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P R O C E E D I N G S OF THE 15th INTERNATIONAL SHIP AND O F F S H O R E STRUCTURES CONGRESS

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P R O C E E D I N G S OF THE 15th I N T E R N A T I O N A L SHIP AND O F F S H O R E S T R U C T U R E S C O N G R E S S

VOLUME 2

Edited by A.E. MANSOUR

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PREFACE

This volume contains the 6 Specialist Committee and 2 Special Task Committee reports that will be presented and discussed at the 15th International Ship and Offshore Structures Congress (ISSC 2003)in San Diego, USA, 11-15 August, 2003. Volume 1 contains the 8 Technical Committee reports. Volume 3 will include the discussions of the reports, the chairmen's reply, the text of the invited lecture and the congress report of ISSC 2003, and it will appear in 2004. The Standing Committee of the 15th International Ship and Offshore Structures Congress in San Diego is: Chairman:

Secretary:

Prof. A.E. Mansour Prof. J.L. Armand Prof. B. Boon Dr. M. Dogliani

USA France The Netherlands Italy

Prof. W. Fricke

Germany

Dr. P.A. Frieze

UK

Prof. Prof. Prof. Prof.

Korea Poland Denmark Norway

C.D. Jang T. Jastrzebki J.J. Jensen T. Moan

Prof. H. Ohtsubo Dr. N. Pegg Prof. Y.S. Wu Prof. R.C. Ertekin

Japan (ex officio) Canada China USA

On behalf of the Standing Committee and members of the ISSC, I would like to thank the American Bureau of Shipping and the Ship Structure Committee for their financial support of ISSC 2003. The support of the City of San Diego is also gratefully acknowledged. Berkeley, USA March 2003

Alaa E. Mansour Chairman

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CONTENTS

Preface

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V

REPORT OF SPECIALIST COMMITTEE V1:

RISK ASSESSMENT

REPORT OF SPECIALIST COMMITTEE V2:

INSPECTION AND MONITORING

REPORT OF SPECIALIST COMMITTEE V3:

COLLISION AND GROUNDING

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71

REPORT OF SPECIALIST COMMITTEE V4:

STRUCTURAL DESIGN OF HIGH SPEED VESSELS . , . . . .

109

REPORT OF SPECIALIST COMMITTEE VS:

FLOATING PRODUCTION SYSTEMS

149

REPORT OF SPECIALIST COMMITTEE V6:

FABRICATION TECHNOLOGIES

189

REPORT OF SPECIAL TASK COMMITTEE V1.l: FATIGUE LOADING

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37

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235

REPORT OF SPECIAL TASK COMMITTEE V1.2: FATIGUE STRENGTH ASSESSMENT 285 Indexes

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15th INTERNATIONAL SHIP AND OFFSHORE STRUCTURE CONGRESS 2003 11-15 AUGUST 2003 SAN DIEGO, USA VOLUME 2 /EGO,

SPECIALIST C O M M I T T E E V.1

RISK ASSESSMENT

MANDATE Concern for the development of rational procedures for qualitative and quantitative risk assessment of ships. This shall include assessment of probability and consequence of accidental situations as well as evaluation of measures to control and mitigate the risk. Particular attention shall be paid to fire and explosion, extreme environmental condition, human element, traffic and obstructions, and operational hazards.

MEMBERS Chairman:

Dr. William Moore Professor Y. Chert Mr. A. Dinovitzer Professor O. Litonov Dr. Marc Prevosto Dr. Angelo Tonelli Professor Y.S. Yang Mr. Koichi Yoshida

KEYWORDS Risk assessment, risk analysis, formal safety assessment, hazard, accident, consequence, frequency, probability, cost, benefit

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CONTENTS

1 INTRODUCTION

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5

2 R E V I E W OF R I S K A S S E S S M E N T ACTIVITIES IN THE M A R I T I M E INDUSTRY 2.1 Regulatory . . . . . . . . . 2.1.1 Formal Safety Assessment (FSA) for the Maritime Rule Making Process 2.1.2 Risk Acceptance Criteria for IMO . . . . . 2.1.3 Bulk Carriers . . . . . . . . 2.1.3.1 F S A Study on Bulk Carrier Safety by Japan . 2.1.3.2 F S A Study on Life-Saving Appliances by Norway . . 2.1.3.3 FSA Study on Bulk Carrier Safety by Internationally Collaborated Group . . . 2.1.3.4 Decision Making at IMO . . 2.1.3.5 Other F S A Studies for Bulk Carriers 2.1.4 Passenger Ships . . . . . . . . 2.1.5 Maritime Security . . 2.2 Industry . . . . . . . . . 2.2.1 International Association o f Classification Societies (IACS) F S A Training 2.2.2 Guidance Publications on FSA . . . 2.2.3 Incorporation of Safety Assessment into the Rule Making Process 2.2.4 Application o f Risk Assessment to Icebreakers . 2.2.5 Joint Research Team on FSA . . . 2.2.6 Alternative Design and Arrangements for Fire Safety . 2.2.7 Marine Insurance Industry: Risk Assessment and Risk Selection 2.3 Applications . . . . 2.3.1 Risk Based Fire Safety Design . . . . . . . . 2.3.2 Event and Fault Tree Application . . . . . . . . . . 2.3.3 Fuzzy Set Modelling and its Application to Maritime Safety 2.3.4 F S A for Safety o f Coastal Trading Ships in Japanese Waters 2.3.5 Safety o f Ships Carrying Irradiated Nuclear Fuel . . . . . . . 2.3.6 Alert Communication from Small Craft Using Cellular Phones . .

8 8 13 13 13 14 14 14 14 15 16 19 19 21 21 22 24 25 26 26

3 E L E M E N T S OF R I S K A S S E S S M E N T . . . . . . . . . . . 3.1 Uncertainty of Data . . . . . . . . . 3.2 Decision Making Process based on the Results o f FSA . . 3.3 Effect of Safety Measures that have not Appeared in Historical Casualty Data 3.4 H u m a n Element . . . . . . .

27 27 28 29 29

4

Specialist Committee V.1

4 CONCLUSIONS

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5 RECOMMENDATIONS

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29

APPENDICES . . . . . . . . . . . . . . . . . . . . . A p p e n d i x 1: Indices for Cost Effectiveness Analysis ( C E A ) . . . . . . . . A p p e n d i x 2: C o m b i n a t i o n o f R C O s and the Effect . . . . . . . . . . .

30 30 31

REFERENCES

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Risk Assessment

1

5

INTRODUCTION

The application of risk assessment has evolved over 20 years in the offshore industry and within the last 5 years in the marine industry albeit in different directions. The offshore industry has focused on the application of risk assessment to evaluate the safety of individual offshore constructions. The marine industry has primarily focused on the applications of risk assessment to further enhance and bring greater clarity to the international rule making process. ISSC established, at its conference in 1997 in Trondheim, Norway, Risk Assessment. At the time of the first report of this Committee (Yoshida et al, 2000) the International Maritime Organization (IMO) risk assessment methodologies and techniques for the first time. The the application of risk-based approaches in the offshore industry.

Specialist Committee V.1, at ISSC 2000 in Nagasaki was agreeing to the use of 2000 report also addressed

This report provides the status of the application of risk assessment with a specific focus on the marine industry and provides insight into the direction that the industry is following in the research, development and application.

2

REVIEW OF RISK ASSESSMENT ACTIVITIES IN THE MARITIME INDUSTRY

Yoshida et al (2000) provides a review of the fundamentals for the application of FSA to the IMO rule making process. Since the development of formal safety assessment (FSA) approaches at IMO, there has been a wide range of activities associated with applying these techniques. This chapter provides a brief summary of these activities. 2.1

Regulatory

2.1.1

Formal Safety Assessment (FSA) for the maritime rule making process

As mentioned in the Yoshida et al (2000), IMO established interim guidelines for the application of FSA in IMO MSC/Circ.829 and MEPC/Circ.335 (IMO, 2002a). To date, these interim guidelines have been used, as trial basis, in several risk assessment in conjunction with IMO rule making process. Then, IMO decided to improve the interim guidelines whilst taking into account the experiences obtained through trial application. The Maritime Safety Committee (MSC) of IMO established a correspondence group to revise the interim FSA Guidelines. The group agreed to further include the following into the Guidelines. integration of analysis for human element through human reliability analysis (HRA); and 9 risk evaluation criteria. With regard to the human element, the group agreed that the HRA guidance developed by the International Association of Classification Societies (IACS) should be incorporated into FSA Guidelines, as an appendix. With regard to risk evaluation criteria, the group did not reach any firm conclusion. However, this topic was discussed at 74th session of IMO MSC, and it was agreed that Gross Cost of Averting a Fatality (Gross CAF or G-CAF) and Net Cost Averting Fatality (Net CAF or N-CAF) were most relevant for cost benefit assessment and that G-CAF and N-CAF should be used for comparison among risk control options (RCOs) in relation to the safety of life, and were included in FSA guidelines. In addition, it was further agreed that other indices are necessary to consider RCOs for

6

Specialist Committee E1

reducing the affect on property and the environment. This issue are remained for future consideration. The record of the discussion in the correspondence group was presented to 74th session of the MSC (2001a, 2001b). The record of the discussion at MSC 74 is given by paper of IMO (2001c) and further summarised in Gard Services (2001). IMO has since agreed, in both MSC and Marine Environmental Protection Committee (MEPC) to a final set of FSA guidelines as provided in IMO (2002a). 2.1.2

Risk acceptance criteria for IMO

As part of the FSA initiatives, recent efforts have also addressed the issue of risk acceptance criteria. As noted in Skjong (2002) it is difficult to make risk-based decisions without using or disclosing risk criteria. Risk acceptance criteria is of particular importance to IMO and efforts are currently underway to provide 'explicit' acceptable risk criteria. Skjong and Eknes (2001, 2002) provide an outline from which societal risk acceptance criteria may be established based on similar activities within other industries with similar maritime comparisons made for various ship types. Risk acceptance criteria will continue to be on the forefront of IMO related activities in the coming years. 2.1.3

Bulk carriers

IMO, recognizing the importance of enhancing the safety of bulk carriers, had considered and developed provisions, which were adopted as Chapter XII of 1974 International Convention for the Safety of life at Sea (SOLAS 74), as amended, at a SOLAS Conference held in November 1997. The Conference also adopted several resolutions concerning the safety of bulk carriers. Taking the resolutions into account, IMO MSC, at its 69th session in May 1999, agreed that it should further consider safety of bulk carriers. At the 70th session of MSC in December 1999, the United Kingdom offered a plan of conducting an internationally collaborated FSA study regarding bulk carrier safety. At that session, Japan announced that it would also conduct an FSA study on bulk carrier safety by itself. 2.1.3.1 FSA study on bulk carrier safety by Japan Since January 1999, a research committee (RR74BC-WG) in the Shipbuilding Research Association of Japan has been established under the supervision of the Ministry of Land, Infrastructure and Transport (MLIT) in co-operation with participants of the representatives of ship-builders, ship owners and operators, ship masters, officers and crew, the Japanese Coast Guard, National Maritime Research Institute and Class NK, for the purpose of conducting the FSA study on bulk carrier safety. The research committee conducted the FSA study, according to the FSA Guidelines in IMO (2002a), on typical bulk carriers with have topside tanks and hopper side tanks in the cargo spaces. The size of the bulk carriers under study was categorized into 4 groups by deadweight tonnes, (i.e. cape size, panamax size, handy size and small handy size). The casualty data-base was provided by Lloyd's Maritime Incident Service and Class NK was used. The results of the FSA study including final recommendations have been reported to IMO (2002b, 2002c and 2002e). The final recommendations for decision-making from the study are as follows: .1 The risk level of whole bulk carriers in future would stay at a relatively upper part of the 'As low as reasonably practicable' (ALARP) region even after recently adopted RCOs of SOLAS

Risk Assessment

7

chapter XII are implemented. Moreover, it is higher than other types of ships such as tankers and container ships. Therefore, IMO should pursue further cost effective safety measures that could reduce the risk of bulk carriers to ALARP (See Figure 1). .2 The risk level of the bulk carriers less than 150m in length is higher than that of the other size of bulk carriers. RCOs for mitigating consequences after hold flooding as required in SOLAS Chapter XII are not appropriate for those ships because only one hold flooding is fatal for bulk carriers of less than 150 m in length if the number of cargo holds of current design practice for such smaller ships can not be changed. Therefore, measures to prevent flooding is more important for such smaller bulk carriers.

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Risk level after implementation of SOLAS XII and Enhanced survey Figure 1 F-N Curves of each ship type (IMO, 2002b)

.3 SOLAS Chapter XII can be justified based on the comparison of the relative cost effectiveness versus other relevant RCOs such as a mandatory requirement of double side skin. Exemption of double side skin bulk carriers from SOLAS Chapter XII is also justified based on the same comparison and consideration on the magnitude of risk of accidents for double side skin bulk carriers. .4 For single side skin bulk carrier of 150m and over in length, it is expected that preventive measures against water ingress from a breach of the side shell structure would be effective to reduce the risk. According to the cost effectiveness assessment, it is recommended that corrosion control requirements such as an increase of corrosion margin and preventive coating should be considered, since it was found to be more cost-effective than double side skin. In summary, further investigation on following RCOs was recommended: (1) increased corrosion margin (Design Stage) and (2) corrosion control of single side skin for vessels in service.

2.1.3.2 FSA study on life-saving appliances by Norway An FSA project on life-saving appliances for bulk carriers was carried out in Norway by Det Norske Veritas in co-operation with participants from Norwegian Maritime Directorate, Norwegian Union of Marine Engineers, Umoe Schat-Harding, Norwegian Shipowners' Association and International Transport Workers' Federation (IMO 200 ld).

8

Specialist Committee V.1

The hazard identification (step 1 of FSA) was carried out for conventional lifeboats, throw overboard liferafts, davit/crane launched liferafts and free-fall lifeboats. The study was considered representative for all SOLAS bulk carriers, with the exception of bulk carriers of less than 85 meters in length. The risk reduction effects of introducing free-fall lifeboats as a mandatory requirement was quantified, whilst conventional lifeboats were considered the base case in Step 4 of the FSA. The considered RCOs were: 9 shelter mustering and lifeboat area; 9 remote control of the ship from the mustering area; 9 level alarms to monitor water ingress in all holds and forepeak; 9 individual immersion suits to all personnel; 9 free-fall lifeboats; 9 free-fall lifeboat with an additional free float mode; 9 marine evacuation systems for throw overboard liferafts; 9 enclosing open lifeboats for all existing ships with open lifeboats; 9 redundant trained personnel; and 9 improved pick-up function (crane). After carrying out an extensive review of historical data and completing all steps of the FSA, it was concluded that the following RCOs were providing considerable improved lifesaving capacity in a cost-effective manner: .1 Free-fall lifeboats with an additional free float mode. This solution was slightly better than the flee-fall lifeboat solution itself and considered cost effective (new ships). .2 Water level alarm with continuous water level indication in all holds and fore peak (new and existing ships) .3 Personal immersion suits to all personnel (new and existing ships) It was noted that the success rate in evacuation from existing ships remain rather low also after implementing the suggested RCOs. This might call for additional measures, in particular, focused on crew competence and training.

2.1.3.3 FSA study on bulk carrier safety by internationally collaborated group At the 71st session of the Maritime Safety Committee (May 1999) it was agreed that in light of the investigation into the loss of the DERBYSHIRE, the UK proposed leading an initiative to perform a holistic FSA analysis on bulk carrier safety. Many papers were submitted to IMO and within the collaborative group. The final results of their initiative can be found in IMO (2002e). At the 76th session of the MSC (December 2002), a final list of risk control options was prepared for decision to be made by the MSC on how to proceed.

2.1.3.4 Decision making at IMO MSC 76 did not consider the combination of risk control options and their prioritization in terms of risk reduction and cost benefit assessment. Based on the outcome of the International Collaborative Bulk Carrier FSA study coordinated by the United Kingdom the following decisions were made as shown in Table 1 (Gard Services, 2002b). It was agreed that Gross Cost of Averting a Fatality (GCAF) and Net Cost of Averting a Fatality (NCAF) would be used as primary selection critiera. Appendix 1 provides a brief explanation as to the benefits of using

TABLE 1 BULK C A R R I E R S A F E T Y INITIATIVES A G R E E D T O A T T H E 7bTHSESSION O F THE M A R I T I M E S A F E T Y C O M M I l T E E (Gard Services. 2002b)

Application Risk control option Double side-skin construction (hull envelope)

Improve coatlngs (hull envelope)

Steel repair standards (hull envelope)

Iiold frames (hull envelope)

Summary

New ships

Agreed that this will apply to new shlps to be constructed lssucs such as unifonn intemstionnl rtandards for double side s k ~ n construction and coatings, strength of inner s h n s are to be cons~dered. Controls andlor performance For new ships of double side standards lor protective coatings construcuon improved coatings In rclatinn to compatibility with will he requ~redfor dedicated cargoeb. seawater ballast tanks and v o ~ d spaces w~thlndouble hull spaces coatlng. Thls I S to be done In accordance with current SOLAS requirements for ballast spaces. Coating In cargo spaces 1r lcft to the discretion of owner and classtficatton soclety. T~ghtercontrols on grades of stccl No requirements. but a draft and weld~ngrods used for ~n c~rcularreminding owners and operators of their responsib~llt~es service repalr. under SOLAS regulation 11-113-1 concerning provisions that s h ~ p s shall be maintamed In accordance with rhe structural rcqulremcnts of recognised class socteltes and other related ohl~gationsunder the ISM Code. Various measures were cons~dered Not appl~cable. for hold frame5 including reduced d~minutlonallowances. strengthening to comply with IACS UR S12 ( N ~ for S side structures in single skin bulk carriers) and coatings. Concerns regarding thc ingress via loss of slde s k ~ nstructural lntegnty leading to many accidents.

Existing ships

Status

Not applicahlc.

To be further developed by the Design and Equipment (DE) subcommittee and completed by 2004.

No new requirements. Suffic~ent control measures are already in place through thc enhanced survey programme.

T o he funher developed by the De%igiland Equlpment (DE) subcommittee and completed by 2004.

Same as for new ships.

To be further developed by the Destgn and Equipment (DE) subco~r~rn~ttee and completed by 2003.

No requirements. but governments will be urged via an MSC Resolution to ensure that all nonIACS classificaion soc~eties comply with the lACS UR S31 (rencw:~lcrlteria for s ~ d eshell frames for vessels not bull1 in compliance wlth the revised IACS UR S12.

