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HSE Health & Safety Executive Review of issues associated with the stability of semi-submersibles Prepared by BMT Fluid Mechanics Limited for the Health and Safety Executive 2006 RESEARCH REPORT 473

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Page 1: Semi Submersibles Stability Issues

HSE Health & Safety

Executive

Review of issues associated with the stability of semi-submersibles

Prepared by BMT Fluid Mechanics Limited for the Health and Safety Executive 2006

RESEARCH REPORT 473

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HSE Health & Safety

Executive

Review of issues associated with the stability of semi-submersibles

BMT Fluid Mechanics Limited Orlando House

1 Waldegrave Road Teddington Middlesex TW11 8LZ

This review study was undertaken as part of a wider exercise to assess the need for HSE Guidance within the UK Safety Case regime. The study included a comparison between stability standards specified by key regulatory authorities and classification societies for intact and damaged semi-submersible units, a review of relevant published literature and of HSE/ Department of Energy reports, a review of past incidents involving loss of stability of semi-submersibles, and a review of issues associated with alternative uses of semi-submersible units.

A key recommendation coming out of this study is that the HSE should investigate further the practicality of reconciling traditional prescriptive stability standards with a risk-based Safety Case approach.

This report and the work it describes were funded by the Health and Safety Executive (HSE). Its contents, including any opinions and/or conclusions expressed, are those of the authors alone and do not necessarily reflect HSE policy.

HSE BOOKS

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© Crown copyright 2006

First published 2006

All rights reserved. No part of this publication may bereproduced, stored in a retrieval system, or transmitted inany form or by any means (electronic, mechanical,photocopying, recording or otherwise) without the priorwritten permission of the copyright owner.

Applications for reproduction should be made in writing to: Licensing Division, Her Majesty's Stationery Office, St Clements House, 2-16 Colegate, Norwich NR3 1BQ or by e-mail to [email protected]

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EXECUTIVE SUMMARY

This review study was undertaken as part of a wider exercise to assess the need for HSE Guidance within the UK Safety Case regime. The study included a comparison between stability standards specified by key regulatory authorities and classification societies for intact and damaged semi-submersible units, a review of relevant published literature and of HSE/ Department of Energy reports, a review of past incidents involving loss of stability of semi-submersibles, and a review of issues associated with alternative uses of semi-submersible units.

A key recommendation coming out of this study is that the HSE should investigate further the practicality of reconciling traditional prescriptive stability standards with a risk-based Safety Case approach.

Intact Stability

No obvious major deficiencies were found in established intact stability standards. These appear to have been successful in avoiding capsize and loss while units remained intact and watertight, and in accordance with the standards.

Key conclusions from the review study, relating to intact stability standards, were:

• The intact stability standards of most regulatory authorities are outwardly similar, although there are important differences in specific requirements (eg. in values of limiting angles and the minimum value of GZ). It is not obvious how these variations affect stability margins, or levels of uniformity and consistency between different units.

• Ambiguities in definitions of the most critical heel axis and between ‘free trim’ and ‘free twist’ analysis procedures have been criticised. It is important that key procedures and definitions should be understood clearly by all engaged in such work.

• Large heel angles are required by certain designs in order to meet the 1.3 area ratio requirement. Large heel angles would seem to be undesirable for several reasons, including hazards caused by large items shifting, hazards to personnel, and difficulties during evacuation and escape.

• There is a sound basis for retaining existing HSE minimum GM and GZ requirements, because these are likely to reduce the likelihood of abrupt changes in heel during operation, large angles of heel, steady tilt and low-frequency motions.

• Certain authorities (including HSE) state that moorings should be ignored in a stability analysis. It would seem prudent, however, to require detrimental effects of moorings, risers, thrusters and similar items to be considered explicitly.

• Existing standard procedures for calculating the wind heeling moment are simplified, and most such procedures do not take adequate account of the lift-induced moment. This component can be important for heeled semi-submersibles, and may be either over or under-estimated.

• In circumstances where location-specific environmental standards are appropriate, careful consideration should be given to the averaging and return periods used when estimating the wind speed.

• The alternative ABS/ IMO intact criteria should be treated with caution until experience has been gained. It is not known how extensively these criteria are used for semi-submersibles working on the UKCS.

Damaged Stability

There are major differences between the damaged stability standards adopted by different regulatory authorities (see Appendix A). There are in particular major differences between the approaches adopted by the HSE (based on a minimum area ratio) and in the IMO’s 1989 MODU Code (based on

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minimum values of the righting arm and stability range). Discussions leading towards convergence between these different standards should be encouraged.

The damaged stability standards in the IMO MODU Code are based on criteria originally proposed by ABS. The small range of units considered during the ABS JIP, and simplifications adopted during the model tests, are matters of concern. These criteria should be treated with caution until further experience has been gained.

Key conclusions from the review study, relating to damaged stability standards, were:

• There are differences between the areas of damage and flooding specified by different authorities, the consequences of which are not known.

• The level of damage that occurred in most incidents was consistent with levels envisaged by conventional damaged stability standards, and generally involved flooding of a single compartment.

• The traditional 1.5m penetration depth requirement seems somewhat arbitrary, and does not take account of increases in the size and power of supply vessels over the years, the size of the unit itself or its structural loading capacity.

• A number of authorities (including HSE) specify a minimum 4m wave clearance above the damaged waterplane, regardless of sea conditions, the vessel’s behaviour or the location and size of the downflooding point. The rationale for this 4m clearance seems questionable.

• The use of a 50 knot wind speed in damaged stability standards has been questioned on several occasions. At issue is whether the standards should reasonably address the combination of circumstances that led to the Ocean Ranger accident.

Other Issues

Moves towards quantitative risk-based stability analysis procedures, integrated within an overall safety assessment, should be encouraged as a long-term objective, although there are likely to be major difficulties in achieving this objective in the short term. Further review of the RABL work and other developments in this area would be a useful first step.

The review of past incidents suggested that there is a relatively low risk of major damage following collision, although minor damage has been recorded. Blowout is undoubtedly a major hazard, because of fire and explosion risks, but does not seem to be a major issue for capsize. Ballast system failures, errors in operating the ballasting system, and flooding due to variety of causes, including human error, seem to be the most important issues. Storm damage and towing errors are also significant factors. The number of ballasting faults and errors is disturbing, and also the fact that many flooding incidents outside the UKCS are unexplained.

No overlap was found between incidents reported in the WOAD and HSE databases, raising possible concerns about the accuracy and completeness of these records.

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Contents

Page

EXECUTIVE SUMMARY ...........................................................................................................................iii

1. INTRODUCTION ...................................................................................................................................... 1

2. GENERAL APPROACH ........................................................................................................................... 32.1 REVIEW OF EXISTING STABILITY STANDARDS.................................................................... 32.2 LITERATURE REVIEW .................................................................................................................... 32.3 REVIEW OF ALTERNATIVE USES OF SEMI-SUBMERSIBLES.............................................. 42.4 REVIEW OF PAST INCIDENTS ...................................................................................................... 4

3. BACKGROUND TO STABILITY STANDARDS ................................................................................. 53.1 HISTORICAL DEVELOPMENT....................................................................................................... 53.2 ALTERNATIVE APPROACHES ...................................................................................................... 7

4. REVIEW OF STABILITY STANDARDS AND GUIDANCE............................................................114.1 INTACT STABILITY STANDARDS .............................................................................................124.2 DAMAGED STABILITY STANDARDS .......................................................................................14

5. ISSUES ARISING FROM THE LITERATURE REVIEW ..................................................................195.1 DISCUSSION ON INTACT STABILITY STANDARDS.............................................................195.2 DISCUSSION ON DAMAGED STABILITY STANDARDS.......................................................255.3 WIND FORCES AND HEELING MOMENTS ..............................................................................295.4 EFFECTS OF MOORING LINES, RISERS AND THRUSTERS ................................................315.5 EFFECT OF ICING ...........................................................................................................................325.6 STEADY TILT ANGLE AND SUBHARMONIC MOTIONS......................................................325.7 WATERTIGHT AND WEATHERTIGHT INTEGRITY...............................................................335.8 LOSS OF BUOYANCY FOLLOWING A BLOW-OUT ...............................................................345.9 VARIATIONS IN METACENTRIC HEIGHT DURING OPERATION .....................................345.10 STABILITY MANAGEMENT ......................................................................................................36

6. ALTERNATIVE USES OF SEMI-SUBMERSIBLES ..........................................................................396.1 DRILLING UNITS ............................................................................................................................396.2 FLOATING PRODUCTION UNITS ...............................................................................................396.3 ACCOMMODATION UNITS ..........................................................................................................416.4 DIVING SUPPORT VESSELS ........................................................................................................416.5 PIPELAY VESSELS .........................................................................................................................416.6 CRANE VESSELS ............................................................................................................................416.7 OTHER APPLICATIONS.................................................................................................................41

7. REVIEW OF PAST INCIDENTS ...........................................................................................................43

8. DISCUSSION OF KEY ISSUES ............................................................................................................478.1 INTACT STABILITY .......................................................................................................................478.2 DAMAGED STABILITY .................................................................................................................488.3 STABILITY MANAGEMENT ........................................................................................................498.4 OTHER ISSUES ................................................................................................................................508.5 REVIEW OF PAST INCIDENTS ....................................................................................................50

9. CONCLUSIONS.......................................................................................................................................519.1 INTACT STABILITY .......................................................................................................................519.2 DAMAGED STABILITY .................................................................................................................529.3 OTHER ISSUES ................................................................................................................................52

10. REFERENCES........................................................................................................................................55

ABBREVIATIONS AND NOTATION ......................................................................................................61

APPENDIX A: SUMMARY OF STABILITY STANDARDS...............................................................63

APPENDIX B: SUMMARY OF INCIDENT INFORMATION ............................................................73

APPENDIX C: THE ALEXANDER L. KIELLAND AND OCEAN RANGER ACCIDENTS ................81

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Review of Issues Associated with the Stability of Semi-Submersibles

1. INTRODUCTION

BMT Fluid Mechanics Limited (BMT) was commissioned by the UK Health and Safety Executive (HSE) to undertake a review of issues associated with the stability of semi-submersibles which are mainly used as drilling units. This review was undertaken as part of a more general appraisal of the need to retain elements of the Department of Energy/ HSE’s Fourth Edition ‘Guidance on Design,

1Construction and Certification’ [1] within the framework of the UK Safety Case regime.

Most regulatory authorities and classification societies have stability standards for vessels and semi-submersibles operating under their jurisdiction. Whether any particular standard is mandatory or advisory depends on the responsibilities of the organisation in question, on the area of operation and on the unit’s flag status. International Maritime Organisation (IMO) MODU Code standards are not mandatory. Responsibilities for specifying and enforcing stability standards are placed on flag and coastal states. They are also delegated to the classification societies, such as the American Bureau of Shipping (ABS), Det Norske Veritas (DNV) and Lloyd’s Register of Shipping, and there have been moves towards harmonising classification society rules towards the IMO MODU standard.

The Offshore Safety Division (OSD) of the HSE is responsible to the UK Government for ensuring that risks to people from work activities in the ‘upstream’ oil and gas industries are properly controlled, when taking place in UK coastal waters. A key feature of the UK post-Cullen regulatory regime is the requirement for the owner or operator of every offshore installation to prepare a Safety Case, and submit it to OSD for formal acceptance. The Safety Case has to demonstrate that an effective management system is in place to control major accident risks and to ensure that they are ‘as low as is reasonably practical’ (ALARP). These risks specifically include those arising from loss of stability. The responsibilities of the owner or operator are interpreted in a very wide-ranging manner, and are not limited to meeting any particular prescribed standards or criteria.

The HSE’s Fourth Edition Guidance [1] specifies certain minimum standards, but these standards are not mandatory under the Safety Case Regime and represent guidance only. The HSE is now reviewing whether particular sections of the Guidance should be retained, and in particular whether there is any need to retain Section 31 of the Guidance relating to stability, watertight integrity and ballasting.

During the present review study BMT was specifically requested to:

• review issues which may affect the intact and damaged stability of semi-submersibles;

• identify issues which are likely to be of particular concern to the HSE and the offshore industry.

This report describes the review findings, and is structured as follows:

• Sections 2 and 3 outline the general approach adopted during this review, and the historical background to existing and proposed alternative stability standards;

• Section 4 reviews the standards specified by key regulatory and classification authorities for both intact and damaged semi-submersible units;

• Section 5 covers additional issues arising from a review of relevant literature;

• Section 6 reviews issues associated with alternative uses of semi-submersible units;

• Section 7 reviews past incidents involving loss of stability of semi-submersibles;

• Sections 8 and 9 summarise key issues and conclusions coming out of the review study.

A list of references may be found in Section 10 on page 55.

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2. GENERAL APPROACH

The main tasks undertaken during this review study were as follows:

• to review the stability standards of significant regulatory and classification authorities relating to semi-submersible units in both the intact and damaged conditions;

• to carry out a literature review, which involved a search of publicly accessible abstracts databases, and a review of reports and documents made available by the HSE;

• to review recorded incidents and casualties, following a search of the WOAD and HSE incidents databases;

• to review and identify key issues, and then draw conclusions.

Views from a major semi-submersible designer and an operator were also sought to help identify problems experienced in complying with stability requirements during design or in service. These views were taken into account when identifying key issues.

Some of the issues identified during this investigation are similar to those highlighted during an earlier review study on jack-up stability [2]. In both cases conventional intact stability standards seem to have been successful in avoiding capsize, but a number of losses have occurred after damage and flooding. There are important differences between jack-ups and semi-submersibles, however, such as:

• There have been many reported losses of jack-ups, but relatively few semi-submersible losses.

• The causes of two major semi-submersible losses have been thoroughly investigated and are well documented. Significant factors included catastrophic structural failure in the case of the Alexander L. Kielland [3], and loss of control over the ballasting system in the case of the Ocean Ranger [4, 5, 6]. After these initiating events, subsequent flooding and capsize took place relatively quickly (see Appendix C).

• Most jack-up losses occurred after minor damage and a long period of progressive flooding. Green water is a relatively common occurrence on the decks of jack-ups in severe storms, because of their relatively low freeboard. Water impact damage to items on deck proved to be a significant factor in many jack-up losses. Green water can only reach the deck of a semi-submersible, however, after the unit has already developed a large heel or trim angle.

2.1 REVIEW OF EXISTING STABILITY STANDARDS

Section 3 of the report reviews the historical development of stability standards for ships as well as semi-submersibles, including attempts to make the standards take better account of vessel and wave dynamics.

A critical review of current intact and damaged stability standards applying to semi-submersible (column-stabilised) units was then undertaken. Comparisons were made between standards specified in the HSE’s Fourth Edition Guidance, Norwegian regulations, the International Maritime Organisation MODU code, and classification society rules.

Results from this review of current stability standards for semi-submersibles in both the intact and damaged conditions are presented in Section 4, and issues of special concern are identified. Appendix A summarises the stability standards of six organisations.

2.2 LITERATURE REVIEW

A literature search was performed to identify relevant research papers and reports. These documents were obtained and reviewed, and key issues relating to the stability of semi-submersibles were identified.

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Initial searches revealed a large number of projects, reports and papers relating to semi-submersible stability and sea-keeping. It was not practical to carry out an exhaustive and detailed review of all these documents. Attention was therefore focussed on results from relatively recent research, or documents which were considered to be especially relevant.

The literature review found many research papers from the period immediately following the losses of the Alexander L. Kielland and Ocean Ranger, but relatively few papers after about 1990, suggesting that research interest in semi-submersible stability had waned. The literature review therefore includes many papers published before 1990, which are still relevant to the issues in question.

The Norwegian ‘Risk Assessment of Buoyancy Loss’ (RABL) project [7] arose directly out of research undertaken after the loss of the Alexander L. Kielland, and was an important step towards developing more rational, risk-based stability standards. The RABL reports were deliberately excluded from this review study, however, because they raise issues which go well beyond those of a conventional stability analysis, and it was agreed that they should therefore be reviewed separately.

The HSE undertook an initial keyword search of their Herald database, and identified a shortlist of 34 Department of Energy (DEn) or HSE projects which were considered to be especially pertinent to the issue of semi-submersible stability. After carrying out a preliminary review of project summary details, BMT considered seven projects to be of special interest. Reports from these seven selected projects were included in the literature review, and the findings were taken into account when identifying key issues. Many of these reports have not been published, however, and so are not included in the reference list.

A number of further DEn/ HSE projects had addressed relevant topics, but were not considered in detail during the present study. Other projects appeared to have little direct relevance to present objectives, and were considered no further. These included a number of reports concerned with the seakeeping behaviour of semi-submersibles rather than stability issues.

Key issues emerging from the literature review are discussed in Section 5.

2.3 REVIEW OF ALTERNATIVE USES OF SEMI-SUBMERSIBLES

Semi-submersibles have always been used for a wide variety of different purposes, such as drilling, diving support, fire-fighting, crane operations, pipe-laying and accommodation. Semi-submersibles are now being used or proposed for a number of new offshore roles, including combined drilling, workover and floating production units [8, 9]. In some circumstances the need to remain on station for long periods in severe weather may impose special demands on the unit.

Issues associated with alternative applications of semi-submersibles are discussed in Section 6.

2.4 REVIEW OF PAST INCIDENTS

A search of the HSE’s incidents database and the Worldwide Offshore Accident Databank (WOAD) revealed 36 separate incidents involving loss of stability of semi-submersibles. These included not only the well-known and well-documented losses of the Alexander L. Kielland [3] and Ocean Ranger [4, 6], but also many other less well-known incidents.

Results from this review and key issues relating to the stability of semi-submersibles are discussed in Section 7 of this report. Summary details of these incidents are listed in Appendix B, and additional details of the Alexander L. Kielland and Ocean Ranger accidents are given in Appendix C.

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3. BACKGROUND TO STABILITY STANDARDS

Established stability standards for semi-submersibles are based on older criteria developed for naval and merchant ships, and it is instructive to look back at the history of some of these criteria. Conventional quasi-static stability standards have developed in a historical manner, and are intended to provide a practical and reasonably uniform basis for design, so that new designs are at least as safe as existing well-designed vessels. They are less concerned to represent realistic sea conditions and vessel behaviour, and more concerned with maintaining existing standards.

The apparent arbitrariness and limitations of present quasi-static stability standards have long been recognised, and the fact that no explicit dynamic formulation has yet been agreed is a measure of the difficulty of the task. The international search for a universally applicable and acceptable dynamic stability criterion has been long and controversial; indeed Kuo et al. [10] reproduced a comment made at the First International Conference on Ship Stability in 1975: that the task will take ‘another hundred years’.

The present quasi-static standards have the attraction of being explicit, and are easy to apply in practice. History suggests that they also provide an adequate margin of stability, in the sense that no past semi-submersible losses have been attributed to deficiencies in the intact stability standards. The main motivation for any change in the standards would therefore have to be either a reduction in conservatism, resulting in lower construction costs, greater uniformity of design, or else concern that future changes in design might erode the conservatism.

A major problem for researchers on stability is the way in which stability issues impinge on other aspects of design. It has been found, for example, that the design of the mooring system can affect the vessel’s stability detrimentally, and increasing the metacentric height above a certain level may cause undesirable motions in certain sea states. As noted by Vassalos et al. [11], this can easily turn stability work into ‘a long-term research programme making no contribution whatever to immediate needs.’

A major difficulty with introducing explicit dynamic terms into stability standards has been lack of understanding of the physical processes, and therefore of the mathematical equations, and lack of agreement about the critical factors leading to capsize. Much previous research seems to have been driven by the beliefs of individual researchers about the mechanisms that are important, and (equally important) about those that may be neglected. Undoubtedly the capsize process is highly complex and non-linear, and different processes may be at work in different circumstances. All this has led to controversy and uncertainty about how to incorporate these features into design. A measure of the controversy surrounding this whole topic may be gained from the discussion following a paper by Vassalos [12]. Henrickson [13] has also summarised some of the criticisms levelled against certain types of stability standards.

Practical considerations and past experience suggest that improved standards are only likely to gain acceptance on a step-by-step basis, through a process of enhancing existing standards, rather than through some radically new approach.

3.1 HISTORICAL DEVELOPMENT

Henrickson [13], Bird and Morrall [14] have provided useful summaries of the history of methods used to assess stability, and of stability research up to the mid-1980s. An excellent review of stability theory, methods used to assess stability, comparisons between the standards of different regulatory bodies, and special considerations for different types of vessel, may be found in Chapter 4 of the recent CMPT Floating Structures guide [15]. Mills et al. [16] also provided an excellent review of the regulatory position as it stood in 1991, and Martinovich and Praught [17] reviewed the development of stability standards from the viewpoint of a semi-submersible designer.

Established methods for assessing stability may be categorised [14] as follows:

a) static stability methods, such as the Rahola [18] criterion, based simply on minimum righting moment requirements,

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b) moment balance requirements, which, together with:

c) energy balance requirements, are included in all known current semi-submersible codes, as well as in the Sarchin and Goldberg [19] criteria for naval ships, and the International Maritime Organisation’s (IMO) ‘weather criterion’ [20] for merchant ships,

A number of alternative stability standards have emerged in recent years, based on considerations of vessel dynamics and risk, mainly in the aftermath of the Ocean Ranger and Alexander L. Kielland accidents. These included the Norwegian MOPS [21] and RABL [7] projects, and the PRESS project [22] in the UK. The other major recent event has been the development of alternative intact and damaged stability criteria during Joint Industry Projects managed by the American Bureau of Shipping. Further information about the ABS projects and the resulting criteria may be found in Sections 5.1.3 and 5.2.3.

3.1.1 Static Righting Moment Methods

Simple righting moment requirements have a fairly obvious physical interpretation, and require no difficult decisions about wind and wave loading parameters. They give no indication of safety margins, however, and are likely to be inappropriate if vessel forms and dimensions are changed from those originally envisaged, or for vessels with large wind areas. Henrickson [13] presented a detailed critique of such methods.

3.1.2 Moment and Energy Balance Requirements

Simple righting moment criteria have now been largely superseded by procedures based on moment and energy balance. Moment balance methods take account of the steady wind heeling moment and righting moment, together with other static forces such as weights lifted over the side, or tow-lines. These methods take account of the energy imparted to the vessel by a sudden wind gust, causing it to heel down-wind from an initial angle. Energy balance methods involve comparing areas under the wind heeling moment and righting moment curves. The energy balance method is essentially static in its approach, normally considering only the total work done by a steady wind moment as the vessel heels, and ignoring any pre-existing motions, wave-induced roll, damping and changes in the righting moment associated with the wave. These dynamic effects are taken into account through overall safety factors, such as the 1.3 factor applied to the area ratio for semi-submersibles.

Most existing intact stability standards, for all types of vessels including semi-submersibles, specify moment and energy balance requirements, together with a minimum range of angles over which the vessel is stable. Some also specify a minimum value of GM, thus retaining one feature of earlier righting moment criteria.

3.1.3 Stability Philosophies

Stability standards for both ships and semi-submersibles are based on a two-tier approach:

• intact stability requirements, designed to ensure that the unit will withstand all expected environmental conditions when in its normal operating or survival condition, and while it remains undamaged and watertight;

• damaged stability requirements, designed to ensure that the unit will not capsize in foreseeable environmental conditions, after undergoing a limited amount of damage or flooding, and will be capable of returning to the upright condition.

Two alternative approaches are normally adopted when defining damage: damage to any one compartment at any draught, or waterline damage, including breaching of internal watertight divisions between compartments. Mills et al. [16] noted that both approaches have their strengths and weaknesses. An offshore unit designed to meet the any one compartment standard cannot necessarily be guaranteed to meet the waterline damaged standard, and vice versa. The Norwegian Maritime Directorate (NMD) was the first regulatory authority to consider both damage scenarios. Both

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scenarios are now considered (in one form or another) by most regulatory authorities, as discussed in Section 4.2.

After the Alexander L. Kielland accident the NMD adopted a three-tier approach. The first two tiers were the established intact and damaged stability philosophies, and the third was a requirement that the unit should withstand loss of buoyancy from either the whole or a major part of one column, but without any requirement to return to the upright position. The objective in this case was to allow the crew time to evacuate the unit. This requirement was expressed in terms of providing a maximum angle of heel after a large loss of righting moment, and a minimum level of reserve buoyancy above the damaged waterline. The concept of providing some level of reserve buoyancy, beyond that necessary to meet basic code requirements, has since been widely accepted. The NMD’s proposals have not been adopted internationally, however, the IMO preferring an alternative form of reserve stability requirement. The reserve buoyancy concept is discussed further in Section 5.2.4.

The MOPS project [21] considered alternative philosophies which might be adopted when developing future stability standards. While recommending no immediate changes to existing Norwegian regulations, MOPS recommended a move away from rigid prescriptive regulations to simple statements of aim, supported by non-mandatory guidelines. Such moves were considered likely to increase the overall safety of units, while encouraging stability to be integrated within an overall safety assessment. This general philosophy has now been embraced by various regulatory authorities, including the HSE.

