unloading. Crude/product tankers and tankers carrying substances with flash point less than 60° C (Volatile Organic Compounds) should adhere to Class Notation VCS-2.
Other Any incinerators installed should comply with MARPOL Annex VI Reg 16(1), meaning that they should be approved with respect to MEPC.76(40), Standard specification for shipboard incinerators, with amendments MEPC.93(45) Amendments to the standard specification for shipboard incinerators. Use of the incinerators should be noted in garbage record books and oil record books as applicable. It should be noted that the use of incinerators might be illegal under local legislation.
Discharges to Sea Cargo residues: The discharge of oil and chemical cargo residues has obvious negative impacts on the marine environment. The introduction of MARPOL Annex I has greatly reduced the discharge of oil from cargo and machinery spaces, while MARPOL Annex II controls the discharge of chemical cargoes. For Class Notation CLEAN, the Rules aim to limit the discharge of oil cargo residues by the fulfilment of MARPOL Annex I, for both “old” and “new” tankers. In practice, all vessels delivered as new builds today will comply with Annex I. With regards to chemical cargoes, the requirement is that the maximum allowable remaining cargo quantity after stripping is 0.1 m3 for pollution category B and 0.3 m3 for category C under Annex II.
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This is a more stringent requirement than Annex II, and the onboard cargo operational procedures should reflect this. An oil record book or cargo record book should be maintained, also recording the discharges to sea and deliveries to shore of cargo residues, wash water etc. Tankers must also have double skin and segregated spaces and piping for cargo and ballast water. An important reference is MARPOL Annex I Ref. 13F, “Prevention of oil pollution in the event of collision or stranding”. For Class Notation CLEAN DESIGN, the maximum allowable remaining chemical cargo quantities are 0.05 m3 for both pollution categories B and C. In addition to the requirements for double skin, cargo tanks must be designed with smooth surfaces and cargo wells for efficient stripping. Under-deck longitudinals of slab type are acceptable. Horizontal areas on stiffeners and brackets should be avoided. Horizontally corrugated bulkhead plating with maximum angle of 65°, with or without vertical girders, is acceptable. Crude Oil Washing efficiency should have a coverage of minimum 96 %, documented by shadow diagrams.
Cargo handling The requirements for cargo handling arrangements and procedures aim to prevent spillage of oil containing substances and chemicals during cargo operations. For offshore support vessels cargoes with special concern include e.g. liquid mud in addition to various types of oils carried as cargo to the oil rigs. In addition special care shall be taken with respect to mud and liquids returned from the oil rigs which may be contaminated. Requirements for control of such operations are much the same as described under oil bunkering arrangements, below.
Oil bunkering arrangements This requirement applies to fuel oil, lubricating oil, hydraulic oil and all other oily substances which can be filled from a source without the vessel’s crew attending. In practice, this means that oil filled by hand pumps or pneumatic pumps from oil drums on deck are exempt from the requirements, while some cargo substances like oil based mud, oil for both cargo and consumption purposes (as is often the case on e.g. platform supply vessels) are included. This is for practicability, as the systems for cargo and bunkering often use the same piping and tanks, or at least have the filling points and vent heads placed close together. All points where oil spills may occur must be fitted with spill trays. This means that both fill points and vent heads must have trays. The trays must be designed to catch spills coming through vents at high speeds. Tanks are required to have high level alarms to prevent overfilling. Bunker operations procedures must be submitted to ensure that spill protection is considered during bunker operations. Systems and tanks for recovered oil on ships equipped for oil spills at sea only are exempt from the requirements, as such operations are considered as emergency operations.
Ballast water One of the emerging environmental considerations nationally and internationally is the transfer of harmful micro-organisms in the ship’s ballast water. The introduction of invasive species can result in
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depletion of stocks of native species, unwanted algal blooms etc, and can have extreme local ecological and financial consequences. To prevent this, a vessel can either 1. Treat the ballast water mechanically or chemically to kill off the unwanted organisms, 2. Exchange the ballast water far from shore, as theoretically the species living in coastal waters will not survive the conditions in deep sea waters and vice versa. This may be done a. By continuous flow-through of at least 3 x the total ballast water volume or to some predetermined volumetric efficiency, but the efficiency of this method is being debated b. By sequentially emptying and filling each ballast tank in turn, this imposes severe stresses on the hull structure and cannot be undertaken in heavy weather conditions Requirements based on alternatives a) and b) will be phased-in in accordance with the International Convention for the Control and Management of Ships’ Ballast Water and Sediments. In any case, the vessel must carry a Ballast Water Management Plan which is approved separately by other units within DNV. The approval of this plan is outside the scope of the approval work for Class Notations CLEAN and CLEAN DESIGN.
Bilge water Bilge water from machinery spaces contains various amounts and qualities of oil from maintenance operations, minor leaks, machinery “sweating” etc, but also other substances such as detergents. The bilge water must run through a bilge water separator and an oil content monitor, automatically shutting off the discharge if the oil content exceeds 15 ppm (parts per million). In reality, the performance of both the separator and the monitoring units is strongly influenced by the presence of detergents and other contaminants, and the resulting discharged bilge water will often contain too much oil. For CLEAN, the discharge of bilge water is regulated by MARPOL Annex I. The ship must have a valid IOPP (International Oil Pollution Prevention) certificate. It must have oil filtering equipment and 15 ppm alarm with automatic stop, and holding tanks with facilities for delivery to shore. For CLEAN DESIGN the vessel must have bilge water holding tanks as required for the Class Notation OPP-F, which means that they must have required capacities dependent on the engine rating. The machinery space bilges must not be discharged to sea, but be discharged to shore. This requirement should be explicitly expressed in the bilge water management plan. CLEAN DESIGN requires oil content of bilge water to be less than 5 ppm. Bilge water from non-machinery spaces is to be treated separately from the machinery space bilges. Wash water and slop from cargo spaces are defined as “Residues of cargo oil or chemicals” and are covered by Section 2 C 200 and Section 3 C200.
Garbage Waste produced onboard vessels can seriously harm the marine environment, especially sea-living mammals, reptiles, birds and bottom-dwelling animals, and is extremely unsightly when littering beaches, tidal zones, mud-flats etc. It can also be harmful to shipping, especially smaller vessels. Handling and disposal of garbage is regulated by MARPOL Annex V. Garbage is divided into six categories according to its constituent material and hazard potential. The various categories can be treated onboard, and can be disposed of at sea at various distances from nearest land. The exception is plastics, which under no circumstances can be dumped at sea. Various onboard treatment options include grinding, incineration, compaction, shredding, pulping and so on.
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For CLEAN, the vessel must comply with MARPOL Annex V, and have in place a Garbage Management Procedure showing the various garbage treatment options onboard. The discharge criteria should also be explicitly expressed in the plan. Incinerators are not compulsory, but if installed they must be approved with respect to MEPC.76(40), Standard specification for shipboard incinerators, with amendments MEPC.93(45) Amendments to the standard specification for shipboard incinerators. Vessels with Class Notation CLEAN DESIGN must store all garbage onboard prior to delivery on shore, or incinerate it. Only comminuted (ground) food waste can be disposed of at sea. Please note that there might be local restrictions on the use of incinerators in some areas, like for instance the Baltic Sea where the use of incinerators is prohibited. In all cases a Garbage Record book must be kept.
Sewage Sewage is defined as drainage from toilets, medical bays, spaces containing living animals or any drainage mixed with any of these. This is also known as black water. In addition, grey water can be defined as drainage from wash basins, showers, galleys, dishwashers, in short “domestic” drainage which does not pose a health risk due to contamination by pathogens (bacteria, virus etc). It may still contain detergents, food waste, grease and other pollutants. Biologically active sewage poses a threat to human health by transmitting diseases to people swimming in the sea, or via eating fish and shellfish contaminated with sewage. Vessels trading in inland waters might contaminate drinking water by discharging sewage. It also introduces nutrients to inland and coastal waters causing eutrophication or algal blooms. The treatment and discharge of sewage is controlled by MARPOL Annex IV, which is in force from autumn 2003. Annex IV is only concerned with black water. For CLEAN, vessels must either 1. Have in place an approved sewage treatment plant, or 2. Discharge comminuted (ground) and disinfected sewage at a distance from shore of no less than 4 nautical miles, or 3. Discharge untreated sewage at a distance from shore of no less than 12 nautical miles. Note that this option is under special circumstances, only. For options 2) and 3) the sewage must have been comminuted and ground prior to settling in a storage tank, and if discharged at sea, must be discharged at a moderate rate and while the vessel is en route i.e. at a normal trade at a speed of at least 4 knots. For all the options above, any discharge of sewage should be recorded with date, location and quantity. This can be entered into the ship’s log book, or noted in a dedicated Sewage discharge log. Please be aware that more stringent local legislation might apply, such as for the Great Lakes or the Baltic Sea. For CLEAN DESIGN, passenger vessels must have sewage holding tanks large enough to store both black and grey water while in port. All vessels must have a sewage treatment system, and passenger vessels must also treat grey water.
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Anti-fouling The use of various bio-toxic compounds has proved an effective way of limiting marine growth on ships’ hulls, thereby lowering friction losses and fuel consumption. One of the most effective ingredients has been found to be a group of chemicals called tributyl-tin, TBT. However, the TBT released into the marine ecosystems has been found to cause a number of problems such as reproduction problems in marine organisms. TBT-based anti-fouling systems are therefore not allowed to be applied on vessels from 1. Jan. 2003, and must be removed from ships’ hulls or sealed properly within 1. Jan. 2008. Vessels with the Class Notations CLEAN or CLEAN DESIGN must carry a “Statement of Compliance with International Convention on the Control of Harmful Anti-Fouling Systems”.
Protective design – CLEAN DESIGN only. The requirements for Protective Design are valid for the Class Notation CLEAN DESIGN only, and are aimed at limiting the probability and consequences of accidents resulting in discharge of harmful substances. Approval of the protective design must be undertaken by the unit responsible for Environmental Class Notations at DNV Head Office.
Fuel oil tank arrangements Fuel oil is very often heavy oil fractions extremely harmful to the environment. Heavy fuel oil will not evaporate, it disperses poorly in sea water, often forms sticky emulsions, and due to not evaporating or dispersing, it often reaches shore and fouls beaches and kills wildlife. By ensuring fuel oil tanks are protected by a double hull, the likelihood of discharge when an accident occurs will decrease. Tankers for Oil or Chemicals have for some time been regulated such that tanks shall be in protected location. In a revision to MARPOL Annex I, the new Reg. 12, it is specified that also fuel oil tanks shall be in protected location from dates defined in this new regulation. Ships with class notation CLEAN DESIGN are required to comply with requirements for cargoand fuel oil tanks already today. The requirements apply for cargo oil tanks even if the ship is not defined as a “Tanker”. It is also worth noting that the requirement is extended to apply for all oil containing tanks. For supply vessels and other vessels carrying liquids such as e.g. base oil, oil based mud (and return mud from oil drilling rigs) the double hull requirement will also apply for such tanks.
Nautical equipment and arrangement In order to avoid accidents at sea it is imperative that 3 truths are known; own position, where to go and what lies in between. In addition it is of course important that the navigators’ working environment and communication is best possible. In DNV we have developed requirements for this described by our class notation NAUT-AW. For special offshore vessels the same is achieved through the separate class notation NAUT-OSV(A). In consideration of the positive effects we require that CLEAN DESIGN vessels shall have the applicable of these nautical class notations.
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Emergency propulsion By offering redundancy in propulsion and steering systems, the vessel should be able to reach a safe haven even if one propulsion or steering system is out of order. This reduces the probability of accidents. Separate engine rooms and propellers, drop-down thrusters or “take-me-home-devices” ensure that at least one main engine can fail with the vessel still being able to manoeuvre. Given the range of possible design solutions, approval for the purposes of CLEAN DESIGN is to be done by the unit responsible for Environmental Class Notations at DNV Head Office. Vessels with the Class Notation RP (redundant propulsion) fulfil the criteria with margins. Class Notation RP is, however, not required. Vessels using a double boiler to drive a single steam turbine and propulsion plant are also accepted.
Approval of operational procedures Of required documentation it is worth noting that improved environmental performance is dependent on operational procedures in addition to only requirements to equipment and structural arrangements. The following issues need be covered either as independent documents, or as amendments to existing procedures: 1.
2. 3.
4.
5.
6.