To be funher developed by the Design and Equ~pmcnt(DE) subcommittee and completed by 2003

TABLE 1 (cont.) BULK CARRIER SAFETY INITIATIVES AGREED TO AT THE 76THSESSION OF THE MARITIME SAFETY COMMIlTEE (Gard Services, 2002b)

Application Risk control option Forecastle fittings (hull envelope)

Fore deck fittlngs(hul1 envelope)

envelope)

(hull envelope)

covers)

covers (hatch covers)

I

Summary Forecastle requirements to reduce green sea loads on forward end of vessel and protect foredeck fittings. Strengthen stud pipes for alr and vent pipes to be sufficient to withstand horizontal forces of green sea loading. Closing devices and strength of small hatches to be sufficient to withstand vertical and horizontal green sea load~ngsin accordance with standards being developed by IACS.

detecting water ingress into cargo holds and dry spaces forward with visual and audible alarms in permanently manned spaces.

double side construction.

rmnimum bow height and reserve

1988 Load Line Protocol on hatch cover loads.

New ships lACS ongoing development of Unified Requirement (UR) S28 requiring fitting of forecastles on bulk caniers contracted on or after 1st January 2004. No requirements, since it is already being addressed via two lACS URs: S26 (strengthening and securing small hatches on exposed foredecks) and S27 (strengthening requirements for foredeck fittings and equipment). Governments will be urge ships flying thelr flag to comply via an MSC Resolution to ensure that all non-IACS classification societies comply with the lACS URS. SOLAS regulation X11112, Hold, ballast and dry space water ingress alarms will apply to bulk carriers regardless of their date of construction and will enter into force on 1st July 2004. Shim of 150111in lenah and upwards of double side construction should also comply with all structural strength provision of regulation XI15 requiring that the ship shall have sufficient strength to withstand flood~ngof any one cargo hold. Amendments were made to Regulation 39 to enhance the requirements for minimum bow height and reserve buoyancy. A revised simplified formula for design wave loads on hatch covers for bulk caniers (Regulation 16.1 of the 1998 LL Protocol).

I

Existing ships Not applicable.

. -

requirement and.w!ll be delivered

1

I

Same as for new shivs. Design and Equipment (DE) subcommittee and completed by 2003.

2I?. s -. 2

1

Same as for new shivs. ingress alarms and these will be adopted at the next session of the MSC in June 2003.

1

..

Not av~licable.

Not applicable.

Design and Equipment (DE) subcommittee and completed by 2003.

Action taken at MSC 76 and the 1988 LL Protocol amendments were adopted.

Not applicable. 1988 LL Protocol amendments were adopted.

I

TABLE I (cont.) BULK CARRIER SAFETY INITIATIVES AGREED TO AT THE 76T" SESSION OF THE MARITIME SAFETY COMMITTEE (Gard Services, 2002h) Application Risk control option Redes~enlreinforcementof hatch covers (hatch covers)

lmmerslon suits (evacuation)

Frce fall lrfeboats wrth float-free mode (evacuation)

Early abandonment (evacuation)

Summary

I Redeslen of hatch covers and

New ships Not amlicable.

hatch cover securing arrangements for e x ~ s t ~ nshins g

securlng mechanisms to withstand both ven~caland horizontal loads. Personal immersion suits for all personnel onboard shlp

Slngle free-fall survlval craft wlth float-free capability enabling rapid evacuation of crew from ship. Consideration of guidel~nesfor when and ~fbulk carrlers should be abandoned vessels at an earller stage in the event of flooding.

Terminal interface improvement (operational)

Improvement of ship-shore communications, training of stevedores and terminal operators and better control of loading capabilities.

Port State Control training (operational)

Prov~sionsof special~sedtraining for pon State control Inspectors in bulk carrier design and operation, with panlcular emphasls on areas of vulnerab~l~ty.

Existing ships

I Standards are to be developed tor

be provided.

free fall llfcboats wlth lloat-free capability. Not applicable.

No requirements, but Governments are to be urged via an MSC circular to apply the Bulk Load~ngCode of Practice for Safe Loading and Unloading of Bulk Carriers (BLU) Code. In addition, a manual for terminal representatives is to be developed. No requirements, but an MSC circulir would be developed on this recommending that the various Port State Control Memoranda of Understanding (e.g. Paris MoU, Tokyo MoU, 1 etc.) develop speclalised tra~ning for pinpointing vulnerability within structures panicularlv for

be prov~ded.

Status

1

To be further developed by. the . Design and Equipment (DE) subcommittee and com~letedby

develop draft amendments to SOLAS chapter 111 andfor the Llfe

develop draft amendments to SOLAS chapter I11 and/or the L ~ f e

( No requirements, but a circular urglng shipowners to issue guldance to ship's personnel on the possible need for early abandonment of a bulk carrier that has any single hold flooded and in particular for vessels which may not withstand flooding of any one cargo hold. Same as for new ships.

Same as for new s h ~ p s

I Saving Appl~ances(LSA) Code. I To be further developed by the Design and Equipment (DE) subcommittee and completed by 2003.

T o be funher developed by the Dangerous Goods, Solld Cargoes and Containers (DSC) subcommittee and completed by 2003. The manual will be developed In conjunct~onwith the Sh~p-PonInterface Woking Group. To be developed by the Flag State (FSI) ' ~m~~ernentatrbn subcommittee and completed in 2004.

TABLE 1 (cont.) BULK C A R R I E R S A F E T Y INITIATIVES AGREED T O A T T H E 76T11SESSION O F T H E MARITIME S A F E T Y C O M M I T T E E (Gard Services, 2002b)

Application Risk control option

Summary

New ships

Improved loading/stability ~nformation(operat~onal)

Provision of detail, comprehensive and user-friendly information covering stability and stress characteristics of the ship's hull.

Alternative hold loading (operational)

Alternative hold loading has been observed to ~ncreasethe longitudinal stresses on ships panicularly for ships over a defined age.

Guidelines are to be developed for the provision of detail, comprehensive and user-friendly information covering stability and longitudinal stress characteristics of the ships hull during loading and unloading. Not applicable.

Availability of pumping systems (operational)

Availab~lityof draining and pumping ballast tanks forward of the collision bulkhead, and bilges of dry spaces any part of which extends forward of the foremost cargo hold shall be capable of being brought into operation from a readily accessible enclosed space accessible from the navigational bridge or propulsion machinery control position without transversing exposed freeboard or superstructure decks.

Not applicable.

Existing ships

Status

Same as for new ships.

To be developed by the Stability and Load Line and F~shingVessel Safety (SLF) and DE subcommittees and completed in 2004.

Consideration is to be given as at what age alternative hold loading will be banned. This will be based on the types of cargoes and may include requirements based on a successful completion of a condition assessment. Bulk carriers constructed on or before 1st July 2004 shall comply with the requirements of this regulation no later than the date of the first intermediate or renewal survey of the ship to be carried out after I July 2004, but in any case, no later than 1st July 2007.

To be developed by the DSC subcommittee and will repon their findings and recommendations in 2004.

Agreed to by MSC 76.

Risk Assessment

13

GCAF and NCAF for the RCO selection process. In addition, criticisms were made regarding the lack of proper consideration of RCOs chosen in combination. Appendix 2 provides some description of how RCOs chosen in combination should be considered in the selection process.

2.1.3.5 Other FSA studies for bulk carriers Further papers addressing various FSA related activities regarding the IMO study on bulk carriers and other related ship reliability aspects include Bitner-Gregersen et al (2002), Lee, J.O. (2001), Lee, J.O. et al (2001), Skjong (2002), Skjong and Wentworth (2001), and Yeo et al (2000).

2.1.4

Passenger ships

FSA activities are also being considered for application to the safety of large passenger ships as discussed in Lee, J.K. (2002). The Netherlands addressed the issue of finding the root accident causes through systematic events analysis in IMO (2002d). The investigation identified operations and management as being the main causal category whilst navigational and watchkeeping figured as the main sub-category within the operations and management segment. The pattern for large passenger ships was found to be a clearer version of the general pattern. Particular efforts are being progressed in the area of evacuation studies for safety of large passenger ships as described in Boer and Skjong (2001), ICCL (2000), Park et al (2002). Due to the limitations of regulations, many research teams are developing new evacuation models based on microscopic simulation. Korea Research Institute of Ships and Ocean Engineering (KRISO) launched 2-year project to develop a new maritime evacuation model, IMEX (Intelligent Model for Extrication simulation). In the IMEX project an evacuation model was defined and briefly discusses some models and their limitations. Also, the report focuses on describing the configuration, feature, and status of IMEX that is designed to overcome those limitations. SOLAS Chapter II/2 (fire protection) was recently amended including the possibility, through regulation 17, to deviate from prescriptive fire protection requirements of SOLAS (Regulation II2/17 "Alternative design and arrangements"). Proposed alternative design and arrangements must achieve a fire safety level at least equivalent to the prescriptive design criteria in SOLAS. In order to provide this demonstration, a risk-based fire engineering analysis is to be carried out according to IMO's Guidelines as laid down in MSC Circ./1002.

2.1.5

Maritime security

In December 2002, the IMO adopted amendments to the SOLAS Convention to address maritime security in wake of the terrorist attacks in the United States in September 2001. These amendments require security assessments for both ships and port facilities using the requirements developed in the International Ship and Port Security Code. Although no specified criterion has been established on how these assessments are to be performed, the Code does specify criteria that should be considered in making these assessments. The USCG (2002) and Gard Services (2002, 2003) outline the current status of these activities and provide guidelines on outlining these criteria. Gard Services (2002, 2003) highlights that ports facility assessments and security plans with three proposed levels of security have been suggested: Level 1 (low level security measures required), Level 2 (additional security measures required) and Level 3 (high security measures required). Authorised Port Security personnel in accordance with general requirements for the appropriate security level will carry out assessments. These requirements will include a security assessment, a security plan, a designated port facility security officer and security training and drills. It is

14

Specialist Committee V.1

generally agreed that the assessment of an appropriate security level is a matter for national administrations and the ship and port facility security plans should allow for changes in security levels. Owners will be required to obtain an International Ship Security Certificate issued by the flag Administration for each ship indicating compliance with the mandatory sections of the ISPS Code. Compliance with the ISPS Code will require: 9 development of a ship security assessment and plan; 9 documenting training, incidents, breaches of security, maintenance and calibration of security equipment records, etc.; 9 a designated and properly trained ship security officer; 9 a designated and properly trained company security officer to co-ordinate the company security plan; and 9 training and drills for ships and companies to respond to terrorist threats. Companies will be expected to follow established procedures in keeping copies of all ships' documents, certificates and plans ashore as already required by the International Safety Management (ISM) Code for other types of documentation. The company or ship's personnel will conduct ship security assessments. The assessment addresses the security risk level to be levied for each ship or each class of ships as a prerequisite for the development of the ship security plan (SSP).

2.2

Industry

2.2.1

International Association of Classification Societies (IACS) FSA training

In 2001, IACS developed a standardised training course on FSA in order to establish a common understanding of FSA within the maritime community, provide a basis of information of the sequence of analysis steps and demonstrate the uses of various FSA techniques. This training course is provided in two levels: Level 1 provides a high-level overview of the FSA process and Level 2 (9 modules) provides a detailed overview that includes exercises and case studies. More information on the IACS FSA training course can be found at the following website http://www.i acs.org.uk/t'sw'wlp5/fsatrainin g.htm.

2.2.2

Guidance publications on FSA

The American Bureau of Shipping (2000) has developed the Guidance Notes on Risk Assessment Application for the Marine and Offshore Oil and Gas Industries. These Guidance Notes provide an overview for managers and technical professionals for application of risk assessment to the maritime and offshore industries. This guidance introduces the concept of risk and introduces risk assessment tools that can be used in risk determination.

2.2.3

Incorporation of safety assessment into the rule making process

In 2001, the Russian Register (RS) published the Rules for Classing, Building and Equipping Offshore Drilling Units and Sea Stationary Platforms. For the first time in the world practice RS introduced the chapter "Safety Assessment", incorporating the following factors: 9 risk identification; 9 the concept of analysing hazard situations; 9 methods for analysing accident situations; 9 methods for risk quantitative assessments;

Risk Assessment 9 9 9 9 9 9 9 9 9

15

statistic models of accident situations; assessment of individual and social risks; control of risks; selection of risk control options; cost evaluation associated with measures for reducing risks; platform sufficient safety criteria; recommendations for accepting solutions on reducing accident risks; principle of as low as reasonable practicable level (ALARP); and negligible and unacceptable risk levels.

In the future, RS intends to introduce a similar chapter in the Rules for Classing and Building Transport Ships.

2.2.4

Application of risk assessment to icebreakers

Appolonov et al (2001) inform that in the Rules in force of Russian Maritime Register of Shipping (2001) the following two basic principles using safety notions are applied to ice-going ships (lOS): 9 the ice category is considered as a ship safety guarantee in specified permissible ice service conditions (safety guarantee principle); and 9 within the ice category all ships independently of their main dimensions, hull lines and configuration should have equal permissible ice service conditions (unified safety standard principle). The analysis of the data of accident rate for the Russian ice fleet has permitted to disclose the following types of emergency conditions: ship wreck in ice; breach in outer shell; mass damage of ice strengthening structures. As a result of processing the ice damage statistics, a statistical model including the main service factors as parameters has been developed. The statistical model is based on the assumption that IGS service under impermissible ice conditions (IIC) is a main cause of mass ice damage. The probability of service under IIC was determined by the following relationship:

Pc,,c = P~,c "Pc

(1)

where P~,c=probability of ship getting in IIC and Pc =conventional probability of the situation that after ship getting in IIC a ship navigator decides to continue ship operation on the specified route. The following expression was obtained for function P1~c, approximated by two-parametric Weibull's law:

P1,c(K)=exp(ln(Pl,,,g,/3XK-Kp~r) 2' )

(2)

where K = index of the actual navigation type: K=I - e a s y navigation" K = 2 - middle navigation; K = 3 - heavy navigation; K > 4 - extreme navigation" Kpe r = K perO . ~ f " ~ f = function considering additional strength reserves in comparison with the ones required by the RS Rules" ~/i > 1" KperO = index of the permissible navigation type by Table 2; Plon~ = probability of long-range forecast for one type" Pt,mxl =Pitc (K,)"

K 1 -- Kper_

. .

16

Specialist Committee V.1 TABLE 2 VALUES OF PARAMETER KperO (Appolonov et al, 2001) Category LU4

Way of operation IO

Winter-spring navigation B 1

K 0

L 0

ES 0

2,5 1 0 0 2 1 0 0 3,5 1,5 1 1 3,5 1 1 1 IO LU6 4 2,5 1,5 1,5 4 2 1 1 IO LU7 4 4 3,5 3,5 4 4 2,5 2,5 IO LU8 4 4 4 4 4 4 4 4 IO LU9 4 4 4 4 B - The Barents Sea; K - The Kara Sea; L - The Laptev Sea; VS - The East-Siberian Sea; Ch - The Chuckchee Sea; IO - independent operation; PI - pilotage by icebreaker. LU5

IO

Ch 0

0,5 0 1,5 1 2,5 2 3,5 3,5 4 4 4

B

4 4 4 4 4 4 4 4 4 4 4 4

Summer-fall navigation K L ES 2 1 1 3,5 2 2 3 2 2 3,5 3,5 3,5 4 3 3 4 4 4 4 4 4 4 4 4 4 4 4 4 4 4 4 4 4 4 4 4

Ch 2 2,5 2 3,5 3 4 4 4 4 4 4 4

A quantitative estimate of the ship's design risk and assessment of the effectiveness of measures for risk monitoring can be made and thus decrease the scope the following FSA steps.

2.2.5

Joint Research Team on FSA

In 1995, the China Classification Society (CCS), Nippon Kaiji Kyokai (Class NK), the Indian Register of Shipping (IRS) and the Korean Register (KR) have formed a Joint Research Team (JRT) to address matters related to FSA. The JRT/FSA tasks were assigned to focus on: 9 safety assessment of passenger ship fire; 9 study on bulk carrier safety due to structural defects of the hull structures in cargo area; 9 A New Idea Considering FSA on a new tool or methodology applying FSA; 9 Application of formal safety assessment to rule making process; 9 Application of formal safety assessment to electro-hydraulic steering gear; 9 Application of FSA methodology to ship structure (especially for bulk carriers); 9 Safety assessment of engine room's fire of ships; 9 Safety assessment of collision and stranding of ships; 9 Safety assessment of fire in cargo holds; and 9 Database related to FSA. The main objectives of the above tasks were: i) identification of risk to be considered; ii) procedure of probabilistic approach to ships; iii) analytical method of risk advancement to accident; iv) acceptable risk level for ships; v) data base for all risks to be considered; and vi) case study of assessment The above research achievements were introduced into the Guidance of Application of FSA drafted by Dasgupta et al (1998, 1999) on behalf of the four classification societies above and the Guidance was made public in 2002 in order to indicote application of FSA at maritime community in Asia. Some achievements were applied to many aspects in China, Japan, Korea and India, e.g.

Risk Assessment

17

the methodologies adopted in the submissions to IMO provided by these countries mentioned above were from the Guidance. The Guidance included the following aspects: 9 The Guidance for Application of FSA (Version 1) containing details on the process of actual application of FSA as supplement to IACS Training Modules. It also addresses the issues on Quality Assurance of the FSA process. 9 Considering that IMO has formally issued amended Guidelines for FSA application (MSC Circ. 1023/MEPC Circ. 392) incorporating the HRA issues. This aspect is to be incorporated in the Guidance. 9 IACS has also evolved a basic Glossary of Terms on FSA that has been put up to IMO (MSC 76/INF.3). 9 The IACS AHG/FSA unanimously agreed that the IACS and the Japanese studies in which the main methodologies come from are more effective RCOs for bulk carriers. 9 A few issues have raised some controversy about the acceptability of the RCOs after the cost-benefit assessment. There has been an opinionated discussion on the use of GCAF and NCAF as the criterion of acceptance. Similar issues and other anomalies observed in the study reports strongly suggest the need for the introduction of process oriented quality assurance (QA) approach during the FSA study. Along with HRA study and updating the Guidance document, the application of QA to FSA can form a direction of future JRP of classification societies in Asia. Chen et al (1996, 1997, 1998a, 1998b) has been carried out a series of research projects in fire protection area for passenger ships sailing in the water area in China including the fire protection mathematical model, fuzzy theory application in fuzzy database as well as combination application of fuzzy theory and neural network theory. In the papers, they discussed the acceptable criteria of fire risk for passenger ships by means of fuzzy methods. This method describes the fuzzy property of the objects considered the structure property, knowledge property and other properties of objects. For that, they also try to reconsider the problem on the basis of Neural Network Method, which can describe the properties of objects synthetically. Chen and Lin (2001) carried out study on bulk carrier safety due to structural defects of the hull structures in cargo ships using FSA. The main purpose is to develop a method to carry out hazard identification and a corresponding database in which the fuzzy features in the process of application of FSA will be not avoided, so it is necessary to seek a new method to deal with the fuzzy feature in the practicable collection of data. The new ideas are applied to the process of FSA. In the paper, a comprehensive fuzzy method in the application of hazard identification of FSA is described, in the method the ship system is regarded as a whole that covers: 9 9 9 9 9 9 9

ship set; crew set; environment set; relationship between relationship between relationship between relationship amongst

ship and environment; ship and crew; crew and environment; and ship, crew and environment.