3.2 ALTERNATIVE APPROACHES

3.2.1 Weather Criteria

The Sarchin and Goldberg [19] criteria and the related IMO ‘weather criterion’ [20] attempt to place stability standards on a more physical and rational basis, by including an explicit dynamic term representing the effects of waves and wind gusting. This dynamic term is represented by a quasi-static adjustment to the unit’s initial angle of heel. Conventional stability criteria effectively assume that the initial angle of heel is zero.

Sarchin and Goldberg’s criteria were first adopted by the US Navy and subsequently by the US Coast Guard. The vessel is assumed to be initially heeled in the up-wind direction at an angle of 25o relative to the static wind heel angle. This 25 o angle represents the maximum likely roll response in waves. The criteria include an area ratio requirement with a safety factor of 1.4. It should be noted that this 1.4 factor was applied to the net areas between the righting and heeling moment curves, rather than the sums of areas used in most other standards.

Sarchin and Goldberg’s criteria also include a moment balance requirement, and separate criteria for lifting weights over the side, for crowding of passengers against the side, for high speed turns and icing.

The IMO ‘weather criterion’ [20] is intended to allow for rolling motions due to waves, and involves a complex formula based on the ratio of vessel breadth to length, its block coefficient, the ratio of bilge keel area to transverse area, the vessel’s roll period, a bilge keel factor, and the height of the centre of gravity. The basis of this formula is unclear, and its complexity makes its effects difficult to assess. It is understood that this criterion was only agreed after several years of discussions, and is a compromise between Russian and Japanese proposals. It is limited to merchant ships greater than a certain size and of conventional hull form, and is inappropriate, in its present form, for application to semi-submersibles.

Kuo et al [10] and Vassalos [12] put forward an alternative form of ‘weather criterion’. Their analysis allows the righting moment parameter GZ to vary with time, in order to represent changes in waterplane area as waves pass the vessel. These variations may cause a sudden loss of stability, and had been found to be important for small ships. These variations are likely to be less significant for semi-submersibles.

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Henrickson [13] noted various criticisms of the ‘weather criterion’ approach. The inherent assumption, that capsize follows a large quasi-static roll in the up-wind direction, is known to be false for ships; capsize seems to occur in a more random manner [23]. Numata [24] noted that model tests on semi-submersibles had shown no tendency to capsize in extreme conditions, even when the area ratio was less than 1.0. Numata’s criticisms were directed mainly at traditional calculation procedures used to estimate the wind heeling moment, however, rather than at the stability standards themselves.

Various alternative ‘weather criteria’ have been proposed. Bush and Ahilan [25], for example, proposed an analysis procedure for jack-ups which is similar to the Sarchin and Goldberg [19] method, but the fixed 25o angle was replaced by a maximum roll angle based on results from relevant model tests and numerical calculations.

This approach has the merit of being applicable to any type of unit, rather than specific hull forms for which sets of empirical formulae have been derived, and could in principle be applied to semi-submersibles. It requires an estimate to be made of the unit’s maximum roll angle, however, and an associated choice has to be made about the design sea state. Numata [24] reported that the ABS rules committee did in fact consider a ‘weather criterion’ for semi-submersibles, similar to that developed by Sarchin and Goldberg, but was reluctant at that time to specify any particular correlation between wind and waves, and felt that there was insufficient data on actual semi-submersible motions to establish roll excursions appropriate to different configurations. The proposal was therefore withdrawn, although the rules at that time did allow for alternative criteria based on a documented and rational approach.

Vassalos et al. [11] put forward a ‘weather criterion’ based on the energy balance equation, which took account of the unit’s dynamic response. They proposed that the net area under the energy curve should be positive over an extreme half-cycle of motion, and presented a sample calculation based on an Aker H-3 design. A difficulty with applying this criterion in practice is in defining an appropriate half-cycle of response.

The alternative stability criteria, proposed by the American Bureau of Shipping [26], also represent a form of weather criterion, based on the dynamic responses of semi-submersibles in extreme sea conditions. These criteria were developed after carrying out systematic numerical simulations and physical model tests on several generic series of semi-submersibles, and will be discussed in detail in Section 4.1.1.

3.2.2 Risk-Based Procedures

Established stability standards have been criticised on the grounds that it is difficult to relate them to levels of safety and reliability. Caldwell and Yang [27] outlined a methodology for a risk-based approach to stability analysis. A risk-based approach is attractive because it is consistent with the ‘safety case’ methodology. The level of risk that is regarded as acceptable varies between different industries and circumstances, and can only be judged in relation to other comparable hazards.

-5Caldwell and Yang suggested that a target annual risk of capsize should be no greater than 10 , and

-4the annual risk of death due to capsize should be no greater than 10 . These risks were considered to be comparable with the accident rate in the home, and better than the average annual risk from car travel.

Further recent research on the development of risk-based capsize standards has been reviewed by Kobylinski [28, 29].

A risk-based methodology for assessing the stability of semi-submersibles was developed during the RABL (Risk Assessment of Buoyancy Loss) project [7] in the aftermath of the Alexander L. Kielland accident.

The RABL project concluded that collisions are the most likely cause of damage, and estimated that -2 -3

the probability of severe damage to a semi-submersible following a collision is between 10 and 10per year. In terms of risk of loss, collisions were followed by burning and blow-out, although the probabilities associated with these events were very vessel-dependent. The next event in terms of importance was ballast control failure. The probability of this type of failure decreased, however, with

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the appearance of semi-submersibles designs in which ballast pumps are located at both ends of each pontoon. The fourth most probable causes were considered to be fatigue and fracture.

Despite these conclusions, it is disturbing that the only two semi-submersibles losses which resulted in major loss of life (the Alexander L. Kielland and Ocean Ranger) were associated with events ranked low in terms of risk significance. The incident review (see Section 7) also suggested a somewhat different order of importance.

Risk-based procedures will only be practical if they are based on a relatively small number of well-defined scenarios and environmental conditions. The MOPS and RABL projects made significant progress in identifying such scenarios, but the results were criticised (in an unpublished review for the HSE) for being too general, and the recommendations too subjective. A risk-based approach also needs be closely tied into an overall risk evaluation, which should consider the circumstances of past losses and the risks associated with combinations of events. All of this will add to the complexity and difficulty of carrying out such an analysis.

A quantitative risk-based approach is nonetheless a worthwhile long-term objective, in order to improve the basis for assessing priorities, for placing stability analysis within a overall safety assessment framework, and for reducing overall risks. Quantifying risk allows the most likely routes to failure to be identified, and overcomes preconceived or intuitive notions of risk. This type of approach can also help to identify the joint probability of two events, one of which affects the other. One example is damage in severe weather leading to mooring line failure, and causing an additional heeling moment through tripping. A broad risk-based approach should also help to identify other implications of severe weather, such as delays and difficulties in evacuation and rescue after loss of stability.

Quantitative risk assessment techniques have not yet become established as a practical, routine alternative for assessing the stability of a semi-submersible, and are unlikely to become practical in the short term. The present review study is concerned primarily with established techniques, used by the industry today. It was therefore agreed that a detailed review of the RABL work and other risk-based procedures should remain outside the scope of the present study.

3.2.3 Model Testing and Numerical Simulations

Systematic numerical simulations and model tests undoubtedly have a role in helping to develop and enhance stability standards. Systematic numerical simulations have been used, for example, in conjunction with model tests, to help develop survivability criteria for naval vessels [30]. The Norwegian MOPS project [21] compared predictions from a numerical simulation model with model test measurements, and found that it was possible to carry out reliable motion simulations of heavily listed, as well as intact, units. Similar conclusions were reached during the ABS JIP [26], where numerical simulations and model tests were used to develop alternative ABS criteria, which are discussed in Sections 5.1.3 and 5.2.3.

Model tests have the merit of including most of the relevant physical processes, such as vessel and wave dynamics, large-amplitude waves and roll motions, water on deck, and can also be made to represent effects of flooding.

Numerical simulations are analogous to model tests, and can include a large number of complex non-linear and coupling phenomena. There are a number of difficulties, however, in applying such techniques to capsize. The equations only represent physical processes which are understood well enough to be expressed in mathematical and numerical terms. The dynamics of capsize are often poorly understood, and the confidence attached to the resulting simulations will therefore be limited.

There are also practical difficulties in incorporating requirements for numerical simulations into standard stability assessment standards. It would be difficult, for example, to define an appropriate level of complexity for such simulations, bearing in mind that these simulations might have to be performed using standard design-office computers.

The main disadvantages of both model tests and numerical simulations are the lack of any explicit procedures for selecting relevant test conditions, and the need to carry out a very large number of test

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runs in order to investigate the stability boundaries and statistics of capsize. This approach is therefore likely to be costly.

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4. REVIEW OF STABILITY STANDARDS AND GUIDANCE

Most regulatory authorities and classification societies have stability standards for vessels and semi-submersibles operating under their jurisdiction. There are a number of important differences between these standards, however, which are highlighted below. These differences are likely to affect the design, operation and inherent stability of a given unit, but it is not easy to draw general conclusions about how these differences affect safety margins, or levels of consistency and uniformity between different units.

No attempt was made during this review study to carry out an exhaustive review of all applicable stability standards, and no attempt was made to look retrospectively at earlier stability standards which still apply to many older units. The objective was simply to identify key areas of similarity and difference between the HSE’s Fourth Edition Guidance [1] and the current stability standards of five significant organisations, which are known to have been especially active in developing and applying standards for mobile and offshore units:

• The Norwegian Maritime Directorate (NMD) [31],

• The Canadian Coast Guard (CCG) [32],

• The International Maritime Organisation (IMO) [33, 34],

• The American Bureau of Shipping (ABS) [35, 26],

• Det Norske Veritas (DNV) [36].

Lloyd’s Register of Shipping [37] specifies no particular stability standard. Section 1.3.1, Part 1, Chapter 2 of the Rules states that:

• The Rules do not cover certain technical characteristics, such as stability, trim, hull vibration, etc., but the Committee is willing to advise on such matters although it cannot assume responsibility for them.

The stability standards of the US Coast Guard (USCG) [38] were also reviewed briefly at a late stage in the project. These standards seem to be less stringent than others studied, and are not included in Appendix A. USCG intact stability criteria require that the area ratio should be greater than 1.3, the metacentric height (GM) should be greater than 50mm over the full range of draughts, and the righting arm (GZ) should be positive between upright and the second intercept angle. There are no limitations on either the static heel or second intercept angles. USCG damaged stability criteria state that the final equilibrium waterline in operating and storm conditions should be below the lowest downflooding point. There is no minimum area ratio requirement, nor any limitations on the static heel angle or stability range. Wind speeds are identical to those specified by other authorities.

Appendix A summarises the stability standards of the HSE, NMD, CCG, IMO, ABS, and DNV in tabular form. These tables show key aspects of the standards only, as they stand at the time of writing, and as interpreted by the present authors. There are many subtle differences between the standards of different authorities, the implications of which would only become apparent in relation to a specific design. Reference should be made to the relevant organisation to obtain an authoritative interpretation, and to obtain full and up-to-date details.

This comment applies particularly to weathertight and watertight integrity requirements, which are very complex and detailed. Table A2 summarises only those integrity requirements which have a direct bearing on meeting stability standards.

Tables A8 and A9 summarise requirements relating to KG limit curves. Certain organisations (e.g. HSE, NMD, CCG and DNV) have specific and detailed requirements, whereas other bodies (e.g. IMO and ABS) outline general objectives only.

All authorities permit alternative stability standards to be considered. Alternative criteria developed during the ABS JIP will be discussed at some length in Sections 5.1.3 and 5.2.3, firstly because the ABS criteria represent a major new departure from established methods, secondly because they have

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� �

been adopted as either primary or alternative standards by the IMO, and thirdly because of interest expressed through earlier review studies commissioned by the HSE.

Authorities willing to consider alternative standards make statements similar to Paragraph 3.3.3 of the IMO MODU Code [34]: that alternative standards may be considered provided an equivalent level of safety is maintained, and if they can be demonstrated to afford adequate positive initial stability. It may prove difficult, complex and expensive in practice, however, to meet these conditions. In order to demonstrate the acceptability of such standards the IMO states that the following issues should be taken into account:

• environmental conditions representing realistic winds (including gusts) and waves appropriate for worldwide service in various modes of operation;

• dynamic responses of the unit, including the results from wind tunnel tests, wave basin tests and non-linear simulations, where appropriate;

• potential for flooding, taking into account dynamic responses in a seaway;

• susceptibility to capsize, considering both the mean wind speed and maximum dynamic response;

• an adequate safety margin to allow for uncertainties.

4.1 INTACT STABILITY STANDARDS

Established stability standards for all types of vessels, including semi-submersibles, are quasi-static in character. It has nonetheless been recognised for some years that the actual performance of a semi-submersible in severe conditions depends to a large extent on its dynamic behaviour. Conventional stability standards take no explicit account of wave conditions, or of the response of the vessel to waves. They compare the steady wind heeling moment curve with the hydrostatic righting moment curve, both curves being plotted as functions of the vessel’s heel angle about its ‘most critical axis’.

A AA C

A B

Moment arm

Righting moment

Wind heeling moment

rea rea

rea

1 Angle of heel � R � 2

Figure 4.1Wind heeling moment and righting moment arm curves,and parameters used in an intact stability analysis.

Figure 4.1 shows typical heeling and righting moment arm curves, together with parameters considered during a conventional intact stability analysis. The HSE’s intact stability standards [1] for column-stabilised units (i.e. semi-submersibles) may be summarised as follows:

a) The ratio of areas under the heeling and righting moment curves has to satisfy the criterion:

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+ . ( B + C )A B � 13

where the areas A, B and C are integrated up to an angle �R , which is either the second intercept angle, �2 , or minimum downflooding angle, �D , whichever is the smaller.

c) The static heel angle, �1 , must be no greater than 15o.

d) The metacentric height, GM, must be not less than 1.0m in the operating, transit and severe storm conditions, and not less than 0.3m while changing draught between these conditions.

.e) The righting arm parameter, GZ, must satisfy the relationship: GZ � 05 GM sin� over the 0

range 0 � � � (the minimum of �D or �M or 15o ), where �M is the angle of maximum righting

lever, and GM0 is the minimum permissible value of GM specified above.

The area ratio criterion compares the potential energy gained by the vessel through its restoring moment with the work done by the wind heeling moment during a steady wind gust which inclines the vessel from its upright position to the angle �R . The area A + B represents the potential energy of the vessel, and the area B + C represents the work done by the wind heeling moment. The safety factor 1.3 is common to the rules of all regulatory bodies considered during this review, and is intended to make allowance for uncertainties and dynamic effects which are not taken into account in the analysis.

Key differences between the intact stability standards of the various regulatory bodies are:

• The HSE and CCG require the static heel angle to be less than 15°, whereas NMD specifies a maximum static heel angle of 17°. NMD also requires the second intercept angle to be not less than 30°. Other authorities specify no limits on either the static heel angle or second intercept angle.

• All authorities (except IMO) require the metacentric height, GM, to be positive. The HSE, NMD, CCG and DNV all specify higher minimum values of GM, which are common to all four authorities, and depend on whether the unit is in a normal operating, transit or survival condition, or in a temporary condition between draughts.

• The HSE specifies a minimum value of the restoring arm parameter, GZ, based on the minimum permissible value of the metacentric height. This requirement provides a minimum level of stability over a range of small heel angles. Other authorities (IMO, NMD, ABS and DNV) state instead that the righting moment must be positive over the range from the upright condition to the second intercept angle.

Most authorities require the area ratio and other intact stability parameters to be calculated up the second intercept or first downflooding angle, whichever is the lesser. NMD does not consider the first downflooding angle to be an upper limit, and allows compartments to flood sequentially during the analysis, giving the righting moment curve a ‘stepped’ appearance. This procedure is questionable because it implies that these compartments fill, and stability is lost, during a single large roll excursion. Subsequent effects of having already flooded compartments are not considered. It would seem more rational to assume that these compartments remain flooded throughout.

The HSE refers to the first downflooding angle, explaining that this term denotes the angle at which any opening which does not meet requirements for closure reached the water surface. Other authorities refer to the limit of weathertight integrity. There seems to be little obvious practical difference between these two terms. As noted by the HSE [1], the importance of downflooding will depend in practice on the size of the opening and the frequency with which it is submerged. These factors are not taken into account in established stability standards.

Most authorities (including HSE) state that mooring lines should be disregarded in a stability analysis, although ABS also states that detrimental effects of moorings should be taken into account. CCG and DNV require the effects of mooring line weight on the unit’s displacement and KG to be taken into account. HSE and NMD also require the effects of thrusters to be taken into account. The influence of moorings and thrusters will be discussed further in Section 5.4.

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4.1.1 Wind Speed Requirements

There is general agreement on the wind speeds to be considered during a conventional intact stability analysis.

A wind speed of 100 knots (51.5 m/s) is used universally to represent the extreme storm or survival condition. This event is intended to represent an extreme wind gust, the duration of which is long enough to heel the unit. The gust speed may be perhaps 10% to 20% higher than the corresponding hourly-mean wind speed, and is reasonably consistent with 100-year return period hourly-mean wind speeds typically used in design for the north-east Atlantic area.

A wind speed of 70 knots (36 m/s) is used when assessing stability in transit, normal operational and intermediate conditions.

A reduced wind speed of 50 knots (25.8 m/s) is used when assessing stability in the damaged condition. Authorities other than the HSE also permit a 50 knot wind speed to be considered when assessing intact stability at a sheltered location.

Most authorities require the analysis to consider wind from any direction. Wind heeling and righting moment curves then have to be defined with respect to a ‘most critical’ axis or axes. Only the HSE and CCG define what this critical axis is, in terms of the lowest value of the maximum KG. Only the NMD and CCG state that the unit should be free to trim during the inclination process.

Most authorities accept wind heeling moment curves that have been calculated using an approved method, or else measured in a wind tunnel. Many authorities (including the HSE) specify standard ‘building block’ procedures for estimating wind heeling moments, together with force coefficients for individual structural components. The reliability of these standard procedures will be discussed further in Section 5.3.

4.2 DAMAGED STABILITY STANDARDS

Damaged stability standards for semi-submersible units have been in a state of flux for several years, following the losses of the Alexander L. Kielland and Ocean Ranger, and the research activity that followed. At least two different approaches have emerged among the major regulatory authorities (one based on a conventional area ratio requirement, and the other based on minimum values of the righting arm and stability range). A number of alternative or additional criteria (e.g. reserve buoyancy) have also been adopted. Some of the differences between these criteria are obvious, whereas others are more subtle, and the effects would only become clear during a detailed analysis of a specific unit.

Two alternative types of damage are envisaged in most damaged stability standards: either peripheral damage to the outer hull surface, close to the waterline, or internal one-compartment flooding from some cause other than peripheral damage.

4.2.1 Extent of Peripheral Waterline Damage

Peripheral waterline damage is likely to be the result of a low-energy impact with an attendant vessel, and is therefore considered at operating and transit draughts only (not generally in the survival condition). Damaged areas of the hull surface are specified by the authorities, and lie around the outer shells of the pontoons and legs at levels likely to suffer a vessel impact. The regulatory authorities agree on the principles involved in selecting areas of damage, but differ in the zones of damage specified. Peripheral damage zones are summarised in Table A4 of Appendix A.

All regulatory authorities require the horizontal zone of damage penetration to be 1.5m, and assume that watertight subdivision will limit further ingress. This penetration depth was intended to be consistent with the type of damage that might follow a low-energy collision from a supply vessel. The damage from such a collision is likely to be small because vessels of this type tend to be moving at low speeds. Stability standards are not intended to protect units against a collision from a large passing ship or a major structural failure.

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The traditional 1.5m penetration depth nonetheless seems somewhat arbitrary, and does not take account of increases in the size and power of supply vessels over the years, the size of the unit itself or its structural loading capacity. This requirement might therefore have to be reviewed in certain circumstances. Reference [42] proposed that the standards should take account of the design characteristics of the semi-submersible, in a similar manner to criteria relating to collisions between conventional ships.

The standards also specify maximum vertical and horizontal extents of damage across the hull surface. Most authorities (except CCG) have similar requirements about the vertical extent of the damage, which is assumed to extend a vertical distance of 3m at any level between 5m above and 3m below the draught under consideration. The CCG does not specify a height range, referring instead to exposed portions of columns on the periphery, and lower hulls and footings in light and transit draught conditions.

The assumed horizontal extent of peripheral waterline damage varies between different authorities:

• IMO states that the horizontal extent of damage should be one eighth of the perimeter of the column,

• ABS does not specify a horizontal extent of damage, but one eighth of the perimeter of the column may be inferred from bulkhead damage requirements,

• whereas HSE, NMD, CCG and DNV specify that the peripheral extent of damage should be 3m.

NMD requires that watertight subdivision should be sufficient to allow at least one bulkhead to be assumed damaged. The IMO and ABS require no vertical bulkheads to be considered as damaged, unless they are spaced closer than a distance of one eighth of the column perimeter at the draught under consideration, in which case one or more of the bulkheads should be considered as damaged.

The IMO, CCG, ABS and DNV have a further requirement. When any watertight flat surface is situated within the above vertical extent (1.5m above or below the waterline in the case of CCG), then flooding is assumed to occur in compartments both above and below the surface in question.

Other authorities have no specific requirements about damage to bulkheads or flat surfaces, but this possibility would normally have to be considered when the bulkhead or surface lies within a specified damage zone.

4.2.2 Flooding of Internal Compartments

Flooding of internal compartments may occur as a result of failure of a hatch cover, leakage or ballasting errors. These would be regarded as causes of non-peripheral damage, which may affect internal compartments which have no direct contact with the sea. Non-peripheral damage zones and standards are summarised in Table A6 of Appendix A.

• The HSE is more stringent in this respect than other authorities, requiring any one watertight compartment (even those with no direct access to sea water) to be considered damaged. This requirement is intended to provide a basic standard of watertight subdivision throughout those volumes on which buoyancy and stability depend.

• The IMO, ABS and DNV require a watertight compartment to be considered damaged if it lies wholly or partially below the waterline, and is either a pump-room, a room containing machinery with a salt water cooling system, or a compartment adjacent to the sea. NMD has similar, though not identical, requirements.

• The CCG has no requirements of any kind relating to internal one-compartment damage.

4.2.3 Damaged Stability Standards

The HSE’s damaged stability standards [1] for column-stabilised units (i.e. semi-submersibles) may be summarised as follows:

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a) The ratio of areas under the heeling and righting moment curves has to satisfy the criterion:

A B � B C + +

where the areas A, B and C are now integrated from the static angle of heel after damage (without wind) up to an angle �R , which is either the second intercept angle, �2 , or minimum downflooding angle, �D , whichever is the smaller. The downflooding angle for the damaged condition has to allow for 4m clearance above any damaged waterplane.

b) The static heel angle after damage, but without wind, �1 , must be no greater than 15o.

There are important differences between the damaged stability standards of different regulatory authorities, which fall into two distinct groups. The HSE, NMD and CCG base their standards around a minimum area ratio requirement, whereas IMO, ABS and DNV base their peripheral damage standards around maintaining a minimum stability range and a minimum value of the ratio of the maximum righting moment to the wind heeling moment at the same heel angle. The peripheral damage standards specified by the IMO, ABS and DNV are based on recommendations made during the ABS Joint Industry Project, and are discussed in Section 5.2.3. NMD has additional reserve stability requirements, which are discussed in Section 5.2.4.

Key points of similarity and differences between the authorities may be summarised as follows:

• The wind speed is always assumed to be 25.8 m/s (50 knots).

• A most critical heel axis and associated wind direction have to be selected, as in an intact stability analysis. The HSE and CCG specify how the most critical axis should be defined; other authorities do not.

• Authorities generally state that the ability to reduce heel angle by the use of water ballast, pumping, mooring forces, etc., should be disregarded.

• HSE, NMD and CCG require that the area ratio should be at least 1.0. Other authorities specify no minimum value of the area ratio.

• IMO, ABS and DNV specify that the ratio of the righting moment to the wind moment should reach a value of 2 at some angle within the range of positive stability, when considering waterline damage. No other authorities have this requirement.

• IMO, ABS and DNV specify that the range of stability should be at least 7o up to the minimum of the limit of weathertight integrity and the second intercept, when considering waterline damage. Other authorities specify no minimum range of stability.

• The HSE, IMO, ABS and DNV state that the zone of weathertight integrity should extend at least 4m above the relevant damaged waterplane, whereas NMD and CCG do not require a 4m clearance. ABS imposes a further requirement: that this zone should extend 7o beyond the damaged waterplane.

• HSE and CCG specify a 15° maximum static heel angle without wind. IMO, NMD and DNV specify a 17° maximum static heel angle with wind and with peripheral damage. ABS specifies no upper limit on the static heel angle.