Fuel Oil Management plan: In order to control the emission of SOx the vessel procedures must specify the maximum sulphur content of fuel to be used in various areas. A fuel oil log must be in place to document the qualities of fuel ordered. Sampling of fuel, and the retainment of samples and bunker delivery notes, should be specified. Sampling should comply with DNVPS guidelines “Marine Fuel Management”. Bunkering procedure: This should reflect all precautions made to prevent spillage at bunkering operations. It should also make provisions for the sampling of bunker at transfer. Refrigerant management procedure: In order to prevent leakage of refrigerants to the atmosphere, provisions should be in place for checking and logging refrigerant levels at regular intervals. This must be recorded in a log, and the procedure should detail the limits for when action must be taken, what actions should be taken and responsible personnel. Maintenance is also to be recorded. If a leak detection system is installed, the procedures must include means to use this system efficiently for leak detection with environmental issues in mind. Whenever refrigerants are drained or added, this should be noted in the refrigerant log so that it can be verified that the annual leakage is less than the 10 % stipulated in the Rules. Garbage management plan: Handling of garbage is regulated by MARPOL Annex V. A procedure for onboard handling of waste categories defined in MARPOL must be in place. This should include discharge criteria, responsibilities and logging in Garbage record book. Any onboard waste management devices such as incinerators, compactors, shredders etc. should be taken in to consideration by the plan. Sewage management plan: Handling of sewage is regulated by MARPOL Annex IV. Discharge criteria, responsibilities and discharge logging procedures should be laid down in the sewage management procedure. If the vessel is equipped with sewage treatment plants, holding tanks for black/grey water etc, this should be reflected in the Plan. Ballast water management plan: The ballast water management plan is approved separately from the Class Notations CLEAN and CLEAN DESIGN. Proof of approved ballast water management plan must be submitted.
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7.
Bridge and engine room operation control procedures. These will show compliance with the technical requirements for Class Notations NAUT-AW or NAUT-OSV(A) (CLEAN DESIGN only) and E0 as applicable. These must be approved by authorised personnel. If vessel has any of the Class Notations, the procedures need not be submitted for CLEAN.
Ship in Operation (SiO) Survey During operation all ships are subject to regular surveys to confirm compliance with applicable rules. Specifically requirements in CLEAN and CLEAN DESIGN will be checked during annual surveys. The main part of such annual surveys will be through checks of log books to verify that operations have been carried out as required by the procedures. Any deviation will have to be explained and the surveyor will evaluate whether deviations from the rule requirements have acceptable explanations.
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References DNV’s Rules for Classification of Ships, specifically: Rules Pt. 6 Ch. 12 Environmental Class Rules Pt. 1 Ch. 1 General regulations Rules Pt. 6 Ch. 1 Sec. 6 Additional Oil Pollution Prevention Measures – Fuel Oil Systems (Class Notation OPP-F) Rules Pt. 6 Ch. 2 Redundant Propulsion (Class Notation RP) Rules Pt. 6 Ch. 3 Periodically unattended machinery space (Class Notation E0/ECO) Rules Pt. 6 Ch. 8 Nautical Safety (Class Notation NAUT-AW) Rules Pt. 6 Ch. 10 Vapour Control Systems (Class Notation VCS-2) Regulations and guidelines MARPOL 73/78 EU Directive 99/32/EC (EU Sulphur Directive) IMO NOx Technical Code (IMO MP/Conf. 3/35 Resolution 2) US EPA air emission standards proposal Tier 2 IMO Resolution MEPC.46(30) Measures to control adverse impacts associated with the use of Tributyl-tin compounds in anti-fouling paints International Convention on the Control of harmful Anti Fouling Systems. International Convention for the Control and Management of Ships’ Ballast Water and Sediments. IMO Standards for Vapour Emission Control Systems, MSC/Circ.585 and MARPOL Annex VI, reg. 15 IMO Resolution MEPC.76(40) on Standard specification for shipboard incinerators USCG 46, CFR 39 on cargo handling vapour emission control systems USCG 33 CFR 159 on marine sanitation devices Montreal Protocol on substances that deplete the ozone layer ISO 8217, Petroleum Products – Fuels (Class F) ISO 3170/ISO 3171 (or equivalent national standard), Code of practice for bunkering by bunker barges/tankers ISO 8754, test method, fuel sulphur content ISO 7934/ISO 7935/ISO 11632, test method, emission sulphur content DNV Petroleum Services Guidelines, “Marine Fuel Management”
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OSV Singapore 2007 Jointly organized by Joint Branch of RINA-IMarEST Singapore and CORE. 24-25 September 2007
Dynamics of Single Point Mooring Configurations Thomas E. Schellin* *Rule Development, Analysis of Hull Structures & Damages Germanischer Lloyd, 20459 Hamburg, Germany
ABSTRACT This paper presents a theoretical-numerical model to investigate the dynamics of single point moored vessels subject to current, wind, and waves. The time-domain analysis described is based on an experimentally validated, comprehensive mathematical maneuvering model of ship maneuvering equations in three degrees of freedom (surge, sway, and yaw) that considers five sets of forces: (1) nonlinear quasi-steady hydrodynamic response and control forces, (2) linear memory effects due to radiated waves, (3) nonlinear mooring restoring force characteristics, (4) empirical wind actions, and (5) first-order wave forces and second-order wave drift forces. The significant nonlinearities inherent in single point mooring systems, mainly due to hydrodynamic response and control forces as well as mooring system restoring forces, may lead to multifarious dynamic phenomena, such as self-sustained oscillations. Under certain conditions, three practical measures, namely, rudder deflection, reverse propeller action and asymmetric mooring, are shown to stabilize the moored vessel, thus reducing motion amplitudes and mooring hawser tensions. Computer simulations for a supertanker are experimentally validated by comparison with model experiments and, as an extension, specifically focused on the effects of water depth. Effects of different single point mooring configurations on the dynamic behavior of a supertanker are examined, comprising buoy mooring, articulated tower mooring, bow turret mooring, and internal turret mooring. In addition, based on the use of an alternative frequency-domain method, results of a mooring analysis used for design approval are presented for an offshore coal transshipment terminal that includes single point mooring systems of a transshipper and two coal barges. INTRODUCTION At unprotected offshore locations, tankers and other vessels are increasingly moored using socalled single point mooring (SPM) systems. The purpose of this paper is to present rational methods for analyzing SPM systems. Different configurations are in use. A single buoy mooring (SBM), an internal turret mooring (ITM), and an articulated tower mooring (ATM) are schematically shown for a tanker in Fig. 1. The obvious advantage of these configurations is that the vessel is free to assume a favorable alignment to the prevailing current, wind or waves, thereby substantially reducing tensile forces in the mooring hawser compared to forces possible if the vessel’s heading were constrained. However, under certain conditions, the vessel may not attain a stable equilibrium even in a seemingly innocuous steady environment, but indulge in large amplitude low-frequency oscillations of periodic or aperiodic nature. The problem is of practical importance due to the formidable environmental implications of an accident. The crucial aspect is the sometimes unpredictable peak forces in the mooring hawser threatening a possible line break (Sharma et al. 1994).
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During the last decades, a considerable amount of research has been devoted in the development of theoretical-numerical methods to predict full-scale behavior of an SPM vessel. Significant nonlinearities inherent in an SPM system, mainly due to hydrodynamic response and control forces as well as mooring system restoring forces, are responsible for difficulties in making accurate predictions of the vessel’s response, displaying many features typically associated with nonlinear dynamics and instability. A computational analysis calls for a reliable mathematical model incorporating sufficiently accurate descriptions of hydrodynamic and mooring restoring forces, a realistic representation of environmental effects (current, wind and waves) as well as a precise account of all relevant physical interactions. Apparently, the earliest studies of SPM ships were carried out by Wichers (1976), Faltinsen et al. (1979), and Owen and Linfoot (1985). To ensure economic offshore exploration, SPM systems for large floating production, storage and offloading (FPSO) vessels are vital as exemplified by recent investigations by, for example, Matter et al. (2007), Munipalli et al. (2007), and Vazquez-Hernandez (2007). Single point moorings are used primarily to position tankers. For tankers, parameters required for the specification of all forces in the mathematical model are usually available. Generally, parameters (coefficients) that depend on hull geometry, propeller, and rudder are obtained from planar motion tests with a captive model in a towing tank. The resulting timedomain simulations must be of sufficient length to establish reasonable confidence bounds for the expected maximum response in the storm duration. Typically, specified storm duration will be simulated Fig. 1 Schematic of SPM configurations several times, using statistical fitting techniques to (Jiang et al. 1995) establish the expected maximum response. The timedomain mathematical model is here presented in some detail, including experimental results that validated the model. Using this model, computer simulations are performed to demonstrate the effects of different SPM systems on the dynamic behavior of a supertanker. For offshore type structures and other vessels, on the other hand, hull dependent parameters are usually not available. In such cases, a rational method documented, for example, by the American Petroleum Institute (API 1997) is frequently used to analyze, design or evaluate mooring systems. This method, most conveniently performed in the frequency-domain, provides an alternative approach to determine the adequacy and safety of mooring systems. It calls for first defining a mean offset of vessel displacement due to a combination of mean wind, current, and wave drift forces. The maximum offset is then the mean offset plus appropriately combined wave-frequency and low-frequency vessel motions, and a probabilistic approach defines the chance of exceeding the combined high- and lowfrequency motions on average of once in the specified storm duration. Results from the use of this method, performed in the frequency-domain, are presented for the mooring analysis of an offshore coal transshipment terminal that includes single point mooring systems of a transshipper and two coal barges.