So this is a common method, which can be used in assessment of many events in ship system, such as the strength assessment for the number 1 cargo hold of bulk carriers, fire protecting, HRA and other nature assessments.

18

Specialist Committee V.1

Dasgupta et al (1998) studied both prescriptive and performance based regulations. The difference between the two forms in terms of the flexibility and their application were studied. The current indications of developments of performance based regulations as noted in the new SOLAS Chapter III (Life Saving Appliances) and SOLAS Chapter II-2 (Fire Safety) were highlighted. The authors proposed use of "success tree" (a similar as the fault tree; except that the probability of the success of a regulation is used instead of the failure probability) in the identification of regulatory requirements. Two issues were investigated: collisions as per SOLAS Chapter II-1 and for fire as per SOLAS Chapter II-2. The procedure enables a qualitative estimation of the effectiveness of a set of regulations as a safeguard against a specific hazard both in terms of its preventive and the mitigating clauses. The procedure can also be used to quantify the regulatory effectiveness once the individual clauses can be quantified. The procedure can also possibly be used in the quantification of the regulatory impact in future. The paper also identified the various sources of databases useful in the application of the FSA. Dasgupta et al (1999) studied a genetic model that was established for the steering gear system comprising of the following subsystems: main and auxiliary steering gear, the steering gear control system, and the steering gear power. The steering gear was identified to interface with main and emergency electrical powers source and switchboard. Based on the published failure statistics the relative reliability of the steering gear components were chosen with respect to the operational life. Where marine data were lacking failure data of components was chosen from the 'NPRD-95' handbook of "Reliability Analysis Center". The human factor affecting operator function was not considered separately through HRA. Arima et al (1996) carried out a trial application of FSA to the safety assessment of engine room fire of cargo ships was carried out. Its main purpose is to examine the applicability of FSA methodology, especially risk assessment and risk control options (Step 2 and 3 of the IMO FSA Guidelines). The fuel oil and lube oil systems in the engine room of a specific typical bulk carrier was taken as an example in this study instead of making its generic model. In addition, the occurrence and effects of fire due to combustible oil leakage from potential leakage locations were assessed. This paper proposed a technique for assessing the relative frequency ratio at each probable location where a fire may occur, and for assessing the effects of the layout, and number and types of safety equipment such as fire detectors and fire extinguishers on the scale of the fire. This trial application demonstrated that the methodology of probabilistic risk assessment could be applied to ship safety issues. Arima et al (1996) studied application of FSA on collision and stranding. At the same time related database were investigated. Base on the survey of data and information, fault tree analysis (FTA) and event tree analysis (ETA) was performed. Based on the casualty data, the basic patterns of FTA and ETA were constructed selecting macroscopic factors affecting collision and stranding, taking normal ships as the object of investigation with the objective of creating a framework for analysis in the future. It is convenient to classify collision and stranding casualties into light, heavy and total loss for the purpose of analysis and assessment in line with the categorization used by classification societies. A light casualty does not necessitate dry-docking after the accident or require the assistance of salvage tugs. Depending on the stranding of ship, the ship can move under power either by

Risk Assessment

19

waiting for the high tide or by shifting cargo on board the ship. A heavy casualty is an accident that requires the assistance of tugs, and total loss refers to a disabled ship, which can no longer be used; total loss includes a sunk or ships lost at sea. The human element comes into play in all cases of collision and stranding casualties. Based on actual accident reports, ETA and FTA were developed including human elements. Concerning collision and stranding, management factors related to human elements are very important. In the study, it can be concluded that the quantitative risk assessment can be carried out including human element, but is difficult to find universal accident measures based on survey about a small number of actual accidents. In addition, it is effective to survey minor accident because there is a possibility that those events could result in catastrophic accidents so we can prepare counter measures before catastrophic accident occur. Arima et al (1998) performed a safety assessment for fires in cargo holds, as the result may depend on ship type and size, they focused on a typical Aframax double-hull tanker. In this research, ships during unloading, tank cleaning and gas freeing are considered because the IGS is active and not a few accidents have been reported in this period. First, five initial events and scenarios were selected through a literature survey and discussion within the research group including consultants outside NK and preliminary risk assessment of them were carried out. Yeo et al (2000) studied trial applications of FSA methodologies to flooding of the number 1 cargo hold of bulk carriers. In this trial application, it was intended to: 9 develop the structural safety assessment system by applying the FSA methodology; 9 apply the developed system to risk assessment of bulk carriers whose dead weight is greater than 50,000 tones; and 9 suggest possible risk control options (RCOs) aimed at reducing the potential loss of life (PLL) from the viewpoint of structural integrity by implementing the proposed RCOs.

2.2.6

Alternative design and arrangements for fire safety

The new SOLAS Chapter II-2 entered into force on July 1st 2002, accordingly Registro Italiano Navale (RINA) developed class requirements as well as an expanded version of the FSA Guidelines where worked examples of all the steps in the analysis are included (RINA, 2002). For example, a fire event tree is provided in Figure 2 (reproduced from RINA's Rules).

2.2.7

Marine Insurance Industry: risk assessment and risk selection

The marine insurance industry has for assessing risks financial and economic, technical and operational factors to properly select and price risks. Insurance protection against risks and perils are handled differently amongst the well-known insurance conditions for hull and machinery. Hull and machinery policies are written for named risks and perils (International Hull Clauses, 2002) as well as for all risks and perils (Norwegian Marine Insurance Plan, 1996). Hull and machinery insurance is primarily covered by commercial insurance companies with a few notable exceptions. Therefore, risk selection is very important prior to accepting the risk. Significant effort has been placed on determining risks based on a number of criteria that include: 9 recent claims records; 9 classification society and current class records; 9 flag State; 9 ship type;

Specialist Committee 1/:.1

20

YES I NOT

I

A

~

Bl B2 Ct C2 D1 02 D3 ,.~MC~EINCIDENTCLA,..~ LOCATED LOCALt~ED MAJOR MAJOR MAJOR

MAJOR

LOCAL~ED

LOCALISED LCICAL~EO LOC&USEO'

MAJOR

MAJOR

ED

E~_~_

R~REII'r LOCAUSEO LOCALtSED LOCAUSEO LOCALLSED

LOCAUSEO LOC,~U~EO' LOCAUSEO

MAJOR

MAJOR

t.OC~Lt~SED

LOCAL~EO LOCAUSEO LOCAL~EO MAJOR

MAJOR

LOCALL~EO LOCAL~$EO

MAJOR

MAJOR

L~JOR

LOCAL~ECI LOCAL~EO

MAJOR

M~IOA LOCAL~EO LOCAL~ED MAJOR

MAJOR

LOC&LtSED MAJOR

JM,JC~

LOCAL~EO.

MAJOR

LOCAL~SED LOCAL~EO

~ $ V O G ~ I

MAJOR ~,~r

MAJOR LOCAUSEO LOCAL~$ED

L~MAJC~

Code EI B1 B2 C1 C2 D1 D2 D3

Event Ignition Event Rapid self termination Manual detection Automatic detection Forced ventilation shutdown Natural ventilation prevented Local manual suppression Automatic suppression On board manual suppression

Description Fire self terminates in the first instants subsequent to ignition People awake in the space are able to detect fire Equipment is able to detect fire Forced ventilation shutdown Whether door and/or windows are open/closed People in the space are able to extinguish fire Fixed equipment is able to extinguish fire People external to the space (fire brigade on board) are able to extinguish fire

Figure 2: Figure: event tree for a fire in a cabin (RINA, 2002) 9 9 9 9 9 9

ship size; ship technologymparticularly machinery type and manufacturer; ship age; port State inspection record; nature of trade including types of cargoes; domicile country of seafarers; and

9

trading regions.

The majority of ships with P&I cover are entered with P&I mutual clubs. The concept behind mutuality is insurance 'at cost'. Therefore, risk pricing is primarily based upon payment for claims incurred by the individual club member and, in some cases, shared payment of claims of other club members. P&I insurance costs are primarily based on cost of claims, cost of reinsurance and

Risk Assessment

21

performance of the club's investment portfolio. Risk selection also includes those elements described above. For further information, Gold (2002) provides an excellent summary of property and casualty insurance (hull and machinery) and protection and indemnity insurance (P&I) as well as the North of England P&I Club et al (1998).

2.3

Applications

2.3.1

Risk based fire safety design

Lee, J.H. et al (2001) provide a fire safety assessment about ship's fire protection design and classification society rules, statistical information and modeling techniques for the fire safety engineering are investigated and probabilistic safety assessment methods in the structural reliability engineering. A fire safety evaluation module (FSEM) developed in this paper calculates the probability of fatality, which can be used as an index of fire safety. FSEM is used to calculate the probability of fatality of the evacuees in a small room installed according to the rules for fire protection. Sensitivity analysis is executed to investigate the FSEM's applicability to ships. From the results, the necessity of new criterion for ship's fire safety design, the need to study the human behaviour in the evacuation from fire, and the development of new fire progress model considering special situations in ships are acknowledged. Yang et al (2001) summarises an FSEM that quantitatively evaluates the risk of evacuees in case fire occurs in ship has developed based on the research works done at Lurid University. The developed FSEM is applicable to multi-room structures as well as one-room structures. The necessary input data for the FSEM are obtained from a fire model, CFAST, and an evacuation model, MonteDEM. The MonteDEM evacuation model is developed by combining the Monte Carlo method for the random simulation of evacuee behaviour with the distinct element method (DEM) for the deterministic prediction of evacuee's movement. Compared with other evacuation models, the advantage of the MonteDEM evacuation model is that it includes the effects of ship motions to handle transverse inclinations. To verify the extended MonteDEM evacuation model, some numerical examples are demonstrated using the improved FSEM. The effects due to rolling motion should be considered to correctly evaluate the safety of evacuees in fire evacuation program. Through the numerical example, the quantitative estimation method for the latent risk of evacuation program is verified to be applicable and effective. Perhaps the most significant and thorough application of risk based fire safety design for ships is being completed within the 3 years duration SAFETY FIRST R&D project (European Community DGXII - 1998-2002 Growth Programme "SAFETY FIRST: Design For S a f e t y - Ship Fire Engineering Analysis Toolkit" Contract G3RD - CT99 - 0031). The aim of SAFETY FIRST is to ensure that a tried and tested fire ship engineering analysis toolkit, enabling ship designers to comply with IMO's revised fire safety regulations in place by the 1st July 2002. Now, it is possible for shipyards and ship designers to use a new approach where alternative, performancebased fire safety design and arrangements are accepted in lieu of traditional prescriptive designs as allowed by SOLAS regulation II-17 (alternative design arrangements). The analysis toolkit developed within the project, according to IMO's alternative design guidelines, is based on three pilot case studies of practical engineering relevance involving the fire protection design of passenger and crew accommodations and technical spaces on cruise ships as well as vehicle decks on ro-ro passenger ferries. Achieving these goals involves a good degree of scientific and applied research and development activities being performed by a team of 9 partners from 4 European Countries. Due to the novelty

22

Specialist Committee V.1

and complexity of SAFETY FIRST objectives, experience from civil buildings, nuclear and transport industry (railway and aircraft applications) is being considered, since fire engineering science is more developed in other industrial fields rather than in shipbuilding. To find an acceptable way of overcoming the limitations imposed by current fire safety regulations, that are not able to keep the pace with technological developments and customers' demand, the SAFETY FIRST project is structured into two main parts. The first one is based on scientific research and risk assessment techniques, aimed at assessing the applicability on ship design of performance-based analyses of fire and smoke models derived from other industrial fields. The second part of the project is devoted to practical simulations - with fire consequence modelling, qualitative and quantitative analyses of the selected Case studies - involving therefore a higher degree of applied research. The two parts of the project are linked to ensure that the models and the tests are always focused on real design alternatives. The expected achievements of the SAFETY FIRST project are to: 9 assess the practical applicability of IMO guidelines on the alternative design and arrangements with significant case studies; 9 develop a library of risk models, products and materials to identify representative fire scenarios to be readily available for application on the alternative performance-based approach to fire safety design; 9 provide a comprehensive toolkit for designers for practical application, including a costbenefit analysis to assess whether the alternative design is economically competitive; and 9 allow EU shipbuilding industry and ship owners to take immediate advantage of the new regulations leading to both enhanced competitiveness and improved safety.

2.3.2

Event and fault tree application

Assessment of risks is made through statistic models including those based on the full probability formula, Bayes' theorem, Monte-Carlo's and Delphi's methods, etc. Event and fault trees play important part in developing statistic models. Event trees and fault trees are used actively in investigation of different hazardous situations. Risk evaluation performed within the investigation floating production storage and offioading (FPSO) systems is a good example of such usage. In Wolford et al (2001), various events and fault trees including process loss of containment, mooring, transient induced leak frequency are explained. The function of the mentioned trees in specific analysis is of interest. There is information that marine event scenarios were represented with 89 unique initiating events, 12 frontline systems event trees, one support tree and 141 marine fault trees of which 89 developed specific initiating events and 52 modeled system response function (see Table 3). Over 2 billion unique event sequences were evaluated. Fire initiating event frequencies were developed for 70 individual hazard zones combined with an assessment of initiator density. The modeling of structural failures also followed a broadly similar approach. Structural event scenarios were represented with 46 unique initiating events, 13 frontline system event trees and one support tree. In the paper of Karsan et al (2001) it is said that risks associated with a FPSO system differ from those for the existing systems such as the conventional steel jacket, compliant tower, TLP and Spar. In the first phase of the existing Joint Industry Project (JIP), focus was made on evaluation of risks associated with the production operations from a Gulf of Mexico (GOM) FPSO. The JIP objectives included demonstration of the acceptability of a FPSO in the GOM: identification of

Risk Assessment

23

accidental events and F P S O components with high environmental pollution, loss of life, and financial risks, and recommendation of reasonably practicable risk reducing measures. Ten (10) oil companies, 3 F P S O contractors, 4 certification agencies and M M S sponsored the JIP. The risk concept developed for the G O M by a major oil company. TABLE 3 DESCRIPTION OF MODEL PARTITIONS (Wolford et al 2001) System Category Process

Marine

Initiating Events

171 Loss of Containment 9 Fault Tree modeling each of the 171 initiating events (Parts Count) 9 57 Escalation Events 9 89 (70 fires by zone)

Event Trees

9

Structural

9

46 Structural Damage

Total

363 Initiating Events

9 9 9 9 9 9 9 9 9 9 9 9 9 9 9 9 9 9 9 9 9 9 9 9 17

Comprised of 54 process section Each of the 2 phases in 3 separators is modeled as a separate initiating event 3 Hole sizes (small, medium, large) Model Functions 9 Ignition/Explosion 9 Isolation 9 Blowdown 9 Fire Suppression Failure of cargo management Ballast control failure Flooding from seawater system Flooding from cargo oil system Rupture of marine pressure vessels Energetic Release - turbine breakup Marine fire Crude oil spill Pump/engine room explosion Inadvertent discharge of oily waste due to bilge system failure Inadvertent discharge of oily waste due to surface runoff Single mooring line failure Corrosion holes or fatigue crack turret shell Vessel impact with turret Hull damage following vessel impact or helicopter crash Reduced weather vaning Turret superstructure or foundation damage Turret superstructure underdeck damage Process support damage Process support underdeck damage Transverse bulkhead damage Longitudinal bulkhead damage Ship hull damage during extreme weather Turret events following fire and explosion Billion individual event sequences

The use of quantitative risk analysis technique provided a tool to identify potential escalation scenarios and probabilities of terminating events, and consequences. Similar issues are discussed in the paper of Nesje et al (1999). Here the qualitative and quantitative methods to assess F P S O risks, including event trees and fault trees, are presented. The risks addressed cover those originating from subsea releases from risers and flowlines, leaks from process systems possibly causing fires and explosions, crude oil storage tanks and engine rooms, offloading to shuttle tankers, collisions with different types of vessels, and escape and evacuation operations. Another aspect associated with F P S O problems where the event trees are used is reflected in the work of M a c D o n a l d et al (2001). The question is about Collision Risks Associated with F P S O in

24

Specialist Committee V.1

Deep Water Gulf of Mexico. The paper focuses on three main collision scenarios: route based traffic, random traffic, area density traffic. Formulae for evaluation of the passing vessel collision risks and for calculating the drift vessel collision frequency using the COLLIDE methodology are cited. At that the input parameter include the following: number of vessels using the route; vessel type category on the route, e.g. merchant, tanker, supply, standby, etc; vessel size categories on the route (divided by vessel type and deadweight tonnage); Closest Point of Approach (CPA) of the route to the installation; beating of the route to the installation at its CPA; standard deviation of the route (related to route width). The paper of Xu et al (2001), in which the principles and strategies of in-service inspection programs for FPSO's are explained, could be joined to the mentioned papers by its content. The paper summarizes the technical basis for three levels of inspection strategies: 1) probability-based inspection method, 2) risk-based inspection method, and 3) "optimum" inspection method.