• Most authorities apply the same standards regardless of whether the damage is peripheral (around the waterline), or flooding due to causes other than waterline damage. IMO and DNV have less stringent requirements in the latter case, however, requiring only that the maximum static angle without wind should be 25o, and the minimum range of positive stability should be at least 7o. CCG has no requirements whatsoever for one-compartment damage.

• The NMD has a number of additional requirements for reserve buoyancy. These requirements are treated by the IMO as a possible alternative to conventional damaged stability standards. No other authorities have any requirements whatsoever for reserve buoyancy.

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4.2.4 Wind Speed Requirements

The choice of 50 knots wind speed for the damaged condition has been criticised on several occasions because both the Ocean Ranger and Alexander L. Kielland were lost following storm damage. The wind speed at the time of the Alexander L. Kielland accident was in fact reported [3] to be only 16 to 20 m/s (31 to 39 knots), and therefore less than 50 knots. Reports on the Ocean Ranger accident [4] indicate that the mean wind speed was about 70 knots, however, with gusts up to 90 knots. These speeds are well in excess of the 50 knots prescribed in the standards.

The choice of 50 knots wind speed is based on the implicit assumption that damage is most likely to be caused by a collision with another vessel, and will therefore be only slightly correlated with severe storm events. At issue therefore is whether the combination of circumstances that led to the Ocean Ranger accident was a ‘reasonably foreseeable’ event, and therefore one which the standards should reasonably address.

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5. ISSUES ARISING FROM THE LITERATURE REVIEW

5.1 DISCUSSION ON INTACT STABILITY STANDARDS

Established intact stability standards have developed as a result of a long historical process, and seem to have been completely successful in avoiding capsize of semi-submersibles when in the intact condition. This success has led certain investigators to ask whether established stability standards may in fact be too conservative, and may perhaps cause the industry unnecessary costs. It is not easy to assess the amount of conservatism within existing standards, however, and studies have repeatedly indicated that the standards do not provide consistent and uniform levels of conservatism in all circumstances. The authorities have therefore generally taken the prudent view that any relaxation of existing standards should be approached with extreme caution.

Experimental and numerical simulation studies by Takarada et al. [39] showed that an intact semi-submersible unit, satisfying conventional intact stability standards, might get into a ‘dangerous state’ in certain environmental conditions, which might lead to capsize. The ‘dangerous state’ was deemed to occur when the mean heel angle exceeded 15o, or when wave crests hit the underside of the upper deck. In certain circumstances the experimental model took up a large steady tilt angle, which was largely attributed to differential vertical forces acting on the two pontoons: a mechanism which is discussed further in Section 5.6. The investigators concluded that the tilt angle was not affected by green water on the deck. Numerical simulations indicated that wind, waves, current and mooring line forces all affected the unit’s stability. The authors put forward a new stability criterion based on achieving a minimum value of a net righting arm parameter, which was designated GM, but was in fact based on the net righting moment, as a function of heel angle, and took account of wind and current forces, wave-induced steady heeling moments and mooring forces. It is not clear how specific these results are to the particular unit and environmental conditions considered. The experiments and numerical simulations were performed in regular waves only, and the conclusions may only be valid in these artificial conditions. It is therefore not possible to assess how significant these results are for an actual semi-submersible unit in realistic sea conditions.

Chen, Shin and Wilson [43] noted that simplifications in the conventional area ratio criterion provide a certain level of conservatism. This conservatism arises because the model ignores dissipation of energy associated with the motion of the dynamic system, and because the wind gust force is not steady and does not develop instantaneously. These effects represent only part of the uncertainty in the quasi-static model, however, and the overall factor 1.3 is intended to take account of all such uncertainties.

Sample calculations undertaken during the MOPS project [21] indicated that the 1.3 area ratio criterion more than accounted for the effects of environmental loading. In the absence of sufficient data, however, the MOPS project made no recommendations about how far this ratio might be reduced.

Model tests undertaken during the MOPS project indicated that capsize was governed by hydrostatic processes, rather than dynamic effects from waves. MOPS nonetheless recommended that a dynamic analysis should be performed in order to determine proper locations for floodable openings.

Several authors have criticised the fact that the first downflooding angle, used in an intact stability analysis, is normally defined relative to a still water surface. The angle at which openings enter the water will in practice depend on the motions of the vessel and of the water surface. Whether intermittent submergence is important will depend on the size of the opening, and the frequency and duration of immersion, together with the ability of bilge pumps to de-water. Existing quasi-static intact stability standards take no account of these factors, although the HSE’s damaged stability standards make some recognition of this fact by requiring a 4m allowance for the effects of waves. The Ocean Ranger accident drew attention to deficiencies in the standards, because chain lockers became flooded long before the quasi-static angle of downflooding was reached. The rationale of assuming calm water in the intact condition, and a 4m clearance in the damaged condition, is not clear, considering that more severe environmental conditions have to be considered in an intact analysis.

Certain successful designs utilise the buoyancy of the upper deck to gain a rapid increase in the area under the righting moment curve, in order to meet the 1.3 area ratio criterion. Mills et al. [16] noted that use of upper deck buoyancy conflicts with established design practice for fixed platforms, where

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the aim is to maintain an air gap between the deck and water surface. The principle of carrying out wave-in-deck analyses is becoming accepted for fixed structures, however, as an alternative to maintaining the air gap, and there would seem to be no logical objection to eliminating the air gap provided structural integrity and other safety requirements are met [40].

The upper deck will only provide buoyancy at large heel angles. If such large heel angles were to occur in practice, they would probably cause non-fixed items to move, risking major structural damage to the unit, or causing a shift in its centre of gravity, and thus compromise its stability. Large heel angles would also cause hazards to personnel, and would hinder evacuation and escape. Although these large angles would only occur in extreme conditions, it seems undesirable that such large angles are needed at all in order to meet basic stability standards.

Mills et al. [16] also noted that the righting moment curves for semi-submersibles often have discontinuities associated with diving tube surge tanks or external appendages. These discontinuities can cause the unit to take up an angle of loll, even though the unit strictly complies with the standards. Mills et al. noted that additional stability requirements would normally be applied in these circumstances, even though these are not specified in standard criteria.

5.1.1 Free Trim or Fixed Trim Calculations

Van Santen [41] criticised the authorities for being ambiguous about the definition of the ‘most critical’ heel axis (or axes), and about the way in which the unit should be heeled during a stability analysis. The HSE and CCG define the ‘most critical’ heel axis as the one for which the maximum KG satisfying the criteria takes its lowest value. The NMD and CCG state that the vessel should be able to trim freely while it is being heeled. Other authorities leave these interpretations to the designer. Unlike a ship, the ‘most critical’ axis and heel angle for a semi-submersible or jack-up are not always obvious, and the analysis must be performed for all wind directions in order to find the worst one in terms of the unit’s self-restoring capacity.

The problem arises because heeling a unit about one axis will generally give rise to a trimming moment about the perpendicular axis, and may also give rise to a vertical heave force. There are two alternative ways in which the trimming moment may be eliminated:

• by applying an appropriate trim angle (‘free trim’), or

• by rotating the vessel about its yaw axis (‘free twist’).

The difference between these procedures is that, in the first case the unit rotates about an initially horizontal axis, whereas in the second it rotates about an initially vertical axis. The definition of the heel angle varies in consequence. Either procedure is acceptable, but care is needed in interpreting the results. If the ‘free twist’ approach is adopted, the heel angle of the vessel is the angle of steepest slope of an initially horizontal plane. In this case, however, the wind direction has to be adjusted so as to be consistent with the twisted heel axis. If the ‘free trim’ approach is adopted, the definition of the heel angle is ambiguous, and is always less than the angle of steepest slope.

Procedures used by the industry vary, depending on the capabilities and limitations of individual hydrostatics programs. Some programs allow the vessel to change its trim and draught automatically, whereas others require manual adjustment.

8

Van Santen [41] found in sample calculations that, depending on the interpretation, differences of 5o to o in the stability range might result, and this might make the difference between passing or failing the

acceptance criteria. The differences follow no obvious trend, and may be a function of the design and condition of the vessel under consideration [42]. Van Santen expressed the view that definitions of the ‘most critical’ axis and heel angle, and the calculation procedure, should be specified by the regulatory authorities.

Organisations routinely involved in stability assessment currently apply a variety of analysis techniques, including both the ‘free trim’ and ‘free twist’ methods, and important aspects (such as the definition of the heel angle relative to the absolute vertical, and the definition of the most critical axis)

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seem to be well understood by the experts. It is nonetheless important that key procedures and definitions should be understood clearly by all engaged in such work.

5.1.2 Minimum Metacentric Height and Righting Moment

The metacentric height, GM, has been used for many years as a simple measure of a vessel’s stability, and minimum values of GM in the intact condition are still specified by most authorities. ABS only requires GM to be positive in the intact condition, whereas HSE, NMD, CCG and DNV require the metacentric height to exceed certain positive values, depending on the unit’s condition. Minimum values of GM are intended, firstly, to provide a margin for error in the vertical centre of gravity (VCG) position, and, secondly, to prevent excessive heel due to wind [42].

A further justification for requiring a minimum value of GM is that steady tilt angles and low-frequency rolling motions (see Section 5.6) only seem to have been observed in model tests where the metacentric height was low.

The metacentric height provides a simple indication of the vessel’s stability at small heel angles, but provides no indication of its stability over a range of heel angles. Most authorities therefore also require the righting moment to be positive between the upright condition and second intercept angle. The HSE has a more stringent requirement, however, which specifies a minimum value of the righting arm parameter, GZ. Minimum GM requirements alone have been found to be insufficient for certain non-wall sided units, and a minimum GZ requirement provides a minimum level of stability over a range of small heel angles.

As noted in Section 5.1, large heel angles are occasionally needed in order to comply with certain stability standards. Large heel angles are unlikely to develop, however, if the unit provides sufficient righting moment at small heel angles, providing further justification for specifying minimum values of both GM and GZ.

There are several possible reasons, however, why a designer or operator might wish to keep the metacentric height as low as practical:

• to reduce the cost of the unit,

• to improve its motion characteristics (by increasing its natural roll and pitch periods),

• to increase the unit’s carrying capacity.

Model tests have shown that vessels with very low metacentric heights can fulfil the stability requirements, and never show any tendency to capsize in the intact condition [24]. Model tests were performed on a semi-submersible in a 140 knot wind, with a metacentric height of 0.6m, area ratio of only 0.82, maximum wave height of 36.6m, with the moorings over-ridden to increase the wind moment effect. The model did not overturn, ‘nor was any capsizing situation approached, even though the top of the leeward caisson was frequently several feet under the water surface’. Significant wave impacts occurred under the platform deck, and the lower the available righting moment, the more pronounced the impacts were seen to be. Although deck impacts would not directly lead to capsize, they might cause damage and flooding, and might therefore be an indirect cause of loss.

Numata [24] concluded that there are good reasons for maintaining a positive metacentric height, sufficient to facilitate safe operations. Abrupt changes in the unit’s heel or trim angle could, for example, jeopardise the safety of personnel or cause damage to equipment and the vessel’s structure. Maintaining a minimum value of GM should also avoid the development of a permanent tilt angle (see Section 5.4). Abrupt changes in heel or trim angles might be expected to occur in the following circumstances [24]:

• when the drilling derrick is not above the centre of flotation, and has to lift a heavy weight,

• failure of a mooring line,

• heavy lift crane operations (not covered by Safety Case regulations).

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Numata [24] nonetheless suggested that the choice of a suitable minimum value of GM should be a matter of judgement and experience, and should only depend on the type of operation required.

The MOPS model tests [21] demonstrated that GM was not an important parameter for determining capsize. An intact model with negative GM survived extreme waves and wind. The report noted, however, that no downflooding openings were represented in the model tests, and that small or negative GM values are unacceptable for operational and safety reasons. The project therefore recommended that minimum GM requirements should be retained in the regulations.

5.1.3 Alternative ABS and IMO Intact Criteria

A major new development in the past 15 years has been the publication of alternative intact stability criteria for semi-submersibles by the ABS [26], based on results from a Joint Industry Project (JIP). These alternative criteria take account of certain dynamic aspects of semi-submersible behaviour, and were based on results from systematic series of numerical simulations and model tests on a number of generic semi-submersible types.

In a pilot study for the ABS JIP, Chen et al. [43] compared results from a limited range of numerical non-linear time domain simulations with results from model tests, and concluded that dynamic motion analysis procedures provide a ‘reasonably accurate’ prediction of semi-submersible motions in waves. Similar conclusions were reached during the Norwegian MOPS project [21]. The ABS pilot correlation study appears to have considered a stable rig model in regular wave conditions only. It is not clear, therefore, whether the time-domain simulation model was capable of predicting large-amplitude and capsize of the unit accurately in random waves. The ABS pilot study nonetheless identified a number of parameters which affected the unit’s behaviour, including the metacentric height, draught, wind unsteadiness and effects of moorings. From this investigation Stiansen et al. [44] concluded that neither the area ratio nor the metacentric height guarantees a uniform safety margin against downflooding.

Further systematic numerical simulations and model tests took place within the ABS JIP [44]. The JIP considered a parametric series of 27 generic units, representing a range of 13 recent twin-hull drilling semi-submersible designs. These units included small, medium and large displacements, with varying waterplane areas. After an initial appraisal 15 of these units were selected for a frequency domain analysis. Detailed numerical simulation studies were then performed on seven of these units.

The new criteria included specific terms representing the dynamic effects of wind and waves, and the investigation considered combinations of wind and wave parameters representing extreme (100-year return period) hurricane and sustained storm conditions:

• Set I: an extreme hurricane in the Gulf of Mexico or South China Sea, represented by a 100-knot wind speed and 12.5m Hs ,

• Set II: a severe storm in the North Sea, represented by a 75-knot wind speed and 16.8m Hs .

The simulations considered alternative forms of wave spectra, and the wind gust spectrum was chosen to provide a conservative input of energy at low frequencies. The peak spectral wave period was varied, and the wave spectrum considered in each simulation was chosen to be critical to the unit in question.

Key results from the JIP and proposed new intact stability criteria were reported in references [44, 45]. The JIP concluded [44] that downflooding comes about as a combined effect of vessel heave, pitch, roll and wave surface elevation. Furthermore the effects of downflooding during the simulations were found to be less severe than conventional standards suggested. According to the authors, these inconsistencies arise from two so-called ‘erroneous assumptions’ within the existing standards:

• the 30% margin of safety, based on the wind heeling energy, accounts for dynamic responses due to wind gusting and wave forces;

• downflooding is assumed to be a function of roll and pitch only.

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These investigations suggested that established stability standards provide an inconsistent margin of safety between different units, in terms of avoiding both downflooding and capsize.

Shark et al. [46] also noted that deck submergence tended to change the unit’s natural roll and pitch periods. In some circumstances this change increased the relative motion between the unit and the water surface, and increased the risk of downflooding.

ABS [26] subsequently developed stability criteria which were intended to be more closely related to the physical phenomena, but at the same time were straightforward and unambiguous to apply. The criterion for downflooding took into account the fact that downflooding was highly influenced by the unit’s heave response and the profile of the seaway, whereas capsize was influenced by the vessel’s rotational motions, and therefore by the wind. The new ABS downflooding criteria were therefore based on results obtained in Set II (extreme North Sea) conditions, whereas the capsize criteria were based on results obtained in Set I (Gulf of Mexico hurricane) conditions.

ABS found that downflooding generally occurred well before capsize. Downflooding would not occur on a unit with complete watertight integrity, however, and a separate capsize criterion was therefore considered to be necessary.

The ABS downflooding criterion [26, 44, 47] takes the form of a requirement that the reduction in the downflooding distance, RDFD, due to both static wind heel and dynamic vessel response, should be not more than the initial downflooding distance, DFD0 :

DFD0 � RDFD � 0 0.

where: RDFD = S k QSD f ( 1 + RMW )

The static wind heel component, QSD1 , should be calculated using a wind speed of 75 knots (i.e. Set II conditions). The correlation factor k was intended to represent the effects of wind gusting, and is to be calculated using a simple empirical formula. The safety factor, Sf , was set equal to 1.10. This factor was intended to allow for uncertainties associated with the numerical predictions, including both extreme value prediction and non-linear effects.

The capsize criterion is based on a higher wind speed of 100 knots (i.e. Set I conditions). The criterion is based on the ratio of two areas under the righting moment curve, and is therefore based on energy considerations. Two areas under the righting moment curve have to be calculated:

i) an area A � between the static heel angle (the first intercept) in a 100-knot steady wind and a so-called ‘maximum dynamic response angle’, �MAX ,

ii) an area B � between �MAX and the second intercept angle.

The reserve energy ratio, RER: .RER = Area B � Area / A� � 010

provides a measure of the additional energy required to capsize the vessel compared with the energy required to heel it to its maximum dynamic angle. The area under the wind heeling moment curve is not included in the criterion, but the wind heeling moment curve determines the first and second intercept angles.

ABS proposed simple fitted formula for calculating the dynamic response component of the downflooding distance, RMW, and the maximum dynamic response angle, �MAX . These formulae were based on various dimensions, areas and volumes of the unit. These formulae may only be used within certain ranges of these parameters.

Beneficial effects of under-deck buoyancy and slamming forces were not included in the simulation model, and the maximum dynamic heel angle obtained from the simulations was increased after carrying out the model test correlation study.

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It is not obvious, however, whether these formulae will provide conservative results under all possible combinations of parameters within their specified ranges. Nor is it obvious whether the procedures and assumptions adopted in the simulation model and subsequent analysis are valid throughout the entire range of possible conditions. It is noted that ABS permits use of specific calculations for a specific design, rather than the simple fitted formulae, provided these are based on procedures similar to those adopted during the original JIP study.

ABS [45, 47] commented on the need for further research into non-linear effects, such as wave drift forces and wave diffraction, effects of upper deck submergence and pontoon emergence, the representation of wind gust effects, and extreme value prediction techniques to account for non-linearities, and combined random processes. Having acknowledged these uncertainties, however, ABS considered that there were adequate safety margins and sources of conservatism in their proposed criteria.

The new criteria were assessed in terms of their effects on 31 existing semi-submersible designs. Conventional standards were considered to provide an inconsistent margin of safety against downflooding. The change in the criteria resulted in a reduction in the minimum value of GM for 19 designs, an increase for 10 designs, and no change in the remaining two cases. ABS nonetheless claimed that the new criteria preserved the appreciable margin against capsize afforded in existing designs [47]. ABS also found that the downflooding criterion determined the minimum GM value in 75% of existing designs. Designs with upper deck buoyancy more readily met the downflooding criterion, and designs with a greater initial air gap readily met the capsize criterion.

Only one of these semi-submersible units was model-tested [48, 49]. This unit was regarded as the median model among the 27 generic units considered during the study. The model tests were intended to calibrate and validate the numerical model, and to provide hydrodynamic coefficient values for use in the simulations. Over 200 separate tests were performed, covering a range of six draughts, a range of moderate to extreme regular and irregular wave conditions, together with specialised tests to investigate low-frequency subharmonic responses. The majority of the tests were in the intact condition, but a small number of tests took place with weights added to represent a flooded corner column compartment. The test programme also investigated the effects of an impulsively applied wind, and of an unsteady wind modelled by wind fans.

It was not possible to review the model test and correlation reports in detail. The summary reports and published papers [44, 45, 47, 48, 49] indicate, however, that the correlation study was primarily concerned with comparing predicted and measured response amplitude operators (RAOs) and statistical relationships. After the hydrodynamic coefficients had been calibrated using the test data, the heave, roll and pitch RAOs generally seem to have agreed well. Shin and Shark [45] noted discrepancies, however, between predicted and measured root mean square responses in random waves. These discrepancies were attributed to differences between the waves used in the numerical model and those occurring in the test basin. There was acceptable correlation between the measured and predicted subharmonic responses [45], although the predicted low-frequency motions in large-amplitude regular waves were of somewhat lower amplitude. The numerical model was considered to represent second order effects ‘to a degree’ [47]. The most probable maximum values of roll, pitch and heave in a three hour extreme storm were predicted well. Extreme relative motions (which are important for estimating downflooding) were found to be sensitive, however, to the phasing of vessel motions. Some problems were reported in representing wind gusting, where the turbulence level from the wind fans was found to be too low [48].

The results from the model test correlation study seem to have given reasonable confidence in the ability of the numerical model to predict the behaviour of an intact unit in moderate and large sea states, although it was not clear whether the model is capable of predicting the unit’s behaviour in near-capsize conditions.

Song, Petty and Bone [50] compared results obtained using conventional and alternative ABS intact stability criteria with results obtained from time-domain simulations of the motions of a moored Aker H-3 semi-submersible. The simulations represented the effects of a wind gust spectrum and mooring lines, although the ABS criteria require that mooring lines should be ignored except when they produce detrimental effects. The alternative criteria allowed a 500 ton increase in the deck load when compared with conventional standards. The authors also concluded that results obtained using both sets of ABS

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criteria were more conservative than the results obtained from the computer simulation study. It is not clear, however, whether any general conclusions can be drawn from these specific results.

Takaishi et al. [51] reported the results from a Japanese study which compared results from numerical simulations with the ABS alternative criteria. Numerical simulations were performed on three generic semi-submersible units in both the intact and damaged conditions, varying both the unit’s draught and metacentric height. The survival wind and wave conditions used in this study were similar (but not identical) to the ABS Set II conditions. The intact units showed no tendency to capsize in survival conditions. The authors concluded that the ABS empirical formulae represented the vessels’ dynamic responses well, although the results showed significant (up to ±30%) variations in the level of agreement. They noted that one model failed the capsize criterion with GM equal to zero, and two models failed at maximum draught. One model also failed the downflooding criterion at deep draught. The authors made no attempt to draw general conclusions.

Further unpublished studies, commissioned by the HSE, found that similar levels of effort were required to apply the new ABS criteria and more established methods, and (in specific cases) both procedures seemed to result in similar margins of safety against capsize when the unit was intact. Concern was expressed about uncertainties in the safety margins, however, and about possible variations in safety margins between different units.

A number of other issues, such as the adequacy of the ABS empirical formulae in representing resonance effects, may require further investigation. It is also disturbing that the ABS/ IMO alternative criteria ignore the overall shape of the wind heeling moment curve, and only consider the first and second intercept angles. It is in principle possible to satisfy these criteria with a very low area ratio and very little reserve energy.

The new ABS intact stability criteria [26] were later proposed for adoption by the IMO in a Ship Design and Equipment Sub-Committee resolution [52]. They were eventually adopted in paragraph 4.6.5 of the IMO Code on Intact Stability [33], with their validity restricted to units with area and volume ratios lying within specified ranges. The IMO Code describes them as an ‘example’ of alternative intact stability criteria, and makes the rig designer responsible for justifying the adoption of these or any other alternative standards.

5.2 DISCUSSION ON DAMAGED STABILITY STANDARDS

Semi-submersible units have been lost on a number of occasions after damage and flooding (see Section 7), but only after major structural damage has already occurred or serious errors in ballast control have been made by the crew. Deficiencies in the stability standards were not identified as a key factor in any known major incident. This fact alone does not mean that existing standards are entirely satisfactory, however, and several research studies have identified deficiencies in the standards, or circumstances in which units could experience undesirable dynamic behaviour even though they apparently comply with the standards. Results from some of these research studies are reviewed below.

Concern after the losses of the Ocean Ranger and Alexander L. Kielland resulted in many suggested new ways in which stability might be assessed, and damaged stability standards for semi-submersibles are still in a greater state of flux and uncertainty than intact standards. Two distinctly different approaches have emerged among the regulatory bodies:

• the established approach, taken by the HSE, NMD and CCG, based on meeting minimum area ratio and maximum static heel angle requirements, together with a maintaining a minimum clearance between the water surface and downflooding points,

• the approach first put forward by ABS, but since adopted by the IMO and DNV, based on meeting minimum righting moment to heeling moment ratio, static heel angle and stability range requirements.

It is not easy to justify either approach in an entirely rational manner, and not easy to decide which approach offers greater margins of stability, or better levels of consistency and uniformity between

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different units. The established approach has been in use for many years, and there is little evidence to suggest major failings. This alone gives it a certain level of credibility. The new approach was based on numerical simulation studies and physical model tests, but on a very limited number of units and damage conditions, using a highly simplified physical model. These simulation and model test studies nonetheless had the merit of explicitly representing the dynamic behaviour of the unit, and of wind and wave loading, in a realistic manner.

The established approach makes no explicit allowance for any of these dynamic effects, or treats them in a very simplified manner (e.g. as a quasi-steady load).

The traditional minimum area ratio in the damaged condition is 1.0, and therefore includes no safety margin to allow for uncertainties, or for vessel and wave motions. Some allowance for these effects is made when calculating the minimum downflooding angle, by maintaining a certain minimum clearance above the damaged static waterplane. The specified clearance (4m in the case of most authorities, but zero in the case of NMD and CCG) seems to be somewhat arbitrary. Downflooding will in practice be affected by the sea state, the vessel type and behaviour, and the location and size of the downflooding opening.