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MATHEMATICAL MODEL Two coordinate systems, see Fig. 2, describe the motion of the vessel singlepoint moored in current, wind, and waves. First, an earthbound coordinate system P x0 y0 , conveniently centered at the effective mooring point P, describes the position of the vessel’s midship point O and heading angle ψ as well as the direction of the mooring hawser ψA. It also defines the the reference directions Fig. 2 Coordinate systems of current, wind, and waves (ψC,ψW,ψS). Second, a vessel-bound coordinate system O x y is used to simplify the description of external forces acting on the vessel. The center of gravity G of the vessel and the attachment point A of the mooring hawser have fixed coordinates xG ,yG and xA ,yA, respectively, in the vessel-bound coordinate system. The relation between horizontal vessel velocity components x& 0 , y& 0 and u, v resolved along earthbound and vessel-bound coordinates, respectively, is contained in the trajectory equations:
x& 0 = u cosψ − v sinψ y& 0 = v cosψ + u sinψ ψ& = r
(1)
r&I zz + (v& + u r ) xG m = N
(6)
(2)
(3) which also define yaw rate r. The vessel is treated as a transversely symmertic rigid body having three degrees of freedom: surge sway, and yaw. Effects of heave, pitch, and roll are believed to be small on extremely slow horizontal motions and hence neglected. The dynamic equations of maneuvering can then be written in standard form: (u& − v r − r 2 xG ) m = X (4) (v& + u r + r& xG ) m = Y (5) where m is vessel mass and Izz its moment of inertia about a vertical axis through O. The net external time-varying horizontal force components X,Y resolved along axes x,y and their moment N about O result, in general, from a complex interaction of various physical phenomena. A simple linear superposition of five effects is considered: F = { X , Y , N } = FQ + FM + FA + FW + FS T
(7)
where subscripts Q,M,A,W,S stand for quasi-steady, memory, anchoring, wind, and sea waves, respectively; superscript T denotes transpose. The quasi-steady hydrodynamic response and control force couple FQ is calculated according to the four-quadrant model of Oltmann and Sharma (1984). The force couple elements are synthesized as follows: ⎧⎪ X HI + X HL − RT + X P + X R ⎫⎪ FQ = ⎨YHI + YHL + YHC + YP + YR ⎬ ⎪⎩ N HI + N HL + N HC + N P + N R ⎪⎭
(8)
where subscripts H,P,R stand for system elements hull, propeller, rudder and I,L,C for physical mechanisms ideal fluid, lifting, cross-flow effects, respectively; the odd term RT simply denotes ordinary resisitance to pure longitudinal motion. This force couple depends linearly on accelerations u& , v&, r& and in a highly nonlinear way on velocities u,v,r as well as on control parameters propeller rate n and rudder angle δ. The explicit formulations are fully documented in the work cited. However, three features are worth mentioning. First, the four-quadrant model, unlike many others in common use,
29
does not break down near speed reversals (u = n = 0) and is, therefore, specially suitable for simulating slow vessel motions. Second, this model painstakingly accounts for three-way hull-propeller-rudder interactions. Third, it incorporates simple empirical corrections for the main Reynolds number associated scale effects on hull resistance and wake with important ramifications for propeller and rudder forces. Standard ITTC (1984) corrections to ship resistance are considered without allowance for the roughness scale effects. To obtain full scale propeller and rudder forces, the wake fraction for the ship condition is needed. Again, the ITTC (1984) correction formula is used. A linear response force couple associated with hydrodynamic memory is calculated by means of a finite state space model fully described by Jiang et al. (1987). The final result can be summerized as follows:
[
]
(9)
s&n−k = sn+1−k − Ak s0 − Bk v
(10)
FM = [a (0) − a (∞)]v& + b (0) − b (∞) v + s 0
with k = 0,1, · · · , n, s n +1 = 0 and v = (u , v, r ) T . The state vectors sk effectively store memory effects of motion history in the time domain. Parameter matrices Ak , Bk of an (n + 1) state space model can be identified by a least squares method to fit theoretically calculated added mass and damping matrices a (ω ), b (ω ) in the frequency domain. Thus, this formulation accounts for linear memory effects due to waves radiated from an oscillating hull. ENVIRONMENTAL EXCITATION For slow motion of the moored vessel, the main part of hydrodynamic forces is due to the prevailing current. Therefore, current forces are treated specially. Unlike others, who treat current forces by means of modular modeling, the vessel’s velocity relative to the ambient water is substituted into the four-quadrant maneuvering model, thereby accounting for the entire interaction between vessel motion and current speed. For a steady uniform current of magnitude VC and direction ψC, the relations between hull velocity components u,v over ground and urel ,vrel relative to ambient water are u rel = u − VC cos (ψ C − ψ )
(11)
vrel = v − VC sin (ψ C −ψ )
(12)
The wind-generated force couple, due to a wind acting on the above-water parts of the hull and superstructure, is calculated as usual by means of the following empirical formula:
⎧ 0.5 ρ AVW2 AT C XW ⎫ ⎪ ⎪ FW = ⎨ 0.5 ρ AVW2 AL CYW ⎬ (13) ⎪ 0.5 ρ AVW2 AL LC NW ⎪ ⎭ ⎩ where ρA is mass density of air, VW is wind speed (accounting in standardized ways for vertical profile and turbulance level), AL and AT are longitudinal (broadside) and transverse (head-on) projected above-water areas, L is vessel length between perpendiculars, and CXW,CYW,CNW are ship-type dependent force and moment coefficients as functions of wind angle of attack (π – ψW – ψ), see OCIMF (1994). Wave forces generated on the hull by the action of ambient sea waves are approximated as the following sum: FS = FS(1) + FS( 2) (14) where FS(1) is the first-order force and FS( 2) the second-order slowly varying drift force. The firstorder force is constructed by superimposing responses to N individual wave components:
FS(1) = Re
N
∑ H (ω j =1
ψ S − ψ ) A j exp[ − ik j (x0 cosψ S + y 0 sinψ S ) + iω j t ]
j,
(15)
30
with wave amplitude Aj and wave number kj corresponding to wave frequency ωj. The hull-form dependent complex frequency response vector H is also referred to as the linear transfer function of wave forces. The second-order slowly varying force is assumed proportional to the local wave envelope profile squared:
FS( 2) = ξ 2 G (ω1 ,ψ S −ψ )
(16)
where G is the hull-form dependent drift force coefficient vector at the mean wave frequency ω1, also known as the quadratic transfer function of wave forces. The local wave envelope profile ξ is computed as follows:
ξ = ς 2 +η 2
(17)
with wave elevation ζ and its Hilbert transform η defined as follows:
ς + iη = ∑ A j exp[− ik j (x0 cosψ S + y0 sinψ S ) + iω j t ] N
(18)
j =1
Numerical values of linear H and quadratic G transfer functions for a given direction of wave propagation ψS can be obtained effectively using three-dimensional potential boundary element methods in the frequency domain. ALTERNATIVE MOORING CONFIGURATIONS For a given vessel, hull hydrodynamics are assumed to remain unchanged in different mooring configurations, i.e., by neglecting hydrodynamic interactions between the mooring system and the hull. This is the case for single buoy mooring (SBM) and articulated tower mooring (ATM) configurations. For bow turret mooring (BTM) and internal turret mooring (ITM) configurations, hydrodynamic interactions between turret and hull are assumed small and, therefore, neglected. Thus, the difference of alternative mooring systems manifests itself only in the mooring retsoring characteristics. Three different polynomial functions define the total quasi-static horizontal mooring system retsoring force FA in the SBM, ATM, and BTM (ITM) configurations. First, a forth-order polynomial represents the SBM configuration: FA =
where
1 [1 + sgn (Δ L A )]C 4 (Δ L A ) 4 2
(19)
Δ L A = L A − L AU Here FA arises from the elasticity of the the hawser and the catenary action of the anchored buoy, and LA is the instantaneous horizontal distance of attachement point A from effective mooring point P, whereas LAU is the hawser’s reference length corresponding to the unstretched no-load condition. Second, a second-order polynomial represents the ATM configuration:
(
FA = 0.5 (1 + sgn (Δ L A )) C1 (Δ L A ) + C 2 (Δ L A ) 2
)
(20)
Here FA results from the elasticity of the hawser and the buoyant restoring force of the inclined articulated tower. Displacement ΔLA has the same meaning as above. Third, a third-order polinomial represents the BTM and ITM configurations: 2
FA = C1 Δ L A + C 2 Δ L A + C3 Δ L A
3
(21)
Here ΔLA is the displacement of turret position A from its no-load reference position. The mooring restoring forces act horizontally either at attachement point A on the vessel in the direction from A to effective mooring point P or at instantaneous turret position A on the vessel in the direction from A to its no-load reference position. The coefficients C1, C2, C3, C4 can usually be obtained by using the least squares method to fit statically calculated or empirical load-elongation data. The instantaneous horizontal distance of attachement point A from effective mooring point P is
31
given by the following relationship:
[
L A = ( x0 + x A cosψ − y A sinψ ) 2 + ( y 0 + x A sinψ + y A cosψ ) 2
]
1 2
(22)
Finally, the formulation for force couple FA is as follows:
⎧⎪ FA cos (ψ A −ψ ) ⎫⎪ FA = ⎨ FA sin (ψ A −ψ ) ⎬ ⎪⎩ x A FA sin (ψ A −ψ ) − y A FA cos (ψ A −ψ )⎪⎭
(23)
where ψA is the direction of horizontal mooring retoring force relative the x-axis in the vessel-bound coordinate system O x y . COMPUTATIONAL PROCEDURE The basic principles underlying the computational procedure can be stated symbolically in a concise form, representing system dynamics by equations (1-6) and (9-10) in a canonical form as the following generalized state equation: x& = f ( x , c , t ) (24) expressing the rate of change of dynamic state vector x = (u , v, r , x0 , y 0 ,ψ ) of order nd = 6 without memory and x = u , v, r , x0 , y 0 ,ψ , s0T , s1T , ⋅ ⋅ ⋅ , s nT T of nd = 6 + 3(n+1) with memory. Vector f is a function of the state vector x , of a time independent parameter vector c , and of time t, which is the independent variable. If the governing state equation does not contain time explicitly, that is, there exists no timedependent external excitation, then the system is autonomous; otherwise, the system is nonautonomous. In the autonomous mode, the vessel is subject to a steady current, a constant wind, and to wave forces that are reduced to their time-averaged components by disregarding the purely oscillatory components. In the nonautonomous mode, time-varying wave forces, comprising oscillatory first-order wave forces and slowly-varying wave drift forces, are included. Due to the strong nonlinearity of equation (24), there generally exists no analytical solution, and results can only be obtained numerically. Vector c in equation (24) comprises all system parameters required for the specification of forces, including control parameters (rudder angle and propeller rate), operational parameters (attachment point location and mooring hawser length), and environmental parameters (current speed, wind velocity, and incident wave identifiers). Essentially three groups of parameters (coefficients) specify all forces in our mathematical model. First, parameters are required for the inertial terms and the empirical formulas for mooring restoring and wind forces; mostly, these comprise vessel main dimensions and certain coefficients. Second, depending on hull geometry, propeller and rudder, more than 50 parameters are needed to quantify the forces in the four-quadrant maneuvering model; generally, these are obtained from planar motion tests with a captive model in a towing tank. Third, parameters must be identified for the memory associated forces as well as for the first- and second-order wave forces; usually, these are calculated from threedimensional boundary element methods.
(
)
T
STABILITY ANALYSIS In the autonomous mode, the moored vessel will have equilibrium positions x E , which are mathematically defined as follows: f (xE , c ) = 0 (25) Physically, this equation means that the time rates x& of all state variables have to vanish at system equilibria. The equilibrium positions are determined as iterative solutions of this nonlinear algebraic equation. Depending on parameter values chosen, there may exist one or more equilibrium positions with their associated basins (domains of attraction). The boundary between those basins can be
32
complicated, sometimes fractal. To classify the equilibria, stability analyses are performed in the vicinity of each equilibrium, Therefore, the nonlinear state equation about the equilibrium x E is linearized as follows: y& = A y
(26)
The perturbation vector y = x − x E is then amenable to analytical solutions. The Jacobian matrix is
A=
∂f (x E , c ) ∂x
Local stability of any examined state is usually assessed in the sense of Liaponov by solving a classical eigenvalue problem.
A − Iσ = 0
(27)
where I is the unit matrix having the same dimension as A , and σ denotes the eigenvalues. If A has no eigenvalues with a zero real part, then the equilibrium position is hyperbolic or nondegenerate. Otherwise, it is an elliptical or degenerate equilibrium. The classification of hyperbolic equilibrium is straightforward. If all real parts are negative definite, the equilibrium state is stable, and the autonomous system has to asymptotically return to it after a sufficiently small, arbitrary initial disturbance. If one or more real parts are positive definite, the equilibrium state is unstable. Even though the initial disturbance is arbitrarily small, the system will almost never return to such an unstable equilibrium. It may asymptotically wander away to a neighboring stable equilibrium, enter a periodic orbit (limit cycle), get trapped in a quasi-periodic orbit on a torus, or execute chaotic motion. In a degenerate case the linear stability criterion is not sufficient for classification of equilibrium. Such elliptical equilibria define a center manifold in state space, and the corresponding parameter values define a bifurcation point in parameter space. NUMERICAL SIMULATIONS The global behavior of the nonlinear differential equation (24) can generally be approximated by numerical integration. The operation is symbolized as follows:
{x (0) , c } ⇒ x& = f (x , c , t ) ⇒ x (t )
(28)
where initial states x (0 ) and system parameters c are the input, while system response x (t ) is the output. Strong nonlinearities in the dynamics, specially in the mooring restoring forces, may cause this differential equation to become numerically stiff. Therefore, numerical computations call for integration with controlled accuracy. To interpret the simulated motions, it is useful to know whether the subject system is conservative or dissipative. The divergence (div) of the nonlinear flow (equation 24) is as follows: div f ( x , c , t ) = spur A
(29)
The trace (spur) of Jacobian matrix A is defined as follows: spur A =
nd
∑A
jj
(30)
j =1
with A=
∂f (x , c , t ) ∂x
If the time averaged mean values of equation (29) are negative, then the system is dissipative or nonconservative. However, the distinction between a conservative and a dissipative system is not a trivial numerical task, due to the required long-term predictions of the trajectory. From the physical point of view the systems dealt with here have both potential and viscous damping terms that cause the system to be dissipative. The asymptotic behavior of such dissipative systems can be designated as an
33
attractor. Typical attractors are a fixed point, a limit cycle, a torus, and a strange (chaotic) attractor. One characteristic feature of chaos is the extreme sensitivity to small change in initial conditions, implying long-term unpredictability. NUMERICAL PREDICTIONS The two exemplary supertankers TOKIO MARU and ESSO OSAKA were chosen for computational studies. Their principal particulars are given in Table 1. Specific system parameter values constituting the four-quadrant maneuvering model of TOKIO MARU in deep water were documented by Oltmann and Sharma (1984); corresponding parameter values of ESSO OSAKA in deep and shallow water, by Jiang and Sharma (1993). Tanker ESSO OSAKA meanwhile has been recommended by the ITTC as a standard ship for comparative maneuvering studies of all kinds. The mooring restoring coefficients are quantified as C1 = 176.69 kN/m, C2 = –2.44 kN/m² and C3 = 0.11 kN/m3 for BTM or ITM, as C1 = 15.92 kN/m and C2 = 0.68 kN/m² for ATM, as C4 = 0.0113 kN/m4 (TOKIO MARU) and C4 = 0.025 kN/m4 (ESSO OSAKA) for SBM. The corresponding horizontal mooring hawser forces are graphically shown in Fig. 3. Table 1 Principal particulars of subject tankers Length between perpendiculars Length at waterline Beam Draft Block coefficient Number of propellers Screw sense Number of rudders Rudder area
Seven environmental parameters were selected as varying system parameters: current speed (VC) and direction (ψC), wind speed (VW) and direction (ψW), significant wave height (H1/3), mean wave period (T1) and wave direction (ψS); two control parameters: rudder angle (δ) and propeller rate (n); and two operational parameters: mooring hawser attachment location (xA, yA) and hawser length (LAU).