2.3.3

Fuzzy set modeling and its application to maritime safety (prepared by Dr J. Wang of Liverpool JMU, UK)

With the cost of construction, operation and maintenance estimates in the multi-millions of dollars, the marine industry is seeking ways of reducing both the time and money spent to provide the high-quality marine engineering systems. Decision-making based on conventional mathematics that combines qualitative and quantitative concepts always exhibit difficulty in modeling actual problems. The successful selection process for choosing a design/procurement proposal is based on a high degree of technical integrity, safety levels and low costs in construction, corrective measures, maintenance, operation, inspection as well as preventive measures. However, the objectives of maximizing the degree of technical performance, maximizing the safety levels and minimizing the costs incurred are usually in conflict, and the evaluation of the technical performance, safety and costs is always associated with the uncertainty, especially for a novel marine system at the initial concept design stage. Furthermore, the safety of a large marine system is affected by many factors involved in its design, manufacturing, installation, commissioning, operation and maintenance. Consequently, it may be extremely difficult, if not impossible, to construct an accurate and complete mathematical model for the system in order to assess the safety because of inadequate knowledge about the basic failure events. This leads inevitably to problems of uncertainty in representation. In probabilistic risk assessment, probability distributions are used to describe a set of states for a system and to deal with uncertainty in order to evaluate potential hazards and assessment system safety. In many cases, however, it may be difficult or even impossible to precisely determine the parameters of a probability distribution for a given event due to lack of evidence or due to the inability of the safety engineer/designer to make firm assessments. Therefore one may have to describe a given events in terms of vague and imprecise descriptors such as "likely" or "impossible", terms that are commonly used by safety analysts/designers. Such judgments are obviously fuzzy and non-probabilistic, and hence non-probabilistic methods such as fuzzy set modeling may be more appropriate to analyse the safety of systems with incomplete information of the kind described above. .1 Use of fuzzy set modeling in risk assessment of offshore support vessels and fishing vessels (Sii et al, 2001; Pillay et al, 2002). A rule based fuzzy set modeling method was developed to carry out risk based design/operation decision making for offshore support vessels. A fuzzy set modeling method was combined with the fault tree analysis to deal with fishing vessel safety.

Risk Assessment

25

.2 Use of approximate reasoning approach for the design of offshore engineering products (Sii et al, 2002; Sii and Wang, 2002). Three different modeling frameworks were developed for safety-based design evaluation and decision support. These include: (1) a framework for risk analysis of offshore engineering products using approximate reasoning and evidential reasoning methods, (2) a decision support framework for evaluation of design options/proposals using a fuzzy-logic-based composite structure methodology, and (3) a design-decision support framework for evaluation of design options using a composite structure methodology based on approximate reasoning approach and evidential reasoning method (Sii et al., 2002). The first framework is designed for risk analysis of an engineering system having a hierarchical structure involved in safety assessment. The other two frameworks are used for design-decision support, using a composite structure methodology grounded in approximate reasoning and evidential reasoning methods. Fuzzy set modeling can also be used together with multiple attribute decision-making (MADM) methods to assist decision makers in selecting the winning design/procurement proposal that best satisfies the requirement in hand. It can also be used together with Analytical Hierarchy Process (AHP) and the Delphi method in carrying out design support evaluation (Sii et al, 2002). In maritime risk assessment, the application of numerical risk criteria may not always be appropriate because of uncertainties in inputs as discussed by Wang and Kieran (2000). Accordingly, acceptance of a safety case or formal safety assessment is unlikely to be based solely on a numerical assessment of risk. Where there is a lack of safety data for analysis or the level of uncertainty in safety data is unacceptably high, maritime safety analysts to facilitate risk modeling and decision-making may effectively use fuzzy set modeling as a useful alternative approach. Application of fuzzy set theory to risk assessment can also be found in Zolotukhin and Gudmestad (2000), Chert and Lin (2001), Wang et al (1995a, 1995b) and Wang and Kieran (2000).

2.3.4 FSA for safety of coastal trading ships in Japanese waters The Shipbuilding Research Association of Japan has been conducting a series of wide range of FSA studies for safety of coastal trading ships in Japanese water. The main topics are collision, grounding and fire casualties. The studies have been conducted in principle with the IMO FSA Guidelines and have used other analytical techniques for determining probability of collisions, groundings and fires. Progress reports of the studies have been issued annual basis (Shipbuilding Research Association of Japan, 2002). The main contributors to the studies are National Maritime Research Institute (NMRI), Nippon Kaiji Kyokai (Class NK), Shipowners Association of Japan (JCS), representatives of shipbuilding companies, Japan Ship-masters Association, professors of naval architecture and Ministry of Land, Infrastructure and Transport (MLIT) of Japan. The studies have derived casualty data from the records of Judges taken at Maritime Casualty Coat in Japan. In addition, statistical data of rescue record of Japanese Coast Guard have been used. However, it has been found that such data do not necessarily contain information on ship operations leading to casualties. Therefore, investigations on operations in near miss cases were conducted by way of questionnaires and interviews with seafarers. Hydrographical data on main sea routes and density of traffics on the routes were also used. Based on these data sources, hazards have been identified and several main casualty scenarios have been developed. Then, several event trees have been developed based on the scenarios.

26

2.3.5

Specialist Committee E1 Safety of ships carrying irradiated nuclear fuel

Irradiated nuclear fuel, Plutonium and high-level radioactive wastes are categorized as "B" type irradiated cargo by IAEA safety standards, and carries in flasks in accordance with IMO's Code for the safe carriage of irradiated nuclear fuel, plutonium and high-level wastes in flask on board ships (INF Code). The requirement for the flask is "to withstand a fire of 800~ for 30 minutes", which has been developed based on risk assessments for land transport. Concern was expressed that the fire resistance level would not be sufficient for maritime transport. Therefore, IAEA evaluated the fire safety of such flask carried on board ships (IAEA, 2001). In this project, the Shipbuilding Research Association of Japan conducted a risk based assessment study for the flask for the carriage of high-level radioactive wastes and irradiated nuclear fuel (Shipbuilding Research Association of Japan, 2000). The assessment comprised (1) collection of ships' fire casualty data and real scale fire test data, (2) establishment of fire scenario based on the fire casualty data, (3) estimation of temperature and fire conditions in the cargo space during fire casualty scenario, and (4) risk evaluation of fire around the flask. Two fire scenarios, under which the cargo space for the flask would be affected, were derived as: (a) engine room fire, and (b) fire after collision with a tanker or gas carrier. An event tree analysis was carried out for engine room fires, and two major serious fire scenario were considered: (i) Oxygen rich fire where closure of engine room fails and fire spread rapidly and reach high temperature. The fire continues for few hours and decades. (ii) Oxygen controlled fire where closure of engine room succeeds, but fire continues for longer hours in relatively low temperature. The temperature history in the engine room was simulated based on various real scale fire tests. It was concluded that temperature in the aft-most cargo hold during such engine room fire does not reach the temperature of 800~ in any case. An event tree analysis was carried out for fire resulted from collision with tankers. A scenario of cargo oil fire on the surface of sea around the ship was derived as the major serious fire scenario. Temperature of and heat flux from the fire was estimated based on several real scale oil pool fire test data. Then, temperature in the cargo hold where the flask was stored was calculated. Because the flask carrier has thick double sided shell and the double side spaces are ballast water tanks capable of filling water, it was estimated that the inside temperature of the cargo hold did not reach 800~ in nay case.

2.3.6 Alert communication from small craft using cellular phones A risk based evaluation of FSA on the use of personal cellular phone in board small craft was conducted in National Maritime Research Institute of Japan by Mitomo et al (2002). As an example of an application of FSA for the maritime field, the event tree analysis was applied to assess the effectiveness of cellular phone for reducing number of fatalities or missing people in maritime accidents happened on smaller crafts. Casualty data on small crafts, statistics of such crafts and statistics of population of cellular phone available in Japan were used. Scenarios of casualties of small crafts were identified and an even tree was developed for the case in which a cellular phone was carried in such craft, and another case in which cellular phone was not carried. The conclusion of the report shows that cellular phone would be a valuable means of communication when the craft stays in up to about 6 nautical miles from the coast and can reduce the number of fatalities by about 60% which have no sufficient standard marine communication facilities installed, or in casualties which are unexpectedly rapid with little or no time or

27

Risk Assessment

opportunity to communication with installed communication facilities. This study provides a good example of application of FSA for smaller study items, in which FSA study is relatively easily conducted.

3

ELEMENTS OF RISK ASSESSMENT

3.1

Uncertainty of Data

A report of a study on treatment of uncertain casualty data during FSA was submitted by Japan to IMO (2002g). It became clear, during discussions within the international collaborating FSA study group and FSA team of Japan for bulk carrier safety (see section 2.1.3), that the existence of casualty cases where the causes of the accident are unknown results in a discrepancy between the calculated probabilities of hatch cover failure reported. Transparency and neutrality are of paramount importance for FSA. Therefore, judgment of cause of such casualties needed to be done in a transparent and neutral manner where expert judgment should be avoided as far as possible. So, this study provides a method of estimation of number of casualties and fatalities caused by hatch cover failures using Bayesian theorem as follows. Two separate causes of casualties about a group of ships each of which is the cause of a certain number of casualties were used. However, there were a certain number of casualties whose causes remain unknown and cannot be investigated or solved due to lack of relevant data. During the investigation of bulk carrier casualties by FSA the cause of casualties of bulk carriers was classified into 'hatch cover failure', 'side shell failure' and 'unknown'. However, even in such a case, it would be possible to estimate the true number of casualties fatalities of each category by a probabilistic approach based on Bayesian theorem as follows. Following symbols are used in following sections. nh is the number of casualties cause of which is obviously hatch cover failure; ns is the number of casualties cause of which is obviously side shell failure; nu is the number of casualties cause of which is unknown; nt is the total number of all casualties, i.e. nt = nh + ns + nu; th is the true number of casualties cause of which is hatch cover failure; ts is the true number of casualties cause of which is obviously side shell failure; then ts = n t - th.

Here, let nh = H, ns=S, nu=U, th -- X. Then, next equation is produced from Bayesian theorem.

P((th : X ) [ (n h = H )

From the theory of probability,

(n s : S )) : P((nh : H )

(n' : S ) [ (t h : X )) . P(t h = X ) = : S))

(3)

Specialist Committee V.1

28

P(,~ : x ) :

~ c~

~ c~

_

2n,

nt

Z nCi

9

i=0

H+U ZiCH P((M h = Y)(~(/~/s

"- S))--

i=H

~

Cs

n,

--

H+U Z i CH On,-i CS i=H n,

Z ,,, Ci i=0

P((,,~ : H)~ (n, : s)l (t~ : x)):

~ c.

.._~

c,

nt C x

Therefore,

P((,~ = x) I(.~ = H ) ~ (.~ = s)):

* c..~

H+U

ca

(4)

Z i CH Ont-i CS i=H

Equation 4 means the conditional probability of the number of hatch cover related casualties (th = X) when nh - H, ns=S, nu=U. The next step is to obtain the probability of the number of fatalities (fatalities). Let the number of fatalities for each casualty i (i=1, nu) cause of which is unknown denote as R(i), the number of fatalities of all unknown caused casualties denoted as Nu and the number of fatalities of all obviously hatch cover related casualties denote as NH. Then the probability of the number of fatalities of all hatch cover related casualties NffNH
P(NF): E

P(x I(.~ : H)~(.s : s ) )

Z

X =H x-It NF=N"+Z R(ij) j=l

X-H

(s)

Zx,,c~ k=O

In conclusion, analysis to deal with uncertain data is very important for FSA study, and the method for such purpose should be established. The idea given this paper is a good start for such development. 3.2

Decision making process based on the results of FSA

A report on study on use of GCAF and NCAF and on consideration to combination on the effect of independent RCO conducted by Arima et al (1999) was submitted from Japan to IMO (20021). It is essential for FSA that the process of analysis is transparent and neutral. It should be noted that decision-making using different final recommendations from various independent FSA studies would be a very challenging thing because there is no well-established methodology to do so. The decision-making process which might proceeded based on different recommendations from various FSA studies needs thorough review and consideration for final recommendations for decision making to ensure transparency. In this regard, this paper present some review of methodology of giving recommendations based on casualty data and Cost Effective Analysis (CEA) for RCOs as follows (see Table 2 for decisions made based on bulk carrier FSA initiatives).

Risk Assessment 3.3

29

Effect of safety measures that have not appeared in historical casualty data

The Japanese FSA study on bulk carrier safety has tried to estimate the safety effects of SOLAS Chapter XI1 by combination of various kinds of information and knowledge available. This kind of trial is essential for an FSA to be transparent and traceable. According to the experience obtained the FSA study, followings should be recommended for improving the FSA methodology and its applicability in future. It is found possible to obtain risks quantitatively through analysis on historical database on casualties and statistical data on world fleet. It is also possible to reduce the use of expert judgment by utilizing objective data based on engineering analyses as far as possible. However, in order to avoid discrepancies on interpretation of ‘unknown’ casualties, it had better improve the data gathering system regarding casualties and near miss cases within maritime industry as far as possible. 3.4

Human element

It has been found during various FSA studies that casualty data, in many cases, do not provide sufficient operational information carried out on ships in casualties. On the other hand, it has been found also that operational data in hazardous situations can be obtained through near-miss case report. IMO has recognized this fact and issued a MSCiCircular (IMO 2001e) by which it recommends member government to consider the establishment of near-miss reporting scheme. It is also recognized that questionnaire and interview to seafarers for accident cases and/or nearmiss cases are very useful tools for obtaining necessary and sufficient data for risk analysis on human element related risk assessments.

4

CONCLUSIONS

In recent years there has been an increase activity in the use of risk based approaches for the development of rules/regulations and standards and designs for ships. Two of the most interesting challenges are the area of risk acceptance criteria and how it will be applied within the marine industry in a consistent fashion and taking into consideration uncertainty into the risk modelling. It should be also recognized that analytical results of risk based approaches would have a level of uncertainty. Techniques to deal with such uncertainty should be developed.

5

RECOMMENDATIONS

1. The ISSC continue to monitor risk assessment and Formal Safety Assessment activities within the maritime industry and how that activity affects the application of risk based approaches to marine industry.

2. The ISSC should endorse the activities of the industry in developing risk assessment approaches to the development of rules and standards for design, maintenance and operations within the marine industry. 3. There is a significant need to develop approaches to risk acceptance that are consistent and agreeable within the maritime industry.

4. Further experiences gained through the application of the IMO FSA Guidelines should be used to refine the Guidelines for effective and efficient use of risk assessment in the marine industry.

Specialist Committee V.1

30

5. Techniques to deal with such uncertainty in risk based approaches should be developed. 6. With regard to the use of historical data, the decision-making process should require that: 9 interpretation of historical data and risk pictures should keep their best transparency; 9 any proposal of RCOs should be in accordance with the results of risk analysis; and 9 any cost-benefit analysis that indices values such as AR and AC should be clearly presented. It is also recommended that proposals of multiple RCOs should require special consideration on their CBA methodology and that the decision-making process should be based on thorough review and comparison of them taking technical issues described in Section 3 into account. In conclusion, followings are recommended: .1 Interpretations of historical data and risk picture should be thoroughly reviewed in a transparent and neutral manner. .2 Methodology and criteria for cost-benefit analysis should be thoroughly reviewed. .3 GCAF should be prioritized rather than NCAF. .4 Bearing in mind the controversial nature of net cost, net cost should be thoroughly reviewed and estimated for each stakeholder, and NCAF values of each stakeholder should be calculated accordingly. .5 Selection and combination of RCOs should be based on the FSA methodology, and recalculation of GCAF and/or NCAF for the combination of RCOs is essential for making decision.

APPENDIX 1:

INDICES FOR COST EFFECTIVENESS ANALYSIS (CEA)

It is also necessary to consider indices of CEA and their usage. GCAF and NCAF are defined as the indices for CEA in the FSA Guidelines. Definitions of these indexes are given as:

Where,

AC AB AR

AC GCAF = ~ AR

(A 1)

AC-AB NCAF = ~ AR

(A2)

is the cost of the Risk Control Option is the economic benefit resulting from the implementation of the Risk control Option is the risk reduction implied by the Risk Control Option

According to the experience obtained from several projects on application of FSA study, the paper has recommended that the first priority should be given to GCAF because of following reasons. While figures of GCAF and NCAF are positive, their meanings are understandable in the context of balance between the effect and the cost of RCOs. However, when a figure of NCAF becomes negative, the figure or absolute value of NCAF becomes meaningless except the meaning that such RCO can be justified only by an economical reason. When comparing RCOs whose figures of NCAF are negative, we should use absolute values of (AC - AB) instead of negative NCAF. Let's consider, for example, following five cases shown in Table 4. The 'Case 1' has the smallest NCAF values, but it is apparently not the best solution. Which case is better between 'Case 2' and 'Case 3'? The NCAF value of 'Case 2' is smaller than that of 'Case 3', but 'Case 3' is preferable

Risk Assessment

31

'Case 3'? The NCAF value of 'Case 2' is smaller than that of 'Case 3', but 'Case 3' is preferable because of larger AR value. In this example, 'Case 4' would be recommended because of the largest AR and the smallest Net Cost while its NCAF value is neither smallest one nor largest one among five cases. TABLE AI.1 AN EXAMPLE OF IMAGINARY RESULTS OF COST EFFECTIVENESS ASSESSMENT WITH NEGATIVE NCAF Case 1 Case 2 Case 3 Case 4 Case 5

AR 0.002 0.01 0.02 0.20 0.20

AC (US$) 1,000,000 1,000,000 1,000,000

AB (US$) 1,100,000 1,200,000 1,200,000

AC- AB (US$) -100,000 -200,000 -200,000

NCAF (Million US$) -50.0 -20.0 - 10.0

1,000,000

2,000,000

- 1,000,000

-5.0

1,000,000

1,200,000

-200,000

-1.0

Let's consider two stakeholders who are mainly concerned with some RCO GCAF value of which is marginal but NCAF value is small enough considering the criteria if we see figures in 'Total' case as shown in Table 5.