The inclusion of dynamic effects in the downflooding assessment can be especially important. Model tests undertaken after the Ocean Ranger accident [4] showed that the static angle of downflooding to the chain lockers, based on the still water level at 80 foot draught, was 27 degrees of trim by the bow. The model tests demonstrated that the rig was susceptible to downflooding, however, when only trimmed by 10 to 12 degrees by the bow. The substantial difference between the downflooding angle indicated by static calculations and that observed during the model tests was probably caused by motions of the water surface and vessel.

5.2.1 Dynamic Behaviour when Damaged

Vassalos et al. [11] listed the following effects which, alone or combined, may lead to the capsize of the unit:

• listing due to damage or incorrect ballasting,

• progressive flooding,

• resonant rotational motions due to swell or non-linear effects,

• some loss of stability due to the local wave system,

• storm waves impinging on the vessel,

• wind and current effects,

• the influence of mooring lines.

Sample calculations by Vassalos et al. [11], based on their own proposed new stability criteria, suggested that conventional intact standards were conservative, but conventional damaged standards might be unconservative in certain respects.

In a series of model tests, reported by Stone et al. [53], moderate and higher levels of damage to a single column seemed to cause significant subharmonic motions. The location of damage also influenced the unit’s behaviour. There were substantial differences between the roll and pitch motions of the unit, depending on whether the damaged column was on the windward or leeward side of the unit. Asymmetries in the subharmonic response were also reported by Takaki and Higo [73], when studying a submersible with a large list angle.

Model tests during the MOPS project [21] indicated that a platform damaged at the bow is at greatest risk when waves come from astern, whereas waves onto the bow have a stabilising effect.

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5.2.2 Transient Motions

Breakage of a mooring line or sudden flooding can cause transient motions, and the maximum displacement and inclination during this transient motion are likely to be larger than either the initial or final at-rest conditions, and in some circumstances may result in further damage or may endanger personnel. The rate of flooding is a key factor influencing transient behaviour. If the rate of flooding is low, the unit will respond in a quasi-static manner, and no transient will be observed. If the damage is sudden, however, then the transient response may be very large.

Adachi and Kagemoto [54] carried out model tests to investigate the behaviour of a semi-submersible following breakage of a mooring line and flooding of tanks. Mooring line breakage caused a fairly large transient roll response of several degrees. Flooding occurred more gradually, and the results appear to be quasi-static in character, with little obvious transient response. Adachi and Kagemoto nonetheless recommended that transient motion after damage should be considered in safety regulations.

Collins and Grove [48] also carried out model tests to investigate transient behaviour following failure of a mooring line, as part of the ABS JIP. The model test set-up was unable to reproduce mooring line tripping, but represented a severe failure condition because the single failed line on the model represented two lines on the actual unit. The authors reported that the unit never approached the capsize condition during the transient, even when the model was subjected to a large wave in combination with a 100 knot steady wind.

Söylemez and Incecik [55] presented results from numerical simulations of the behaviour of a semi-submersible in waves before, during and after flooding. The authors claimed that transient motion displacements during progressive flooding can be significantly higher than those during the post-flooding stage. They also claimed that as the flooding time decreased, the motion amplitudes increased. These changes appear to be fairly small, however, and the results contained in the paper seem to show little transient response.

The above results indicate that flooding is unlikely to occur quickly enough to cause a significant transient response. Transient behaviour following failure of a mooring line may cause the vessel to roll unexpectedly by a few degrees. This might endanger personnel, but seems unlikely to cause capsize. These issues should nonetheless be addressed as part of a safety case.

5.2.3 ABS/ IMO/ DNV Damaged Criteria

The ABS JIP put forward an alternative set of criteria for residual stability after damage [46, 47]. The ABS investigation was based on numerical simulations on three generic damaged semi-submersible units, and considered variations in the initial shape of the righting arm curve, upper deck buoyancy, flooded compartment volume and heading. The results indicated that any tendency to capsize could be assessed using two parameters: the ratio of the maximum righting moment to the heeling moment at the same heel angle, and the range of positive stability [46].

The ABS residual stability criteria are as follows:

• The ratio of the maximum righting arm to the corresponding 50-knot wind heeling arm should be greater than or equal to 2.0, and a minimum 7o range of positive stability should be provided.

• A zone of weathertight integrity should be maintained 4m above, and 7o beyond, the 50 knot damaged equilibrium waterline.

The range of positive stability appeared to be the controlling criterion in beam seas, whereas the minimum ratio of maximum righting arm to heeling arm was the controlling criterion in diagonal seas. The numerical simulations showed that semi-submersibles possessing a maximum righting arm greater than twice the heeling arm would not capsize in the environmental conditions considered. In cases where capsize did occur, the range of positive stability was very small.

Model tests were performed on a model with 17o heel after damage. ABS [47] stated that the time-domain simulation results agreed well with the model test data, although few details were available in

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the documents that were examined [47, 48, 56]. The hydrodynamic coefficients used in this study remained unchanged from the earlier intact stability study. So-called one-to-one comparisons were made between short segments of time histories obtained from the model tests and from the corresponding simulations.

ABS [56] concluded that:

• The initial GM can influence the amount of residual buoyancy, but was not in itself an adequate indicator of the amount of residual buoyancy possessed by an semi-submersible, especially at large heel angles achieved after damage.

• Upper deck buoyancy can have a direct effect on residual stability.

• Upper deck buoyancy when submerged appeared to increase the relative motion of the unit by reducing its natural periods in roll and heave, possibly increasing its susceptibility to intermittent downflooding.

• Capsize occurred in the numerical simulations in damaged cases having very little residual stability. Pontoon emergence and column/upper deck submergence within the same angular proximity tended to cause a multi-peaked, flat GZ curve, and large dynamic responses.

A number of criticisms have been levelled against these new criteria. The most serious is that they were based on a very limited modelling programme: numerical simulations on three generic units only, and model tests on only one unit with only one leg damaged, the damage being represented by solid weights [47, 48].

The limited number of cases considered during this phase of the ABS JIP, and the simplified way in which damage was represented in the physical model, are particular aspects of concern. The margins of stability, levels of consistency and uniformity provided by the resulting criteria are also far from clear.

The ABS recommendations have nonetheless been adopted as the basis for the IMO’s 1989 MODU Code [34] and by DNV, but have not so far been adopted by the HSE, NMD and CCG. It is suggested that these criteria should be treated with extreme caution until they have been validated against a wider range of different units and damage conditions.

5.2.4 Alternative IMO and NMD Reserve Buoyancy Criteria

Reserve buoyancy criteria were developed in the aftermath of the Alexander L. Kielland accident, based on recommendations made during the Norwegian MOPS [21] and RABL [7] projects. The NMD took the lead in setting reserve buoyancy requirements for semi-submersibles, in addition to standard intact and damaged criteria. The aim was to provide additional buoyancy in the above-water structure, sufficient to avoid capsize after a major loss of righting moment, thereby giving the crew time to evacuate.

Initial proposals to the IMO were rejected on the grounds that they were too restrictive, and imposed cost penalties on many designs, because safety features inherent in those designs were not taken into account. Further Norwegian proposals were placed before the International Maritime Organisation’s Sub-Committee on Design and Equipment in the form of two papers (DE 31/4/7 and DE 31/4/8). These recommendations were also rejected by the Sub-Committee, on the grounds that they were too prescriptive, and not the best way in which to achieve the desired objective.

The basic principle of providing some level of reserve buoyancy, beyond that necessary to meet basic code requirements, was nonetheless widely accepted [16]. The IMO Committee concluded that the intended effect of having a reserve of buoyancy was to provide a margin of righting lever and righting energy, and that this would be a better way in which to formulate requirements for reserve buoyancy.

A follow-up study on behalf of the HSE concluded, however, that it was not possible to derive a general relationship between reserve buoyancy and reserve stability.

IMO subsequently adopted Resolution A.651(16) [57], the requirements of which are essentially the same as current NMD reserve buoyancy criteria [31]. There is an important difference, however, in the

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way IMO and NMD apply these requirements. Whereas Resolution A.651(16) is described as an ‘example’ of an alternative to the MODU Code 7o stability range and minimum GZ requirements, applying only to units which have buoyant volumes contained in watertight upper-deck structures, NMD requires all semi-submersible units to satisfy both their own conventional damage (based on a minimum area ratio) and reserve buoyancy criteria simultaneously.

The NMD and IMO reserve buoyancy requirements assume the same extent of damage as the relevant standard criteria, together with 50 knots wind speed. The criteria are as follows:

• the positive range of stability after damage should be at least 10 degrees between the first and second intercepts,

• for one and the same angle within the positive range: the righting lever after damage should reach a value of at least 2.5m, and

• at least 1.0m of this righting lever should come from enclosed watertight volumes above the lowest continuous deck.

• Various means of watertight closure are specified for openings in areas providing reserve buoyancy.

Follow-up studies for the Norwegian authorities concluded that the alternative IMO criteria would be at least as safe as the standard 1989 MODU Code procedure, and in most cases should be much safer.

The margins of stability, and the levels of consistency and uniformity provided by these alternative IMO/NMD criteria are not clear. It is also not clear how these margins of stability compare with those provided by the HSE standards. As noted earlier, the HSE’s damaged stability standards differ from the IMO’s 1989 MODU Code in several important respects. It would only be possible to establish the relative margins of safety between the HSE’s and other authorities’ standards by carrying out systematic calculations and model tests on a number of different units.

5.3 WIND FORCES AND HEELING MOMENTS

Wind heeling moments are of crucial importance in a stability assessment, because they are the only environmental factor included in the analysis. Standard ‘building block’ calculation procedures, as specified by many regulatory authorities (including the HSE) for estimating wind heeling moments, have been criticised on a number of occasions.

Numata [24] found that standard ABS procedures overestimated the drag force on one unit by 50%, and the heeling moment by 20%, compared with wind tunnel tests. Numata’s investigation also revealed that lift forces and lift-induced moment were significant.

Miller and Davies [58] compared results from wind tunnel tests on semi-submersibles with an estimation procedure. Errors in the estimated heeling moment varied between ±30% at level trim and heel, and in quartering conditions the drag force was consistently underestimated by 20%. Miller and Davies concluded that the lift-induced moment was significant, and that wind tunnel tests were the only satisfactory method for establishing wind moments with any precision.

Yu and Won [59], Ogiwara and Sakata [60] reached similar conclusions from comparisons between wind forces and moments obtained using regulatory authority formulae and wind tunnel measurements. Their conclusions may be summarised as follows:

• The calculation procedures always provided conservative estimates of the drag force, compared to wind tunnel tests. Ogiwara and Sakata suggested that the calculations were conservative because interaction effects, such as shielding, were not taken into account correctly.

• The calculated heeling moments were not always conservative however. Lift forces made a significant contribution to the heeling moment, even at zero angle of heel.

• The DNV procedure, which represented lift forces on large surfaces such as the main deck, predicted the heeling moment better than the ABS procedure.

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Freathy and Taggart [61] also compared results from standard calculation procedures with wind tunnel test data. They concluded that the ABS and British Standard procedures predicted drag forces at an even keel with acceptable accuracy, Two other codes (DNV and API) gave consistently high estimates. Agreement was improved if wind areas were calculated orthogonal to the flow, and if shielding was taken into account. Moments were predicted much less accurately by the ABS method. At even keel, the calculations did not seem to represent the distribution of loads between different levels accurately, and the procedure did not take vertical lift forces into account when calculating moments on an inclined platform. Use of a semi-empirical moment correction method was proposed.

Singh [62] compared a number of alternative calculation procedures, and reached the following conclusions:

• Recommendations on gust factors are not always clear. The implication is that one-minute mean speeds should be used, but 15-second or 3-second values should be used for certain items. Clearer and consistent guidelines on gust durations and assumptions are needed.

• There is an inconsistency between approaches adopted towards fixed and mobile structures in respect of wind shear. There are also inconsistencies in the formulation of the shear profile.

• There is an abundance of data for isolated elements, such as circular cylinders, in uni-directional steady flow, but little for more complex structures or circumstances.

• Methods based on an assessment of the complete structure, rather than on the interaction of individual elements, seem to be promising for truss-type sections. Methods such as the lattice tower code should be validated, and then adopted by the authorities.

• Wind tunnel measurements appear to be the only viable method for estimating loads on complex offshore structures. These measurements should be made towards the later stages of the design process, in order to refine and calibrate/ validate calculation procedures.

• There are limitations on what can be achieved in wind tunnels, but, provided certain basic similarity requirements are met, wind tunnel techniques can provide reliable data.

Despite the importance of wind heeling moments in a stability analysis, the ‘building block’ procedures commonly used to calculate these moments still seem to be a matter for concern. Substantial differences have been reported between wind heeling moments obtained using standard recommended procedures and those measured during wind tunnel tests. Calculation procedures generally seem to provide conservative estimates of drag forces, but this is not necessarily true of the overturning moment. Calculation procedures proposed by most authorities (including the HSE) are based on drag forces only, using projected areas in the vertical plane, and do not represent the lift-induced moment. Several investigators have shown that lift makes a significant contribution to the heeling moment on a semi-submersible, even when on a level keel, and the calculations may either over or under-estimate the true heeling moment.

At this stage wind tunnel tests seem to offer the only reliable means of estimating the overturning moment.

5.3.1 Dynamic Wind Loading

Established stability standards are based on calculating a quasi-steady wind heeling moment, and dynamic wind gust loading is not normally considered in the analysis. It is nonetheless worth noting that the conventional area ratio concept is in fact based on the work done by a simple quasi-steady wind gust which heels the unit from an initially upright condition.

Model tests undertaken during the MOPS project [21] seemed to indicate that the effects of wind gusting were small.

Various other studies have shown, however, that wind gusting can significantly affect a semi-submersible’s dynamic behaviour. A study by Chen, Shin and Wilson [43] showed that wind unsteadiness can have an appreciable influence on the motion of a semi-submersible in heavy sea conditions. Under certain conditions, the roll motions caused by the unsteady wind could be of similar

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magnitude to those caused by waves, and could also equal the mean roll angle caused by the mean wind.

Similar conclusions were reached by Rowe, Brendling and Davies [63], and during a follow-up study by Standing, Wills and Singh [64]. Rowe, Brendling and Davies found that large wind-induced pitch motions of a semi-submersible were predicted by a numerical simulation model. They also pointed out that the platform’s roll and pitch motions could have a significant non-linear effect on the wind force and heeling moment, as the area and shape presented to the wind vary. They suggested that use of constant aerodynamic coefficients, independent of the roll and pitch angles, might lead to either under or over-estimation of the unit’s maximum pitch response.

Standing, Wills and Singh reported results from model tests in unsteady wind and waves, which showed that the responses to wind and waves could generally be calculated independently, and then superimposed. This procedure underestimated the mean pitch angle, however. They also reported large differences between forces and moments obtained from measurements in steady and unsteady flow, which might have been due to non-linearities in the forcing process.

Several alternative wind gust spectral formulae have been used in offshore design. These formulae represent significantly different amounts of energy at low frequencies. Many of these spectra were derived from measurements over land, and intended for the design of land-based structures with relatively high natural frequencies. These formulae are generally inappropriate for representing conditions at sea. Standing et al. [64] recommended use of a modified Kaimal spectrum for offshore design, based on measurements at the West Sole Field.

The Norwegian Petroleum Directorate (NPD) [65] specified a wind gust spectrum ‘for structures and structural elements for which the wind fluctuations are of importance’, and the NPD wind spectrum has now become a de facto standard in certain areas of offshore design. Vessel stability standards are outside the jurisdiction of NPD, however.

The ABS JIP [47] used an alternative wind gust spectrum, based on offshore measurements, which was claimed to be more conservative than others then in use. It is not clear, however, whether this spectrum is more conservative than the modified Kaimal or NPD spectra.

The above investigations showed that wind gusting can have a significant effect on a semi-submersible unit’s dynamic behaviour. It is also clear that non-linearities, associated with variations in the presented area and shape of the deck structure, can influence wind loading and heeling moments, and that there are significant differences at low frequencies between wind gust spectra used in design.

It is not clear, however, whether wind gust loading has any significant influence on stability. There are no known and documented instances where dynamic wind loading has been linked to stability problems.

5.4 EFFECTS OF MOORING LINES, RISERS AND THRUSTERS

Several regulatory authorities state that moorings should be ignored when carrying out a stability analysis. The unit is then regarded as free-floating, and the wind heeling moment is calculated about the centre of lateral resistance. ABS specifically requires certain detrimental effects of moorings to be considered, however, and the HSE would also expect such effects to be taken into account nowadays under the Safety Case regulations. Moorings can affect stability adversely if the mooring line fairleads are lower than the centre of lateral resistance, increasing the effective wind heeling moment.

Vertical loads from mooring lines may affect the righting moment. DNV and the CCG specifically require vertical components of the mooring and riser force to be taken into account when calculating the unit’s displacement and KG values.

Nishimoto et al. [66] reported the results from an extensive study of the influence of mooring lines and risers on the area ratio. They carried out an integrated mooring and stability analysis with the unit in both its intact and damaged conditions, combined with failure of each of the two upwind mooring lines.

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The analysis also considered variations in parameters including the fairlead height, water depth and line length. The main conclusions from this study were as follows:

• Raising the fairlead height increased the area ratio, but resulting variations in the maximum allowable KG were small.

• The area ratio showed some sensitivity to water depth, and between intact and failed line conditions, but it was difficult to draw clear conclusions.

• In certain ranges of fairlead height and line length, they found that the area ratio was less than obtained when the unit was free-floating.

It is not clear how specific the conclusions from this study are to the particular unit and mooring configuration under investigation, and the probabilities of various line and vessel failure modes are not considered. Nishimoto et al. nonetheless successfully demonstrated that an integrated mooring and stability analysis is possible, and that mooring lines can have a detrimental effect on stability. As noted by the authors, this type of approach might have a significant role (possibly as part of a risk analysis study), when a semi-submersible is used as a floating production system.

Takarada et al. [39] also showed that mooring lines can have a significant effect on the stability of a semi-submersible. They also showed that current forces can make a significant contribution to the total heeling moment when the unit is moored.

The MOPS model tests [21] demonstrated that major damage in survival conditions could cause mooring lines to fail. These particular tests found that tripping of the mooring lines had no significant influence on the dynamics of the platform.

In summary: a number of research studies have found that mooring lines can have a detrimental effect on stability in certain circumstances, and conversely that vessel damage can lead to failure of mooring lines. These results suggest that it would be prudent for regulatory authorities to require detrimental effects of mooring lines to be taken into account explicitly in stability analyses, and that a combined risk analysis involving loss of buoyancy and mooring line failure may sometimes be justified.

Certain authorities (HSE and NMD) require the effects of thrusters to be taken into account, with the thruster force acting at the level of the thrusters. HSE also state that any difference between the wind force and the maximum thruster force should be assumed to act at the centre of lateral resistance. Thrusters lower the effective centre of lateral resistance of the vessel, thus increasing the wind heeling moment. This effect could be significant in certain circumstances.

5.5 EFFECT OF ICING

Certain authorities (e.g. the IMO and CCG) require the effects of ice and snow to be taken into account where units are liable to operate in such conditions.

The effects of icing were assessed during the MOPS project [21]. Numerical models were developed to represent the cumulative effects of frozen sea spray and wet snow, with input from Norwegian meteorological observations. The results were regarded as approximate, but showed that ice loads can have a significant effect on the unit’s displacement and centre of gravity, and should be taken into account in a stability analysis. The importance of icing clearly depends on the unit’s location, and on the use of anti-icing or de-icing equipment. Continuous monitoring of the unit’s metacentric height and draught would allow instantaneous effects of icing to be quantified.

5.6 STEADY TILT ANGLE AND SUBHARMONIC MOTIONS

Model tests performed during the 1970s and early 1980s [24, 39, 67, 68, 69, 71] showed that semi-submersible units may sometimes adopt a steady angle of tilt when subjected to regular waves of an appropriate amplitude and frequency. If this phenomenon were to occur on actual units in the sea, then it might cause the upper deck to be struck by waves, mooring difficulties, and possible capsize. Miller

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[70] reported that this phenomenon seemed to be highly tuned, however, and only occurred near one particular regular wavelength for any given model.

Research showed that the steady tilt phenomenon could be explained by vertical second-order wave forces acting on the two lower hulls [24, 71, 72]. In certain circumstances these forces caused a de-stabilising moment, increasing any initial tilt caused by the wind, current or mooring lines.

Takaki and Higo [73] also observed low-frequency subharmonic rolling motions during model tests on a semi-submersible in regular waves. These motions were related to the steady tilt angle of the unit, increasing as the steady tilt angle increased. Similar motions were observed in numerical simulations, but did not always occur in exactly the same conditions because of differences in the viscous damping. These motions occurred when the wave period was close to half the natural roll period of the unit, and were attributed to a combination of:

• periodic variations in the righting moment arm caused by heave motions,

• variations in the mooring reaction force,

• hydrodynamic forces acting on the asymmetric hull form when the unit had a large list angle,

• the difference between vertical hydrodynamic forces on the two lower hulls.

The first of these mechanisms (a form of parametric instability) has also been investigated by Rainey [74].

Takaki and Higo [73, 75] demonstrated that the steady tilt angle, and the occurrence of low-frequency roll and heave motions, could all be reduced by attaching the mooring lines to a device which allowed the mooring fairlead height to vary. These particular model tests and numerical simulations were performed in irregular waves, and demonstrated that steady tilt and low-frequency motions could occur in irregular wave conditions.

They also considered that this device could generate a larger righting moment than could be obtained by using ballast water.

Takaki and Higo did not state the metacentric height of their model. Evidence from tests and numerical simulations elsewhere [24, 39, 71, 72] have shown that the steady tilt phenomenon occurred only when the metacentric height, GM, was small.

The HSE requires that the metacentric height, GM, must not be less than 1m in normal operating, survival and transit conditions, and this lower limit is in line with recommendations made during the PRESS project [11]. Morrall [71] nonetheless recommended, on the basis of model tests and numerical calculations, that GM should be at least 2m in order to avoid a steady tilt or occasional long-period roll motion occurring in certain sea conditions.

Despite the high level of research interest in steady tilt and low-frequency motions, there is little documented evidence of such problems occurring in practice. There is also strong evidence that such problems are unlikely provided the unit meets certain minimum GM requirements.

5.7 WATERTIGHT AND WEATHERTIGHT INTEGRITY

An earlier review of stability issues for jack-ups [2] found that progressive flooding in storms, following damage to vent pipes, covers and similar openings, had been major factors in many incidents, leading to eventual loss. Damage of this type and consequent loss of watertight integrity seem to be rare on semi-submersibles, however, because of their substantially higher freeboard.

Loss of watertight integrity is nonetheless implicated in a number of incidents reviewed in Section 7. In most of these cases the damage has been attributed to a single event, such as a snagged anchor fluke, or a valve or hatch left open, rather than storm damage, and the flooding was confined to a single compartment. This is the type of damage envisaged by the standards, and the problem was easily rectified without endangering the unit.

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Concern was expressed about flooding through non-weathertight hawse pipes and chain lockers, following the Ocean Ranger accident, where flooding of the chain lockers was considered to be an important factor leading to its final capsize [4]. The HSE’s Guidance includes a specific requirement that chain locker openings in column-stabilised units should be fitted with weathertight and remotely operable closing appliances. The possibility of failure of such devices should, of course, be considered.

Most current regulations, framed after the Ocean Ranger accident, include extensive guidance on the provision of weathertight and watertight closing appliances, and there is no evidence that watertight integrity is a major issue of concern for modern semi-submersible units.

5.8 LOSS OF BUOYANCY FOLLOWING A BLOW-OUT

A review of the WOAD accidents database (see Section 7) revealed two incidents associated with blow-outs. The first occurred in 1985 when the West Vanguard encountered an unexpected pocket of shallow gas. The gas ignited, and an explosion was followed by a fire on board the unit. The unit developed a list as a result of a leak into a leg. The loss of buoyancy in this case seems to have been a direct consequence of structural damage and flooding.

The second incident occurred in 1993 when the Actinia also encountered a pocket of shallow gas. There was no explosion or fire, but, according to the WOAD report, ‘the rig rolled over and settled with a 15 degrees list with its BOP stack damaged’. The cause of the list is not made clear, but the report mentions ‘a circular brown patch almost two kilometres wide’. The nature of this brown patch is also not clear, but it probably consisted of mud from the sea-bed mixed with gas. The unit’s list may have been caused by either the strong flow velocities or the reduced density of the water/gas mix in this region.