TOKIO MARU 290.0 m 296.4 m 47.5 m 16.1 m 0.81 1 righthanded 1 73.5 m²
ESSO OSAKA 325.0 m 335.0 m 53.0 m 21.8 m 0.83 1 righthanded 1 124.7 m²
Fig. 3 Horizontal restoring forces of four mooring configurations (Sharma et al. 1994)
Hopf Bifurcations For the two exemplary supertankers numerous stability analyses were performed in parameter space and carried out as corresponding time domain simulations in state space, intending to cover a wide practical range of parameter values. For the tankers subject to a steady current without wind and waves under single buoy mooring configurations, linearized stability analyses demonstrated that dynamic (Hopf) bifurcations occur. Mathematically, this means that there is a complex eigenvalue pair whose real part crosses the stability limit Re[σ] = 0 in parameter space. As current speed is increased (around VC = 0.45 m/s), the real part of the critical eigenvalue of the Jacobian matrix crosses the stability limit, indicating transition from stable to unstable equilibrium. The global behavior near this Hopf bifurcation was investigated by performing two numerical simulations having identical fixed parameter values except for current speed. At a current speed of VC = 0.2 m/s (before Hopf bifurcation), the tanker asymptotically returned to its stable equilibrium after initial disturbance. Such an attractor is
34
considered a fixed point. As current speed is increased to VC = 0.5 m/s (just after Hopf bifurcation), even though the initial disturbance is arbitrarily small, the tanker never returned to its unstable equilibrium but asymptotically entered a limit cycle. The two simulated trajectories of the SBM tanker ESSO OSAKA subject to current, showing the transition from a fixed point to a limit cycle, see Sharma et al. (1994). If the SBM tankers are subjected to the additional effects of wind and waves, subcritical Hopf bifurcations occur in certain parameter subspaces. A critical feature of subcritical Hopf bifurcations is the coexistence of a fixed point and a limit cycle of system behavior. Simulated trajectories of the SBM tanker TOKYO MARU subjected to wind and waves directed opposite to current were carried out with the following parameter values: VC = 1.5 m/s, LAU = 75 m, xA = 145 m, yA = 0, n = 0, δ = 2.0°, H1/3 = 2 m, T1 = 10 s, ship condition. They show the coexistence of two attractors before the subcritical Hopf bifurcation (Sharma et al. 1994). For identical parameters values, depending on initial conditions chosen, the tanker asymptotically returns to either a fixed point or it enters a stable limit cycle. Global Bifurcations and Chaotic Motion A common feature of dissipative systems is that global behavior changes qualitatively if one parameter value Fig. 4 Qualitatively different attractors of SBM tanker changes. This phenomenon is shown in ESSO OSAKA subject to current in shallow water Fig. 4, displaying the rich variety of (Sharma et al. 1994) attractors attainable by the tanker ESSO OSAKA in a steady current. Results shown here are for the ship condition in shallow water with the following parameter values: LAU = 78 m, xA = 162.5 m, yA = 0, n = 0, δ = 0. As current speed increases, the global behavior is first characterized by a fixed point (a), then a limit cycle of period close to the linearized eigenperiod (b), next a limit cycle with approximately twice the linearized eigenperiod (c), followed by chaos (d), reverting surprisingly to a shoe-shaped limit cycle of single eigenperiod (e), and moving on to an oval limit cycle (f). At the local level, after the supercritical Hopf bifurcation, all cases differ only quantitatively, i.e., they are all dynamically unstable. The possible pathway from a limit cycle (case b) to chaotic motion (case d) can be understood here as a sequence of global bifurcations. Influence of Alternative Mooring Configurations One of the most important nonlinearities in the dynamic system arises from the mooring restoring characteristics. They are totally different in alternative mooring configurations; see Jiang et al. (1995). To demonstrate the influence of mooring restoring force characteristics on stability of the SPM tanker, linearized stability analyses were performed of tanker TOKYO MARU subject to steady current for identical conditions, but under buoy mooring (SBM) and articulated tower mooring (ATM) configurations. The stability analysis revealed that increasing current speed is destabilizing and increasing rudder deflection is stabilizing. But the primary effect here is that restoring force
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characteristics resulting from the different mooring configurations have no remarkable influence on Hopf bifurcations. To investigate effects of changing the mooring restoring force characteristics on global behavior, two simulations were carried out of tanker TOKYO MARU subject to steady current for identical conditions, and again under buoy mooring (SBM) and articulated tower mooring (ATM) configurations. These simulations revealed that the mooring restoring force characteristics affect the form of the asymptotic motion trajectories as well as their amplitudes (Jiang et al. 1995). Influence of Water Depth Hydrodynamic response and control forces are the other most important nonlinearity in the system. They vary strongly with water depth. The influence of water depth on system behavior is manifested by two simulations shown in Fig. 5. The tanker ESSO OSAKA was moored in deep (solid lines) and shallow (dashed lines) water under otherwise identical conditions: VC = 1.5 m/s, LAU = 78 m, xA = 162.5 m, yA = 0, n = 0, δ = 0, model condition. Simulated time histories show that hawser tension peaks are an order of magnitude higher in shallow water than in deep water. Furthermore, the asymptotic trajectories are qualitatively different. Whereas the tanker enters a periodic limit cycle in deep water, it executes a chaotic attractor in shallow water with larger excursions.
Fig. 5 Simulated time histories and trajectories of SBM tanker ESSO OSAKA subject to steady current in deep and shallow water (Sharma et al. 1994)
Stabilizing Measures The practical consequences of self-sustained oscillations or even chaotic motions of the SPM tanker are that they cause large motion amplitudes and high line tensions peaks. The large motion amplitudes impair safe operations of the tanker, and high hawser peaks threaten a possible hawser break leading to a tanker accident. Simple practical measures have been found to stabilize the equilibrium states, thereby reducing motion amplitudes and mooring hawser tensions. They are static rudder deflection, reverse propeller rate, and asymmetric fairlead location or their suitable Fig. 6 Simulated time histories and trajectories of combinations, see Jiang et al. (1987) and Sharma SBM tanker ESSO OSAKA subject to current with et al. (1988). Figure 6 confirms the stabilizing and without rudder application (Sharma et al. 1994) effect of static rudder deflection for the SBM tanker ESSO OSAKA in ship condition in shallow water. Simulated time histories show strongly reduced hawser tension peaks when rudder is applied (dashed line). Trajectories asymptotically lead to a fixed point with rudder hard to starboard (δ = –35°), but to a chaotic attractor with rudder amidships (solid line). EXPERIMENTAL VALIDATIONS To validate the numerical predictions, model experiments were conducted in the main towing tank
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(200 m x 9.8 m x 1 m) of the shallow water research facility DST in Duisburg (Development Centre for Ship Technology and Transport Systems, formerly VBD), which has the feature of controllable water depth from zero to one meter. A 1:65 scale model of tanker ESSO OSAKA was used. The main purpose of model experiments was to physically simulate and record the dynamic behavior of an SPM tanker in a steady current. Specifically, the following quantities were measured as functions of time: horizontal motions of the model in three degrees of freedom (surge, sway, and yaw) and mooring hawser tension as well as elongation. Froude dynamic similarity was assumed, i.e., no attempt was made to compensate for the difference in Reynolds number between model and full-scale. Thus all results of model experiments correspond to numerical simulations for the model condition. Current speed was simulated by towing the model through still water. The nonlinear load-elongation characteristics (see Fig. 3) were dynamically modeled by employing, in series, a cascade of several linear springs of successively increasing stiffness and individually bounded extension. For details of model experiments and measured results, see Jiang and Sharma (1992).
Fig. 7 Measured and calculated time histories of SBM
Self-Sustained Oscillations tanker ESSO OSAKA in shallow water A calculated time history (dashed lines) is (Jiang and Sharma 1992) compared to a corresponding measured time history (solid lines) in Fig. 7 for the SBM tanker subject to a steady current in shallow water. The parameter values are as follows: VC = 2.0 m/s, LAU = 78 m, xA = 162.5 m, yA = 0, n = 0, δ = 0. For this locally dynamically unstable case identified by stability analysis, the corresponding horizontal motions from numerical simulation as well as from model measurement are characterized by slow selfsustained oscillations (fish-tailing motions), which are often observed in practice when tankers are moored to SPM systems. Furthermore, these slow motions cause high loads in the mooring hawser as seen in the time histories of hawser tension. The agreement between calculation and measurement is remarkable for motions as well as for the high peaks in hawser tension. Similar agreements were also achieved for other compared cases, thus experimentally validating the mathematical model. Influence of Water Depth To experimentally demonstrate the influence of water depth, model tests were conducted for the SBM tanker subject to a steady current in deep and shallow water under otherwise identical parameter values, but different initial conditions. The corresponding time histories for deep and shallow water are documented by Sharma et al. (1994). Both cases reflect dynamically unstable equilibria (which can be numerically verified by stability analysis); however, motion amplitudes are much larger and hawser tension several times higher in shallow water than in deep water. Stabilizing Measures To validate the effectiveness of the numerically found stability measures, measured time histories with rudder deflection were compared with those with rudder amidships. Results showed that when rudder is applied to a static value of δ = –35°, self-oscillations of tanker motion and high peaks in hawser tension both disappear after the initial disturbance. Reverse propeller rate, as another stabilizing measure, works in the same manner.