TABLE A1.2 AN EXAMPLE OF TWO STAKEHOLDERS RELATED CASES ]ag Stakeholder A Stakeholder B Total

[NCAF [~xC laB I Net Cost I GCAF 0.1 1,000,000 500,000 500,000 lO.O I 5.0 0.1 0 600,000 -600,000 0 -6.0 0.1 I 1,000,000I 1,100,000I -100,000I 10.0[ -1.0

It is clear that the RCO should not be implemented in this case unless a scheme in which the appropriate portion of the cost of Stakeholder A is compensated by Stakeholder B. It is not justified by negative NCAF value of this 'Total' case while only the stakeholder pays the cost for the benefit of both. APPENDIX

2:

COMBINATION

OF RCOS AND THE EFFECT

Both GCAF and NCAF use the magnitude of risk reduction from base risk level as a denominator. If implementation of other RCOs changes base risk level, figures of GCAF and NCAF would also be changed accordingly. In this case, re-calculation or re-assessment is straightforward. In case where two or more RCOs are implemented simultaneously as one combined RCO, the situation is more complicated. Let's consider, for example, two RCOs shown in Table 6. In this case, it is assumed that combination of RCO A and RCO B gives a combined risk reduction and benefit, but cost will take place individually. Each of the RCOis recommended separately if we see GCAF and NCAF values separately, but GCAF and NCAF values of combined RCO 'A+B' might not comply with the criteria. Hence the combination of RCO A and RCO B might not be recommended. It is recommended for improving FSA methodology and its applicability that it is necessary to establish usage and criteria of cost effectiveness. It should be noted that the acceptable cost would be a function of, and dependent on, the level of risk. A further presentation of recommendations regarding the decision making process is provided below in the RECOMMENDATIONS section.

32

Specialist Committee V.1

TABLE A2.1 AN EXAMPLE OF TWO RCOS WHOSE COMBINATION IS NOT RECOMMENDED RCO A RCO B RCO A+B

ZLR 0.5 0.5 0.6

AC (US$) 1,000,000 1,500,000 2,500,000

AB (US$) 500,000 500,000 600,000

AC- AB (US$) 500,000 500,000 1,900,000

GCAF 2.0 2.0 4.2

NCAF 1.0 1.5 3.2

REFERENCES

American Bureau of Shipping (2000). Guidance notes on risk assessment application for the marine and offshore oil and gas industries. American Bureau of Shipping. Arima et al (1996) Safety Assessment of collision and stranding of ships, 20-21 Nov. 1996, the first meeting of JRT, Tokyo, Japan. Arima et al (1998) Safety Assessment of engine room's fire of ships, 29-31 Oct. 1997, the second meeting of JRT, Beijing, China. Arima et al (1999) Safety Assessment of fire in cargo holds, 20-22 Oct. 1999, the forth meeting of JRT, Taejon, Korea. Appolonov E.M. et al (2001). Problems of Safety of Ships and Floating Structures Improvement. Proceedings, Russian Maritime Register of Shipping. 4. St.Petersburg. Bitner-Gregersen, E., Hovem, L., Skjong, R. (2002) Implicit reliability of ship structures. Proceedings, 21 st International Conference on Offshore Mechanics and Arctic Engineering, Oslo. Boer, Louis C., and Rolf Skjong (2001) Emergency Evacuation: How Better Interior Design Can Improve Passenger Flow. Proceedings, Cruise & Ferry Conference. Lloyd's List/Informa. London, May 8 t h - 10th, 2001. CCS, IRS, KR and NKK (2002) The Guidance for Application of FSA, Joint Research Team for Formal Safety Assessment. Chen, Y., Du yaojun (1996) Risk Assessment on Fire Protection Structure for Passenger Ships, 20-21 Nov. 1996, the first meeting of JRT, Tokyo, Japan. Chen Y., Du yaojun (1997) Fire Risk Research for Passenger Ships-Ranking and Fire Risk Level of Spaces, 29-31 Oct. 1997, the second meeting of JRT, Beijing, China. Chen Y., Du yaojun (1998a) Neural Network Method for Risk Research of Passenger Ships, 20-22 Oct. 1999, the forth meeting ofJRT, Taejon, Korea. Chen Y., Du yaojun (1998b) Acceptable Criteria of Fire Risk for Passenger Ships, 18-20 Aug. 1998, the third meeting of JRT, Mumbai, India. Chen Yingqiu and Lin, S. (2001), Comprehensive Fuzzy Method in Hazard Identification of Formal Safety Assessment (FSA), Proceedings, PRADS2001, Shanghai, China. 245-252. Dasgupta et al (1998) Application of Formal Safety Assessment to Rule Making Process, 18-20 Aug. 1998, the third meeting of JRT, Mumbai, India. Dasgupta et al (1999) Application of Formal Safety Assessment to Electro-Hydraulic Steering Gear, 20-22 Oct. 1999, for the meeting of JRT, Taejon, Korea. Gard Services (2001) Report from the 74 th session of the IMO Maritime Safety Committee. Gard Services publication, Arendal, Norway. Gard Services (2002) Report from the 75 th session of the IMO Maritime Safety Committee. Gard Services publication, Arendal, Norway. Gard Services (2003) Report from the 76 th session of the IMO Maritime Safety Committee and SOLAS Conference. Gards Services publication. Arendal Norway. Gold, E. (2002) Gard Handbook on P&I Insurance (5th ed.). Assuranceforeningen Gardgjensidig. Royal Corporate Print, London. Gudmestad O.T., (2000). Risk Assessment Tools for Use During Fabrication of Offshore Structures and in Marine Operations Projects, Proceedings, Offshore Mechanics and Arctic Engineering.

Risk Assessment

33

ICCL (2000) Recommendations on evacuation analysis for passenger ships and high-speed passenger craft: passenger vessel evacuation analysis. FP 45/3/1. London. IAEA (2001) Severity, Probability and Risk of Accidents during Maritime Transport of Radioactive Material (TECDOC 1231). International Maritime Organization (2001a) Report of the correspondence group Part 1: Minutes of the discussion, MSC 74/16 submitted by Japan, London. International Maritime Organization (2001b) Report of the correspondence group Part 2: draft guidelines of Formal Safety Assessment of use in IMO rule-making process, MSC 74/16 submitted by Japan, London. International Maritime Organization (2001c) Report of the Joint MSC/MEPC Working Group on the Human Element and Formal safety Assessment, MSC 74/WP. 19, London. International Maritime Organization (2001d) Bulk Carrier Safety: Formal safety Assessment for life-saving appliances for bulk carriers FSA/LSA/BC, MSC 74/5/5 submitted by Norway and ICFTU (International Confederation of free Trade Unions), London. International Maritime Organization (2001 e) Reporting near misses, IMO/Circ. 1015, June 2001, London. International Maritime Organization (2002a) Guidelines for formal safety assessment (FSA) for use in the IMO rule-making process. MSC/Circ.1023 and MEPC/Circ.392. London. International Maritime Organization (2002b) Bulk Carrier Safety: Report on FSA study on bulk carrier safety, MSC 75/5/2 submitted by Japan, London. International Maritime Organization (2002c) Bulk Carrier Safety: Step 2 of FSA on bulk carrier safety (risk analysis), MSC 75/INF.6 submitted by Japan, London. International Maritime Organization (2002d) Systematic incident analysis--finding the causes of dangerous occurrences. MSC 75/Inf. 14. London. International Maritime Organization (2002e) Bulk Carrier Safety: International Collaborative FSA Study--Final Report. MSC 76/5/5 submitted by UK. London. International Maritime Organization (2002f) Bulk Carrier Safety: Consideration on decision making process using independent FSA studies, MSC 76/5/12 submitted by Japan, London International Maritime Organization (2002g) Bulk Carrier Safety: Estimated number of casualties caused by hatch cover failures and those caused by side shell failures and fatalities in them of bulk carriers, MSC 76/INF. 17 submitted by Japan, London Karsan D.I. et al (1999). Risk Assessment of a Tanker Based Floating Production Storage and Offloading (FPSO) System in Deepwater Gulf of Mexico, Proceedings, Offshore Technology Conference, Houston. Lee, J.H. et al (2001) A Study on the Fire Safety Assessment of a Ship. Journal of the Society of Naval Architects of Korea, 38(1). Lee, J.K. et al (2002) IMO Trends on Large Passenger Ship safety. Journal of Ships & Ocean Engineering, 33. Lee, J.O. (2001) A Trial application of FSA Methodology to Hatchway Watertight Integrity of Bulk Carriers. Marine Structures, 14(6), 651-667. Lee J.O., et al (2001) Quality assessment of hatchway of a bulk carrier using QFD & FMEA, (2001), Journal of Safety & Reliability Society, 21(1), 7-19. MacDonald, A. et al (1999). Collision Risks Associated with FPSO's in Deep Water Gulf of Mexico. Proceedings, Offshore Technology Conference, Houston. Mitomo, N, et al (2002) Assessment for the Effectiveness of Cellular Phone for Reducing Number of Dead or Missing People in Maritime Accidents, Annual report of Institution of Electronic, Information and Communication Engineers, Tokyo. Nesje J.D. et al (1999). Risk Assessment Technology and its Application to Tower Based Floating Production Storage and Offloading (FPSO) Systems. Offshore Technology Conference, Houston. North of England P&I Club and South Tyneside College (1998) P&I Insurance--A training course. (3 rd ed.)

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Specialist Committee E1

Norwegian Marine Insurance Plan (1996) version 1999. Det Norske Veritas. Elanders Publishing Company AS. Oslo, Norway. Park, J.H., et al (2002) Development of an Intelligent Evacuation Model for Ships and Marine Structures. Journal of Ships & Ocean Engineering, 33. Pillay A., Wang J., Wall, A., Ruxton T. (2002) Risk assessment of fishing vessels using fuzzy set approaches, in press, International Journal of Safety, Quality and Reliability, June 2002. RINA (2002) 2002 Rules for the Alternative Design and Arrangements for Fire Safety. Registro Italiano Navale. Genova, Italy. Russian Maritime Register of Shipping (2001) The Rules for Classing, Building and Equipping of Drilling Units and Sea Stationary Platforms. St.Petersburg. Shipbuilding Research Association of Japan: (2000) Research for safety of marine transport of irradiative materials, March 2000, Tokyo. Shipbuilding Research Association of Japan (2002) Report of investigation of total safety evaluation of ships RR49: March 2002, Tokyo. Sii H. S., Ruxton T., Wang J., (2001) Fuzzy-logic-based approach to qualitative safety modeling for marine systems. Engineering Reliability & System Safety, 73(1), 19-34. Sii H. S., Wang J., Ruxton T., (2002) A safety based decision support system for the design of large offshore engineering products. The United Kingdom Health & Safety Executive website can be found at http://www.hse.gov.uk/research/frameset/offshore.htm, 2002 (final report submitted to the United Kingdom Health & Safety Executive for research contract D3474). Sii H. S., Wang J., (2002) A safety model for risk analysis of offshore engineering products using approximate reasoning and evidential reasoning approaches, in press, Structural Safety. Skjong, R. (2002) Risk acceptance criteria: Current proposals and IMO position. Proceedings, Surface transport technologiesfor sustainable development. Valencia. Spain. Skjong, R., Eknes, M.L. (2001) Economic activity and societal risk acceptance. Proceedings, ESREL 2001, 16-20 September, 2001, Turin, Italy. Skjong, R., Eknes, M.L. (2002) Societal Risk and societal benefits. Risk Decision and Policy (2002), v.7, pp 1-11, Cambridge University Press. Skjong, R. Ronold, K.O. (2002) So much for safety. Proceedings, 21st International Conference on Offshore Mechanics and Arctic Engineering. Olso, Norway. Skjong, R. and Wentworth, B.H. (2001) Formal safety assessment of lifesaving appliances for bulk carriers. DNV Report 200-0539. Hovik, Norway. Skjong R. and Wentworth B.H., (2001). Expert Judgment and Risk Perception, Proceedings, ISOPE. 537-544. USCG (2002) Security guidelines for vessels. Navigation and Vessel Inspection Circular No. 1002. Washington, D.C. Wang, J., Yang, J. B., Sen, P. (1995a) Safety analysis and synthesis using fuzzy sets and evidential reasoning, Reliability Engineering and System Safety, 47, 103-118. Wang, J., Yang, J. B., Sen, P. (1995b) Multi-person and multi-attribute design evaluations using evidential reasoning based on subjective safety and cost analyses. Reliability Engineering and System Safety, 52(2), 113-128. Wang J., O. Kieran (2000) Offshore safety assessment and safety based decision making- the current status and future aspects. Journal of Offshore Mechanics and Arctic EngineeringTransactions of the ASME. 122(2), 63-69. Xu T. et al (2001) Risk Based "Optimum" Inspection for FPSO Hulls, Proceedings, Offshore Technology Conference, Houston. Yang, Y.S. et al (2001) Development of fire safety assessment method based on reliability concepts. Proceedings, ICRAMS, Dalian, China. Yeo, et al (2000) Trial Application of FSA Methodology to Flooding of No. 1 Cargo Hold of Bulk Carriers. Proceedings, Offshore Mechanics and Arctic Engineering Conference. New Orleans.

Risk Assessment

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Yoshida, K., Loland-Eknes, M., Ludolphy, W.L.H., Moore, W.H., Tikka, K., Tonelli, A., Vinnem, J.E., (2000) Risk assessment. Proceedings, 14 th International Ship and Offshore Structures Congress. Nagasaki, Japan. Zolotukhin, A.B. and Gudmestad, O.T. (2000). Use of fuzzy sets theory in qualitative and quantitative risk assessment. SPE paper 61048.

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c,~oRES TRUt~

15th INTERNATIONAL SHIP AND OFFSHORE STRUCTURES CONGRESS 2003 AUGUST 11-15, 2003 SAN DIEGO, USA

U m oZ]31

VOLUME 1

SPECIALIST C O M M I T T E E V.2

INSPECTION AND MONITORING

C O M M I T T E E MANDATE Concern for the development of rational procedures for structural monitoring of ships and offshore structures. This includes methods for on-board and remote surveillance of loads, responses and structural deterioration for operational decision-making using information technology. Consideration shall be given to guidelines for inspection, assessment criteria for damage, and cost-risk based decision procedures for remedial action.

COMMITTEE MEMBERS Chairman"

Mr George J Bruce Prof. M. Duan Dr. G.V. Egorov Dr. R. Folso Prof. Y. Fujimoto Prof. Y. Garbatov Mr. J-C. Le Hire Dr. B.-C. Shin Mr. Ole T. VSrdal

KEYWORDS 'ships', 'offshore', 'structure', 'in-service', 'inspection' 'monitoring', 'measurements', 'assessment', ' procedures', 'fatigue', ' corrosion ', ' guidelines', ' survey', 'crack', 'buckling'

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INTRODUCTION

Since the previous, similar committee report to the 1991 Congress there have been changes in the issues, and many developments in methods and procedures have been reported. Perhaps the most significant is that the ship classification process is much more analysis oriented. For growing numbers and types of vessels, the consequences of in-service structural wastage and degradation can be better quantified. If the results of classification inspections identify damage, wastage and/or cracks, then there is an increased possibility of evaluating their implications in a rational analytical manner. This is a significant benefit of analysis based processes and there is potential feedback into the design process. For offshore structures, of both movable and fixed-site types, rational analysis based processes have long been used in the design and classification activities and, again the results from inspection activities can be monitored in a rational and predictive manner. There are also similarities. The 1991 report identified issues related to the world fleet at that time, and despite changes in technology, and large numbers of new built ships, the average age in 2001 is the same as in 1991. Many old offshore structures continue in use, and are also an ageing population. Given the size of the investment in any marine structure, a long life is to be expected, and the requirement to inspect and monitor many older structures is still the challenge identified in 1991. There has been no significant move towards standardisation, rather a proliferation of new types of fast ships in particular. The marine environment is still harsh and its effects on structures difficult to predict. Recent research is identifying effects not noted earlier, for example the security of ship hatch covers. Increasing cost pressures are reducing ship crew sizes, making monitoring more difficult and limiting in-service inspection. Faster port turn round times also make inspection difficult. Downtime costs continue to demand minimum time off hire which might be available for inspection purposes.

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OPERATIONAL BACKGROUND

Structural Problems and Issues

The key issues are safety at sea, pollution avoidance, integrity of structure, comfort, (avoidance of vibration and noise levels), and minimisation of maintenance costs. These issues concern all stakeholders, especially owners, operators, classification societies, regulatory authorities, flag and port states, designers, builders, underwriters, brokers, banks, passengers and crew, (and the cargo). Figures 1 and 2 illustrate alternative principles for securing structural integrity during operations. Figure 1 is typical for the classification regime, where follow up forms (general for given class of structure), define the inspections to be performed. The unit operator plans when and how to perform the required inspection, repairs, etc. The classification society accepts the operator activity by Letter of Compliance. The follow up condition presented in figure 1 is normally to be extended by special renewal requirements for the condition of service life beyond initial design service life. Figure 2 shows the system within the offshore sector in the North Sea. The regulatory body defines general rules and guidelines and the operator is responsible for safety, defines inspection and maintenance need, reports planned activity, findings and evaluated condition annually and every four year. The procedure for condition assessment, inspection and maintenance is available to the regulatory

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body, which audits the Operator. This principle is more flexible in including the effect of the inspection history of a given structure, and is thus better suited for operation beyond initial design service life. Rational condition assessment (CAP procedure, hull renovation) enables charterers or buyers to identify the best vessels, and to evaluate an existing vessel for possible future maintenance costs before considering its purchase, or to assess a vessel before a major conversion is contemplated.

Figure 1 Traditional Inspection and Maintenance process in the Classification regime. This report focuses on conventional vessels and structures, as the majority of those currently inservice, but some consideration is also given to less conventional structures. Arguably these need more monitoring and inspection as innovation often results in higher demand and response uncertainty and potentially creates new inspection problems, until operator confidence is gained. Innovation includes construction materials, e.g. FRP, concrete, forms of construction, e.g. cellular panels, and methods of construction, e.g., laser through-thickness welding. The failure rate may follow a "bath tube" curve where design and fabrication failure are important for the first years in-service, and deterioration is more important for ageing structures. For the bum-in period a random occurrence may be assumed. For the wear-out period knowledge of the deterioration processes may be used to define inspection and maintenance programmes that significantly reduce the failure rate (figure 3). The failures in the wear-out period will mainly be due to the failure mechanism fatigue and corrosion. The required strategies for prediction and determination of the degree of deterioration with a reasonable level of confidence may therefore vary with time in-service.

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Figure 2 Inspection and Maintenance process for structures based on use of Quality Assurance.

Figure 3 The failure rate of a structure. Alternative curves in the Wear-out period are presented for indication of effect of inspection and maintenance efforts.