Concern about the possible effects of a gas bubble plume on the stability of a vessel moored directly above led Moros and Dand [76] to undertake numerical simulation studies of both the gas plume itself and the behaviour of a vessel moored above the plume. The plume took the form of a strong, narrow jet with a maximum velocity of about 6 m/s. The simulations indicated that the worst conditions, from the point of view of vessel stability, would occur at high flow rates in deep water. In shallow water the gas reached the surface very quickly, and did not have time to diffuse very far laterally. The dynamic effects of turbulent fluctuations in the water were considered to be important for the vessel’s stability. Strong currents caused by the blow-out tended to move the vessel, breaking mooring lines during some of the simulations. The paper notes, however, that the lines in these cases were both stiff and short because the water was shallow. The simulations were performed using a ship-shaped drilling vessel, and there was no hint of the vessel sinking. The authors concluded that the density difference did not significantly affect the stability of the vessel. They noted, however, that the simulations did not take account of water taken on board as a result of upwelling through the moonpool, and they considered that this might change the stability behaviour of the vessel.

This investigation suggested that a ship-shaped vessel caught in a gas plume would suffer little direct loss of buoyancy or stability. It is possible that loss of buoyancy of one leg of a semi-submersible might have a more serious effect, although fire and explosion are the most obvious hazards. Dynamic loads on the unit, caused by strong local velocities and turbulence in the water jet, and upwelling through the moonpool might also cause hazards.

Swan and Moros [77] subsequently investigated the two-phase flow associated with an underwater gas release, and compared results from experiments with both theoretical and numerical predictions. High radial velocities occurred at the water surface.

5.9 VARIATIONS IN METACENTRIC HEIGHT DURING OPERATION

In order to avoid the occurrence of tilt, low-frequency motions, and the possibility of abrupt changes in the unit’s heel and trim, it is important to maintain a minimum value of the metacentric height during

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service. The metacentric height is sensitive to changes in the distribution of mass and ballast on the platform, and may also be sensitive to the contents of tanks containing liquids.

Regulatory authorities (including the HSE) therefore require vessel operators to carry out inclining tests at the beginning of the vessel’s life, and during each major survey. They also require operators to maintain an accurate record of all changes that affect the vessel’s weight or centre of gravity, and there are provisions for further surveys and inclining tests after such changes.

After the Alexander L. Kielland accident, the Norwegian Maritime Directorate re-inclined 20 column stabilised units [16]. The results showed serious discrepancies, mostly in the unfavourable direction. These reductions in GM suggested that unrecorded items had gradually accumulated on the rigs, or that initial inclining tests had not been properly executed. The Department of Energy subsequently required all units operating in the UK sector of the North Sea to be re-inclined.

Increases in lightship weight are not unique to semi-submersible units, but they are especially sensitive to such changes, because the weight tends to be added at main deck level, significantly above the unit’s centre of gravity, and the waterplane area is relatively small.

The accuracy of results from an inclining test will depend on the following:

• the equipment used,

• the accuracy of the observations,

• the prevalent weather conditions.

Even if the inclining test has been performed competently and with high accuracy, there is no guarantee that GM will remain known accurately between tests. According to Bradley and MacFarlane [78], the lightship weight typically accounts for less than 50% of a vessel’s total displacement, and the final result depends heavily on the accuracy with which weight records are maintained on board, and are used subsequently to update estimated GM values. Furthermore, no matter how diligent and experienced the crew may be, the deadweight records will always contain certain assumptions and estimates, with no means of checking entries in the record once they have been made.

Bradley and MacFarlane [78] pointed out that it is relatively easy to control the vessel’s displacement, and this is often known to a high level of accuracy. The displacement is not a sufficient parameter for checking the stability condition of the unit, however, and there is little correlation between having the correct displacement and a correct VCG.

Lewis [79] made an early attempt to quantify the errors and uncertainties associated with a ship design. He estimated that a well conducted inclining test could determine the metacentric height of the ‘as inclined’ vessel to an accuracy of around 2% (an optimistic assumption, according to Bradley and MacFarlane [78] ), but that subsequent calculations, based on lightship and deadweight estimates, would be no better than 11%. This increased error came mainly from omissions and inaccuracies in the vessel’s deadweight.

A review of methods for measuring the stability of semi-submersibles [80] found that natural roll and pitch periods could be estimated from motions in moderate sea conditions, using spectral analysis techniques, with an accuracy of about 3%. The metacentric height could then be inferred from the natural period and an estimate of the vessel’s moment of inertia, with an error estimated to be in the range 5 to 10%. This review also considered errors in estimated values of the metacentric height caused by a mooring system With a relatively low mooring stiffness the error in GM remained less than 10%, whereas with a stiffer system the error was between 10% and 20%. Accuracy might be improved by monitoring the mooring line tensions.

Naess [81] also commented on the difficulties of maintaining accurate records of weights loaded onto and off a unit, or else re-positioned. He estimated that the error in estimating the metacentric height from measurements of the unit’s natural roll and pitch periods would be less than 10% in moderate sea conditions, provided the unit’s motions are linear. He noted that this technique could not be used at transit draught, when non-linear effects are significant.

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A number of procedures and systems for monitoring the metacentric height of semi-submersibles have since been developed. Römeling and Jacobsen [82] described the early development and trials of one such procedure, which used parameter identification techniques to estimate the metacentric height. This procedure was tested in the laboratory and then on a large semi-submersible at sea. Estimates of the metacentric height of the laboratory model rapidly converged to a stable value, which was within 1% of the correct value. Estimates obtained from the full-scale measurements did not converge so rapidly or satisfactorily, however, and their accuracy was not known because of uncertainties in the unit’s condition.

Bradley and MacFarlane [83] subsequently discussed some of the difficulties of using conventional parameter and system identification procedures to obtain estimates of the metacentric height from measurements of the vessel’s motions alone, with no knowledge of the wave forcing function. The analysis either requires estimates to be made of the vessel’s moments of inertia and added inertia, which are subject to considerable uncertainty, or else the analysis has to be based on an assumed ‘cost function’, the physical basis of which is uncertain. They put forward a ‘model optimisation’ procedure, claiming that this has advantages over alternative methods when the input forcing function is unknown. This procedure depends on assumed relationships between one response of the system and another. Few details of the procedure are given, but the authors describe the in-service application of this procedure to determine the metacentric height of two semi-submersible units at sea. In one case corrections had to be made for mooring forces. This procedure was considered to be at least as accurate as a conventional inclining experiment (one benchmark against which it was tested), and appeared to provide accurate estimates of the metacentric height while the unit was in service.

This procedure was subsequently developed into the MOSIS system. Bradley and MacFarlane [78] described operating experience with MOSIS, and drew conclusions about weight and KG control procedures on board semi-submersible units. At that stage MOSIS had been installed on 12 rigs. The authors claimed that MOSIS always estimated the same value of the righting arm as was obtained from an inclining experiment. On one occasion the system revealed inaccuracies in an earlier inclining test.

Comparisons with in-service records showed that operators were generally maintaining accurate records of the total displacement of semi-submersible units, within an accuracy of about ±100 tonnes. Estimates of the KG value were more variable. There was a positive bias of about 0.4m, with KG, on average, higher than expected. The number of positive offsets was also significantly larger than was expected from a simple random process. This bias was believed to point to a consistent and serious underestimation of the centre of gravity height by certain operators. A semi-submersible has a relatively small waterplane area, and its the draught is therefore sensitive to changes in weight, which are therefore readily detectable. Unlike weight changes, vertical movements of the position of the centre of gravity are not directly observable, and errors may persist for long periods undetected.

5.10 STABILITY MANAGEMENT

The literature and the accident statistics reveal several examples of errors in maintaining accurate deadweight records, or serious errors in the use of the ballast control system. Of the 36 incidents listed in Appendix B, flooding was a major factor in about 19 incidents, and human error (either in operating the ballast control system, or errors of omission) appears to be a significant factors in at least five. Two of these incidents, the Ocean Developer and the Ocean Ranger, resulted in loss of the unit, and the second in major loss of life.

The loss of the Ocean Developer and the Ocean Ranger [4, 5, 6] illustrate the importance of having clear operating procedures and a well-trained crew. The Ocean Ranger accident occurred because of an initial ballast control system failure, exacerbated by errors made by a poorly trained crew in operating an obscure and poorly-understood manual over-ride system. These accidents illustrate the relatively delicate nature of stability control on semi-submersibles, and the fact that stability can be lost through error or incompetence in a matter of minutes.

Well-designed, reliable, automated monitoring systems [84, 85] can make it easier to maintain accurate deadweight records, and can provide immediate estimates of the vessel’s weight, vertical centre of gravity position and metacentric height, by monitoring the weight and position of the deck loads, the

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position of the cranes and tank levels. If proper attention is given to the way in which information is presented to the operator, and to the design of the system as a whole, these tools can make a real contribution to maintaining overall safety.

Computerised weight management systems can also assist with training and planning. They may, for example, have an off-line simulation mode, in which the operator can become familiar with ballasting techniques, hydraulic network topology, the way in which the vessel will respond to loading/ unloading and crane operations. They may also be used to analyse future loading operations. Before undertaking a crane manoeuvre, for example, a simulation of the operation will show which ballast compartments have to be filled in order to control the unit’s trim and heel.

These systems are only effective, however, if used correctly and carefully, if suitable systems are in place to ensure quality control and checking, and if there are adequate and well-understood back-up procedures in case the primary system fails. Maintaining accurate control over the weight and stability of a semi-submersible requires a high level of competence and diligence from the workforce, and a high level of commitment from the management. This requires a broad management commitment to, and understanding of, safety and training needs. Bradley and MacFarlane [78] commented that good operators will manage their rigs well, regardless of the regulatory regime, whereas bad operators may continue to operate unsafely.

Section 31.5 of the HSE’s Guidance [1] contains extensive and detailed recommendations on ballast control systems for column-stabilised units, and Section 31.7 provides guidance on operating procedures. A review of these sections of the Guidance was outside the scope of this study, but no major inadequacies have been brought to the authors’ attention. These requirements, and those introduced by other regulatory authorities in the wake of the Ocean Ranger accident, led to significant improvements in the design and operation of ballast control systems. Key improvements included the introduction of FMEA methods, greater redundancy in valves, pumps, power sources and control equipment, better operator information and safety precautions, and improved ability to right the unit when heeled. The RABL project [7] estimated that the risk of ballast system failure on so-called ‘first’

-3 -4and ‘second’ generation units was 10 and 5�10 respectively, with ‘third’ generation (post Ocean

-5Ranger) units having a risk of 10 . The lower risk for third generation units was attributed largely to having ballast pumps located at both ends of each pontoon.

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6. ALTERNATIVE USES OF SEMI-SUBMERSIBLES

In common with most other regulatory authorities, the HSE’s stability standards make no distinction between different operational modes of a semi-submersible: for example, whether it is used for drilling, production, diving, pipe-laying or accommodation. This common approach to stability issues may be justified on the grounds that capsize is a catastrophic event, which endangers the lives of those on board, regardless of the purpose of the unit or the operation being undertaken. Certain operations may nonetheless have special features which influence the exposure risk, for example, or the time required to evacuate. Certain operations are also associated with greater risks of pollution, fire or explosion, where there may be a need to consider the combined risks to the system as a whole.

6.1 DRILLING UNITS

In the past semi-submersibles have been used mainly as mobile offshore drilling units (MODUs), and the regulatory requirements generally reflect the way in which drilling vessels operate. Such vessels have to be mobile, and may be required to work world-wide, and not at any specific location. A MODU is able to undergo regular special surveys and tests, either in dry-dock or in sheltered water at a light draught.

Gueguen [86] pointed out that drilling units differ in one respect from semi-submersibles used for accommodation, production or other purposes. Part of the load on a drilling unit (cement and mud) may be dumped into the sea in an emergency. Certain authorities (e.g. NMD, HSE) note that dumping may be taken into account, subject to environmental regulations, when deballasting from operating to survival draught.

6.2 FLOATING PRODUCTION UNITS

One authority (ABS) has a separate code [87] specifically for floating production, storage and offloading system. The sections on stability simply refer to the appropriate sections of the MODU rules, however, indicating that in this respect ABS treats floating production units in the same way as a conventional drilling unit.

An unpublished review study for the HSE suggested that it would be reasonable to expect more stringent standards for semi-submersibles used as floating production units than are presently accepted for drilling units, in order to achieve the same overall level of risk. When used for floating production, the unit has to stay at one location for perhaps 20 years, and may never be brought inshore or deballasted to a light transit draft for inspection and survey. Inclining tests may also be impractical in offshore conditions. Accurate weight control and monitoring of the centre of gravity position are therefore matters of crucial importance for semi-submersibles used for floating production. Automatic stability monitoring systems may have an important role here, and have been installed on a number of fields, but have not yet become fully accepted by the authorities. This means that each application has to be considered and approved individually.

A floating production system will be exposed to a single operating environment for its entire life. This environment may range from benign to severe, depending on the location. Several regulatory authorities allow operators to use location-specific environmental criteria, but few provide detailed guidance on how to select these criteria. In particular, it is not obvious how to select an appropriate return period or averaging period when calculating the wind speed.

Bowie and Richardson [88] suggested that wind criteria based on 100-year return period conditions are acceptable for the intact unit, provided the wind speed is not less than 50 knots. The ABS alternative criteria [47] are based on 100-year return period wind and wave conditions in the Gulf of Mexico and North Sea. The averaging period associated with the wind speed is not stated, but hourly mean values seem to be implied. One has to question, however, whether an hourly mean wind speed with a 100-year return period provides a sufficiently low level of risk.

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Conventional stability standards may be interpreted as modelling the effects of a sustained wind as well as a single severe wind gust, which is of sufficient duration to cause the unit to roll to its maximum extent. This implies that a gust speed with an averaging period of about one minute might be appropriate for the analysis. The wind speed associated with such a gust might be 15 to 20% higher than the hourly mean value (see Table 11.4 of [1] ).

The choice of return period may also be questioned. The choice of a 100-year return period wind speed -2

does not necessarily imply that the risk of capsize is as high as 10 per year. This high level of risk would obviously be unacceptable. Risk depends on many other factors, however, including the safety margins implicit in the standards, and these factors are often difficult to quantify.

The Norwegian Petroleum Directorate [65] recommends that structural loads in an abnormal event situation should be based on environmental loads with an annual exceedance probability of 10-4 when

-2the structure is intact, and 10 when damaged. 10,000-year return values of wind speed and wave height are used increasingly in design in circumstances where the consequences of overload will be catastrophic. By analogy, one might consider using the 10,000-year return value of the one-minute mean wind speed in an intact stability analysis, and the corresponding 100-year wind speed in a damaged analysis. Considerable work would have to be done, however, to place the choice of environmental criteria, return periods and safety factors on a sufficiently firm basis for a quantitative risk analysis.

Bowie and Richardson [88] noted that existing semi-submersibles converted to floating production systems may have difficulty in complying with the latest MODU stability criteria, when in the damaged condition. ABS therefore decided that the Rules in effect at the time when the unit was built and classed may be used, provided a safety management system or equivalent can ensure similar safety as would be obtained using the latest MODU Rules. This approach is in line with current safety management practice, but obviously places major responsibilities on the operator to ensure that there is strict compliance with the safety management system.

Most existing semi-submersible production units have little on-board storage, but larger units are now being built or proposed [8, 9], and these are likely to carry significant quantities of produced oil. A conversion of the DB101 derrick barge [89] was expected to provide capacity for up to 130,000 barrels, and the MPSS [8] was designed to hold up to 500,000 barrels. On-board storage may pose special stability problems because of the need to consider a wide range of tank contents, and the possibility that a storm might occur while the unit is fully loaded.

The volume of hydrocarbons is likely to be significantly higher on a production unit than on a drilling unit, and its long-term presence could also increase local pollution risks. Dyer [90] recommended that a probabilistic approach should be adopted to meet both the marine pollution and damaged stability requirements for FPSOs. Dyer’s paper outlines a probabilistic methodology, based on the risk of multiple compartment damage. The calculations involved in carrying out such an analysis are substantial, requiring many more damage cases to be considered than at present, and it is not clear whether such complex and extensive calculations are realistic.

An unpublished review study for the HSE noted that ship impact damage to ballast tanks used for oil storage would have different consequences from damage to water ballast tanks, because of the different density of oil. The damage zones of drilling and floating production units are also likely to be different, and additional piping and drains may penetrate bulkheads and decks. These issues have implications for a stability assessment.

It was also noted that floating production systems are non-standard products, tailored to a specific field and application, and have to satisfy rules for both production vessels and mobile units. The fact that one set of rules was developed for fixed installations, and the other for drilling units, makes it inevitable that interpretations have to be made during the certification and classification process, which can make the process more difficult, costly and open to inconsistency.

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6.3 ACCOMMODATION UNITS

Stability standards for drilling and accommodation units are generally the same. A common approach seems reasonable because both types of unit are likely to be moved from one location to another, and should generally satisfy similar worldwide environmental criteria.

If an accommodation unit is moored in close proximity to another floating structure or fixed platform, due consideration has to be given to any mooring lines or walkways linking them. Certain regulatory authorities specifically require detrimental effects of mooring lines to be taken into account. NMD also requires the unit to satisfy damaged standards in survival conditions if the two structures are connected by a gangway and cannot pull away. HSE would normally expect such problems to be addressed as part of a Safety Case, although there are in fact no semi-submersibles of this type currently on the UKCS which cannot pull away.

A further consideration is the number of personnel onboard. A typical drilling unit is likely to have no more than 100 people onboard, whereas an accommodation unit may have several times that number. The number of personnel on-board should not be an issue in deciding what is an acceptable risk of capsize, but may affect the time taken to evacuate the platform in the event of an accident.

6.4 DIVING SUPPORT VESSELS

There are no obvious special requirements for diving support vessels, other than the need to take lifting loads and thruster systems into account when assessing stability.

6.5 PIPELAY VESSELS

Semi-submersibles have been used for many years as pipelay vessels, and the stability standards are no different from those applying to drilling units. In certain circumstances it may be appropriate to take account of tensions in the pipe, in much the same way as loads in moorings and risers are currently included in stability analyses for drilling units, where these items have an adverse effect on the unit’s stability.

6.6 CRANE VESSELS

The outreach and position of the crane, and the maximum load it will carry, have a significant effect of the unit’s stability, and need to be considered when carrying out the stability analysis, setting limiting KG values, and planning lifting operations. The effects of dynamic ballasting on the unit’s stability should also be considered, although there are no specific stability requirements for such systems. The crane will also have to be suitably stowed and secured in severe weather. The stability of a crane vessel should otherwise be treated in essentially the same manner as the stability of a drilling unit.

6.7 OTHER APPLICATIONS

Patel and Walker [91] described a novel type of semi-submersible in which the main buoyancy columns were articulated at the base. The purpose, so the authors claimed, was to achieve the high metacentric heights required for a large deck payload without the penalty of structure weight. Patel and Walker proposed a possible analysis approach for a unit of this type. As far as is known, however, no such structure has been built.

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7. REVIEW OF PAST INCIDENTS

Two databases were consulted in order to obtain information about past incidents involving loss of stability or flooding of semi-submersibles:

• the Worldwide Offshore Accident Databank (WOAD), operated by DNV Technica, which holds details of over 2000 offshore accidents worldwide,

• the HSE’s own in-house database of incidents in the UK sector.

Keyword searches were undertaken to find incidents involving semi-submersibles in accidents involving either capsize, flooding, leakage or listing. Both searches revealed useful information about a significant number of incidents (27 from the WOAD database, and 9 from the HSE database). It is interesting that there was no overlap between the incidents obtained from the two databases. The objectives and types of incidents recorded in these two databases are, of course, significantly different: the information in the WOAD database is reported voluntarily and is primarily concerned with accidents to installations, whereas accidents in UK waters endangering personnel have to be reported to the HSE. Some overlap between the two databases might nonetheless have been expected, and this leads one to suspect that other incidents worldwide may not have been recorded in either database. Details of all the incidents discovered during these searches are contained in Appendix B.

Appendix B shows the wind speeds and wave heights recorded in the WOAD database. These values do not always agree with values published elsewhere, and it is not clear whether they represent significant wave heights and mean wind speeds, or maximum values. They should therefore be treated with some caution.

The losses of the Alexander L. Kielland in 1980 and the Ocean Ranger in 1982 were major landmarks in the development of stability standards for semi-submersibles. Major research investigations followed these two losses, resulting in a major re-think of stability standards. Details of these two losses and of subsequent enquiry findings have been widely published elsewhere. Brief summaries are nonetheless presented again in Appendix C for readers unfamiliar with these accidents, because of their importance in the context of this review study. A succinct description of both accidents may also be found in a paper by Mills et al. [16].

Both accidents resulted in major loss of life, and both occurred in severe storm conditions. Inability to launch lifeboats successfully from the listing rig (in the case of the Alexander L. Kielland), and inability to rescue crew members from lifeboats which had been launched successfully (in the case of the Ocean Ranger) were important factors contributing to this loss of life.

The UK Department of Energy [92] investigated the loss of Transocean III east of Orkney in January 1974. The unit was newly built and of an unusual design. It was lost in bad but not extreme weather. The primary cause of loss was structural failure of one of the four main support columns. Inadequate training of key personnel was identified as an important factor. A ring of wedges was intended to transmit bending movements between the columns and cross girder. The investigators found that these wedges were inadequate for the task, the rig operators did not fully appreciate the significance of these movements and their effect on the structure, and the barge engineers had not been instructed on the need to keep these wedges firmly in position. The wedges began to move under wave action, then started to deform and break up. This allowed additional movement of the leg, resulting in damage to the locking pins and local plating, and further movement of the wedges, which eventually caused the leg to break away. The unit capsized, but had already been evacuated without loss of life or injury.

Very little information could be found about any of the other incidents summarised in Appendix B, despite carrying out literature searches. In most cases the only available source of information was the brief report contained in the relevant accident database. Very few of these incidents resulted in loss of life.

An earlier review of jack-up stability incidents [2] revealed a typical, but fairly complex, sequence of events leading eventually to capsize and loss. The semi-submersible incident reports indicated a relatively simple sequence, however, with a small number of different types of initiating events. Key

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events in each incident are summarised in Table 7.1. The most complex sequence of events was probably that leading to the loss of the Ocean Ranger.

Initiating events may be classed as follows:

• ballasting and equipment failures, leading to inadvertent flooding (e.g. the Ocean Developer, Ocean Alliance, Ocean Nomad, Diamond M Epoch, New Era and other incidents). Errors in operating the ballast control system were also a factor in the loss of the Ocean Ranger;

• storm damage, leading directly to major structural failure (e.g. the losses of the Alexander L. Kielland and Transocean III);

• storm damage (e.g. the Ocean Ranger and Ocean Bounty incidents), which led to flooding;

• blowout (e.g. the Actinia and West Vanguard incidents), which led to damage, fire and/or explosion;

• operational errors during anchor handling and mooring (e.g. two almost identical incidents involving the Sedco 712, and incidents involving the J.W. Mclean and Maersk Vinlander), causing minor structural damage and flooding;

• towing errors (e.g. the Safe Petrolia and Ross Isle incidents), which led to grounding and damage;

• collisions with vessels servicing the rig (e.g. the MSV Tharos incident), causing minor structural damage.

The results from this review seem to be at variance with findings of the RABL project [7], which identified collisions as the most critical safety hazard, followed by blowout, and then ballast system failure. The RABL reports were not included in the present review study, and it is therefore difficult to comment on the reasons for these differences.

This review suggested that major damage following collision is a relatively low risk event, although minor damage has occurred after collisions. Blowout is undoubtedly a major hazard, because of fire and explosion risks, but does not seem to be a major issue for capsize. Ballast system failures, errors in operating the ballasting system, and flooding due to variety of causes, including human error, seem to be the most important issues. Storm damage and towing errors also appear to be significant factors.

The number of ballasting faults and errors is disturbing, and also the fact that many flooding incidents are unexplained. In the case of the Ocean Ranger and Ocean Developer ballast control errors and failures were major factors leading to total loss. In most cases, however, the flooding was noticed at an early stage and was confined to one compartment, so that ballast pumps were effective.

The level of damage that occurred in most incidents was consistent with levels envisaged by conventional damaged stability standards, and generally involved flooding of a single compartment. The losses of the Alexander L. Kielland and Transocean III occurred after major structural failure during storms. It is important to recognise, however, that the key issues in both cases were ones of structural failure rather than loss of stability. Both units were also of unusual designs.

The Alexander L Kielland accident enquiry [3] noted that established damaged stability rules do not envisage major structural failure, such as loss of an entire column, and consider damage/ flooding in no more than two compartments. Submergence of the deck was not considered during the original stability analysis of this rig. The enquiry specifically recommended a re-evaluation of the damaged stability standards, leading to requirements for reserve buoyancy. Improved systems for weight control during operation were also recommended. These recommendations led subsequently to the MOPS and RABL research projects [7, 21], and NMD requirements for reserve buoyancy discussed in Section 5.2.4.

The enquiry also made a number of recommendations and comments relating to structural design, noting that the structure and moorings should be designed to provide adequate safety against failure, in order to limit the direct consequences of that failure. Lack of structural redundancy in the bracing of the Alexander L. Kielland, which caused rapid overload of the five remaining bracing members, seems to have been a key design fault, which received relatively little attention in the official report.