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SPM MOORINGS FOR AN OFFSHORE TRANSSHIPMENT TERMINAL In 2002, Oldendorff Carriers (2006) installed the world’s largest offshore coal transshipment terminal in the Bay of Iskenderun, Turkey. Around 20 capesize shipments with about 3.2 million tons arrive annually from Columbia and South Africa to feed the coal to a stockpile located adjacent to a recently commissioned power station. The transshipper ISKEN, a twin-hulled, nonpropelled floating cargo terminal, and two gravity type selfunloading hopper barges of 118 m length and 27 m breadth are part of the terminal. While one barge is loaded by the transshipper, the other barge is busy with Fig. 8 The transshipper loading coal from a bulker unloading at the shoreside pier. The water into a hopper barge depth at the pier is only 6.0 m, so the barges are designed to carry about 10,000 t on a shallow draft of 4.95 m. With a length of 107 m, a breadth of 44 m, and a total height of 53 m, the transshipper is an impressive steel structure displacing 5700 t. For weathering severe storms, the transshipper and the two hopper barges can be towed to three single point mooring buoys. Each buoy is moored the sea bed by six catenary anchor leg mooring (CALM) chains. A mooring hawser holds the vessels captive to a turntable mounted on top of the buoys by means of a slewing bearing. This bearing allows the turntable to freely weathervane, and the moored vessel can take up the most favorable alignment of least resistance to the prevailing current, wind and waves. The mooring analysis was based on three-hour storm duration. The Deutsche Wetterdienst (DWD), Fig. 9 One of three mooring buoys under Hamburg, furnished wave statistics for the site. A construction constant one-minute average wind speed of VW = 34.1 m/s, a current speed of VC = 0.75 m/s, and a seaway with a significant wave height of H1/3 = 6.5 m and a mean wave period of T1 = 6.9 s characterized the environmental conditions. The severity of these conditions relative to the shallow nominal water depth of 42 m called for a mooring system flexible enough to tolerate large horizontal motions of the moored vessels. The CALM buoy by itself did not supply sufficient horizontal flexibility. A long stretchable hawser was required to produce the needed overall flexibility of the system. The mooring systems comprise six 76 mm diameter Grade 3 stud link chain cables of 275 m length and six Stevpris anchors of adequate holding power. Each mooring chain, pretensioned to 834 kN, extends from the 7.0 m diameter mooring buoys in a spread pattern at 30° intervals around each buoy. Figure 9 shows one of the mooring buoys under construction. A quasi-static analysis method was performed to evaluate the performance of the SBM systems, following the recommendations for design and analysis of floating production systems (API 2001) and station keeping systems for floating structures (API 1997). The effects of line dynamics were accommodated through the use of a relatively conservative safety factor. To account for possible unsteady yaw motions, the moored vessels were analyzed under a 15° angle of attack, i.e., ψA = 15° and ψW = ψC = ψS = 180°. Wind, current, and wave forces were assumed to act collinearly on the moored vessels, subjecting
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same to a mean steady-state environmental force in the horizontal direction. These mean forces cause the vessels to be displaced from their initial no-load equilibrium position to a new equilibrium position, referred to the mean horizontal offset. The wave forces cause the vessels to oscillate about this mean offset. The vessels’ maximum horizontal offset is thus the sum of the steady-state mean offset and the wave-induced horizontal displacement. The wave-induced displacement comprises first-order (highfrequency) motion response at wave frequencies as well as second-order (low-frequency) slowly varying wave drift motion response. Tensile forces in the mooring hawser and the anchor chains are then a function of the maximum horizontal offsets. These forces were evaluated quasi-statically, with appropriate safety factors accounting for dynamic effects. Standard spectral techniques determined oscillatory first-order wave-induced motions, the mean horizontal wave drift forces, as well as the associated drift displacements of the moored vessels. Dedicated computer codes yielded transfer functions of first-order motions (Papanikolaou and Schellin 1992), second-order (quadratic) drift force coefficients (Clauss et al. 1982), and drift damping coefficients (Schellin and Kirsch 1989). The JONSWAP spectrum with a peakedness parameter of 3.3 defined the wave spectral energy density of the seaway. A cosine Fig. 10 Schematic of the SBM systems squared spreading function described of wave energy distribution about the principal wave direction. Another computer code (Schellin and Scharrer 1981) evaluated restoring forces of the SBM systems, based on catenary equations modified for elastic stretch and bottom friction of the anchor chains as well as elastic stretch of the mooring hawsers. The mooring systems as analyzed are schematically depicted in Fig. 10. Table 3 summarizes the resulting mean environmental forces acting on the moored vessels. For the transshipper and the barges, laden as well as in ballast, the steady sate force due to wind turned out to be the dominant part of the total mean environmental force acting on the moored vessels. Table 3 Mean environmental forces Mean force [kN] Total mean environmental force Mean wind force Mean current force Mean wave drift force
Transshipper 2496 2302 21 173
Barge (laden) 1038 777 36 225
Barge (in ballast) 1135 984 27 124
The expected maximum horizontal offset was obtained by adding to the mean offset a combined value of high- and low-frequency horizontal displacements. This combined value was the sum of the maximum value of the dominant component and the significant value of the other component, where the dominant component was that with the highest value. For all three SPM systems investigated, high-frequency vessel motions were the dominant component as they always exceeded low-frequency motions. Maximum horizontal offset xmax was thus obtained as follows: xmax = x E + x HF max + x LF sig
(31)
where x E is the mean offset, x HF max is the statistical maximum amplitude of the high-frequency horizontal motion, and x LF sig is the significant amplitude of the low-frequency horizontal motion. The statistical maximum and significant values were calculated as follows:
x HF max = x HF sig [ 0.5 ln (T ⋅ f 0 )] 1/ 2 and
x LF sig = 2σ LF
(32)
where T is the time in seconds of the three-hour storm duration, f 0 is the average up-crossing frequency of the high-frequency motion, and σ LF is the standard deviation of the low-frequency
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MBL = 4218 F O R C E [k N ]
motion. The method used to yield maximum tension in the most loaded anchor chain and maximum force in the mooring hawser is illustrated in Fig. 11, here for the transshipper single-point moored using a 270 m long hawser with an elastic coefficient of 100 kN/m. (In this figure, FE stands for total mean environmental force, TCHAIN for chain tension, THAWSER for hawser force, MBL for minimum breaking strength, xE for mean offset, and xMax for maximum offset.) For the transshipper, a laden barge, and a barge in ballast, Table 4 summarizes the resulting maxima of chain tension, hawser force, horizontal offset, and hawser stretch. Safety factors against minimum break load for the most loaded anchor chain and the hawser are given as well. According to design conditions for a severe storm, the quasi-static method applied here required a safety factor of 1.8 applied against the minimum rated breaking strength of the anchor chains and the hawser (Germanischer Lloyd 2007).
4000
THAWSER = 2926
3000
FE = 2496
TCHAIN = 2274
2000
1000
XE = 28.19
XMax = 32.70
0 0
10
20
30
40
50
60
70
80
OFFSET [m] MOST LOADED CHAIN
HAWSER FORCE
Fig.11 Chain tension and hawser force for the moored transshipper
Table 4 Maximum chain tension, hawser force, horizontal offset, and hawser stretch for the transshipper, a laden barge, and a barge in ballast Maxima Chain tension (SF)
Transshipper
Barge (laden)
Barge (in ballast)
2274 kN (1.9)
1388 kN (3.0)
1556 kN (2.7)
Hawser force
2926 kN
1313 kN
1632 kN
Horiz. Offset
32.3 m
15.4 m
18.9 m
Hawser stretch
29.3 m
13.1 m
16.3 m
CONCLUSIONS An experimentally validated, comprehensive mathematical model is presented and applied to investigate the dynamics of different configurations of single point moored vessels. Local linearized stability analyses and nonlinear numerical simulations of an SPM tanker reveal that multifarious dynamic phenomena may occur, leading to large motion amplitudes and high line tension peaks. Simple practical measures, e.g., static rudder deflection and reverse propeller rate can be applied to stabilize the tanker’s equilibrium states, thereby reducing motion amplitudes and mooring hawser tensions. Restoring force characteristics resulting from alternative mooring configurations have no noticeable influence on local stability, but they strongly affect the global behavior. Water depth is a relevant parameter; motion amplitudes and hawser tension peaks are higher in shallow than in deep water. For SPM systems where some of the parameters required for the application of the presented mathematical model are unavailable, the rational analysis method recommended by API (1997), combined with an understanding of the environment, the characteristics of the vessel being moored, and other relevant factors, can be used to evaluate single point mooring systems.
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REFERENCES API (2001). “Recommended Practice for Planning, Designing, and Constructing Floating Production Systems,” American Petroleum Institute, RP 2FPS, Washington, D.C. API (1997). “Recommended Practice for Design and Analysis of Station Keeping Systems for Floating Structures,” American Petroleum Institute, RP 2SK, Washington, D.C. Clauss, G., Sükan, M., and Schellin, T.E. (1982). “Drift Forces on Compact Offshore Structures in Regular and Irregular Waves,” J. Applied Ocean Research, Vol. 4, pp. 208-218. Faltinsen, O.M., Kjaerland, O., Liapis, N., and Walderhaug, H. (1979). “Hydrodynamic Analysis of Tankers at Single-Point Mooring Systems,” Proc. 2nd Int. Conf. on Behavior of Offshore Structures, London, Paper No. BOSS 59. Germanischer Lloyd (2007). Rules for Classification and Construction IV, Industrial Installations, 6 Offshore Technology, Hamburg. ITTC (1984). Proc. of the 17th Int. Towing Tank Conf., Göteborg. Jiang, T., Schellin, T.E., and Sharma, S.D. (1995). “Horizontal Motions of an SPM Tanker Under Alternative Mooring Configurations,” J. Offshore Mechanics and Arctic Engineering, Vol. 117, pp. 223-231. Jiang, T. and Sharma, S.D. (1992). “Investigation of Horizontal Motions of an SPM Tanker in Shallow Water Through Computation and Model Experiments,” Proc. 19th Symp. on Naval Hydrodynamics, Seoul, National Academy Press, Washington, D.C., pp. 405-424. Jiang, T., Schellin, T.E., and Sharma, S.D. (1987). “Maneuvering Simulation of a Tanker Moored in a Steady Current Including Hydrodynamic Memory Effects and Stability Analysis,” Proc. Int. Conf. on Ship Manoeuvrability, RINA, London, Vol. 1, Paper No. 25. Matter, G.B., Sphaier, S.H., and Sales Jr., J.S. (2007). “Definition of the Best Position for the Turret in an FPSO Based on a Hydrodynamic-Structural Analysis,” Proc. 26th Int. Conf. on Offshore Mechanics and Arctic Engineering, ASME, San Diego, Paper No. OMAE 2007-29474. Munipalli, J., Pistani, F., Thiagarajan, K.P., Winsor, F., and Colbourne, B. (2007). “Weathervaning Instabilities of an FPSO in Regular Waves and Consequence on Response Amplitude Operators,” Proc. 26th Int. Conf. on Offshore Mechanics and Arctic Engineering, ASME, San Diego, Paper No. OMAE 2007-29359. OCIMF (1994). Prediction of Wind and Current Loads on VLCCs. Oil Companies Int. Marine Forum, London, Witherby & Co. Ltd. Oldendorff Carriers (2006). “Transshipper ISKEN in Turkey – 40,000 tpd,” Update on our Coal Transshipment Activities, Lübeck, Germany. Owen, DG. and Linfoot, B.T. (1985). “Theoretical Analysis of Single Point Mooring Behavior,” Trans. SNAME, Vol. 85, pp. 315-324. Papanikolaou, A.D. and Schellin, T.E. (1992). “A Three-Dimensional Panel Method for Motions and Loads for Ships with Forward Speed,” J. Ship Technology Research, Vol. 39, pp. 147-156. Schellin, T.E. and Kirsch, A. (1989). “Low-Frequency Damping of a Moored Semisubmersible Obtained from Simulated Extinction Tests and Mean Drift Forces,” J. Applied Ocean Research, Vol. 4, pp. 202-211. Schellin, T.E. and Scharrer, M. (1981). “Design Principles to Select Effective Mooring Systems, Hansa, Vol. 118, No. 6, pp. 432-438 (in German). Sharma, S.D., Jiang, T., and Schellin, T.E. (1994). “Nonlinear Dynamics and Instability of SPM Tankers,” Fluid Structure Interaction in Offshore Engineering, S.K. Chakrabarti (ed.), Computational Mechanics Publications, Ashurst, UK, pp. 85-123. Sharma, S.D., Jiang, T., and Schellin, T.E. (1988). “Dynamic Instability and Chaotic Motions of a Single-PointMoored Tanker, Proc. 17th Symp. on Naval Hydrodynamics, The Hague, National Academy Press, Washington, D.C., pp. 543-563. Vazquez-Hernandez, A.O. (2007). “FPSO Conceptual Design System Tools Considering Hurricane Data Base and Production Requirements,” Proc. 26th Int. Conf. on Offshore Mechanics and Arctic Engineering, ASME, San Diego, Paper No. OMAE 2007-29102. Wichers, J.E.W. (1976). “On the Slow Motion of Tankers Moored to Single-Point Mooring Systems,” Proc. Offshore Technology Conference, Houston, Paper No. OTC 2548.
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OSV Singapore 2007 Jointly organized by Joint Branch of RINA-IMarEST Singapore and CORE. 24-25 September 2007
Experiences from Hardware-in-the-loop (HIL) Testing of Dynamic Positioning and Power Management Systems Tor A. Johansen, Asgeir J. Sørensen, Ole J. Nordahl*, Olve Mo, Thor I. Fossen Marine Cybernetics, Vestre Rosten 77, NO-7046 Tiller, NORWAY *Statoil ASA, NO-4035 Stavanger, NORWAY
ABSTRACT The complexity of control systems on offshore ships and rigs is rapidly increasing, and the correct operation of these systems is becoming even more critical for the safe and efficient operation of the vessels. The cost and risk related to incidents and losses due to problems involving control system software, hardware and operators errors are of significant concern for vendors, yards, ship owners, contractors, class societies and oil companies. HIL simulator testing of control systems on ships and rigs has been commercially available since 2004 bringing this technology from aviation and automotive industries into the offshore and maritime industries. The main idea is enhanced system verification using advanced simulators capable of simulating vessel response with its installed systems and equipment for a wide range of operational conditions and single and multiple failure modes in order to verify correct functionality and performance. In particular, redundancy, alarm and failure handling functions of the systems can be tested in detail. The experience from testing of a ten-fold offshore service and construction vessels has proven that Hardware-in-the-loop (HIL) simulator testing provides a significant more in-depth and costeffective testing compared to conventional test methods. HIL testing contributes to reduced risk for unwanted surprises due to design flaws and errors that eventually lead to delayed delivery of the vessel. In ordinary operations HIL testing contributes to improved safety and reduced number of incidents that lead to expensive down time and vessel off-hire costs. HIL testing is organized such that extensive testing is done early in the design and building process, when the cost implications for finding and correcting errors are small. Later in the building process testing of systems integration is provided.