2.2

Conventional Vessels

Since 1991 significant experience has been accumulated of condition assessment, based on inspection and monitoring hulls of vessels of different types, purposes and dimensions, in different external environments. Experience with conventional vessels is a basis for development of procedures for new vessel types and offshore structures. There has been a focus on bulk carriers, and tankers because of concern about pollution. Internal structural and tank monitoring gas carriers is difficult and essential. Because of service experience on conventional vessels, monitoring can be focussed on specific regions of the hull structure. Also, design approval is more likely to be based on FE models and high stress

Specialist Committee V.2

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regions in the hull can be more readily identified. There has been increased use of on-board strain and motions monitoring equipment, but little use of equipment to simultaneously determine the wave environments. Lack of correlation between environmental demands and ship response is a weakness in the overall monitoring process. This adds considerable uncertainty to future damage predictions. Following recent high profile loss investigations the scope of monitoring operations may increase. E.g. there may be more monitoring of spaces in order to detect flooding or dislodgement of hatch covers. These issues are receiving specific attention from current IACS committees, and monitoring is slowly being extended to determine the state of a ship over time not simply based on intermittent inspections.

2.3

High Speed vessels

These include displacement and dynamic lift (planing) types, which present particular problems. They have high power to weight ratios and the lightest weight possible, often using aluminium alloys or high strength steel, with innovative arrangements and fabrication. This may increase possible fatigue, buckling, vibration and crew or passenger discomfort from increased slams and acceleration levels. Uncertainties in demand and response activities, perhaps due to limited design experience may lead to the need for at-sea measurements and long term monitoring. This will validate the design for safety and cost-effective operations well as help improve the design of future similar vessels.

2.4

Multi-hulls

There may be greater uncertainty in both demands on the vessel and its response. Designs tend to use rational analyses, giving inspection and monitoring some rational focus. On-board strain and motic is monitoring equipment, with potentially more data points, particularly for cross-deck structures is required. High-speed craft may need additional inspections, e.g. because of bow and wet deck slams.

2.5

Offshore Structures

Mobile structures are periodically inspected on-shore, normally using classification society rules as a basis for inspection needs. Semi-submersibles need inspection and monitoring of various moorings and fittings in addition to structure and motions like other structures and vessels. They have recently had implemented a leak detection system in the underwater horizontal bracing. Some consideration needs to be given to jack-up structures, particularly with regard to both local seabed monitoring, for example for the possibility of tide- or current-induced scour, and leg jacking/translation mechanisms. Both factors can have effects on the structural integrity. Fixed site structures are generally inspected offshore, with reduced access and more cost, which also applies to repair or modification. Thus it is normal to implement a more structure specific program for inspection and monitoring. Normally there is a distinction between below water and above water inspection programmes. Below water inspection is done by use of diving or ROV supported vessel. Traditional fixed site structures are Jackets and Gravity based, others being steel tripod, articulated tower, guyed tower and special designs of jack-up structures. Examples of floating structures are semi submersibles, tension leg platforms and buoy-type (spar) platforms and ship shaped FPSOs. Periodically attempts are made to use remote sensing equipment on the structures. Leakage tests are now routine for monitoring. For floating structures such as tether leg structures and FPSOs, monitoring may also be based on movements. However, generally monitoring and inspection by remote sensing techniques has not been implemented. Periodical monitoring of motions and stress level has been performed to calibrate response analysis models and procedures. Techniques which

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may be applicable include acoustic emission for crack detection. Changes in movements under static environmental conditions may indicate seabed changes, for example anchor dragging or scour. The gravity based structures in the North Sea have concrete leg structures. These structures have shown to be robust with respect to wear and tear. The major or important findings given from the inspection have mainly been related to the MSF (Main Support Frame) that is, the steel deck structure that also is the connection between the concrete leg structures.

2.6

Other Marine Structures

Many other diverse forms of marine structures, have specific monitoring and inspection problems and solutions. FRP structures have generic problems because of a wide range of designs. Sandwich structures pose inspection problems since surface condition does not indicate possible internal problems, as is the case of steel sandwich panels coming into use in passenger ship superstructures. Ships for restricted navigation areas and inland waterways require specific inspection for damage due to frequent contacts with their infrastructure, or caused by cargo and handling equipment impacts. Bottom structures have loads from short-time groundings. (Egorov G.V., Kozlyakov V.V. 2001a, Egorov G.V. 2000). Further special inspections are needed in ships in ice, to identify indents, and mechanical wear due to ice load on the side shell, its frames and bulkheads, in contact with the ice. Floating steel and composite docks are intended for a life up to 50 years. The main defects are general and localized corrosion (sometimes with reduction of thickness to below 50% of the as-built state), deformation of pontoon decks because of usage of special engineering, and cracks in the under-wall hull plates, obtained during passage to destinations (Kozlyakov V.V. and Stankov B.N. 2000)).

2.7

Naval Vessels

These may be called into service at any time, possibly regardless of known structural problems. This creates a need for monitoring for risk assessment. These vessels also frequently are the subject of major refits or life extension programmes, and again accurate information on condition is a major benefit.

2.8

Pipelines and Risers

Offshore structures also include risers and pipeline, with high consequence of failure due to the potential for pollution and explosion. Regular inspection is performed, internally by pigging to detect corrosion and fatigue cracks before critical size is achieved. External inspection is by ROV(Remote Operation Vehicle) to detect potential damage of riser clamp supports and pipeline free span.

3.

3.1

PURPOSE AND SCOPE OF M O N I T O R I N G AND INSPECTION

Purpose

Inspection consists of periodic, scheduled checks of a structure to supply discrete measurement at points in time. Monitoring is constant measuring or surveillance, to give actual time histories. The primary purpose is to be aware of what is happening to a structure, in order to: assess structural degradation, at the time of inspection and in the future, and to manage the gradually degrading condition as it approaches a stage at which, the level of risk or the degree of unreliability become unacceptable, (Brooking and Barltrop 1993). verify design assumptions, in particular for innovative structures where design loads and structural response are based on advanced calculations and model tests, but where there is no

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experience available. (Iaccarino et al, 2000 and Grossi and Dogliani, 2000). assess potential failures due to gross errors in the design and fabrication. assist the operation of the structure. Examples on high speed ferries are systems to monitor ship motions, which are then transformed into motion sickness incidences, communicated to the bridge in real time, (Folsr & Torti 2001). On many oil tankers a system is installed to supply ship motion and hull stress information. Ship's officers can make informed course and speed changes based on ship behaviour in a seaway to minimize the likelihood of structural damage, (Witmer and Lewis, 1995). In more details it can be stated that the difference between monitoring and inspection is that whereas inspection consists of periodic and scheduled checks of a structure, monitoring implies a constant measuring or surveillance. Thus inspection supplies discrete points of measurements in time whereas monitoring provides the actual time histories. Monitoring can be divided into two categories, passive and active. Passive monitoring records condition over time and provides a basis for trend analysis. Active monitoring uses real-time sensors and can provide information about developing problems as they occur. This is of increasing interest in the field of ship and crew safety. These inherent differences result in different applications. Inspection is typically applied to critical areas, after identifying regions of a structure where degradation could have serious safety effects or be difficult and expensive to repair. Regular, scheduled inspection can minimise the need for repairs, as deterioration will usually be discovered before failure. Inspection therefore needs to identify different classes of damage, e.g. intermittent, non-propagating damage from operations, and damage which needs immediate attention. Monitoring is often used onboard ships where actual vessel behaviour must be assessed. Many vessels are large and or complex with relatively small crews. In heavy weather it is difficult for the crew to know what is happening in remote parts of the vessel. Sensors, possibly also CCTV, can help significantly to inform the crew of actual vessel behaviour, and any developing structural problems. Monitoring of cargo is also of growing importance. This includes physical characteristics, especially possible liquefaction in granular material, any cargo movement, and loading and unloading operations. The latter is of particular concern in bulk carriers where transient over-loading, and excessive stresses, could occur at the berth if the proper loading procedure is not followed.

3.2

The Scope of Operations for Monitoring and Inspection.

Meaningful monitoring is the collection of information of a specified type and form at specified intervals as part of an established plan to store, collate and subsequently use such information in a organised manner and for a specified purpose. The information that is collected may include current structure physical condition, demands on the structure, and the structure' s operational environments. Information will be collected during physical inspection and may require specialised equipment. Further inspection should occur immediately after the structure has experienced a major event, e.g. a large wave. It may be necessary to specify the environmental conditions which require an inspection. Inspection and monitoring must take into account a number of key features, including, what basic variable is to be measured, (e.g. thickness), how reliable is the measurement process, (e.g. access problems), how large an area needs to be measured and is the speed important. Other questions then

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become important, for example, what qualifications must the inspector have, what are the cost and resource demands, and what are the environment requirements to facilitate access to components, Much of the inspection regime is established by classification requirements or the operators CAP, (figures 1 and 2). The environments for the structure are important (Caridis, 2000). Examples are cargo type, time in ballast condition, protective coating system and cathodic protection, navigation routes (sea states, speed), compartment humidity, tank washing, inerting of cargo tanks.

3.3

Regulatory and Classification Requirements

All vessels and offshore structures have routine inspections and planned maintenance activities, for classification and regulatory needs, undertaken in accordance with established practices and schedules. (Figures 1 and 2 show alternative systems) The Classification Societies that are member of IACS must as a minimum include all IACS Unified Requirements (UR, group "Z"), in their rules A Classification Society is free to require more surveys or more detailed inspections, as it deems necessary, but cannot go below the UR level. Besides requirements for the surveys to be carried out, the class societies often also give guidelines on e.g. monitoring of particular details or recurring problems such as flooding. For the North Sea oil and gas industry there has been a continuous reduction of detailed specified requirements, with the Operator responsible for inspection and monitoring to document structural integrity with reasonable confidence and the regulatory body auditing the operator's procedures. For ageing structures the regulations highlight the importance of using rational procedures, and inspection and repair history in the current assessment of the structure against future demands. The traditional follow up scheme from the classification society and regulatory authorities based on assessment performed in the design process or use of design methods for the assessment will normally have shortcomings with respect to including the effect of inspection and repair history.

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4.1

DEVELOPMENT OF RATIONAL PROCEDURES AND METHODOLOGIES

General Aspects.

The requirements from the Classification society and Regulatory Bodies have traditionally been specifically related to failure mechanism and condition of operation. There has been a development towards requirements defined as a function of limit states. These are ULS (Ultimate limit State), FLS (Fatigue Limit State), ALS (Accidental Limit State) and SLS (Service Limit State) In principal the requirements of these limit states are to be deduced from common reliability requirements defined as function of degree of consequence. However, the present requirements defined for the different limit states are more likely to be deduced from accumulated experience or practice. The future should bring more calibrated reliability requirement among the different limit states. Rational procedures and methodologies for condition assessment during operation are based on the requirements of the limit states and information achieved from the inspection and monitoring activity.

4.2

Alternative strategiesfor Inspection and Monitoring,

Different assessment criteria will be used for different forms of damage, based on two main criteria of strength of aged ship hulls - ultimate strength (Egorov and Kozlyakov, 2001b) and fatigue strength (Guedes Soares C., Garbatov Y. 1998b). As long as theory can not predict the existence and extent of

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deterioration and defects with reasonable confidence, there will be a need to inspect and/or monitor. For documenting structural integrity with sufficient confidence two main strategies may be used. Measurements from inspection and monitoring are used to define with a degree of confidence the existence and extent of deterioration damages and defects. This strategy is used in merchant shipping and is based primarily on experience. It is less useful when novel structures are used. Measurements are used to verify theoretical predictions so these define deterioration and damage with a degree of confidence. Used on offshore structures and the models used. have significant costs. The inspection method and classification society schedule follow up scheme is mainly based on the first strategy, even if theoretical predictions are used to define the area to which the inspection applies. The increased use of detailed structural analysis to predict structural capacity and probable deterioration such as fatigue crack growth, make it possible to apply the second strategy. This will require rational procedures to combine the information from detailed analysis and the results of inspection/monitoring.

5. 5.1

INSPECTION - GENERAL ASPECTS Guidelines

IMO (1993), IACS and Classification societies provide inspection and data recording guidelines. Evaluation processes typically require comprehensive data, e.g. current gaugings. Specific vessel and structural types may need special guidelines, especially where there is structural and/or configuration innovation. There are official inspection guidelines (IMO, US Coast Guard, IACS, classification societies, etc) and those of unofficial associations (TSCF, OCIMF, IIMS, SIGTTO, etc). 5.2

Data management

Data must be compatible with subsequent monitoring and evaluation processes. It must be crossreferenced to the as-built condition of the structure and the design analysis. The inspection and repair history will be important for a rational condition assessment procedure. The use of information technology (IT) to report and store inspection and repair history may ensure reliable reporting and availability of information. However, not all database applications are suitable for reporting and retrieval of information. This should not be a general problem, but associated with initial use of IT. 5.3

Access and Inspection Constraints and Limitations

Many studies have reported on access physical requirements, including safe access for humans and the needs of any specific equipment, e.g. X-ray equipment, ultra-sonics, cameras. Consideration is needed of the local geometry of the structure and the 'geometry' of both humans and equipment at the inspection site and the access route. Hazardous areas require additional considerations. Holzman (1992) and Caridis (2000) summarized inspection access methods. These include methods such as, physical climbing, raffing, conventional temporary staging, portable staging, binocular with high intensity beam, divers and other less-common ones. Each method has particular advantages and disadvantages. Inspection effectiveness depends on inspection method and accessibility (Ma K-T. 1998). Most ships only have tank bottom access ladders, so access to some critical structural details such as side shell longitudinals is poor. It can be greatly improved by adding climbing bars, additional horizontal girders, or catwalks with handrails, but these also need inspection before use to ensure safety.

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Innovative inspection techniques include remote controlled lights, video cameras, flat plate inspection techniques, imaging systems, thermography (Goodwin and Hansen 1995). The offshore industry has reduced scaffolding during the last 10 years by use of used Rope Access. 5. 4

Plans and Schedules

Existing practice to maintain safe and effective structures is based on periodic inspections of the complete ship and its systems. Routine ship control procedures are important as the main sources of knowledge about the real condition of structures, and their capability. Ship inspections are presented in Fig. 3. (Caridis 2000), (Egorov 2002a), split into two basic groups, formal, essentially 'minimum' requirements of classification societies and authorities, and optional requirements recommended by some classification societies and supported by some owners. (e.g. CAP, hull renovation). Formal surveys by Classification Societies or Flag Administrations confirm ship safety. Optional surveys by shipowner representatives and others (e.g. insurers) confirm efficiency and safety. Classification societies have developed procedures to decide repair need (some computational) to estimate the hull condition of old ships to see if they conform to the requirements of younger ships. In recent years, a Condition Assessment Program (CAP) has been used to survey ageing ships based on unified rating scale (Straumann P., 1997). CAP, to review hull and other parts, independent of classification, but popular with stakeholders, being conducted by classification surveyors. Structural members may be ranked by a weighting coefficient (Egorov (2002b). This permits a total CAP rating as the sum of separate CAP ratings. Weighting is assessed by estimating the consequences for crew safety, ecology, the ship and the cargo of strength violations (danger) and their frequency. Plans and schedules should allow for the probability of detection of new flaws/damage occurring during the interval between inspections and may be divided into the following items: rational methods or performance procedures for in-service structural inspection and monitoring, rational procedure for information handling and decision making based on inspection and monitoring results by use of information technology including onboard vs. remote based procedures, guideline review and evaluation for inspection planning, performance and use of information, guideline review and evaluation for damage assessment criteria, guideline review and evaluation of cost-risk based decisions for remedial actions, including schedules.

5.5

Manual and Robotic Processes

The scale, time available and access problems are such that small structural defects may go undetected, as may problems from as-built defects or inadequate cleaning or coating of areas of structure. There are two possible approaches to inspection, manual, and robotic. A manual process requires safe human access to structure and cleanliness for accurate visual examination. The inspector may need to use special equipment items as well as having adequate lighting. Inspections may also focus on areas of structure based on experience of similar vessels and critical items. Hence, manual inspection is a time consuming process, based on experience and preferred to be performed in-shore. Many efforts have been and will be made to develop robotic inspection to avoid manual inspection. Robotic equipment may have advantages for monitoring damage, special areas and those with limited access. The use of robotic devices within double hulls, has possibilities, particularly during voyages. However, most inspections are expected to be manual processes in the near future.

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Specialist Committee V.2 Fig. 4 Classification of ship inspections

Manual inspections are usually undertaken in harbour, but if a serious crack is detected at sea then it may be visually monitored at regular intervals, perhaps using fibre optic devices, to detect growth in length. There are remote sensing devices to monitor crack growth in an inaccessible location. Some forms of degradation, such as buckling and dishing are visually obvious in good lighting conditions but difficult to detect in some circumstances, which may limit robotics in such cases. A computer product model is now commonly used at the design and fabrication stages. It is very useful for the safety management of ships in service to link the hull structure model to inspection and maintenance information, A ship inspection simulator (Hamada et al. 2000, 2001) could identify probabilities of damage detection, appropriate inspection, time and cost in advance. It would need subsystems for information management, inspection state verification and easy information input. Virtual reality support systems using a wearable computer have been proposed (Kawamura et al. 2000), (Yamato et al. 2001) with a head mounted display to show the hull structure graphics, a virtual reality sensor to show the inspector's location, voice recording for results, and a space ball to input location of damage. The graphical view of the HMD is adjusted to the view of the inspector.

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A database system to record inspection procedures, deterioration and damage, and repair methods has been proposed. (Kulesh et al 1996), (Kawamura et a1.1998). The database is used to decide safety of a ship hull and the appropriate maintenance schedule. In the future, computer supported systems using IT and multimedia can improve inspection activities, and reduce reliance on subjective judgment. However, careful attention is needed to avoid a large, complicated structure. Underwater inspections, on both ship and offshore structures, are of considerable concern, particularly regarding both safety for the inspectors and the reliability of the inspections and the data so collected. Automated Equipment is developed for use in areas that are inaccessible to humans, e.g. underwater, insulated regions of gas tanks, pipelines. Alarm equipment, for example for flooding, gas detection, temperature excursions, also has a role. Such equipment can be used to examine the structure in physical terms, and to measure the operational and environmental demands and structural response. Physical assessment of changes in the condition of the structure, e.g. coating system breakdown, corrosion, excessive deflections includes measurements of local stress levels and sea states. Different equipment is available and has been tested in-service. Sudden changes in structural response, apparently unrelated to the current environmental conditions, can indicate a potential problem.