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The Alexander L. Kielland enquiry made a number of recommendations. These included: better identification of hazards which may cause injury, loss of life or economic loss, use of risk-based procedures, a more unified approach to safety standards, and a complete safety evaluation at the planning stage. Many of these issues emerged again during the Ocean Ranger and Piper Alpha accident enquiries.

The Alexander L. Kielland enquiry report paid particular attention to deficiencies in safety training, lack of available life-jackets and survival suits, and difficulties encountered during the launch of the lifeboats and life-rafts. The standby safety vessel was designated to serve a number of installations on the Ekofisk Field, and did not reach the scene of the accident until almost an hour later. These issues were significant factors in the loss of life, but are outside the scope of the present study.

The Royal Commission into the Ocean Ranger accident identified 12 key factors contributing to the loss. These included the location of the ballast control room, failure to protect the room and control panel against ingress of water, lack of sufficient training and understanding by the crew of the way in which the ballast system operated, lack of well-understood alternative means of controlling the ballast valves, lack of protection for the chain lockers, and failure of watertight integrity of the upper hull.

One consequence of the Ocean Ranger accident has been greater recognition of the importance of chain locker flooding. Martinovich and Praught [17] noted that certain authorities allow chain lockers to be assumed flooded in an intact stability analysis. Some units now have devices attached to the hawse pipes in order to provide effective means of closure and weathertight integrity up to the second intercept angle, but the designer should then consider the possibility of failure of the closing device.

Another consequence has been recognition of the need to provide effective de-ballasting arrangements, especially when the rig is trimmed or heeled. The HSE’s Guidance contains extensive and detailed requirements on such issues, including a section specifically relating to ballast systems for column stabilised units. The ballast system also has to be designed and constructed so that, in the event of failure of any single component, the remainder of the system continues to be capable of effective operation.

The conclusions from the Alexander L. Kielland and Ocean Ranger accident enquiries received considerable attention from the industry and from regulatory authorities, and the recommendations have largely been acted upon. Both accidents resulted in further major research investigations, both in Canada and worldwide, and international reviews of stability standards for semi-submersibles.

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Table 7.1 Summary of events associated with each reported incident

Unit Date Event sequence

Cecil Provine July 1998 Co, Sd

Dan Princess December 1997 St, Sd, Fl

J. W. Mclean December 1997 An, Sd

Sedco 712 November 1996 An, Sd, Fl

Ocean Developer August 1995 Ba, Ca

Sedco 712 June 1995 An, Sd, Fl, Li

MSV Tharos March 1993 St, Co, Sd

Actinia February 1993 Bl, Sd, Li

Sonat Arcade Frontier June 1992 Fl

Ocean Alliance February 1992 Ba, Li

Ocean Bounty October 1991 St, Sd, Fl

Maersk Vinlander October 1991 An, Sd

Transocean Discoverer May 1991 Fl

J. W. Mclean January 1991 Sd, Fl

Treasure Prospect December 1990 Fl

Bow Drill 2 November 1986 Fl

Sedco J April 1989 Ca

Diamond M Century March 1987 Fl

Ocean Nomad November 1986 Ba, Li

West Vision August 1986 St, Co, Gr, Sd

West Vanguard October 1985 Bl, Ex, Sd, Li

Safe Petrolia August 1983 Gr, Sd

Diamond M Epoch March 1983 Ba, Li

Safe Felicia March 1983 Fl, Li, Gr

Ross Isle February 1983 Gr, Sd

Ocean Ranger February 1982 St, Ba, Li, Fl, Ca

Henrik Ibsen April 1980 Fl, Li

Alexander L. Kielland March 1980 St, Sd, Li, Fl, Ca

New Era June 1977 Ba, Fl, Li

Santa Fe Mariner 2 May 1977 Fl

Santa Fe Mariner 2 August 1976 Fl

General Enrique Mosconi January 1976 St, Sd

Ocean Queen September 1975 St, Li

Louisiana May 1975 Fl, Li

Sedco 702 June 1974 Fl

Transocean III January 1974 St, Sd, Fl, Ca

Key

An damage by anchor

Ba ballasting fault/ error

Bl blow-out

Ca unit capsized/ lost

Co collision

Ex fire and explosion

Fl flooding

Gr grounding

Li list developed

Sd structural damage

St storm

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8. DISCUSSION OF KEY ISSUES

From the foregoing discussion it will be clear that there is no easy way in which to assess the relative margins of stability, and the uniformity and consistency provided by different stability standards. The amount of analysis and detail required to make such a judgement go well beyond the scope of this review, and no such judgement will be attempted.

The stability standards contained in the HSE’s Fourth Edition Guidance are based around established concepts and procedures. Research undertaken in the aftermath of the Alexander L. Kielland and Ocean Ranger accidents led to the development of various alternative standards, and in particular new criteria, put forward by the ABS, based on numerical simulation work and physical model tests on a number of generic units. These new criteria have been criticised for a variety of reasons, but many of these criticisms may also be levelled equally against established stability standards. These criticisms have not prevented established standards from being used successfully in design over many years.

A key difficulty with the new criteria is that they sometimes provide a lower margin of stability than established standards, and it is difficult to know whether this reduction should be regarded as an erosion of unnecessary conservatism or an erosion of necessary safety margins. There is no means of knowing the absolute levels of safety provided either by established techniques or by more novel methods. Established methods have been tried and tested by experience, whereas the success or otherwise of new standards will only become clear after many years of use.

8.1 INTACT STABILITY

Established intact stability standards seem to have resulted in safe designs. No known losses of semi-submersibles have occurred while the unit was in the intact condition and satisfied standard criteria, and there seems to be no major reason for departing from these established standards.

The intact standards of all the authorities considered are based on established 100-knot extreme storm and 70-knot transit or operational wind speed conditions, and on a 1.3 minimum area ratio requirement. The standards differ in a number of other respects (e.g. the maximum static heel angle, stability range, minimum GM and GZ requirements), but it is not clear how these variations affect stability margins or levels of uniformity and consistency between different units.

These differences may cause some inconvenience to operators, however, in cases where specific units do not comply with a particular authority’s requirements.

Ambiguities in the definition of the critical heel angle and between ‘free trim’ and ‘free twist’ methods of stability analysis have been criticised. Practice varies between different organisations and programs, and there is no simple right or wrong way in which to define these parameters or perform the analysis. These differences generally seem to be well understood by more experienced analysts. It is nonetheless important that key procedures and definitions should be understood clearly by all engaged in such work.

Tripping by leeward mooring lines and mooring line weight can sometimes have an adverse effect on intact stability. Existing HSE standards state that mooring lines should be disregarded in a stability analysis. It would nonetheless seem prudent to consider detrimental effects of moorings, risers, thrusters and similar items explicitly in stability analyses.

Several investigators have reported differences between procedures for estimating wind heeling moments. Certain recommended calculation procedures either ignore or represent inadequately the lift-induced moment, and the moment may be either over or under-estimated as a result. The lift-induced moment on a semi-submersible deck is likely to be significant, especially when the unit is heeled, and should be taken into account. Wind tunnel tests seem to offer the only reliable means of obtaining wind heeling moments at present.

The lack of any explicit recognition of dynamic effects when calculating the first downflooding angle has been questioned on several occasions. The HSE follows established practice in taking no account

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of rig or wave motions when calculating the downflooding angle for an intact analysis, but requires a 4m allowance for wave effects to be made in a damaged analysis. The logic behind this is not clear, considering that more severe sea conditions have to be considered in an intact analysis.

Certain successful designs utilise the buoyancy of the upper deck in order to meet the 1.3 area ratio criterion. Use of upper deck buoyancy conflicts with established design practice for fixed platforms, where the aim is to maintain an air gap between the deck and water surface. There would seem to be no logical objection to eliminating the air gap, however, provided structural integrity and safety requirements are met.

The upper deck will only provide buoyancy at large heel angles. If such large heel angles were to occur in practice, they would probably cause non-fixed items to move, risking major structural damage to the unit, or causing a shift in its centre of gravity, and compromising its stability. Large heel angles would also cause hazards to personnel, and difficulties during evacuation and escape. Although large heel angles would only occur in extreme conditions, it seems undesirable that such large angles are needed at all in order to meet basic stability standards.

Model tests have shown that units with very low metacentric heights can be stable, with no tendency to capsize. There are nonetheless good reasons for retaining the HSE’s minimum GM and GZ requirements:

• to avoid abrupt changes in heel and trim during operation;

• large heel angles are unlikely to occur if the unit provides sufficient righting moment at small heel angles;

• steady tilt angles and low-frequency rolling motions only seem to have been observed in model tests where GM was low.

The ABS/ IMO alternative intact criteria are intended to avoid both capsize and downflooding. They were based on an extensive programme of numerical simulations and model tests, but the margins of safety provided by these criteria are not well-established, and the use of these procedures should be approached with caution. A number of other issues, such as the adequacy of the ABS empirical formulae in representing resonance effects, may require further investigation. It is also disturbing that the ABS/ IMO alternative criteria ignore the overall shape of the wind heeling moment curve, and only consider the first and second intercept angles. It is in principle possible to satisfy these criteria with a very low area ratio and almost no reserve energy.

Although wind gust loading might in principle affect a unit’s stability, there seem to be no reported instances where dynamic wind loading has been linked to stability problems.

Ice loading may affect stability, and should be taken into account at appropriate locations.

In circumstances where location-specific environmental criteria are appropriate, careful consideration should be given to the averaging and return periods used when estimating the wind speed.

Existing regulations already include extensive guidance on the provision of weathertight and watertight closing appliances, and there is no evidence that watertight integrity is a major issue of concern for modern semi-submersible units. The risk of failure of such devices should be addressed, however.

8.2 DAMAGED STABILITY

Damaged stability standards for semi-submersible units have been in a considerable state of flux for several years. At least two approaches have emerged among the major regulatory authorities (one based on a minimum area ratio, and the other based on minimum values of the righting arm and stability range). A number of alternative or additional criteria (e.g. reserve buoyancy) have also been adopted.

There are in particular major differences between standards recommended by the HSE, and the new ABS criteria adopted in the IMO’s 1989 MODU Code and subsequently by DNV. The NMD also

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retained area ratio-based criteria, but superimposed a reserve buoyancy requirement. The IMO adopted the NMD reserve buoyancy requirement, but only as an example of an alternative to the 1989 MODU Code.

These differences are likely to cause difficulties to operators who wish to move their units from one regulatory regime to another. Discussions leading towards convergence between these different standards should therefore be encouraged.

All authorities have separate standards for peripheral waterline damage and for internal one-compartment flooding, but there are variations in the definitions of damaged areas and in the compartments to be considered.

The traditional 1.5m penetration depth seems somewhat arbitrary, and does not take account of increases in the size and power of supply vessels over the years, the size of the unit itself or its structural loading capacity. This almost universal requirement might have to be reviewed in certain circumstances.

The rationale for specifying a 4m wave clearance above the damaged waterplane, regardless of sea conditions, the vessel’s behaviour or the location and size of the downflooding point, circumstances, also seems questionable.

The use of a 50-knot wind speed in damaged stability standards has been questioned on several occasions. This speed is based on the implicit assumption that damage is most likely to be caused by a collision with another vessel, and will therefore be only slightly correlated with severe storm events. At issue, therefore, is whether the combination of circumstances that led to the Ocean Ranger accident was a ‘reasonably foreseeable’ event, and therefore one which the standards should reasonably address.

The HSE standards make no provision for reserve buoyancy in the upper deck. Certain other aspects of the HSE standards appear to be more conservative than others, but it is not clear how the overall safety levels compare. Extensive and detailed calculations on a range of different units would be needed in order to establish this.

The IMO’s 1989 MODU standards are based on results from a programme of numerical simulations and model tests undertaken by ABS. The very limited range of cases considered during the ABS JIP (numerical models of only three generic units, and model tests on only one, with fixed weights to represent damage to one leg only) is a matter of concern. The margin of stability provided by these criteria, and the resulting levels of uniformity and consistency between different units, are not well-established, and use of these new procedures should be approached with caution.

The only way to provide this additional level of confidence would be by means of additional systematic numerical simulations and model testing.

8.3 STABILITY MANAGEMENT

Accurate management of the weight and centre of gravity of a semi-submersible are crucial in maintaining its stability margin. There have been a number of major accidents and losses (e.g. the Ocean Ranger) where a poorly designed ballasting system, or errors by operating staff, were significant contributory factors.

Automatic monitoring systems are now available, and have been used successfully on a number of units. These can be useful aids to stability management in certain circumstances, especially where the unit has to remain on location for long periods, but are not a substitute for diligence in keeping accurate weight records. Periodic inclining tests can also provide a valuable check on stability, but again are no substitute for keeping accurate weight records.

Improved regulations, better design and operational practice, have reduced substantially the risks on newer rigs. This may not be true of some older rigs in certain areas of the world, however, and these rigs may require special attention when they come into UK waters.

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8.4 OTHER ISSUES

Existing stability standards do not quantify levels of safety and risk. There are moves towards a quantitative risk-based approach to design, more in keeping with the Safety Case approach, and moves towards incorporating stability assessment into this general methodology are to be encouraged. The RABL project made some initial moves in this direction, but the task of establishing a risk-based approach will be very substantial. Such procedures would also be much more complex to apply, and would probably prove to be impractical in the short term. They are nonetheless a worthwhile long-term objective.

8.5 REVIEW OF PAST INCIDENTS

This review suggested that major damage following collision is a relatively low risk event, although minor damage has occurred after collisions. Blowout is undoubtedly a major hazard, because of fire and explosion risks, but does not seem to be a major issue for capsize. Ballast system failures, errors in operating the ballasting system, and flooding due to variety of causes, including human error, seem to have been the most important factors. Storm damage and towing errors also appear to be significant factors. The number of ballasting faults and errors is disturbing, and also the fact that many flooding incidents outside the UKCS are unexplained.

The level of damage that occurred in most incidents was consistent with levels envisaged by conventional damaged stability standards, and generally involved flooding of a single compartment. The losses of the Alexander L. Kielland and Transocean III occurred after major structural failure during storms. It is important to recognise, however, that the key issues in both cases were ones of structural failure rather than loss of stability. Both units were also of unusual designs.

No overlap was found between incidents reported in the WOAD and HSE databases, raising possible concerns about the accuracy and completeness of these records. It is nonetheless recognised that the two databases serve entirely different purposes, and apply different criteria to decide whether an incident should be included.

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9. CONCLUSIONS

This review study was undertaken as part of a wider exercise to assess the need for HSE Guidance within the UK Safety Case regime. The study included a comparison between stability standards specified by key regulatory authorities and classification societies for intact and damaged semi-submersible units, a review of relevant published literature and of HSE/ Department of Energy reports, a review of past incidents involving loss of stability of semi-submersibles, and a review of issues associated with alternative uses of semi-submersible units.

Established stability standards are quasi-static in character, and do not explicitly represent the dynamic effects of waves and vessel motions. The losses of the Alexander L. Kielland and Ocean Ranger raised questions about the adequacy of these established standards, and subsequent research resulted in a range of new criteria claiming to represent dynamic effects in a more realistic manner. The established standards have nonetheless stood the test of time remarkably well, and are simple and practical to apply during design.

The main challenge now facing the HSE is how to fit established prescriptive stability standards, as defined in the Guidance, into the Safety Case regime. It would be logical to adopt a risk-based approach, but this is likely to be a long-term objective rather than one which is practical in the short term. A key recommendation coming out of this study is therefore that the HSE should investigate further the practicality of reconciling traditional prescriptive stability standards with a risk-based Safety Case approach.

9.1 INTACT STABILITY

No obvious major deficiencies have been found in established intact stability standards. These appear to have been successful in avoiding capsize and loss while units remained intact and watertight, and in accordance with the standards.

Key conclusions from the review study, relating to intact stability standards, were:

• The intact stability standards of most regulatory authorities are outwardly similar, although there are important differences in specific requirements (eg. in values of limiting angles and the minimum value of GZ). It is not obvious how these variations affect stability margins, or levels of uniformity and consistency between different units.

• Ambiguities in definitions of the most critical heel axis and between ‘free trim’ and ‘free twist’ analysis procedures have been criticised. It is important that key procedures and definitions should be understood clearly by all engaged in such work.

• Large heel angles are required by certain designs in order to meet the 1.3 area ratio requirement. Large heel angles would seem to be undesirable for several reasons, including hazards caused by large items shifting, hazards to personnel, and difficulties during evacuation and escape.

• There is a sound basis for retaining existing HSE minimum GM and GZ requirements, because these are likely to reduce the likelihood of abrupt changes in heel during operation, large angles of heel, steady tilt and low-frequency motions.

• Certain authorities (including HSE) state that moorings should be ignored in a stability analysis. It would seem prudent, however, to require detrimental effects of moorings, risers, thrusters and similar items to be considered explicitly.

• Existing standard procedures for calculating the wind heeling moment are simplified, and most such procedures do not take adequate account of the lift-induced moment. This component can be important for heeled semi-submersibles, and may be either over or under-estimated.

• In circumstances where location-specific environmental criteria are appropriate, careful consideration should be given to the averaging and return periods used when estimating the wind speed.

• The alternative IMO/ ABS intact criteria should be treated with caution until greater experience has been gained. The stability margins, levels of consistency and uniformity provided by these criteria

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are not known. It is also not known how extensively these criteria are used for semi-submersibles working on the UKCS.

9.2 DAMAGED STABILITY

There are major differences between the damaged stability standards adopted by different regulatory authorities (see Appendix A). There are in particular major differences between the approaches adopted by the HSE (based on a minimum area ratio) and in the IMO’s 1989 MODU Code (based on minimum values of the righting arm and stability range). Discussions leading towards convergence between these different standards should be encouraged.

The damaged stability standards in the IMO MODU Code are based on criteria originally proposed by ABS. The small range of units considered during the ABS JIP, and simplifications adopted during the model tests, are matters of concern. These criteria should be treated with caution until further experience has been gained, and the stability margins, levels of uniformity and consistency are better established. The only way to provide this additional level of confidence would be by means of additional systematic numerical simulations and model testing.

Key conclusions from the review study, relating to damaged stability standards, were:

• There are differences between the areas of damage and flooding specified by different authorities. The consequences of these differences are not known, and would only become clear in relation to specific designs.

• The level of damage that occurred in most incidents was consistent with levels envisaged by conventional damaged stability standards, and generally involved flooding of a single compartment. Key issues in the Alexander L. Kielland and Transocean III accidents were ones of structural failure rather than loss of stability.

• The traditional 1.5m penetration depth requirement seems somewhat arbitrary, and does not take into account increases in the size and power of supply vessels over the years, the size of the unit itself or its structural loading capacity. This almost universal requirement might have to be reviewed in certain circumstances.

• A number of authorities (including HSE) specify a minimum 4m wave clearance above the damaged waterplane, regardless of sea conditions, the vessel’s behaviour or the location and size of the downflooding point. The rationale for this 4m clearance seems questionable.

• The use of a 50 knot wind speed in damaged stability standards has been questioned on several occasions. At issue is whether the standards should reasonably address the combination of circumstances that led to the Ocean Ranger accident.

9.3 OTHER ISSUES

Moves towards quantitative risk-based stability analysis procedures, integrated within an overall safety assessment, should be encouraged as a long-term objective, although there are likely to be major difficulties in achieving this objective in the short term. Further review of the RABL work and other developments in this area would be a useful first step.

The review of past incidents suggested that there is a relatively low risk of major damage following collision, although minor damage has been recorded. Blowout is undoubtedly a major hazard, because of fire and explosion risks, but does not seem to be a major issue for capsize. Ballast system failures, errors in operating the ballasting system, and flooding due to variety of causes, including human error, seem to be the most important issues. Storm damage and towing errors are also significant factors.

The number of ballasting faults and errors is disturbing, and also the fact that many flooding incidents outside the UKCS are unexplained.

No overlap was found between incidents reported in the WOAD and HSE databases, raising possible concerns about the accuracy and completeness of these records. It is nonetheless recognised that the

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two databases serve entirely different purposes, and apply different criteria to decide whether an incident should be included.

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10. REFERENCES

1. The Health and Safety Executive, Offshore Installations: Guidance on Design, Construction and Certification, Fourth Edition with amendments, 1993.

2. BMT Offshore Limited, Stability Criteria for Jack-ups in Transit. Phase 1: Review of Casualties, Seakeeping Data and Numerical Methods, project no. P45459, 18 July 1994.

3. The Norwegian Royal Commission Report on the Alexander L. Kielland Disaster, 1981.

4. The Royal Commission on the Ocean Ranger Marine Disaster, Report no. 1, The Loss of the Semisubmersible Drill Rig Ocean Ranger and its Crew, 1984.

5. The Royal Commission on the Ocean Ranger Marine Disaster, Report no. 2, Safety Offshore Eastern Canada, 1985.

6. National Transportation Safety Board, Capsizing and Sinking of the U.S. Mobile Offshore Drilling Unit OCEAN RANGER off the East Coast of Canada, 166 Nautical Miles East of St. John’s, Newfoundland, February 15, 1982, Marine Accident Report no. NTSB-MAR-83-2, 1983.

7. Vinnem, J.E., Risk Assessment of Buoyancy Loss - Summary Report, RABL Report no. 11, 1988.

8. Miles, J.D., Lewis, G., and Jolley, H., An FPSO with Storage, Drilling and Workover th

Capability, Proc. 4 Intl. Conf. FPS’97, IIR Energy Conferences, London, 1997.

9. Semisubmersible with Storage Tank Proposed for FDPSO Operations, Offshore, pp. 72-73, May 2000.

10. Kuo, C., Vassalos, D., and Alexander, J.G., Incorporating Theoretical Advances in Usable Ship Stability Criteria, Proc. Intl. Conf. on the SAFESHIP Project, Royal Institution of Naval Architects, London, 1986.

11. Vassalos, D., Kuo, C., and Konstantopoulos, G., An Effective Stability Assessment Procedure for Semi-Submersibles Incorporating the Effects of Wind, Waves and Vessel Motions, in ‘Mobile Offshore Structures’, edited by Boswell, L.F., d’Mello, C.A., and Edwards, A.J., Elsevier Applied Science, pp. 499-518, 1988.

12. Vassalos, D., A Critical Look into the Development of Ship Stability Criteria Based on Work/Energy Balance, Trans. Royal Institution of Naval Architects, London, Vol. 128, pp. 217-234, 1986.

13. Henrickson, W.A., Assessing Intact Stability, Marine Technology, Vol. 17, no. 2, pp. 163-173, 1980.

14. Bird, H., and Morrall, A., Research Towards Realistic Stability Criteria, Proc. Intl. Conf. on the SAFESHIP Project, Royal Institution of Naval Architects, London, 1986.

15. Barltrop, N.D.P., (ed.), Floating Structures: a Guide for Design and Analysis, Centre for Marine and Petroleum Technology/ Oilfield Publications Limited, 1998.

16. Mills, P.J., Stoneman, G.S., and Wilson, T.B., Stability of Ships and Mobile Offshore Units. Recent Developments in Legislation, Lloyd’s Register Technical Association paper no. 7, 1991.

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17. Martinovich, W.M., and Praught, M.W., Stability Requirements for Semisubmersibles - a Designer’s Viewpoint, paper no. 1, Proc. Intl. Conf. on Stationing and Stability of Semi-submersibles, University of Strathclyde, Glasgow, 1986.

18. Rahola, J., The Judging of the Stability of Ships and the Determination of the Minimum Amount of Stability, doctoral thesis, Helsinki, 1939.

19. Sarchin, T.H., and Goldberg, L.L., Stability and Buoyancy Criteria for U.S. Naval Surface Ships, Trans. SNAME, Vol. 70, pp. 418-458, 1962.

20. International Maritime Organisation, Intact Stability Criteria for Passenger and Cargo Ships, 1987 edition.

21. Dahle, L.A., Mobile Platform Stability - the MOPS Project, paper no. 5, Proc. Intl. Conf. on Stationing and Stability of Semi-Submersibles, University of Strathclyde, Glasgow, 1986.

22. Kuo, C., and Vassalos, D., Semi-submersible Research and its Applications, paper no. 6, Proc. Intl. Conf. on Stationing and Stability of Semi-Submersibles, University of Strathclyde, Glasgow, 1986.

23. Miller, E.R., and Ankudinov, V., Evaluation of Current Towing Vessel Stability Criterion and Proposed Fishing Vessel Stability Criterion, Task III Report, Report no. CG-D-4-76, Dept. of Transportation, US Coast Guard, 1976.

24. Numata, E., Michel, W.H., and McClure, A.C., Assessment of Stability Requirements for Semisubmersible Units, Trans. SNAME, Vol. 84, pp. 56-74, 1976.