INTRODUCTION Hardware-in-the-loop (HIL) simulation is a technology for testing of control system software and hardware that is widely used and well proved in the automotive and aerospace industries. Now, it has been recently introduced and adapted to the maritime and offshore industries for independent testing of safety- and mission-critical control systems on advanced offshore vessels, such as dynamic positioning systems (DP) and power management systems (PMS). The HIL simulator technology must accommodate the requirements for the various vessels with corresponding
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configuration and customization reflected in the onboard systems, equipment, and control systems. A HIL simulator test setup involves a real-time dynamic simulator that simulates all signals to, and receives all command signals from, the control system being the test target. For testing of a DP computer control system this means that the HIL simulator will simulate the vessel motion in response to thruster forces commanded by the DP system in addition to wind, wave and current loads. In addition, it must simulate nomal and faulty behaviour of the sensors, position reference systems, power plant, and relevant equipment involved in the marine operations such as hawsers, risers, and mooring lines. The interface between the HIL simulator and the target control system may be the existing signal interface (hardwired analog and digital signal, network protocols, bus protocols etc.) or a dedicated network or bus HIL simulator interface built into the test target in order to facilitate safe and efficient interfacing. Examples of HIL simulator setups for testing of Power Management Systems are given in Figures 1 and 2. Further information can also be found in [2], [3] and [9]. The main motivation for introducing HIL simulator testing to the offshore and maritime industry is due to its potential for increased safety and reduced cost. The objective is to test the control systems earlier, broader and deeper than today’s practice. Earlier testing means that the systems will be more ready when installed on the vessel, leading to faster commissioning time with less risk. Since HIL simulation allows almost any scenario to be realistically simulated well before the vessel is ready for its full scale sea trial program, there will be less surprises during the expensive full scale trials. This means that there will be significant cost savings. Deeper and broader testing means that the control system is tested by simulation for more failure modes, more operational modes, and more environmental and weather conditions. In addition, known failure scenarios or series of events that have been experiences can be simulated. This leads to increased operational availability, less incidents, and less off-hire. HIL testing of dynamic positioning (DP) systems was first demonstrated on the DP class 2 offshore service vessel Viking Poseidon in 2004. Since then a number of new buildings and upgrades involving HIL testing of DP systems have been completed. In 2006 HIL testing of power management systems (PMS) was developed and implemented on several new building projects.
THE PRESENT MARITIME CONTROL SYSTEM TEST AND APPROVAL REGIME The present test and approval regime for control systems in the maritime and offshore industry typically consists of the following activities and milestones: •
Factory Acceptance Testing (FAT). The control system is powered up and configured at the site of the manufacturer. In some cases the FAT may involve a degree of integration testing when several systems (possibly from different manufacturers) are interconnected and tested simultaneously. The equipment manufacturer normally develops an FAT program that is approved by the classification society, who is attending and approving the FAT together with the ship yard and ship owner. Conventionally, for most control systems up to today the FAT is focusing on signal interfaces and functional design, but not on testing of functionality.
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•
•
Commissioning testing and Customer Acceptance Test (CAT). Throughout the construction of the vessel the different systems and interfaces are tested as they are commissioned. As soon as the commissioning of the different systems are completed, the CAT is the key milestone involving the different parties. It usually involves full scale tests. For advanced vessels with redundancy requirements, such as DP equipment class 2 or 3, it is a requirement of IMO and classification societies that a Failure Mode and Effect Analysis (FMEA) is made for the vessel, see ref [5]. The FMEA involves proving trials that for DP class 2 systems shall demonstrate that a single point failure in any active component in the system does not lead to loss of position. For DP class 3 systems there are additional requirements such that failures on passive components, including fire and flooding of a single compartment, shall not lead to loss of position. The DP system FMEA proving trials are normally carried out during sea trials together with the DP system CAT and class approval. This is normally one of the last activities in a new build program since all important systems needs to be fully operational and individually tested and approved before FMEA proving trials commence.
IMO is recommending ship owners to carry out annual DP trials, see reference [5], in order to demonstrate that the FMEA is still valid, i.e. has not been invalidated by degraded or worn equipment, maintenance, or changes in equipment and control system hardware and software. Classification of the vessel is also renewed periodically, e.g. every 5 years. Additionally, for major equipment or software upgrades or retrofits it may be considered necessary to carry out FAT, commissioning testing, CAT and new FMEA trials.
LIMITATIONS OF THE PRESENT TEST REGIME AND BENEFITS OF HIL TESTING Although the present test regime involves many test activities and parties, it is clear that there are limitations and possibilities for improvements with respect to depth, test coverage, and efficiency, see reference [1] for a review. In particular, while FMEA is well suited to analyze equipment and hardware redundancy, it is not a suitable tool for analysis and testing of software and functionality [4]. While equipment and hardware may fail unexpectedly and randomly due to degradation, wear, stress and so on, it is clear that software does not wear or age. On the other hand, software functions may contain hidden errors that cannot be observed or detected until particular conditions arise. Such conditions may be rare, but critical, and may be conditions such as handling of equipment degradation or failure in combination with certain operational modes, conditions or operator input. An FMEA identifies components and systems that may fail, but does usually not go into depth to analyze all the different ways each component or system may fail. FMEA trials are usually conducted by tripping equipment or disconnecting power or signal cables to simulate a circuit break due to practical reasons. On the other hand, it is widely acknowledged that advanced equipment, such as DGPS receivers, electric thrusters, diesel generators, and computer networks, may fail in much more intricate and less easily identifiable manners.
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One area of concern is common mode failures, where two or more apparently independent systems may fail simultaneously or sequentially due to one common cause. Well known examples are satellite failures in the GPS system when using two or more DGPS receivers, certain load sharing systems in electric power plants, and cooling of equipment located together or on the same cooling circuit. Control system software is one particular source of several units failing simultaneously. Even if controller and I/O unit hardware are duplicated or tripled in a redundant control system, it is common practice that the redundant controllers run exactly the same software. Hence, they may fail simultaneously in the same way. Moreover, the correct handling of failures occurring in equipment such as sensors, thrusters, and generators, is usually a software based function. This means that it is not sufficient to have redundant equipment standby unless the control software is able to correctly detect, identify, alarm, and isolate the failure. It is fair to say that in a system with redundant equipment, the weakest links are expected to be the system that integrates them, that is software, computer networks, power distribution and certain auxiliaries such as cooling. HIL testing makes computer control sofware testing more powerful and efficient due to its ability to excitation of the relevant types of failure modes that exist in a computer-controlled system, •
•
Realistic operational scenarios can be simulated with a HIL simulator connected to the target control system in order to verify the functionality of the control system under different operational modes and conditions, in order to verify that there are no hidden failure in the control system software that are triggered by changes in operational conditions or operator input. Failure scenarios can be simulated in order to verify failure detection, identification, alarm and isolation functionality of the software, again under different operational modes and conditions, to ensure that equipment failures can be detected and isolated, that alarms are presented, and that the system switches to a redundant component or function according to the redundancy requirements of the system. HIL simulation allows complex failures such as signal noise, wild points, sensor signal drift, fail-to-maximum, freeze, increased network load, etc. to be reconstructed.
For ships in operation the HIL technology can be used to verify major software and hardware upgrades before they are installed or the accumulated effect of several smaller updates and modifications in periodic or annual tests. The human aspects are important for safe and efficient operation of advanced control systems. HIL simulation technology allows testing to be combined with the development and verification of operational procedures and operator training, ensuring that results of the HIL testing and experiences are taken case of in operations and even to use the simulator as a tool in the operational phase of the vessel for further operator training, with new failure scenarios added over time.
HIL TESTING OF POWER MANAGEMENT SYSTEMS A Power Management System (PMS) is a vital component of a vessel with redundant diesel electric propulsion, such as most DP class 2 or 3 offshore vessels and other vessels such as cruise vessel, passenger ferries, and certain tankers. The main objective of the PMS is to ensure that stable power supply is continuously available, i.e. blackout prevention. This means that no single
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point failure in the power plant will have consequences beyond the worst case single point failure chosen by design, which may typically be short circuit in a main switchboard when operating in a two-split configuration leading to loss of half of the power generation capacity and half of the thruster capacity. In order to achieve this, the PMS functionality may be distributed in several control units such as: • • • •
Switchboard mounted centralized PMS computer system Frequency converters / variable speed drives with load reduction functions Generator protection systems and relays Marine automation system
Typical functionality found in the centralized PMS software may be • • • • • • • • • • • • •
Load sharing (active and reactive power) Load dependent start / stop Mode control Start of standby generator on fault Power reservation Heavy consumer control Blackout prevention Blackout restoration Power plant monitoring Integration with marine automation system Detection, identification, and isolation of failure modes in power generation, distribution, and consumers Detection, identification, and isolation of failure modes in sensors, relays, feedback and command signals Detection, identification, and isolation of failure modes in computers, operator stations, I/O equipment and network
Using a HIL simulator of the power generation, distribution and consumers, it is straightforward to set up scenarios in order to verify these functions. In more detail for failure mode handling, the following common scenarios can be simulated conveniently with a HIL simulator: • • • • • • • • • • • • • • • •
Pre-warning from diesel engines Shutdown of diesel engines Short-circuit of one switchboard Unavailable diesel engine Locked governor – fixed power Loss of fuel supply to one diesel engine Full throttle to one diesel engine Failure in load sharing line of engine governors Reduced max power from engine Loss of generator excitation Full generator excitation Deviating generator excitation Protection trip of generator Protection trip of bus-tie Generator synchronization failure Generator circuit breaker not following command
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• • • • • • • •
Bus-tie synchronization failure Bus-tie circuit breaker not following command Partial blackout Blackout Over / under bus voltage Over / under bus frequency Protection trip of consumers Failure of power reduction function of propulsion/thruster drives
A typical HIL simulator setup is illustrated in Figure 1. CyberSea Simulator
Bus A
Bus B
PMS computer system
Operator Station
PMS controller A
Profibus board 0/1
CyberSea PowerPlant Profibus board 2/3 Simulator NIC
Wago Analog/Digital IO
Ethernet HUB
Operator Station
CyberSea Simulation Manager
PMS controller B
Figure 1: HIL simulator test setup with a CyberSea Power Plant simulator interfaced to a Power Management System. HIL simulator technology may in principle be applied for testing of all control units, but our focus have been to apply it for testing of the switchboard mounted centralized PMS functionality since this is considered the most apparent and valuable application. Typical functionality that are not tested with such a HIL simulator setup may be: • • • • • • •
Wiring in switchboard Protection relay functionality Protection relay settings and selectivity Power system performance such as voltage stability due to the Automatic Voltage Regulator (AVR) tuning Frequency stability due to governor tuning Variable speed thruster drive controller stability and performance Performance of load reduction function in drives. A HIL test verifies that correct load reduction signals are send from PMS, but not that these signals actually are used correctly by the thruster drive.
Technology for HIL testing of the above mentioned functions is certainly viable, see reference [7] and [8]. These functions may, however, be practically tested with the power plant operational.
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In order to support full scale testing with the control system operational, another type of HIL simulator known as an FMEA simulator is available. The objective of the FMEA simulator is to provide a possibility to manipulate control signals flowing between the power plant and the PMS, see Figure 2. Normally this means that the power plant and PMS operates as normal, but it gives us the possibility to simulate the effect of failures by manipulating signals.
PMS computer system PMS HIL FMEA system
Operator Station
PMS controller A
Operator Station
PMS controller B
Power generation, distribution and consumtion
FMEA simulator interface Phoenix contacts on PMS rack
Ethernet HUB
CyberSea Simulation Manager
Phoenix contacts normally connected to PMS rack
NIC
CyberSea FMEA Simulator
Figure 2: HIL simulator test setup with a CyberSea FMEA Simulator interfaced between the Power Plant and the Power Management System.