5.6

Reliability of Inspection Processes

This is very important and factors are involved include human error, usually due to access problems, cleanliness and the sheer volume of structure to be inspected, e.g. the great many potential sites for fatigue crack development. (Ma K-T. 1998). A risk-based approach, priority assessment, provides a basis for inspection strategies to enhance current practice. The probability of defects as a function of the probability of detection has been investigated (Demsetz L., Carlo R., Schulte-Strathaus, R., 1996) Sensitivity analysis for fatigue reliability and inspection of a typical structural member of bulk carrier has been investigated (Chan Kim S., et al 1997). Fatigue properties and probability of crack detection at field inspection are obtained by questionnaire. Changing stress level (fatigue life) of the member, crack growth curve and inspection interval as parameters around a standard condition, failure probabilities of member during 20 years' service are calculated by the Markov Chain Model. The results suggest design solutions to improve reliability level, reduce structural weight and to simplify inspection. Combined inspection and sampling inspection has benefits over sampling, when few cracks in structural members. Inspection frequency is then important, because older similar ships may provide indications of potential crack sites. A focussed inspection programme can be based on records for similar ships defect histories, and a philosophy used which accepts some defects may not be detected at a specific inspection. However, the inspection interval is such that it is assumed that the next inspection will find the defect, considering growth rates and the probability of detection before critical length.

probability of detection

Best estimate of POD m

crack length

Fig. 5 Variation of Probability of Detection curve

95% confidence interval

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Comparison of observed propagating cracks and predicted potential fatigue crack growth have been reviewed (Moan et. al 1999) (HSE OTO 1999- 059, 060 and 061). A probabilistic fracture mechanics crack growth model was used to predict the results of inspections of fixed offshore structures in the North Sea. A correlation between observed crack growth and prediction were documented i.e. it is a potential to improve the inspection planning process by use of the general collected inspection history. The shortcomings of a given inspection planning procedure are presented and potential changes required in the theory for a better fit between prediction and observations is discussed

6.

6.1

MONITORING - GENERAL ASPECTS

Monitoring Processes- Manual

Physical inspections record surface protection break-down, buckling, dishing, high strain related, contact and mechanical damage, severe pitting, local corrosion and general corrosion, and cracking. Much of the inspection is purely visual, but for condition assessment using rational structural analysis techniques, measurement such as ultra-sonic techniques or EC/MPI for crack detection is needed. Manual inspections are not continuous monitoring and are to be defined by frequency and selection of area. Inspection for monitoring may be a barrier to inspection for preventing accidents and pollution. Where cracks are detected, their precise location and a general visual estimate of apparent length may suffice but often measurements are needed of the length and depth of a crack, for example for fracture mechanics based crack growth predictions. Photographic records may also be kept. Some damage, for example that associated with high strain and buckling, may result from a design deficiency, including Rules-related deficiencies. Such damage could indicate low local fatigue lives, and can also degrade local protective treatments, and requires continuous monitoring. 'Dishing' of external surfaces resulting from sea pressures is not normally considered to be a significant problem. Contact and/or mechanically caused damage, e.g. due to berthing, or cargo handling requires case-bycase consideration, as also chafing and wear, particularly of hatch cover seals. Containership hatches may suffer excessive elastic deformations, also affecting seal integrity, leading to frequent monitoring until consequences can be evaluated. While early repair may not be needed, local stress raisers may lead to early fatigue. Any monitoring regime must include an "ad hoc" section to manage these issues. The effects of severe pitting found during inspection are difficult to quantify, but studies indicate it could lead to leakage and possible pollution. It is unlikely to lead to local structural failure unless concentrated along a line of weld material leading to joint failure. The relationships between pitting and fatigue problems are somewhat unclear. In many cases the main problems of hull actual condition are in other type defects - 'pitting' in crude oil tankers, buckling in ice strength hull, where such damages are allowed under classification rules, initial buckling after building and repair. This leads to the three main problems areas, namely general area wastage, cracking and buckling

6.2

General Area Wastage

Generation of corrosion is said to be an unavoidable phenomenon in ships after commissioning and problems of corrosion and wear are important for securing the safety of hull structures. Corrosion can also be accelerated by the type of cargo that is transported. General wastage means diminution of thickness, increasing stress level, appreciably in terms of local bending due to surface pressure forces. The capability goes down, appreciably so in elastic buckling associated with in-plane forces. Thus the safety margin, relating capability to applied forces, reduces appreciably with general diminution. The same effects result from wastage of stiffener scantlings.

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In conventional ships, a corrosion margin and a permissible corrosion level, taking into account past records, are a primary means to cope with problems of corrosion and wear. It is intended that once the margin has been used, the remaining structure will still withstand maximum operational demands. An annual corrosion rate obtained by dividing the thickness diminution of a worn member by a ship's age at a given time has conventionally been used as the basic criterion. This is easy to manage, but has drawbacks, as the time during which corrosion is generated cannot be considered, corrosion rates vary significantly, and it cannot deal with probabilities. A sophisticated corrosion progress model for the painted members of ship hull structure (Yamamoto et a1.1998) assumes corrosion results from three sequential processes: degradation of paint coatings, generation of pitting points, and progress of pitting points. A consistent corrosion model can evaluate the generation and progress of corrosion quantitatively by introducing appropriate simple probabilistic models for each process. The unknown parameters used in the model are determined numerically based on the existing data of plate thickness measurements. The use of a prediction model in combination with inspection results may be used as a monitoring of the general area wastage. Pirker et al (2000) developed a simple corrosion model that can be used easily by engineers without losing the probabilistic nature of corrosion degradation. Corrosion was assumed to be the result of generation and subsequent progress behaviour of pitting points. A closed form relationship between corrosion depth and ship's age is developed, using the distribution property. Wastage is unlikely to be wholly uniform, but thickness sampling in the middle of plate areas is likely to be sufficient in many cases. Such thickness measures imply the use of ultra-sonic equipment. Wastage of stiffener sections should include both the webs and flanges, but detailed measurements in the case of rolled sections may be difficult to undertake. Sampling may be sufficient in most cases.

6.3

Cracking

This is a localised problem, but there may be repetition of cracking at geometrically similar locations, which may indicate design or fabrication deficiency. Cracks typically occur at welded joints, but may occur at other details that result in stress concentrations, e.g. notches, and incorrectly finished welds. It is difficult to assess if a crack is approaching critical length when rapid and unstable growth will occur. Given the large size and complexity of ship and offshore structures, inspections for crack development is a demanding and time-consuming process. It is also a very much an 'eyeball' process, except in selected highly critical regions where ultra-sonic and X-ray methods may be employed. The problem scale is such that use of prior knowledge to anticipate problem areas is most important, and this also causes a difficulty when a novel ship type and structural arrangement is employed. Simultaneous wastage and cracking results in a very complex situation where the development of wastage results in increases in local stress levels and which exacerbates the early onset of fatigue.

6.4

Monitoring Processes- Equipment Based

Hull stress monitoring devices have been used recently, especially for bulk carriers. They use both long- and short-base strain gauges with accelerometers and pressure gauges to process and store results in a PC, and provide real-time graphical output to assist decision making in heavy weather. The results can also be processed statistically in order to understand the vessel's operating environment and a measure of its demands on the hull structure. Unfortunately this monitoring equipment cannot assess changes that may be occurring in the vessel's structure. The output from such devices can also be used in order to determine the vessel's vibration characteristics in various cargo conditions. It would however need major structural degradation to produce noticeable changes in vibration characteristics.

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Acoustic emission crack growth sensing devices have been developed for use in the under-water regions of fabricated steel offshore structures. However it takes a significant crack to create a measurable noise when a tip suddenly extends and this process may be difficult to apply to ship structures. Ultra-sonic methods may be the most feasible approach and high-speed computer processing might be used to examine changing reflection patterns over time to indicate a crack. Some classification societies provide guidance in the technical selection and careful placement of hull stress monitoring equipment, including strain gauges and accelerometers of various types. This includes any intrinsic safety considerations in hazardous areas. Some types of vessels will require additional monitoring equipment, for example temperature sensors on gas carriers, i.e. LNG/LPG vessels, and heated cargoes on others. Both cargo and surrounding structure temperatures may need to be monitored if sudden leaks are to be detected. In some cases information from ships in service can be transmitted directly by satellite link to owners or operators. Equipment based monitoring was used offshore during 1985-90 in the North Sea. After a period with very limited use of equipment based monitoring a new generation of equipment is to be expected.

6.5

Monitoring of the Operational Environment

In many cases it is very important that monitoring of e.g. hull stresses can be correlated with the actual operational environment. This is the case when design assumptions are to be verified, or numerical predictions will be compared to the measurements. When a vessel have been designed, either explicitly or implicitly, for a maximum sea state, it can be useful to have some record of the actual sea conditions that have been experienced, when assessing some forms of damage. However, the monitoring of the operational environment is still today difficult and uncertain (Wang et al. 1999). The operational environment includes waves, wind speed and direction, current, loading condition and speed. In most cases wave induced loads on ship structures matter, but for offshore structures, current and wind loads must also be considered. Monitoring of the wave environment includes a minimum of wave height and period, and usually the wave travel direction. In some cases more complete information such as wave spreading or spectral shapes are provided. A special subset is the monitoring of the wave kinematics, (Nielsen et a1.1995), which is very useful to assess loads on offshore structures. Three possibilities that exist for wave environment assessment are: -The use of on-board wave height and length measurement equipment, together with the travelling directions of the waves. The equipment is usually based on some form of wave radar, -The use of sea area data as recorded from weather satellites for the route that has been recently traversed over, etc. A thorough review of this technology is given by Aage et al. (1998), -The use of wave data based on hind casting, i.e. calculations of the wave climate based on the wind velocity and wind direction records.

7.

7.1

D E V E L O P M E N T S IN T E C H N O L O G Y AND A P P L I C A T I O N S

Optics

Fibre-Optic sensors can be used to measure strain and detect structural damage. Several monitoring methods have been developed for concrete and composite materials. Their applicability for the structural monitoring of steel ships and FRP ships has been reviewed recently. As intrinsically safe devices they are attractive in vessels with hazardous cargoes. The methods can be outlined:

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Fibre-Optic Laser Doppler Vibrometer (LDV) measures dynamic strain of structural members. One end of a long optical fibre is bonded on the member and laser light input from the other end. When dynamic strain is applied to the structural member, the optical fibre is deformed. Light frequency of the laser beam reflected at the end of the fibre is shifted depending on the strain velocity of the fibre by the Doppler effect. As the change of strain velocity occurs within the region at which the fibre is bonded, the measured strain means the average strain of the bonded fibre length. Optical Time Domain Reflectometer (OTDR) detects structural damage e.g. cracking. An optical fibre is bonded to a structural member or embedded in composite material. Pulse laser is input from one end. If the member is cracked, the fibre also cracks, and light is reflected at the surface of the breakage. Measuring the time lag between input and reflected light locates a breakage. (Kageyama et al., 1995) Brillouin Optical Time-Domain Reflectometer (BOTDR) measures the distribution of static strain along the line of optical fibre bonded on the member surface or embedded in composite material. When laser light is input from one end of the fibre, Brillouin backscattering occurs at any locations of the fibre with induced strain. BOTDR is attractive in that the strain distribution along the optical fibre can be measured, but at present dynamic strain cannot be measured due to a long measurement time. Fibre Bragg Grating sensor (FBG) measures static and dynamic strain at a point on a structural member. The Bragg grating of the optical fibre core uses the projection treatment of interference fringe of ultraviolet ray. The refracting angle of the fibre core is changed at the grating. When bonded onto the member and a laser beam including several wavelengths is input from one end of the fibre, only the light with Bragg wavelength is reflected at the grating. When strain is applied to the optical fibre, the pitch of the Bragg grating is changed. The change of the grating pitch induces the change of the Bragg wavelength of the reflected light and strain can be calculated from this. (Murayama et al. 1997) Optic and strain gauge/load cell methods have been compared. Effects of optic types, splicers, coating materials, temperature, strain velocity and load patterns were examined, with the stability and accuracy of the measurement systems. FBG sensors can be effectively used to monitor steel and OTDR for FRP instead of strain gauges. Further, an optic-optic LDV sensor was used to monitor ship bow vibrations. FBG sensors have been used to monitor upper deck strain of a ship at sea. Optic-optic distributed strain sensors (BOTDR) have been applied to the hull strain measurement of a yacht. (Murayama et al. 1998) Laser scanning can accurately establish an existing structural configuration, and is increasingly used in ship construction to maintain an accurate profile during assembly and reduce fairing work (and induced stresses). It is also in use to measure ships during conversion work, to establish the actual shape of the structure. It is difficult to see the technique being commonly used for monitoring purposes, except for special cases, or danger areas, such as hull openings (hatches and large doors), and naval applications (e.g. submarine hull circularity) Changes in measurements over a period of time can indicate permanent strains as a result of structural overload, perhaps due to corrosion. There would be a need to take into account temperature effects.

8.

8.1

EVALUATION AND UTILISATION OF DATA AND RESULTS

Using datafrom Inspections and Monitoring

Ideally data is collected according to a pre-arranged plan, then applied and evaluated in a carefully structured assessment procedure. Several sets of variables, used in the assessment procedure are measured, at intervals, by an inspection activity. Examples basic variables are crack depth, plate thickness, cathodic protection, stress level, deadweight condition, vessel accelerations and slams.

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Inspection and monitoring results are used as input values in assessment procedures or to verify assumptions. It is also necessary to inspect and register the condition on entry into service as the basis for quantifying subsequent changes over time. As-fabricated structures should meet established quality standards, but in critical regions it may be useful to record actual structural details, including dimensions, thicknesses, misalignments and imperfections, before service life commences. It is equally important to assess and record demands placed on a structure, over a period of time, as such can have a significant effect on its degradation. They may not have been allowed for in the design process or may signify a design or construction error. Knowing why degradation has occurred and its quantification will inform changes in design practice or quality control during construction. Classification societies will be concerned with design and both builder and owner with the QC. It is of little value to collect data with no prior plan as to how it will be profitably used. Such a plan will need to consider, organising the information flow from inspection/monitoring through to assessment, making decisions to define the activity requirement or acceptance criteria for the variable to be measured or to make an assessment based on the measured values to establish fitness for purpose. Many classification societies have studied this problem and now have quite comprehensive electronic data processing systems for managing ship condition data and providing detailed evaluations. At one time inspection and monitoring during operation was restricted to examining deterioration to determine if it was below the tolerance included in the design and as specified by classification requirements. This particularly applied to wastage due to essentially uniform corrosion, but is not useful if applied to local pitting, cracking and light buckling, as such deterioration will usually deviate from design assumptions. A better approach is to use direct structural assessment procedures that include the effects determined by the monitoring or inspection process in the actual calculations. Structural Reanalysis System, (SRS) has been applied to offshore structures. If an existing vessel is being converted, for example a tanker or bulk carrier into an FPSO, then records will need to be made of the current quality of the structure for future reference. This also applies to vessels that have been repaired, for example by replacement of steel, and then returned to their normal service operations. Alternative uses of in-service observations are to be included in the historical log of a given hot spot area, as an observation related to present class or type of hot spot i.e. with respect to design, capacity and load effect or as an observation to be used for comparison with theoretical prediction.

8.2

Data Sharing and Exchange

The Tanker Forum and IACS both provide examples of data sharing in the interests of improving future structural reliability. Using the example of corrosion, there is a need for more data to allow a statistical analysis of the characteristics of corrosion. There is a statistical basis for structural capability, for wave and operational environment and for wastage due to corrosion. A probabilistic approach to predicting corrosion requires a very large data set, which can be more easily provided by sharing of information.

8.3

Data for Validation and Improvement of Predictions

The actual observed degradation of marine structures may be viewed as full-scale tests. Reported degradation may potentially be input for adjusting the theoretical prediction model. The collected inspection history from fixed offshore structures may be used to update the theoretical prediction of fatigue crack growth, (HSE OTO reports 059, 060 and 061), (Vardal et. al. 1999, 2000). (Moan et. al. 2000) present how an updated model for presence of fabrication failures in Jacket structures and POD values for MPI inspections by divers may be developed based on a collected inspection history.

Inspection and Monitoring 8.4

57

Damage Assessment.

After damage, a fitness for purpose evaluation (damage assessment) is made. The intention of is evaluation of criticality and whether immediate action or if repair is required. BS7910 published in 2000 is a recent published standard on assessment of reported damage. Damage assessment also includes determination of the extent of damage present in the structure with reasonable confidence. Collected experience may be used as a guide to the level of defects assumed present in a structure: Most fatigue cracks in warships arise from cyclic bending of the hull in waves and stress concentrations resulting from openings in the deck or the endings, (Clarke, (1991). A survey was organised of results of nine tankers, ranging from 35000 to 188000 DWT into a database. More than 40% of the numerous cracks were located in the side shell, more specifically in the connection of longitudinal girders to transverse web frames. (Schulte-Strathaus, 1991) Reviewing the occurrence of general failure in ship structure, (Emi et al., 1993), the occurrence of cracking failures was three times larger than deformation failures and six times that of all others. Many cracks were found in the collars of transverse bulkheads where side longitudinal stiffeners crossed. Most of the cracks were observed in the central area of the ship close to the middle cross section. The report showed a heavy concentration of cracks in the area seven meters below the water line. 40 % of the registered fatigue cracks in ship structures are observed to occur in the side shell in the connections of longitudinal to the transversals web frame, (Mizukami et al., 1994). Fatigue damage is caused partly by vertical and horizontal wave induced hull bending moment and partly by outside water pressure on the side shell. Five vessels were observed, which had been double bottomed or double hulled ranged from 35000 to 188000 DWT and about 1660 cracks were recorded. The data showed that the maximum number of cracks was located in area of the fluctuation of the water line. HSE OTO 1999 -059, The inspection history of 4000 MPFEC inspections from 30 Jacket structures located in the North Sea revealed a total of 511 cracks of 221 classified as propagating cracks. An inspection program is normally defined to determine the extent of structural deterioration. Inspections may be combined with a model to predict the degree and extent of deterioration with a reasonable level of confidence. The required amount of inspection may be defined by the required information for documenting the prediction model to present conservative or significant predictions. Pittaluga et al., (1991) reviewed the use of probabilistic fracture mechanics in fatigue assessment of offshore structures and summarised the practical application of probabilistic load and response analysis models based on fracture mechanics models. A probabilistic cumulative fatigue damage model based on a simple Markov chain approach is proposed by Lassen, (1991). It was developed a tool for a reliability assessment and a strategy for periodic inspection of welded joints in marine steel structures. The properties of materials in structural components can be assumed to vary depending on manufacturing condition, quality control, and the extent of non/destructive testing, (Zhao and Baker M. J., 1995). It was found that the material sensitivity varies depending on the type of failure mode. Probabilistically based guidelines for further rules development, including partial safety factors based on target reliability measures have been proposed, (Guedes Soares et al.,1996a) (Spencer et a1.,1996). A method of fast fatigue life prediction broad band random loading is proposed by Lu et al., (1998), based on statistical theory of the peak distribution of a stationary Gaussian random process and the concept of equivalent stress range.