25. Bush, R.B., and Ahilan, R.V., History, Theory and Extension of Current Jack-up Quasi-Dynamic Stability Criteria, in ‘Mobile Offshore Structures’, edited by Boswell L.F., d’Mello C.A. and Edwards A.J., pp. 81-103, 1988.

26. American Bureau of Shipping, Guide for Application of Dynamic Response Based Intact Stability Criteria for Column-Stabilized Mobile Offshore Drilling Units, 1990.

27. Caldwell, J.B., and Yang, Y.S., Risk and Reliability Analysis Applied to Ship Capsize; a Preliminary Study, Proc. Conf. on the SAFESHIP Project: Ship Stability and Safety, Royal Institution of Naval Architects, London, 1986.

28. Kobylinski, L., Methodology of the Development of Stability Criteria on the Basis of Risk Evaluation, Proc. Fifth Intl. Conf. on Stability of Ships and Ocean Vehicles (STAB’94), Vol. 3, Florida Institute of Technology, 1994.

th29. Kobylinski, L.K., Stability Standards - Future Outlook, Proc. 7 Intl. Conf. on Stability of

Ships and Ocean Vehicles, Launceston, Tasmania, Vol. A, pp. 52-61, February 2000.

30. De Kat, J.O., The Development of Survivability Criteria Using Numerical Simulations, Proc. US Coast Guard Vessel Stability Symp. 93, New London, 1993.

31. Norwegian Maritime Directorate, Regulations for Mobile Offshore Units, 1992.

32. Transport Canada Marine Safety, Standards Respecting Mobile Offshore Drilling Units, ref. TP 6472E, undated.

33. International Maritime Organization , Code on Intact Stability for all Types of Ships Covered by IMO Instruments, Resolution A.749(18), 1995.

34. International Maritime Organisation, Code for the Construction and Equipment of Mobile Offshore Drilling Units, 1989 (1989 MODU Code), Resolution A.649(16), 1990.

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35. American Bureau of Shipping, Rules for Building and Classing Mobile Offshore Drilling Units, Part 3: Hull Construction and Equipment, 1997.

36. Det Norske Veritas, Rules for Classification of Mobile Offshore Units, Part 3, Chapter 2, Special Designs, Equipment and Stability, 1998.

37. Lloyd’s Register of Shipping, Rules and Regulations for the Classification of Mobile Offshore Units, 1996.

38. US Coast Guard, Department of Transportation, Code of Federal Regulations, document ref. 46CFR Ch. 1 (10-1-99 Edition), Special Rules Pertaining to Specific Vessel Types, Subpart C - Special Rules Pertaining to Mobile Offshore Drilling Units, paragraphs 174.030 to 174.100, 1999.

39. Takarada, N., Nakajima, T., and Inoue, R., A Phenomenon of Large Steady Tilt of a Semi-Submersible Platform in Combined Environmental Loadings, Proc. Third Intl. Conf. on Stability of Ships and Ocean Vehicles, Gdansk, pp. 225-238, 1986.

40. Bolt, H.M., Wave-in-Deck Load Calculation Methods, HSE/ E&P Forum Airgap Workshop, London, 1999.

41. Van Santen, J.A., Stability Calculations for Jack-ups and Semi-Submersibles, Proc. Intl. Conf. CADMO 86, pp. 519-550, 1986.

42. Pawlowski, M., Some Inadequacies in the Stability Rules for Floating Platforms, Department of Naval Architecture and Ocean Engineering, Report no. NAOE-87-30, University of Glasgow, 1987.

43. Chen, H.H., Shin, Y.S., and Wilson, J.L., Towards Rational Stability Criteria for rd

Semisubmersibles - a Pilot Study, 3 Intl. Conference on Stability of Ships and Ocean Vehicles, Vol. 2, Gdansk, pp. 61-68, 1986.

44. Stiansen, S.G., Shin, Y.S., and Shark, G., Development of a New Stability Criteria for Mobile Offshore Drilling Units, Offshore Technology Conf. paper no. OTC 5802, Houston, 1988.

45. Shin, Y.S., and Shark, G., A Rational Stability Criteria Format Utilizing Semi-Submersible Dynamic Responses, in ‘Mobile Offshore Structures’, edited by Boswell, L.F., d’Mello, C.A., and Edwards, A.J., Elsevier Applied Science, pp. 227-249, 1988.

46. Shark, G., Shin, Y.S., Grove, T.W., and Stiansen, S.G., Recent Developments on Residual Stability of Semisubmersible Units in Damage Conditions, Offshore Technology Conf. paper no. OTC 6123, Houston, 1989.

47 . American Bureau of Shipping, Final Report - Joint Industry Project on Mobile Offshore Drilling Unit Stability - Phase II. Volume 1 - Executive Summary, Technical Report no. RD-89014, 1989.

48. Collins, J.I., and Grove, T.W., Model Test of a Generic Semisubmersible Related to a Study Assessing Stability Criteria, Offshore Technology Conf. paper no. OTC 5801, Houston, pp. 497-503, 1988.

49. American Bureau of Shipping, Residual Stability of Semi-Submersible Units. Final Report for Special Joint Industry Project, R&D Division Technical Report no. RD-88028, 1988.

50. Song, K., Petty, T., and Bone, T., Real Time Motion Simulation of a Moored Semisubmersible rd

for Dynamic Stability, Proc. 3 Intl. Offshore and Polar Eng. Conf., Vol. III, pp. 558-567, 1993.

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51. Takaishi, Y., Ando, S., Isozaki, Y., Matsumoto, K., Okada, H., and Takai, R., A Study on the th

Stability Criteria of Semisubmersibles, Proc. 10 Intl. Conf. on Offshore Mechanics and Arctic Engineering, Vol. I-B, pp. 703-708, 1991.

52. International Maritime Organisation, Alternative Intact Stability Criteria for Column-Stabilized Semi-Submersible Units, document ref. DE 31/4/2, December 1987.

53. Stone, B.M., Sullivan, M.A., Arunchalam V.M., and Muggeridge, D.B., Model Studies of the Motion Response of a Damaged Four Column Semisubmersible in Regular and Irregular Waves, Ocean Engineering, Vol. 17, no. 3, pp. 235-261, 1990.

54. Adachi, H., and Kagemoto, H., Transient Motions of a Semisubmersible after Damages, Proc. Int. Conf. on Stationing and Stability of Semi-submersibles, University of Strathclyde, Glasgow, pp. 186-196, 1986.

55. S ylemez, M., and Incecik, A., Prediction of Large Amplitude Motions and Stability of Intact and Damaged Mobile Platforms, Offshore Technology Conf. paper no. OTC 5628, Houston, 1988.

56. American Bureau of Shipping, Residual Stability of Semi-Submersible Units. Phase I & II. Final Summary Report, November 1988.

57. International Maritime Organisation, An Example of Alternative Stability Criteria for a Range of Positive Stability after Damage or Flooding for Column-Stabilized Semisubmersible Units, Resolution no. A.651(16), 1989.

58. Miller, B.L., and Davies, M.E., Wind Loading on Offshore Structures - a Summary of Wind Tunnel Studies, National Maritime Institute Report no. NMI R136, Dept. of Energy ref. OT-R-8225, 1982.

59. Yu, B.K., and Won, Y.S., Comparison of Wind Overturning Moments on a Semisubmersible rd

Obtained by Calculation and Model Test, 3 Intl. Conf. on Stability of Ships and Ocean Vehicles, Vol. 1, Gdansk, pp. 253-268, 1986.

60. Ogiwara, R., Yamanaka, N., Tajima, E., Nema, K., and Sakata, R., Stability and Motion Characteristics of Semi-Submersible Drilling Unit SS-4000, Marintec China 85 Conference, 1985.

61. Freathy, P., and Taggart, S., Wind Loads on Floating Offshore Structures, paper no. 3, Proc. Intl. Conf. on Stationing and Stability of Semi-Submersibles, University of Strathclyde, Glasgow, 1986.

62. Singh, S., Uncertainties in the Estimation of Fluid Loading on Offshore Structures with Special Emphasis on Wind Forces, Trans. Institution of Marine Engineers, Vol. 101, part 6, pp. 269-287, 1989.

63. Rowe, S.J., Brendling, W.J., and Davies, M.E., Dynamic Wind Loading of Semi-Submersible Platforms, RINA Intl. Symposium on Development in Floating Production Systems, London, March 1984.

64. Standing, R.G., Wills, J.A.B., and Singh, S., Wind Loading and Dynamic Response of a Floating Production Platform in Waves, SUT Conference on Environmental Forces on Offshore Structures and their Prediction, 1990.

65. Norwegian Petroleum Directorate, Regulations Relating to Loadbearing Structures in the Petroleum Activities, ISBN 82-7257-467-5, 1995.

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66. Nishimoto, K., Brunozi, P.F., and Babadopulos, J.L., Analysis of Mooring Lines and Risers th

Effects on the Stability of Semisubmersibles, 10 Intl. Conf. on Offshore Mechanics and Arctic Engineering, Vol. I-B, pp. 687-693, 1991.

67. Numata, E., and McClure, A.C., Experimental Study of Stability Limits for Semi-Submersible Drilling Platforms, Offshore Technology Conf. paper no. OTC 2285, Houston, 1975.

68. Inoue, R., and Kyozuka, Y., Nonlinear Wave Forces Acting on Submerged Horizontal Cylinders, Proc. Fifth Intl. Offshore Mechanics and Arctic Engineering Conf., Vol. 1, pp. 225-234, ASME, 1986.

69. Inoue, R., An Experimental Study on Reduction of Wave-Induced Steady Heeling Moment Acting on a Lower-Hull Type Semi-Submersible, Hiratsuka Research Laboratory, undated.

70. Miller, N.S., Stability Tests on Two Designs of Semi-Submersibles in Regular Waves, OSFLAG8 report, University of Glasgow, 1977.

71. Morrall, A., The Influence of the Steady Vertical Component of Wave Force on the Stability of Semi-Submersibles, National Maritime Institute Report no. R46, Dept. of Energy ref. OT-R-7833, 1978.

72. Martin, J., and Kuo, C., Calculations for the Steady Tilt of Semi-Submersibles in Regular Waves, Trans. Royal Institution of Naval Architects, London, Vol. 121, pp. 87-101, 1979.

73. Takaki, M., and Higo, Y., A Control of an Unstable Motion of a Semisubmersible Platform th

with a Large List Angle, Proc. 4 Intl. Conf. Stability of Ships and Ocean Vehicles (STAB90), Naples, 1990.

74. Rainey, R.C.T., Parasitic Motions of Offshore Structures, Trans. Royal Institution of Naval Architects, Vol. 123, pp. 177-194, 1981.

75. Takaki, M., Higo, Y., Kohno, M., and Xiu, P., A Control Device for Stabilizing a Semisubmersible Platform with a Large List Angle in Waves, Proc. Workshop on Floating Structures and Offshore Operations, Wageningen, Elsevier, pp. 25-34, 1987.

76. Moros, T., and Dand, I., Two-Phase Flows as a Result of a Subsea Blowout and their Effect on the Stability of Structures, Offshore Technology Conf. paper no. OTC 6478, Houston, 1990.

77. Swan, C., and Moros, A., The Hydrodynamics of a Subsea Blowout, Applied Ocean Research, Vol. 15, pp. 269-280, 1993.

78. Bradley, M.S., and MacFarlane, C.J., Some Lessons to be Learned from the Stability Control of Semi-Submersibles, Offshore 95, Design and Safety Assessment for Floating Installations, Institution of Marine Engineers, pp. 121-133, 1995.

79. Lewis, E.V., Precision in Naval Architecture Calculations, Transactions SNAME, Vol. 49, pp. 122-153, 1941.

80. National Maritime Institute, A Study of Methods of Measuring the Metacentric Heights of Semi-Submersibles, Department of Energy, Offshore Technology Report no. OTH 84 211, 1985.

81. Naess, A., On a Method for Continuous Control of Platform Stability under Operating Conditions, Intl. Conf. on Marine Research, Ship Technology and Ocean Engineering, Hamburg, pp. 133-144, 1982.

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82. Römeling, J.U., and Jacobsen, B.K., At Sea Measurements for Identification of Stability, Offshore Technology Conf. paper no. OTC 4389, Houston, 1982.

83. Bradley, M.S., and MacFarlane, C.J., The In-Service Measurement of Hydrostatic Stability, Offshore Technology Conf. paper no. OTC 5345, Houston, 1986.

84. Van Santen, J.A., and Essens, P.M.J.D., New Generation of Weight and Stability Management Systems for Semisubmersibles and Jack-ups, Schip en Werf, Vol. 55, no. 3, pp. 59-60, 1988.

85. Molinari, R., Penno, C., and Rezzoagli, C., Simulation of Control Functions for Ship Stability: a User Oriented Flexible Tool, STAB 90, Stability of Ships and Ocean Vehicles, Vol. 1, pp. 26-31, 1990.

86. Gueguen, J., Safety of Semisubmersible Platforms, Alsthom Review, no. 4, pp. 51-62, 1986.

87. American Bureau of Shipping, Guide for Building and Classing Floating Production, Storage and Offloading Systems, March 1996.

88. Bowie, R.D., and Richardson, K.L., Classification and Certification Requirements for Floating Production Systems, Offshore Technology Conf. paper no. OTC 7835, Houston, pp. 281-291, 1995.

89. Brown, M.G., Basaran, M.I., and Ramzan, F.A., The Potential to Convert the “Derrick Barge 101” into a Low Cost Floating Production System, Proc. FPSO-Norge ’96, Stavanger, 1996.

90. Dyer, R.C., Regulations and Standards - MARPOL Design and Operating Requirements, E&P Forum, FPSO/FSU Workshop, Heathrow, June 1998.

91. Patel, M.H., and Walker, S., On the Hydrostatics of Floating Bodies with Articulated Appendages, Supplementary Papers, Royal Institution of Naval Architects, London, Vol. 125, pp. 229-236, 1983.

92. Department of Energy, Report on the Loss of the Drilling Barge Transocean III, Petroleum Production Inspectorate, 1975.

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ABBREVIATIONS AND NOTATION

ABS American Bureau of Shipping ALARP as low as is reasonably practical API American Petroleum Institute BMT BMT Fluid Mechanics Limited CCG Canadian Coast Guard COGLA Canada Oil and Gas Lands Administration DEn UK Department of Energy DNV Det Norske Veritas FMEA failure modes and effects analysis FPSO floating production, storage and offloading unit GM metacentric height GZ hydrostatic righting arm Hs significant wave height HSE UK Health and Safety Executive IMO International Maritime Organisation JIP Joint Industry Project KG height of the centre of gravity above keel level MODU mobile offshore drilling unit MOPS Mobile Platform Stability project MOU mobile offshore unit NMD Norwegian Maritime Directorate NPD Norwegian Petroleum Directorate OSD Offshore Safety Division (of the HSE) PRESS Performance Related Efficient Semi-submersible Stability project RABL Risk Assessment of Buoyancy Loss project UK United Kingdom UKCS UK Continental Shelf US United States USCG US Coast Guard VCG vertical location of the centre of gravity

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APPENDIX A: SUMMARY OF STABILITY STANDARDS

The following stability standards are summarised in the tables below:

Health and Safety Executive (HSE) [1],Norwegian Maritime Directorate (NMD) [31],Canadian Coast Guard (CCG) [32],International Maritime Organisation (IMO): intact [33] and damaged [34],American Bureau of Shipping (ABS): standard [35] and alternative [26],Det Norske Veritas (DNV) [36].

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Table A1: Wind conditions specified for both intact and damaged stability standards

Authority

Requirement

HSE 4th Edition NMD CCG IMO ABS DNV

Wind speeds. In operational 70 knots (36 m/s). 70 knots 70 knots 70 knots 70 knots 70 knots conditions. (36 m/s). (36 m/s). (36 m/s). (36 m/s). (36 m/s).

In transit condition. 70 knots (36 m/s). 70 knots 70 knots 70 knots 70 knots 70 knots (36 m/s). (36 m/s). (36 m/s). (36 m/s). (36 m/s).

Survival condition. 100 knots (51.5 m/s). 100 knots 100 knots 100 knots 100 knots 100 knots (51.5 m/s). (51.5 m/s). (51.5 m/s). (51.5 m/s). (51.5 m/s).

Sheltered location 50 knots (25.8 m/s) 50 knots 50 knots 50 knots 50 knots 50 knots and damaged damaged condition only. (25.8 m/s). (25.8 m/s). (25.8 m/s). (25.8 m/s). (25.8 m/s). condition.

Methods and Method for Calculation procedure Either wind tunnel Calculation procedure Calculation Calculation Calculation directions. determining wind

heeling moment. defined. May be supplemented by wind tunnel tests.

tests or calculations. defined. Wind tunnel tests may be considered as an alternative.

procedure defined. Wind tunnel tests may be considered as an alternative.

procedure defined. Wind tunnel tests may be considered as an alternative.

procedure defined. Wind tunnel tests may be considered as an alternative.

Wind and moment Wind from any direction, Wind from any Any direction which Wind heeling and Wind from any Wind heeling and directions. in order to find the one for direction, the vessel gives the most critical righting moment direction in order to righting moment

which the maximum KG being able to trim axis (defined as giving curves need to be find the most curves need to be has its lowest value. freely during

inclination. the lowest KG value). The unit shall be free to

related to the most critical axis (critical

critical (critical direction not

related to the most critical axis (critical

trim when heeled. axis not defined) defined). axis not defined)

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Table A2: Other requirements for both intact and damaged stability standards

Authority

Requirement

HSE 4th Edition NMD CCG IMO ABS DNV

Other Effect of mooring Mooring lines to be Disregard the The displacement To be disregarded. To be disregarded, The displacement and requirements. lines or thrusters. disregarded. Thruster mooring lines, but and KG should although detrimental KG should take

forces to be applied at thruster level, and any

for DP take into account the

take account of mooring line

effects are to be considered.

account of mooring line weight.

difference between the reactive wind weight. Moorings are maximum wind force and thrust at the centre of lateral

force operating at thruster level.

Moorings are otherwise to be

otherwise to be disregarded.

resistance. disregarded.

Minimum Watertight integrity is to be Watertight closing Weathertight Any opening below Any opening below Watertight closing downflooding maintained in compartments means to be fittings to be the static heeled intact the static heeled means are required at angle, watertight required to comply with provided up to the fitted to openings and damaged damaged waterline to least up to an angle of and weathertight intact and damaged stability static intact and leading to spaces waterlines to be made be made watertight, heel equal to the first integrity. standards. Any opening

below the static heeled damaged heeled waterlines, and

required to meet stability

watertight. Openings to chain lockers or

and openings less than 4m above and 7o

intercept in the intact or damaged condition,

intact and damaged weathertight requirements. other buoyant volumes beyond this waterline whichever is the waterlines is to be made closing means up Special to be considered as to be made greater. Weathertight watertight. External to angles consideration will downflooding points. weathertight. closing means are openings which become associated with be given to Stability parameters to Stability parameters required up to an submerged in meeting area stability openings which be calculated up to the to be calculated up to angle of heel equal to ratio requirements must be requirements. cannot be closed lesser of the second the lesser of the the dynamic angle. weathertight. Any opening The righting in emergencies. intercept and the angle second intercept and This also applies to less than 4m above any moment curve is of downflooding the angle of any opening within damaged waterplane (after to be corrected for (intact) or weathertight downflooding (intact) 4m above the final wind heel) should meet the any progressive integrity (damaged). or weathertight waterline. weathertight standard. flooding. integrity (damaged).

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Table A3: Intact stability standards

Authority

Requirement

HSE 4th Edition NMD CCG IMO ABS DNV

Intact stability standards.

Ratio of the area under the righting moment curve to the

� 1.3 � 1.3 � 1.3 � 1.3 � 1.3 � 1.3

area under the wind heeling moment curve.

Static angle of heel due to wind.

� 15° � 17° � 15° Not specified. Not specified. Not specified.

Angle of heel at the second intercept.

Not specified. � 30° Not specified. Not specified. Not specified. Not specified.

Minimum GM. � 1.0 m � 1.0 m � 1.0 m Not specified. � 0.0 m. � 1.0 m

operating, transit and survival operating, transit operating, transit operating, transit and conditions; and survival and survival survival conditions;

� 0.3 m conditions; conditions;

� 0.3 m

intermediate temporary condition.

� 0.3 m

intermediate

� 0.3 m

intermediate all temporary conditions.

temporary condition. temporary condition.

Minimum righting moment or GZ.

Up to the angle of downflooding or maximum righting lever, or up to 15°,

Positive over range from upright to second intercept.

Not specified. Positive over range from upright to second intercept.

Positive over range from upright to second intercept.

Positive over range from upright to second intercept.

whichever is least: GZ � 0.5 GM0 sin �, where GM0 is the minimum permissible GM.

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Table A4: Location and extent of assumed peripheral damage

Authority HSE 4th NMD CCG IMO ABS DNV

Edition Requirement

Location of Number of Any Any Any compartment Any compartment within an Any compartment within an Any compartment assumed compartments compartment compartment within an assumed zone assumed zone of damage. assumed zone of damage. within an assumed zone damage or assumed to be within an within an of damage. of damage. flooding. flooded. assumed zone of assumed zone

damage. of damage

Extent of Circumferencial 3 m along the 3 m along the 3 m along the periphery. 1/8 th column perimeter. Not specified, although 1/8

th 3 m along the periphery.

assumed extent. periphery. periphery column perimeter is implied damage. by the bulkhead damage

requirement.

Vertical extent. 3 m within 5m 3 m within 5m Unspecified exposed 3 m within 5m above and 3 m within 5m above and 3 m within 5m above above and 3m below the specified draught.

above and 3m below the specified draught

portions of columns on the periphery, and lower hulls and footings in light or transit draught conditions.

3m below the specified draught.

3m below the specified draught.

working draught, and 3m below the specified draught.

Horizontal penetration.

1.5 m. 1.5 m. 1.5 m. 1.5 m. 1.5 m. 1.5 m.

Bulkheads. Not specified. At least one Not specified. No vertical bulkhead is to No vertical bulkhead is to Not specified. watertight be assumed damaged, be assumed damaged, bulkhead may be assumed to

except when the bulkhead separation is less than 1/8

th except when the bulkhead separation is lower than 1/8

th

be damaged. of the column perimeter; of the column perimeter; then the bulkhead should be then the bulkhead should be considered as damaged. considered as damaged.

Watertight flat Not specified. Not specified. Assume damaged if the Assume damaged if within Assume damaged if within Assume damaged if the surfaces. flat is located within the specified zone. the specified zone. flat is located within the

1.5m above or below region above. the waterline.

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Table A5: Stability standards for peripheral damage

Authority

Requirement

HSE 4th Edition NMD CCG IMO ABS DNV

Damaged Ratio of righting moment Not specified. Not specified. Not specified. � 2 � 2 � 2 stability to wind heeling moment at standards. some angle within the

range of positive stability.

Zone of weathertight � 4 m above any � 0 m above the final damaged � 0 m above the � 4 m above the � 4 m above and � 4 m above the integrity. damaged waterline. final damaged final damaged 7

o beyond the final damaged

waterplane. waterline. waterline. final damaged waterline. waterline.

Static angle without wind. � 15°. Not specified. �15° 2 . Not specified. Not specified. Not specified.

Static angle with wind. Not specified. � 17° Not specified. � 17° Not specified. � 17°

Range of positive stability. Not specified. � 7° � 7° � 7°

Ratio of the area under the � 1.0 � 1.0 � 1.0 Not specified. Not specified. Not specified. righting moment curve to the area under the wind heeling moment curve.

Additional requirements for reserve buoyancy.

Not specified. Positive range for the righting arm curve of at least 10° between the

Not specified. Not specified. Not specified. Not specified.

first and second intercept, in the damaged condition. GZ should be at least 2.5m within this range, and at least 1.0m of this GZ should arise from watertight volume above the lowest deck level.

2 It is not clear whether this requirement means with or without wind.

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Table A6: Location and standards for one-compartment damage

Authority

Requirement

HSE 4th Edition NMD CCG IMO ABS DNV

Location of Number of Any one watertight Any compartment Not specified. Any one watertight Any one watertight Any watertight assumed compartments compartment. under or partially compartment wholly compartment wholly or compartment wholly or damage or flooding.

assumed to be flooded from causes

under the waterline, or any which is

or partially below the waterline, which is a

partially below the waterline, which is a

partially below the waterline, which is a

other than peripheral damage.

limited by the sea or contains piping

pump-room, a room containing machinery

pump-room, a room containing machinery

pump-room, a room containing machinery

connected to the sea with salt water with salt water cooling with salt water cooling cooling system, or a system, or a system, or a compartment adjacent to the sea.

compartment adjacent to the sea.

compartment adjacent to the sea.

Standards for Static angle without Same requirements Same requirements Not specified. � 25° Same requirements as � 25° flooding from wind. as for peripheral as for peripheral for peripheral waterline causes other waterline damage 3 . waterline damage 3 . damage 3 . than peripheral damage.

Range of positive � 7° � 7° stability to the lesser of the extent of weathertight integrity and the second intercept.