One example of a scenario that can be easily simulated with such an FMEA simulator is testing of blackout restoration after bus short circuit by simulating short circuit on one bus by simultaneously: • Tripping all generator circuit breakers on the bus • Setting short circuit alarm on the above mentioned circuit breakers • Tripping bus-tie breaker, if operating with closed bus-tie normally
HIL TESTING OF DYNAMIC POSITIONING SYSTEMS A DP system consists of a positioning control system (again consisting of a DP computer system, and sensors and position reference systems), a thruster system, and a power system. At a factory test of the DP, only the DP computer system is operative and is the natural HIL test target. A significant part of the test scope may be completed during the FAT, since there is
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usually most time available for extensive testing. At FAT the focus is on testing of the following aspects: • Basic hardware and software configuration testing, including DP computers, and operator stations. • I/O system hardware and software interface, failure handling, barriers to data transmission buffer and interrupt overload • DP and joystick basic modes / functions and mode switches, including thrust allocation modes, forbidden sectors, rotation points and thruster configuration • Special functions related to marine operations such as pipelay, heavy lift, drilling, shuttle tanker offshore offloading, survey etc. • Power and propulsion system configuration. Verification of the correct devices, their location, ratings, that each device can be selected and deselected without reducing the DP performance. Verification of consequence analysis, power limitation and blackout prevention functions. • Position reference and sensor system configuration. Verification of the correct devices, their location, that each device can be selected and deselected without reducing the DP performance. • Alarms and warnings, verification that all simulated single point failures give the expected alarms or warnings, and that correct handling in the DP computer system is made in terms of disabling the faulty device and automatically switching to a correct redundant device if available. A typical test scope may include hundreds of single point failure tests (involving sensors, position reference systems, thrusters, power system etc.). • DP controller tuning, including gain settings, state estimator performance (model / Kalman filter), feed-forward from wind sensor, dead reckoning performance, wave filtering performance, dynamic capability verification, extreme weather performance. Such extensive testing may give useful indications of performance that should contribute to reducing the time needed for tuning and testing during sea trials. • Multiple failures testing, and testing of relevant reported incidents. At dock, quay and sea trials the test scope should be limited to focus more on the items that have been influenced by the installation and commissioning, i.e. re-tuned, upgraded, or re-configured after FAT. In addition, the focus is moved more towards testing of the integrated DP system, rather that the isolated DP computer system. Additional items to be tested include: • •
Testing of integrated network functionality and barriers. This may include network storm testing, monitoring of traffic, robustness to partial loss of network and messages etc. Testing of mode switch between DP, joystick, manual thruster control, and transfer of command between operator stations.
Further information on HIL testing of dynamic positioning systems can be found in [2].
EXPERIENCES FROM HIL TEST PROJECTS In order to support independent HIL testing, DNV has launched a Standard for Certification of HIL testing, see reference [6]. This standard describes the process and responsibilities of the parties involved. It defines the following key roles: •
HIL supplier, the independent party that supplies the HIL test program and the HIL simulator tools. In order to secure high integrity and value of the testing, the standard
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• •
requires that the company involved must be independent of other parties such as the equipment suppliers, ship yard and ship owners with respect to technology and ownership. HIL test organization, the party that performs the HIL testing and reports the results. This could be a different company than the HIL supplier. There are no requirements for independence. HIL test verification, a third party that witnesses and approves the HIL testing and the HIL test package comprising the tools and documentation. This could be a classification society.
Furthermore, the standard provides information and recommendations for the HIL simulator tools, test coverage, and requirements for verification and validation of the HIL testing. Marine Cybernetics has completed ten-folds of HIL test activities on DP and PMS. The following experiences are noticed: • •
• •
•
•
• •
Comprehensive test coverage requires typically one week of testing at FAT. In order to be well prepared for the final sea trials, it is very useful with a follow-up testing at dock or quay focusing directly on closing of findings after FAT, new or changed functionality, and integration testing. It is noticeable that by coordinating the HIL testing at dock and quay closely with the commissioning of the vendors and ship yard, the HIL testing is not on the critical path of the project and does not increase commissioning time. Due to extensive testing at factory, dock and quay, there are usually few new findings at the sea trials. The findings that have been observed are typically due to hardware problems and integrated functionality that requires full scale testing. Typical number of finding at a factory test may range from 5 to 25 critical findings, and up to 25 less critical findings, for each control system on a vessel. Some of these early findings may have been found later also without HIL testing, we expect, but they would have caused higher cost and delay. However, some of the findings made with HIL testing would not have been found without HIL testing, simply because the test scenarios are not commonly done within the traditional test regime for various reasons such as risk or cost, or because it would be impractical. Sea trials with DP HIL testing take 16 to 24 hours, and the number of findings at dock, quay and sea trials are in most cases less than at factory test. Typically, we have found from 5 to 20 critical findings for each control system on a vessel during these activities, and up to 20 less critical findings. In some cases the findings from HIL testing are easy to classify. Sometimes, they lead to questions regarding design philosophy and operational philosophy, and the solution may be clarification of these rather than changes to the control system. In some cases there are design weaknesses in the vessel that cannot be changed, and one benefit of HIL testing is that these weaknesses are exposed and demonstrated to the operators of the vessel by their participation in the testing, and in the documentation. Recommendations for operational procedures and operator training typically results from HIL testing. Many findings are due to configuration of the systems, and integration between equipment of different types. This means that it is necessary to test the control systems on each individual vessel also if the control systems have the same software basis release. For series of “sister” vessels with “the same” control system software the preferred test regime is to have a combination of a core test scope with variations due to a rotational test scope. In this way the testing benefits all vessels and efficient use of test time is
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made. Since sister vessels are delivered over some time and there may be smaller or larger customizations, or intended or unintended differences in each vessel, the control system software tend to differ to some extent, especially over time. In conclusion, this leads to the following key benefits of HIL testing: •
•
•
Increased safety results from fewer incidents during operation since hidden errors in control system hardware and software can be eliminated. For oil companies, contractors and vessel owners this is a great benefit as incidents is considered a major concern that may have consequences for safety, environment and cost due to delays in offshore marine operations. Reduced cost of vessel construction, upgrade and maintenance because at one side HIL testing reduces the need for expensive full scale sea trials, and on the other side it allows to set up simulated scenarios that are too expensive or risky to actually test in full scale. The need for potentially destructive testing is greatly reduced with HIL simulator testing. For the ship yard and ship owner this means that the control systems are tested earlier, deeper and with greater depth. Software upgrades and related equipment modifications or changes on vessels in operation are a serious concern of the industry. HIL testing offers a tool for efficient periodic testing or testing before or during upgrades or retrofits, since a large part of the testing can be done while in dock, at quay, or in transit.
REFERENCES [1] J. Spouge, Review of methods for demonstrating redundancy in dynamic positioning systems for the offshore industry, DNV Consulting, research report 195, HSE, UK, 2004 [2] T. A. Johansen, T. I. Fossen and B. Vik, Hardware-in-the-loop testing of DP systems, Dynamic Positioning Conference, MTS, Houston, TX, 2005 [3] R. Skjetne, O. Egeland, Hardware-in-the-loop simulation for testing of DP vessels, OSV, Singapore, 2005 [4] IMCA, Guidance on Failure Modes and Effects Analysis, IMCA Report M 166, 2002 [5] IMO, Guidelines for Vessels with Dynamic Positioning Systems, IMO Maritime Safety Committee Circ. 645, 1994 [6] DNV, Standard for Certification of HIL testing, Draft, 2005 [7] W. Ren, M. Steuer and S. Woodruff, Applying Controller and Power Hardware-in-the-loop simulation in prototyping apparatuses for future all electric ship, IEEE Electric Ship Technologies Symposium, 2007 [8] S. Palla, A. K. Srivastave and N. N. Schulz, Hardware in the loop test for relay model validation, IEEE Electric Ship Technologies Symposium, 2007 [9] Y. Xie, G. Seenamuni, J. Sun, Y. Liu, Z. Li, A PC-cluster based real-time simulator for allelectric ship integrated power systems analysis and optimization, IEEE Electric Ship Technologies Symposium, 2007
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OSV Singapore 2007 Jointly organized by Joint Branch of RINA-IMarEST Singapore and CORE. 24-25 September 2007
A Survey of Concepts for Electric Propulsion in Conventional and Ice Breaking OSVs Alf Kåre Ådnanes, Dr. Ing., MScEE ABB AS, Business Unit Marine, P.O. Box 94, NO-1375 Billingstad, Norway ABSTRACT The use of electric propulsion in both open water and ice breaking OSVs has become a marine industry standard for a wide range of applications, and is increasing to new applications and into more geographical regions. The technologies develop continuously, and today there are several approaches to reach the "optimal design" that reduces fuel consumption and environmental footprint, simplifies design and construction with better utilization of the on-board space, and creates a better working environment for the crew. This paper presents the various concepts at the market, and summarizes their technical characteristics and limitations. It is aimed to give yards, designers, and ship owner necessary technical information in order to make a proper selection of system topology within the specifics of a vessel design. Important aspects regarding operational safety and availability of the propulsion and station keeping plant is also discussed. INTRODUCTION Since mid 1990’s, OSVs have been equipped with electric propulsion, Fig. 1, where the main propulsors and station keeping thrusters have been driven by variable speed electric motor drives, being supplied from the common ship electric power plant with constant frequency and voltage. Thrusters and propulsors are normally of fixed pitch propeller design (FPP) that reduces the mechanical complexity of the units, and the electric power is normally supplied from fixed speed combustion engines; diesel, gas, or dual fuel. During this period, there has been a continuous development of solutions for the electric power plant for vessels with electric propulsion. The development can be considered as an incremental evolution of concepts, where the building blocks of the electric plant is adapted from the general industry applications, which has a far bigger volume of installations than the marine applications and to a large extent gives the premises for basic technology developments. Even though the suppliers of electric power and propulsion plants utilize building blocks that are based on principally the same fundamental concepts, there is a range of different configurations and preferences in the market. As the technical arguments for the concept appears to be biased and naturally to some extent influenced by a driving force to pursue a sales, it is necessary for ship owners, yards, and designers to be able to evaluate and compare this information to make the decision. As the author represents one of the vendors, this paper should not be considered to be an objective and neutral assessment of competing concepts; however, it is the aim of the author to give a unified description of the most applied concepts. Further, it is the author’s objective that decisions shall be made on understanding of the characteristics and resulting effects of the technology, rather than the technology itself. The author has based his information on own and public available information, and hence, actual solutions may deviate from what has been assumed in this paper. It is therefore necessary to supplement the information herein with the detailed information in each project and from each supplier.
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Aux gen.
Emgc’y gen. 440V, 60Hz
99kVA 99kVA
230V
M
M
Bow Thruster
Bow Thruster Port Side propulsion and shaft gen.
230V Distribution
Stbd Side propulsion and shaft gen.
M
M
M
M
M
Thruster
Main Propulsor
Azimuth Thruster
Main Propulsor
Thruster
Fig. 1: Conventional direct mechanical propulsion, and electric propulsion concept for OSV. THE IMPORTANCE OF TECHNOLOGY DEVELOPMENT For many engineers, researchers, and developers, working with technology gives the daily bread and butter. Technology development is both challenging and interesting for those involved, as well as essential for the continuous improvement of the vessel’s earnings and safety. In evaluation of concepts, the overall performance and characteristics should be assessed rather than the individual component. The advantages of e.g. an efficiency improvement in one component may make no sense if it requires a non-optimal operation of the prime movers in order to function as intended. Utilizing fast changing technologies to obtain some improvement is normal in consumer industry, but in a life cycle assessment of a vessel, it may be painful if that particular technology is obsolete and unavailable after few years. The driving forces for technology development and assessment should be based on the effects and results of the technologies. Generally; the following criterion will be important for the comparison of products, systems, and services, although their weighting and importance may vary over time and between various applications: • • • • • • • • • • • • • • • •
Cost efficient building and installation Flexibility in design that improves ship utilization High safety for crew High safety for operations Continuous availability to propulsion and station keeping systems Reduced fuel consumption Reduced impact on the external environment, lower emissions Improved working environment for the crew Low maintenance costs Availability to maintenance during the life cycle of the ship Availability to maintenance in the region of operation, often world-wide Spare parts availability Remote and on-board support Minimizing constraints of operations leading to non-optimal performance Reduce negative consequences for other equipment High ice braking and ice management performance for ice brakers
Technology can never become the goal itself, but those who are leading in technology development and capable to utilize this competence in a sustainable and commercially viable context should be better fit to meet today’s and future’s solutions that meet the users changing demands. Understanding the needs and requirements of the users requires a continuous cooperation in the global maritime network, and continuous improvements. The “optimal” solution today does not necessarily meet the new requirements in the future.