3 See Table A5.

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Table A7: Alternative stability standards

Authority Requirement

HSE 4th Edition NMD CCG IMO ABS DNV

Alternative General. Not mentioned. Alternative Other standards of Alternative criteria Alternative criteria Alternative criteria stability standards.

criteria may be considered,

stability, and methods for proving

may be considered, provided an

may be considered, provided adequate

may be considered, provided an

provided an equivalent level

that stability criteria are met, may be

equivalent level of safety is maintained.

righting moments are maintained, with

equivalent level of safety is

of safety is maintained.

considered, provided an

sufficient margins to preclude

maintained.

equivalent level of safety is maintained.

downflooding and capsize in intact and damaged conditions.

Intact stability. Not mentioned. Not mentioned. Not mentioned. The booklet gives an Refers to the ABS Not mentioned. example, which may Guide on Alternative be applied to a twin Intact Stability hull semi- Criteria for Column submersible. Separate Stabilized Units. capsize and downflooding criteria are specified.

Damaged stability. Not mentioned. Not mentioned. Not mentioned. Resolution A.651(16): Not mentioned. Not mentioned. an example of alternative damaged stability criteria, applicable to semi-submersible units which have buoyant volumes contained in a watertight upper-deck structure.

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Table A8: KG limit curve requirements

Authority Requirement

HSE 4th Edition NMD CCG IMO ABS DNV

KG limit curve requirements.

Severe storm condition.

Curve of maximum KG for compliance with minimum GM and intact stability criteria for a wind speed of not less than 100 knots.

Similar to HSE, but damaged stability criteria with a wind speed of 70 knots apply if a vessel is alongside, or if other hazardous activities.

Curve of maximum KG for compliance with minimum GM and intact stability criteria for a wind speed of not less than 100 knots.

Not specified. Not specified. Similar to HSE.

Additional requirements for intermediate draughts.

Additional requirements for intermediate draughts.

Additional requirements for intermediate draughts.

Operating A curve of maximum KG A curve of maximum KG for Maximum KG values Not specified. Not specified. Similar to conditions. for compliance with the compliance with the most according to the intact CCG.

most critical of the critical of the following : stability criteria with a wind following :

• Intact stability criteria with a wind speed not

• Intact stability criteria with a wind speed not less than 70 knots.

speed not less than 70 knots and damaged stability criteria with 50 knots wind.

less than 70 knots. • Compliance with the

• Compliance with the damaged stability criteria. damaged stability criteria. • Maintenance of minimum

GM. • Maintenance of

minimum GM.

Temporary Relate to intact stability Maximum KG should comply Maximum KG values Not specified. Not specified. Similar to condition. with 70 knots wind speed with intact stability criteria according to the intact CCG, with

and relaxed GM, and to damaged stability with

with a wind speed not less than 70 knots. Damaged

stability criteria with 70 knot wind.

minimum GM = 0.3m.

single compartment criteria with a wind speed of flooding, but ignoring the 15° limit on heel after

70 knots also apply if a vessel is alongside.

flooding.

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Table A9: Further KG limit curve requirements

Authority Requirement

HSE 4th Edition NMD CCG IMO ABS DNV

KG limit curve Range of conditions. Complete range from light to To cover the To cover the Maximum KG to Maximum To cover the requirements. maximum draughts. Curves relevant range of relevant range of be defined in allowable KG relevant range of

corresponding to: the severe intact operating, intact and damaged relation to draught versus draft curves intact and damaged storm condition; the most transit, survival and operating and or other or equivalent, and operating and transit critical of intact operating and temporary transit conditions, parameters, and associated conditions, intact transit, damaged and minimum GM conditions; and intermediate temporary conditions.

draughts, and separate damaged stability curves for a range of draughts corresponding to damage conditions.

intact temporary and intact severe storm conditions.

based upon compliance with intact and damaged criteria.

limitations or assumptions upon which the allowable KG is based.

temporary and intact survival conditions.

Free surface effects. A correction for free surface effects should not be included

A correction for free surface effects

A correction for free surface effects

Not specified. Not specified A correction for free surface effects

in the KG limit curves, but should be taken into account

should not be included in the KG

should not be included in the KG

should not be included in the KG

in onboard calculations. limit curves, but should be taken

limit curves, but should be taken

limit curves, but should be taken into

into account in into account in account in onboard onboard onboard calculations. calculations. calculations.

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APPENDIX B: SUMMARY OF INCIDENT INFORMATION

Unit/ Location/ Date/ Data source

Weather conditions Brief description of incident Key events

Cecil Provine Wind: not known. A supply vessel hit a water well, causing a small dent. Collision caused minor damage. UK Waves: not known. July 1998 HSE database

Dan Princess Wind: not known. A slow leak into the supporting column void was discovered. The rig had Fender plates lost in a storm. UK Waves: not known. taken on-board approximately 120 tonnes of water following earlier storm Column assumed damaged by loose December 1997 damage. The storm had caused loss of boat fender plates from starboard plate. HSE database columns, and it was assumed that the damage to the column was caused by Flooding of column.

a fender plate while it was partially detached.

J. W. Mclean Wind: 8 m/s. While retrieving an anchor to its bolster, the anchor had to be lowered again Loose anchor fluke punctured the hull. UK Waves: 3m. in order to turn it. While the anchor was hanging below the bolster the December 1997 flukes came into contact with the hull, holing it in two places. HSE database

Sedco 712 Wind: not known. During anchor handling operations the chaser pennant suddenly pulled Hatch cover assumed damaged by chaser UK November 1996

Waves: not known. tight, and then went slack. Almost immediately the rig developed approximately 3� degrees of trim. A ballast tank was found to be flooded,

pennant. Hatch cover torn open.

HSE database and action was taken to restore the trim. An ROV inspection revealed that Ballast tank flooded. a hatch on the ballast tank had been torn open, with the hatch cover bent upwards. Subsequent inspection of the chaser pennant showed that marine growth had been stripped off, but there was no damage to is construction.

Ocean Developer No information Under tow in order to be converted into an FPS. An enquiry suggested that Ballast control error. Off Northern Angola an inexperienced crew member may have caused the accident when Total loss of rig. August 1995 operating the complex ballasting system. The rig sank in 10,800 feet of WOAD database water.

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Unit/ Location/ Date/ Data source

Weather conditions Brief description of incident Key events

Sedco 712 UK

Wind: not known. Waves: not known.

During anchor handling operations the rig suddenly developed 2� degrees of trim. A ballast tank was found to be flooded, and action was taken to

Hatch cover torn off (no cause established).

June 1995 restore the trim. An ROV inspection revealed that a hatch on the ballast Ballast tank flooded. HSE database tank had been torn off. Subsequent inspection of the chain chaser pennant Rig developed a list.

showed no damage to indicate that it had been around the hatch cover, and the master of the anchor handling vessel reported seeing nothing untoward in tension monitoring instruments.

MSV Tharos Claymore Field, UK

Wind: 20 m/s Waves: 7m

The standby/safety/pollution control vessel Cam Sentinel collided with the semi-submersible, which was providing accommodation support to the

Collision caused structural damage.

March 1993 Claymore Platform. The semi-submersible sustained a 3-inch rupture WOAD database above the waterline on one of its eight columns. The collision occurred

after the Cam Sentinel lost power in stormy seas with 40-knot wind and 7m waves, and started drifting. The damage was temporarily repaired on location.

Actinia Wind: none. The rig struck shallow gas, and a blow-out occurred, forming a ‘circular Blowout. Offshore Vietnam Waves: none. brown patch almost 2 km wide, and bubbling up almost 40m high’ 4. The Major list developed. February 1993 rig ‘rolled over’ and settled with a 15 degree list, with its BOP stack WOAD database damaged.

Sonat Arcade Frontier Wind: none. Following an alarm in the starboard pump room, the crew found that the Unexplained flooding in pump room. Offshore Norway Waves: none. room was flooded. The area was isolated by closing all sea chest valves, June 1992 and the water was pumped out using the emergency bilge pump. It WOAD database appeared that a previously blocked/closed hydraulic valve isolating one of

the sea water pumps had been opened, allowing water to enter the room through this pump, which had been partially disassembled for repairs.

Ocean Alliance Wind: not known. A ballast valve failure caused an unintended heel and trim to develop. Ballast valve failure. UK Waves: not known. Heel and trim developed. February 1992 HSE database

As reported in the WOAD database, although it is not clear how this description should be interpreted.

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Unit/ Location/ Date/ Data source

Weather conditions Brief description of incident Key events

Ocean Bounty Wind: 35 m/s. The rig suffered extensive damage in very bad weather. An abnormally Structural damage in a storm. Cormorant Field, UK Waves: 15m. large wave (estimated to be about 30m high) struck the side of the unit, Lower hull punctured. October 1991 causing damage to the anchor winch house, surrounding decks and a WOAD database lifeboat. The lower hull propulsion room was punctured by falling debris,

causing slight water ingress. The ballast pumps coped satisfactorily.

Maersk Vinlander Wind: not known. The rig was under tow between locations, and was ballasting up from 16m Loose anchor fluke punctured a ballast UK Waves: not known. to 10m draught. A loose anchor on the bolster was moved by wave action tank. October 1991 in the splash zone. An anchor fluke is believed to have punctured a port HSE database ballast tank. The tank was kept empty by the ballast pump.

Transocean Discoverer Wind: not known. Sea water ingress at high pressure caused the bilge to fill. The ballast Unexplained flooding in propulsion UK Waves: not known. control system failed safe, closing on the starboard side. The depth of room. May 1991 water in the propulsion room stabilised at about 4 feet, and this was HSE database pumped out using the emergency ballast control system.

J. W. Mclean Wind: not known. A cast iron cover and sea water strainer cracked, allowing flooding. Bilge Failure of cover and sea water strainer. UK Waves: not known. pumps controlled the flooding, but two fuel pumps and a portable auto- Flooding. January 1991 ballast control were damaged. Emergency ballast control was necessary. HSE database

Treasure Prospect Wind: not known. The rig was found to be taking on water while under tow from the Gulf of Unexplained flooding of pump room. Under tow: Gulf of Waves: not known. Mexico to the North Sea. Flooding was limited to a pump room. The US Mexico to North Sea Coast Guard subsequently reported that the rig had ‘internal problems’, December 1990 which had been rectified by the crew. WOAD database

Bow Drill 2 Wind: not known. The port side pump room was flooded to a depth of 4m, submerging all Unexplained flooding of pump room. Offshore China Waves: not known. equipment. November 1986 WOAD database

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Unit/ Location/ Date/ Data source

Weather conditions Brief description of incident Key events

Sedco J Wind: not known. The rig was being towed from Port Gentil to Karachi, via Cape Town, for Cause not known. Offshore Durban Waves: not known. scrapping, when it capsized and sank 500 miles east of Durban. Capsized. April 1989 Total loss of rig. WOAD database

Diamond M Century Wind: none. The rig was moored at Mobile, Alabama. A wooden plug was improperly Plug incorrectly installed. Mobile, Alabama Waves: none. installed by a diver in the port ballast system pipe, which caused the port Thruster room flooded. March 1987 thruster room to flood rapidly when a 12-inch sea suction valve was opened WOAD database for inspection. Only slight damage was reported.

Ocean Nomad Wind: 20 m/s. A malfunction of the ballast control system caused the rig to list 9 degrees Ballast control failure. UK North Sea Waves: 5m. before control was obtained, and the rig uprighted after 90 minutes. Strong Rig listed. November 1986 gale-force winds and 5m waves were reported at the time. WOAD database

West Vision Wind: 23 m/s. The rig broke its moorings during Typhoon Vera, contacting the berth Rig broke moorings in a storm. Daewoo Harbour, Korea Waves: not known. heavily. It then drifted into a bulk carrier under construction before it Collision and grounding. August 1986 grounded. The pontoons were holed, and thrusters were damaged. Pontoons holed. WOAD database

West Vanguard Wind: not known. A blowout occurred after the rig struck a pocket of shallow gas. A BOP Blow-out. Offshore Norway Waves: not known. was not installed, and the diverter system could not withstand the flow. Explosion and fire damage. October 1985 WOAD database

Gas spread around the platform, and first ignited around the cellar deck, followed by a larger explosion ten minutes later, which engulfed the

Leak into a leg. List developed.

platform in a large fireball. The emergency fire-fighting system functioned, One crew member killed. but was unable to put out the blaze. The explosion and fire destroyed part of the deck structure, two legs and the engine room, and killed one of the crew. A 10-degree list occurred as a result of a leak into a leg.

Safe Petrolia Wind: not known. The rig struck bottom while under tow near Bomlo. A leg was punctured. Towing error. Bomlo, Norway Waves: not known. The tow continued after adjustment of ballast. Rig struck bottom. August 1983 Leg punctured. WOAD database Leak developed.

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Unit/ Location/ Date/ Data source

Weather conditions Brief description of incident Key events

Diamond M Epoch Wind: none. A faulty valve caused too much water to enter a leg, and the rig developed a Faulty valve. Bass Strait, Australia Waves: none. 4 degree list. Leg flooded. March 1983 List developed. WOAD database

Safe Felicia Wind: not known. Both forward thruster rooms were flooded, causing the rig to list. The rig Unexplained flooding of thruster rooms. Gothenburg, Sweden Waves: not known. then rested on the river bottom. Rig listed. March 1983 Grounded on bottom. WOAD database

Ross Isle Wind: 1 m/s. Two propellers were damaged, and one pontoon was punctured, after the Towing error. Tönsberg, Norway Waves: none. rig hit the bottom of the fjord near Tönsberg, while under tow. Rig struck bottom. February 1983 Pontoon punctured. WOAD database Leak developed.

Ocean Ranger Wind: 37 m/s. Heavy weather resulted in breakage of ballast control room window(s). Storm damage to ballast control room Hibernia Field, Offshore Waves: 15m. Water caused the ballast control panel to malfunction. Several ballast windows. Canada valves opened, allowing water to enter the forward ballast tanks. A list Ballast control panel malfunction. February 1982 developed, and the chain lockers were continuously flooded. The WOAD Ballast tanks flooded. WOAD database database report then states that the rig overturned, although official reports

pointed to a complex sequence of events (see Appendix C). An List developed. Chain lockers flooded.

unsuccessful attempt was made to save the crew from damaged lifeboats. Crew errors in correcting ballast control system. Rig overturned and was lost. Entire 84-man crew lost.

Henrik Ibsen Wind: none. The rig developed a 20 degree list while trimming for inspection. The Manhole left open. Stavanger, Norway Waves: none. accident was caused by an open manhole, which allowed water into a Column flooded. April 1980 column. Major list developed. WOAD database

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Unit/ Location/ Date/ Data source

Weather conditions Brief description of incident Key events

Alexander L. Kielland Ekofisk Field, Offshore

Wind: 27 m/s. Waves: 8m.

The rig had already been moved away from the Edda Platform because of bad weather. Bracing on one leg failed, and the entire leg failed soon

Fatigue damage during a storm. Structural failure of bracing.

Norway afterwards. The rig rapidly listed 30 to 35 degrees. After 20 minutes the Structural failure of a leg. March 1980 rig capsized. The loss of the leg was subsequently attributed to fatigue A heavy list developed. WOAD database failure in a bracing member, following the installation of a hydrophone, The rig capsized, and was lost.

which had been welded into it. The platform had no redundancy against 123 of the 212 people on board lost their loss of this type of bracing. The large list allowed lifeboats to be launched lives. on one side of the rig only, and three crashed against the platform during launch. 89 out of 212 men were rescued. See also Appendix C.

New Era Wind: none. Failure of a ballast room pump and ballast control system, during trimming Ballast pump and control system failure. Gulf of Mexico Waves: none. operations, caused the rig to list. The rig listed. June 1977 WOAD database

Santa Fe Mariner 2 Wind: not known. Ingress of water through a non-watertight door. No other information Water ingress. Gulf of Mexico Waves: not known. available. May 1977 WOAD database

Santa Fe Mariner 2 Wind: not known. Ingress of water through the engine room exhaust vent. Water ingress. Gulf of Mexico Waves: not known. August 1976 WOAD database

General Enrique Mosconi Wind: 19 m/s. The rig broke its moorings in heavy weather. Damage was caused by Moorings failed in heavy weather. Dunkirk, France Waves: 4m. flooding of machinery inside the flotation feet. The rig ‘rested overside of Flooding of internal spaces. January 1976 harbour’, and was subsequently re-floated successfully. WOAD database

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Unit/ Location/ Date/ Data source

Weather conditions Brief description of incident Key events

Ocean Queen Wind: 30 m/s. The rig listed slightly after being struck by hurricane Eloise. Hurricane damage. Gulf of Mexico Waves: 10m. Slight list developed. September 1975 WOAD database

Louisiana Wind: none. Several compartments were flooded during a transfer operation, causing the Unexplained flooding of compartments. Offshore Trinidad Waves: none. rig to list. The rig listed. May 1975 WOAD database

Sedco 702 Wind: none. The pump room was flooded during the annual inspection and overhaul of Unexplained flooding of pump room. Offshore Netherlands Waves: none. thrusters. June 1974 WOAD database

Transocean III Wind: 21 m/s. Transocean III was a special self-elevating semi-submersible design, which Structural failure, probably due to East of Orkney, UK Waves: 6m. had been modified to operate in the North Sea. Most of the crew had been weather. January 1974 evacuated after the rig developed a ‘fault’, and the remaining crew were Jackable leg broke away. WOAD database evacuated when the weather deteriorated. A jackable leg broke away from Rig capsized.

the machinery house. The rig subsequently capsized, and later sank. Total loss of rig.

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APPENDIX C: THE ALEXANDER L. KIELLAND AND OCEAN RANGER ACCIDENTS

Most of the material in this Appendix has come from relevant accident enquiry reports, and has been publicised widely elsewhere. These details are nonetheless reproduced again here for readers unfamiliar with these two accidents, which stimulated major research investigations into stability loss, and caused a major re-think of stability standards.

C.1 THE ALEXANDER L. KIELLAND ACCIDENT

The semi-submersible accommodation platform Alexander L. Kielland capsized on the Ekofisk Field in March 1980. The Alexander L. Kielland was one of a series of Pentagone type rigs. A small number of such units may still be working, but the design is not typical of modern rigs. At the time of the accident the rig complied with relevant Norwegian Maritime Directorate and Det Norske Veritas standards.

The accident occurred during poor weather, and was initiated by a fracture, where an opening had been cut into a main horizontal bracing member, and a support for a hydrophone had been welded in. Subsequent laboratory investigations showed that fatigue fractures had developed relatively quickly until a sudden residual fracture occurred. After the failure of the lower horizontal bracing, the remaining five bracing members connected to the column failed in quick succession, because they were loaded almost immediately beyond their capacity. The column then floated up, and remained floating at an angle in the sea. The rig heeled rapidly to an angle of about 30 to 35 degrees, with the deck partly submerged, and stabilised temporarily in this position. Water started to flow into two other columns and the partly submerged deck through various openings. The rig continued to list and sink slowly for approximately 20 minutes, until the last anchor wire parted and the rig overturned.

The accident was attributed by the subsequent NMD enquiry [3] to poor design, dimensioning and material quality of the hydrophone support, as well as its connection to the bracing. Quality control during the installation of the hydrophone support failed to reveal defects in the welding.

Deficiencies in safety training, poor availability of survival equipment and problems with the lowering of lifeboats were also noted in the enquiry report. There were difficulties in releasing the lifeboats, three of which were blown against the platform and crashed. Life rafts could not be launched. The storm, the platform’s heavy list and its rapid capsize were all significant factors. Of the 212 men on board the platform at the time of the accident, 123 lost their lives.

C.2 THE OCEAN RANGER ACCIDENT

The semi-submersible drilling unit Ocean Ranger capsized and sank on the Grand Banks, off Newfoundland, Canada, in February 1982. The entire 84-man crew was lost.

The subsequent Royal Commission [4, 5] investigated the accident in considerable detail, and determined the causes and sequence of events with reasonable certainty. Unlike the Alexander L. Kielland, there was no single major initiating event in the loss of the Ocean Ranger, which occurred only after a complex sequence of events.

Weather conditions at the time of the accident were very severe. A sustained wind speed of 70 knots, with a maximum speed of 90 knots, had been forecast [4], and reports from the field that night speak of wind speeds in the 90 to 100 knot range [6]. The wave height was reported to be about 15m.

Normal practice on most rigs is to de-ballast to a lighter draught during severe weather in order to increase the air gap. It seems that the Ocean Ranger had never followed this practice, however, and

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the evidence suggested that it had not de-ballasted before the portlight failed in the ballast control room.

The initiating event was failure of this portlight between 19:45 and 20:00pm. Water entering the control room console caused a malfunction, and it was reported that all the valves on the port side were opening on their own. The control system was designed so that loss of the electric or compressed air supply to the control console would cause the valves to close automatically. The crew would have turned off the electricity after this initial failure, to ensure that the contents of the ballast tanks remained unchanged. Following the initial malfunction, there was no evidence of any change in the rig’s attitude, indicating that, even if the valves had opened by themselves, no significant amount of water entered into, or flowed between, the ballast tanks at that time.

At about 22:00pm the ballast control operator reported that ‘everything was okay’. Conversations suggested that the water and broken glass had been cleared up, the electrical power to the panel had been shut off, and a report to shore indicated that all systems were functioning normally. These reports were clearly inaccurate, and suggested a lack of appreciation of the potential danger. According to the Royal Commission report, ‘it was this ignorance which led to action which contributed to, rather than prevented, the loss of the rig.’

The Royal Commission concluded that electrical power was restored to the control panel between 00:30 and 00:45am, and it is known that the rig acquired a sudden list at the port bow. The Commission concluded that this list was caused by an ingress of sea water into the port pontoon. Whatever the reason for restoring electrical power to the control panel, it would have caused various remotely operated valves to open. Observing this list, the ballast control operator would again have removed power from the control panel. By this time, however, sufficient water would have entered the port pontoon to cause a port heel, and perhaps 4 to 5 degrees trim by the bow. This trim, together with the motion of the rig, would have been sufficient to cause a maximum forward inclination of 8 to 10 degrees, as reported to shore at 01:00am.

An off-duty ballast control operator testified that he had been told, during training that if a valve malfunctioned, it could be operated from the control room by manual insertion of a control rod into a solenoid valve. His training notes stated that insertion of the rod would cause the valve to close, whereas in fact it would open. These rods had been used to test the pneumatic system before the control panel was installed. Although they were stored in one of the panels of the control console, there were no accompanying diagrams or instructions in their use.

A subsequent diver survey showed that the crew had attempted to operate the ballast control system by the manual method, after the electrical system had failed, wrongly believing that insertion of a control rod would close a valve. Insertion of those rods would have resulted in more ballast water moving into the forward tanks, and would have increased the forward trim. The crew also failed to notice that one or more valves to the aft ballast tanks were open, and loss of ballast water as it was pumped unintentionally from these tanks would have increased the forward list.

A major design weakness then came into play. The pump rooms were located at the stern of the vessel. This made de-ballasting increasingly difficult as the rig adopted a forward trim, because of the increasing vertical distance between the suction inlet in the ballast tank, and the pump. The rate at which ballast could be pumped from the bow tanks would have reduced to zero as the trim increased. This design weakness would have seriously hampered attempts to right the rig.

The forward trim was increased further by down-flooding of large chain lockers that had not been weatherproofed. The chain lockers had been treated as down-flooding points in the earlier stability analysis. The static angle of down-flooding, based on an 80-foot draught and still water conditions, was 27 degrees. Subsequent model tests showed, however, that the rig was susceptible to down-flooding when trimmed by no more than 10 to 12 degrees at the bow.

Events developed rapidly after 01:00am. The Ocean Ranger requested its standby vessel to come to close standby at 01:05, and a Mayday was issued at 01:09. At this time the rig reported a severe port list of 12 to 15 degrees, ‘progressing’, and that all counter-measures were ineffective. Subsequent model tests showed that flooding of the chain lockers would have commenced at this point, accelerating the rate of trim. The last communication stated that the crew were going to lifeboat

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stations at 01:30am. As the forward part of the upper deck became lower, its cargo and deck fittings would have been damaged by severe waves or shifting deck cargo. It is likely that water would then have entered the damaged superstructure. The flooding would have created an unstable situation, and the rig capsized and sank at about 03:15am.

The Royal Commission report devotes considerable space to the search and rescue attempts that followed the evacuation. It is clear that a large number of the crew managed to evacuate by lifeboat or life-raft. Repeated attempts were made by a number of vessels in the area to rescue possible survivors using lifebuoys, nets and grapnels but the 60-70 knot winds and 50-60 foot seas rendered rescue efforts futile. The reports describe major damage to the lifeboats and life-rafts by the heavy seas, and no survivors were found. 22 bodies from the 84-man crew were recovered. The autopsies recorded death by drowning while in a hypothermic condition.

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Published by the Health and Safety Executive 07/06

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