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VARIABLE SPEED DRIVES FOR ELECTRIC PROPULSION The variable speed drive (VSD) for propulsors and thrusters is one of the most essential components in a power plant for electric propulsion. The VSD consists of: • Electric motor, normally asynchronous (induction) motors, but also synchronous motors for the high power range. Other types of motors used in special applications; such as permanent magnet motors and DC motors. • Frequency converter, converting the fixed voltage and frequency of the network to a variable voltage and frequency needed to adjust the speed of the electric motor. • Optionally line filters or transformers, depending on configuration for reducing the harmonic distortion of currents flowing into the network and voltage adjustment where applicable. • A control system, consisting typically of a motor controller and an application controller for the propulsion / thruster control, taking care of the control functions as well as monitoring and protection of the VSD. For the power level needed for OSV propulsion, the Voltage Source Inverter (VSI), Fig. 2, is the dominating topology of frequency converters and used by most suppliers to this market. DC drives, Current Source Inverters (CSI) and Cycloconverters are rarely used and being phased out from new buildings of OSVs. Therefore, this paper only considers the VSI in various configurations. The voltage source inverter consists of a rectifier, a DC link with voltage smoothing capacitors, and an inverter unit as the main components. The DC link may where required be equipped with a breaking chopper to dissipate wind-milling power from the propeller in rapid speed variations or in crash stop conditions of the vessel. The motor controller technology has developed from the scalar U/f control system used since the earliest variable speed AC drives. Field oriented control of AC motors was developed already in the 60’s /1/ but did not establish as an industrial standard before in the mid 80’s when digital controllers with sufficient capacity and speed became commercially accepted. The introduction of field oriented control significantly improved performance and efficiency of the VSD, however, the control method was still sensitive to the motor parameters and time variations. A new method of vector control that was based on stator flux oriented control and direct torque control of the motor was developed around 1990 /2/, and was first introduced in large scale commercial drives production in the early 90’s by ABB under the name of DTC™. DTC enables ultra fast control of the torque of the motors, with a robust algorithm that gives high controllability and efficiency with much less sensitivity to variations in motor parameters than the traditional field oriented control. As will be discussed later; the dynamic performance of propulsors and thrusters for OSV is much lower than what the modern VSD may perform, with exception of the need for fast black-out prevention where each fraction of a second is essential. External interfaces - remote control - automation systems - propulsor and auxiliaries Propulsion Controller Motor Controller Measurement and control signals
Rectifier
Supply voltage: Fixed AC voltage Fixed frequency
DC Link
Inverter
Constant DC voltage
Fig. 2: The basic modules for a Voltage Source Inverter (VSI).
MOTOR
Variable AC voltage and frequency
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The key component in the frequency converter is the power semiconductor. Because of the high power, the semiconductors will only operate in the active mode in the transition between being fully off (closed) state and fully on (open) state in order to minimize the power losses. In VSI frequency converters, the components used are either passive components (diodes) or active switching components. Thyristors are only applied in special cases e.g. in configurations where it is a need to control the DC link voltage or for soft charging of the DC link, and in some cases for regeneration of power. Power semiconductors are made for low voltage applications, i.e. system voltages up to 690V between the phases, and for high voltage applications, >1000V system voltage. Depending on configuration, typical voltage levels are 3-3.3kV and 6-6.6kV line-line voltages for high voltage frequency converters. The IGBT (Insulated Gate Bipolar Transistor) is the dominating power semiconductor for low voltage applications and being used by all major suppliers of frequency converters. The low voltage IGBT is normally mounted in compound modules, as shown in Fig. 3(b). Each module may consist of several IGBT switching components, in parallel or separately controlled, together with free-wheeling diodes for reverse currents through the switching elements. For high voltage frequency converters, the GTO (Gate Turn Off) thyristor was for long times used as switching component. In high voltage converters these are used as discrete switching elements with one silicon wafer being installed in press pack (hockey-puck) housing. The GTO is of a robust design, and a highly reliable component itself, although it required a number of auxiliary components to achieve the robustness in operation, and these auxiliary components got quite high stresses from the switching of the GTO. Hence, a further development of the GTO was made by ABB during the 90’s where the basic concept of the GTO was improved to reduce the need for auxiliary components, and hence not only increase the overall reliability but also reduce the power losses in the overall system. This new component is called GCT (Gate Controlled Thyristor), and when integrated with its gate control system, IGCT (Integrated GCT). All ABB IGCTs are press-pack devices. They are pressed with a relatively high force onto heat-sinks which also serve as electrical contacts to the power terminals. The IGCT' s turn-on/off control unit is an integral element of the component. It only requires an external power supply and its control functions are conveniently accessed through optical fiber connections. The IGCT is optimized for low conduction losses. Its typical turn-on/off switching frequency is in the range of 500 hertz. However, in contrast to the GTO, the upper switching frequency is only limited by operating thermal losses and the system' s ability to remove this heat. As the IGBT requires a simpler gate control device, with less power consumption, a large effort has been made to make them available for high voltage. The first IGBTs were of module design, and were not considered reliable enough for demanding propulsion applications. Since IGBTs are made with multiple parallel chips, there is a difficulty - with conventional press-packs - in assuring uniform pressure on all chips; a difficulty which increases with the number of devices in a stack. Today, IGBT in press pack are available from several makers, including ABB /4/, and mainly used in power transmission and static converters, although some makers have eveluated them stable enough to be used in variable speed drives for ship applications.
IGBT Low voltage
(a)
(b)
IGBT High voltage
(c)
GCT/IGCT High voltage
(d)
Fig. 3: (a) The development of power semiconductor switches /3/, (b) low voltage IGBT module, with integrated IGBT switches and freewheeling diodes, (c) press pack high voltage IGBT, and (d) press pack GCT/IGCT.
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Fig. 5 shows the most common configurations for low voltage VSI frequency converters. The rectifier may be of different types, depending on the requirements for each installations and makers’ preference; • 6-pulse diode rectifier is the simplest design, with a full bridge passive rectifier – or several in parallel if necessary to achieve the desired power level. The AC supply voltage is rectified to form a DC voltage, of approximately 1.35 times supply line voltage at full load; i.e. a 690V supply gives approximately 930V DC link voltage at full load; depending on voltage drop and commutation impedance in the supply. The 6-pulse rectifier does not need any supply transformer unless necessary to adapt the voltage. Hence, size and weight is minimized. However, the harmonic distortion from the line currents is high – in the order of 2525% THDi, resulting in a voltage distortion THDu of more than 10%. In order to achieve the limit as specified in IEC and which most classification society now has adapted of 5%, harmonic filtering or clean power supply is necessary. A harmonic filter at the distribution switchboard will reduce the distortion at this voltage level and below, but will hardly have any impact on the distortion at the main switchboard. The main switchboard must hence be designed to tolerate a high level of voltage and current distortion, and be documented accordingly as specified in class rules. • 12- and quasi 24-pulse configurations looks similar with two paralleled diode rectifiers; but the quasi-24 pulse (Q24) VSD transformers are made in pairs of two and two with 15 deg phase shift between the two in each pair. Depending on load conditions in the prospected operation profile and system parameters, 12pulse configuration may meet the requirements to voltage distortion by class. For most cases, however, the worst case conditions will lead to a THDu in the range of 6-8%, which is above the limit of most of classification societies without using harmonic filters. The Q24-pulse rectifier utilizes the same rectifier topology as the 12-pulse rectifier, but since the supply voltages are phase shifted through the supply transformers, the resulting distortion at the main switchboard is reduced. At ideal conditions, where the loads of the VSDs in each pair are equal, the harmonic distortion will be equivalent to a 24-pulse configuration. Under other conditions, a partial cancellation of the largest harmonic components will occur, and the harmonic distortion will in most installations be under the 5% THDu limit in any practical operation mode. Certain constraints in operation may be necessary in order to guarantee this; however, the Q24-pulse configuration is regarded to be a cost efficient way to meet class requirements for most OSVs. • 24-pulse configurations consists of four paralleled 6-pulse rectifiers, each supplied from phase shifted voltages through one 5-winding transformer, or two paralleled 3-winding transformers. This configuration will normally always give distortion under the 5% limit, without constraints in operation. Normally, the transformer will be larger and more expensive than a 3-winding transformer of equivalent rating, and depending on the rating of the diode rectifiers, also the size and price of this may increase. The 18-pulse configuration utilizes the same concept, but with only three paralleled 6-pulse rectifiers and a 4-winding transformer. The THDu will also here normally be within the 5% limit, however, comparing to the 24-pulse topology, the total prize and size is at the same order due to a complex transformer design. 18-pulse rectifiers are therefore rarely in use today. • Active rectifiers with switching elements have for some time been applied in demanding industrial applications, especially where the load characteristics requires regenerative braking to such an extent that it makes it beneficial to use this energy by feeding it back to the network. For propulsors and thrusters, the regenerative energy is normally negligible in a fuel and cost of energy assessment. However, since the rectifier consists of switching devices, the current can be shaped similarly to the motor current and with much lower distortion than the currents of a diode rectifier. Even though the classical harmonic filters then can be avoided without use of drive transformers, one should note that there are ample of harmonic voltages from the switching, with high frequencies and with a high level of electromagnetic noise that must be filtered with high frequency (HF) filters. There is limited experience by use of active rectifiers in weak electric power systems as found on vessels, and in complex systems with many drives and a range of operating conditions, it is challenging to perform a complete system analysis of any modes and configurations in order to detect and avoid possible resonance effects from the numerous combination of paralleled HF filters. As the switching elements are more costly than diode rectifiers, the cost of the frequency converter increases, and its losses will also be higher; reducing the benefit of avoiding losses in the drive transformer.
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The DC link, Fig. 5, consists of a DC capacitor, in order to smooth the DC link voltage to reduce the voltage ripple from the rectifier to an acceptable level for the output stage of the frequency converter, and to filter the high frequency distortion from the switching elements in the inverter and avoid that these are injected to the supply network. If regenerative power is expected from the load, e.g. in crash stop conditions of shaft propellers, the power will be absorbed by the DC link causing voltage rise of the capacitors unless the power can be fed into the supply network by an active or regenerative rectifier, or dissipated in a load resistor bank. For thrusters and Azimuthing propulsors, regenerative voltage is seldom of a concern, as one may restrict these loads to only operate in motoring mode. In shaft line propulsion, or in Azimuthing propulsors where crash stop or equivalent to crash stop maneuvering is made by reversal of the propeller RPM, there are certain conditions where the propeller wind-milling effect may create reverse power; see Fig 4. The inverter module is normally a full bridge IGBT inverter in low voltage VSIs, Fig 5, with 6 IGBT switching elements each with an anti-parallel freewheeling diode for reversing the currents through the switches. The switching elements in each of the three legs of the inverter must operate in inverse mode, where one IGBT always is controlled off to avoid short circuiting the DC link. The objective of controlling the IGBTs is to feed a voltage vector to the stator winding that forces the currents and flux in the machine towards the targeted amplitude and phase angle. Depending on the control scheme, there are different ways to achieve this, such as field oriented control with PWM or hysteresis control, and direct torque control. The characteristics of these methods will be discussed later. High voltage frequency converters, e.g. for 3-3.3kV system voltage, utilizes the same principles as the low voltage VSI frequency converters. Because of the higher system voltage, and maximum voltage availability of power semiconductors, the high voltage VSI frequency converter normally consists of series connected semiconductors, as shown in Fig. 6 for the rectifier and inverter. In order to reduce the voltage over each component, the converters will normally be equipped with a third voltage level, at the middle of the negative and positive DC bus voltage, the so called neutral point. Clamping diodes ensures that each of the switching elements never will be exposed to higher than half of the DC link voltage, and therefore this configuration is called “three level, neutral point clamped (NPC) inverter topology”. By increasing the number of series connected switching elements, the number of switching combinations also increases – and with a specific maximum switching frequency of the components, the harmonic distortion in the load currents and consequentially the torque ripple will be reduced. Some suppliers use this concept to design multilevel inverter topologies, where the ripple current can be further reduced and lower voltage switching elements may be used. The penalty is then the higher number of components in the system, reducing its availability and reliability. Quadrants I and III – Positive Power:
Torque Quadrant II Braking Speed0
Quadrant I Motoring Speed>0, Torque >0
Bollard pull Free sailing Crash stop
Quadrant III Motoring Speed