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For structures cold areas, e.g. platforms in arctic regions, assessment of crack propagation is extremely important (Duan and Liu, 1995). A temperature dependent crack propagation theory (Duan, et al, 1999) is applied to evaluation on offshore structural steels. A critical parameter for the assessment of crack propagation at low temperatures is presented by analysing the pivot point on the FCP diagram (Duan, Li and Li, 1999). For the effect of cracks on the integrity of structures, Duan, et al, (2002) developed a method for determining the effective stiffness of the cracked component by using the finite variation principle and fracture mechanics. The method was used for the safety assessment of a drilling jackup, and it is found that the cracks have little effect on the overall strength of the jackup, even the cracking of one main leg does not result in the collapse of the global structure.

8.5

Corrosion

The ship structure operates in an environment where many kinds of corrosion can attack metals, C02 corrosion, top-of-line corrosion, weld attack, erosion, corrosion fatigue, pitting, microbiological corrosion, stress corrosion cracking. The mechanical process of fatigue crack propagation can be enhanced when the metal is immersed in a corrosive medium. This can be reduced by application of cathodic protection, but if overdone, there may be hydrogen embrittlement, (Fontana, 1992). In general corrosion depth increases non-linearly in a period of 2-5 years of exposure, but afterwards it becomes relatively constant. Experimental evidence shows that a non-linear model is more appropriate. Melchers and Ahammed, (1994) suggested a non-linear model as a function of the steel chemical composition. Time dependent corrosion degradation may be separated into three phases (Guedes Soares & Garbatov, 1998a). In the first, there is no corrosion because the surface protection works, in the second corrosion is initiated when the protection fails, and the third is a phase of stable corrosion depth. In assessing the corrosion rates applicable to the different areas of the ship, due account must be given to the existence of corrosion protection (Emi et a1.,1994). The state of the plating is often verified at repair operations and a probabilistic decision is made about the number of measurements that are necessary for a correct description of the plate thickness (Kmiecik et a1.,1995). Other authors have researched quantifying the corrosion rates in different areas of the ship hull Tanker Structure Cooperative Forum, (1992) as well as of different types of ships Huang et al., (1997). The reliability of plate elements subjected to compressive load and accounting for corrosion was presented in Guedes Soares and Garbatov, (1998b), where the plate collapse strength varies in time as a consequence of corrosion and associated plate renewal. More recently result were presented in Guedes Soares and Garbatov, (2000), where non-linear time dependent model of corrosion and linear fracture mechanics for the reliability assessment of maintained ship hull is applied. Developments in understanding the loading, the widespread use of new materials with higher strength capacity, and the use of refined structural design analyses have made possible weight optimisation of structures. However producing more economical structures with less redundancy has made them more prone to effects of strength degradation phenomena such as fatigue and corrosion. Attention has moved from design considerations to closer monitoring of the effect of maintenance actions. In the offshore industry fatigue strength assessment is an important design issue and fatigue assessment procedures have existed for a long time. Fatigue analysis is required by Classification Societies and the assessments are typically based on specified procedures. In shipbuilding fatigue assessment procedures have attracted more recent attention, and there is no unified requirement of Classification Societies for a fatigue analysis although it is becoming a requirement for some societies. In a recent study Lotsberg, (1997) reassessed the available SN data for tubular Joints and proposed curves for use in the new NORSOK standard. The resulting curves deviate somewhat from those given by HSE, (1995).

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The main difference in the approaches adopted for fatigue in these two closely related industries results the lack of redundancy of offshore structures, so very often the fatigue crack growth is a strength problem in that losing e.g. one tubular member may induce progressive collapse of the global structure. In ships, cracks that develop even in main longitudinal members do not lead to overall collapse of the structure. Widespread growth of cracks in ships leads to expensive repair work and so fatigue crack growth has been considered more an economic maintenance problem than as a safety problem. The simplified procedures adopted in the case of ships are sufficient for fatigue screening of structural details by pointing out the potentially fatigue critical ones, but they cannot be realistically applied for the design of new types of structures. For realistic fatigue life assessment these procedures need to be further calibrated for the different types of ship and the database of fatigue behaviour of ship structural details should be further developed with the results of both tests and theoretical investigations.

8.6

Remedial Actions

The design of any structure should make it easy to inspect, easy to monitor and readily repairable. Typical ship panel structures have these characteristics. Confined spaces with poor access, including novel structural members such as sandwich panels create major difficulties. Some assessment of repairability (and producibility) should feature in any proposed new structural configuration. Upon acceptance of the decision on the need for remedial action, the selection of method of repair of the ship hull is made, using evaluation of the damaged structure according to its significance to the structure for operations and the classification of groups of members by the rules of construction and taking account of requirements for impermeability. Obviously there are two extremes, either a defect must be fixed as soon as possible, and is urgent or it can be left until the time of the next docking.

8.7

Cost and Safety based on Inspection Planning

The offshore industry currently requires periodic inspections of fixed offshore platforms to ensure structural integrity. Past decisions on inspection, repair and maintenance were made by experienced engineers applying judgement along with appropriate deterministic analyses. However recently developed structural reliability techniques and considering the effects of uncertainties, mean inspection and maintenance scheduling can be based more on quantified information. Safety requires structural redundancy and inspection which leads subsequently to repair. Each safety item has a certain cost and it is important to minimise the total expected cost for the lifetime of the structure. Inspection planning needs optimised time intervals and decisions whether to repair or not. MTD 1989-104 "Underwater inspection of steel offshore installations: implementation of a new approach", London 1989 is a summery and introduction to the topic of probabilistic inspection planning methods. The focus is on offshore structures in the North Sea and the HSE reports OTO 1999 -059, 060 and 061 are summaries of validation studies of this class of approach for inspection planning. A probability based optimisation procedure (Cramer, Friis-Hansen, 1992) defines optimal cost, related to design, fabrication, inspection, repair, and failure cost. Fatigue failure is defined as fatigue crack growth beyond a critical crack size. A method (Madsen and Sorensen, 1994) seeks safety against fatigue failure through design, structural redundancy, inspection for fatigue cracks and repair of detected cracks. The total expected cost of different repair strategies were compared. The offshore industry requires rational reliability-based inspection scheduling and suitable software for implementing a chosen strategy, (Dharmavasan et a1.,1994). An inspection, and maintenance planning system incorporates advanced reliability analysis with fracture mechanics based on limit state functions and provides realistic results, using databases with appropriate information for probabilistic modelling.

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For re-qualification of old structures, where fatigue deterioration may be a problem, inspection of all potentially critical joints would be very expensive. A new reliability-based procedure to update the fatigue reliability of structural components of an offshore structure, uses information from inspections of an optimised limited number of joints, (Facciolli et a1.,1995). Fatigue reliability can be updated for inspected and non-inspected joints, for new and old structures. This methodology is being developed to consider structural redundancy effects and class of exposure on the fatigue reliability target of each structural element. An approach describes strength assessment procedures which consist of calibrations of design safety factors according to specific information about the structure, (Di Cocco et al., 1997). One application to ship structures was the study performed by Schall and Ostergaard, (1991). It was based on evaluation of fatigue failure probabilities for different service periods of marine structures. Stresses were modelled as Gaussian random processes, based on calculations using hydrodynamic wave load theory and the finite element method. Emi et al., (1993) studied the assessment of fatigue strength, inspection and maintenance management plans for ships, and formulated guidelines for inspections based on fatigue strength analysis carrying out fracture simulations of cracks and risk assessment. Fujimoto, (1994) presented a method to minimise sequential cost for planning inspection of fatigue deteriorating structures and find an optimal inspection strategy to minimise the total cost between the present inspection and the next. Matoba et al., (1994) suggested a methodology of preventive maintenance of ships, including structural enhancement, s inspection, and monitoring at sea. There is a guideline to control maintenance of ships against fatigue. Optimal inspection procedures based on reliability models were also discussed by Cramer et al., (1995) and a method of inspection planning for ships was presented by Ostergaard et al., (1996). Some technical developments in fatigue research of the ship structural maintenance project were presented by Hu and Bea, (1997). Fatigue reliability and und updating models are addressed followed by a discussion of the effective fatigue damage control. Many ships have experienced fatigue cracking of their critical structural details but repair decisions are often based on previous experience, rarely on traditional fatigue analyses. Only recent have general strategies for critical crack repair been suggested (Kai-tung Ma and Bea, 1995). Fatigue reliability procedures have been developed and applied to estimate the reliability of a single element or joint applying S-N or fracture mechanics. Assessments of ship hull reliability must allow for structural maintenance planning. Account must be taken of fatigue, corrosion, and their interaction, the explicit possibility of fatigue cracks in all joints and corrosion of each midship section plate. A ship repair decision is based not on the status of a crack or corroded plate, but on a generalised state of deterioration. This can be modelled by a global variable such as midship section modulus. The expected value of crack sizes and plate thickness as a function of time is described. The midship section modulus is modelled as a random variable whose mean and standard deviation change with time through degradation. (Guedes Soares and Garbatov, 1996b) Inspection and repair work during a ship lifetime never allows a very dramatic spreading of cracks. This effect has been incorporated in the time variant formulation of ship hull reliability (Guedes Soares and Garbatov, 1996c). That formulation and the corresponding results yield the required information to assess the effect of inspections and repairs at different points in time on the reliability of the hull girder, The effect of repair because of plate replacement in ship hull subjected to corrosion, was modelled in a similar way as the fatigue problem, by equating the repaired state of the structure to the state that the structure had at an earlier time in its life (Guedes Soares and Garbatov, 1996e). Normally both fatigue and corrosion will be present and their combined effect needs to be considered in that the decrease net section due to corrosion will increase the stress levels, which in tum increase the rate of crack growth. (Guedes Soares and Garbatov, 1998b). An improved model of corrosion was

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61

presented in Guedes Soares and Garbatov, (1998a), which accounted for the surfaces that are corrosion, protected initiating corroding at a random point in time in which the protection fails. The side shell pressure loading due to the imbalance between cargo pressure and outside hydrostatic pressure can be important, (Garbatov and Guedes Soares, 1997). The effect on crack growth is such that the rate of repair of these elements in the side shell would be significantly different. The reliability of a ship hull under both fatigue and corrosion has been reviewed (Guedes Soares and Garbatov, 1997, 1999a). Failure is considered possible in components subjected to combined loading as well as the global hull section. The approach was also applied to tanker and bulk carrier structures. Ships are inspected and repaired periodically, and the reliability assessment as a function of time must consider this. One repair policy is plate replacement when its thickness reaches 75% of the original. The time dependent effects of corrosion and crack growth can be modelled as random processes. Reliability is predicted by a time variant formulation and the effects of maintenance actions in increasing reliability are shown. Sensitivity of the reliability estimates with respect to several parameters is also studied. The formulation has been applied to a tanker where four different inspection polices were created for comparative analyses, good, average, bad or without maintenance, (Figure 5). The effects of different maintenance actions appear at 15 years during the second inspection. Inspections at 20 and 25 years show that poor or no maintenance action results in low reliability level. A Bayesian approach can be used to update some parameters of the probability distributions governing the reliability assessment of floating structures. The approach uses a time dependent fatigue reliability formulation with the description of the time to crack initiation, crack growth law and probability of crack detection updated using the information from inspections (Garbatov, Guedes Soares, 2001 a).

Figure 6 Reliability of Floating Structures. R(t) 1 :"'',,, \.. \.

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In most methods, planning is based only on reliability, but the inspection strategy for floating structures has to consider economics. A balance between reliability and economical criteria could be the key to inspection strategy, (Garbatov, Guedes Soares, 200 lb).Repair cost can be used as a criterion in reliability based maintenance planning and in particular how to vary the inspection interval in order to obtain the minimum intensity of repair cost. In some cases to keep the reliability level above a certain acceptable value the costs will not dominate and the reliability criterion will be the governing one. The

Specialist Committee V.2

62

approach was used as a base case to define the optimal strategy for maintenance planning using cost considerations. Simulated strategies for inspection planning showed that the application of repair cost optimisation for floating structures involves many uncertainties, including the costs of the shipyard making the repairs, and the inspection procedures. Evaluation of alternative criteria for maintenance planning in terms of the intensity of repair cost and availability of the platform to perform its intended functions is difficult. Minimum required intensity or repair cost could be related to Classification Societies requirements, but this does not give optimised maintenance effort. When maintenance is intensified inspection and repair costs increase. The search for a maintenance effort that will optimise the use of available resources should consider lifetime cost. 1.00005 r

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9.

F U T U R E NEEDS AND E X P E C T A T I O N S

There is a need to develop inspection methods for FRP and other composite materials, and evaluate the condition of sandwich and other novel structural forms where access is difficult or impossible. Novel ships and structures will continue to be developed and this, together with efforts to minimise structure cost, will create more inspection and monitoring problems. There is also a need to develop improved methods to inspect underwater and other offshore installations, for example flexible risers All one-off structures should have a maintenance record which allows their reliability to be properly assessed. Ideally, the overall technical quality of a structure would be documented using a single parameter. Further guidelines are required to decide on the quality and acceptability of structure lifetime extensions. There is also a need to take a life-cycle approach to structures to a much greater extent that at the present time. All this must be on the basis of an acceptable cost for inspection and monitoring. Formal ways are needed to capture experience into rational analysis where that experience may be at odds with calculations. The use of "expert systems", or artificial neural networks (ANN) may provide suitable approaches, given sufficiently large and reliable data sets. There may be development of further novel materials which indicate stress, or corrosion or whatever is being recorded through features incorporated in the material structure.

Inspection and Monitoring 10.

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CONCLUSIONS AND RECOMMENDATIONS

The development of novel structures, continued existence of sub-standard ships and the occurrence of accidents at sea all demonstrate that problems acknowledged over a decade ago have still to be solved or even fully addressed. There is increased application of rational techniques to inspection and monitoring, but the these will take time to permeate the fleet of ships and other structures. There are significant differences between older, existing structures where limited data is available and newer structures where the design and build process can set up the basis for a whole life history. Increased use of automated techniques is recommended to create an improved database of structural information and to provide increased safeguards against sudden structural failure. Research into materials which can incorporate sensors may lead to more routine monitoring at acceptable cost. More research is also recommended into the mechanisms which lead to sudden failure, and this should include the structure's environment. In creased sharing of data and improved management and use of data to will be a major factor in the development of more reliable predictive techniques, using probabilistic methods. Improvements in corrosion protection are also required.

REFERENCES Aage, C., Allan, T.D., Carter, D.J.T., Lindgren, G. and Olagnon, M., (1998), 'Oceans From Space' IFREMER, Brest, France. Ashcroft A. (1996). Shipboard Monitoring. SNAME Transactions, Vol.104, 1996, pp.549-552. Baerheim, M., Stacey, A., and Nichols, N., 1996, Proposed Fatigue Provisions in the New ISO Code for Offshore Structures, Proceedings of OMAE, Vol. lYI, pp. 513-518. Bai Y and Song R., 1997, Fracture Assessment of Dented Pipes with Cracks and Reliability-based Calibration of Safety Factors. Int. J. Pressure Vessels and Piping, Vol. 74, pp. 221-229 Bai Y, Igland R, and Moan T., 1994, Ultimate Limit States for Pipes under Combined Tension and Bending. Int. J. of Offshore and Polar Engineering, pp. 312-319 Bea R, Farkas B, et al., 1999, Assessment of Pipeline Suitability for Service, Proceedings of the 9th International Offshore and Polar Engineering Conference, Vol.II, pp. 347-354 Bea R.G., et al., 1997, Ship Maintenance Project, US Ship Structure Committee Reports, Vol. 1, Program Summary and Rational Basis for Corrosion Limits on Tankers. Report SSC395V1. Vol. 2, Study of Fatigue of Proposed Critical Structural Details in Double Hull Tankers, SSC395V2, Vol. 3, Repair Management System for Critical Structural Details in Ships, SSC 395V3, Vol. 4, Fatigue Classification of Critical Structural Details in Tankers, SSC395V4, Vol. 5, Fitness for Purpose Evaluation of Critical Structural Details in Tankers, SSC395V5. Bea, R. G., 1994, Evaluation of Alternative Marine Structural Integrity Programs, Marine Structures, Vol. 7, pp. 77-90. Bea, R. G., Pollard, R., Schulte - Strathaus, R., and Baker, R., 1991, Structural Maintenance for New and Existing Ships: Overview, Fatigue Cracking and Repairs, Proceedings of the Marine Structural Inspection, Maintenance and Monitoring Symposium, Vol. II-A, SNAME, pp. 1-25 Bjornoy O H, Rengard O, et al., 2000, Residual Strength of Dented Pipelines, DNV Test Results, Bjornoy O H, Sigurdsson G, et al., 2000, Residual Strength of Corroded Pipelines, DNV Test Results, Proceedings of the 10th International Offshore & Polar Engineering Conference, Vol.II, pp182188, pp. 189-195 Brooking, M., Barltrop, N., 1993, 'Ship Structural Management', Shipbuilding Technology Intl. '93, pp.168-171. Caridis P., 2001, Inspection, Repair and Maintenance of Ship Structures. Witherby & Co., London.

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Chan Kim S., et al., 1997, Sensitivity Analysis on Fatigue Reliability and Inspection of Ship Structural Members, Journal of the Society of Naval Architects of Japan, Vol. 181, pp.367-376. Clarke, J. D., 1991, Fatigue Crack Initiation and Propagation in Warship Hulls, In Advances in Marine Structure - 2, Elsevier Science Limited. Cramer, E. H., and Friis-Hansen, P., 1992, Reliability Based Optimization of Multi Component Welded Structures, Proceedings of the 1l th International Conference on Offshore Mechanics and Arctic Engineering (OMAE'92), Vol. II, New York, USA, ASME, pp. 1202-1223. Cramer, E. H., Schulte Strathaus, R., and Bea, R. G., 1995, Ship Structure Committee, U.S. Coast Guard, Washington, USA. Croll J G A., 1998, Simplified Analysis of Imperfect Thermally Bucked Subsea Pipelines, Int. J. Offshore and Polar Engineering, Vol. 8(4), pp. 283-291 Daidola J.C., Parente J. and Orisamolu I.R., 1997, Strength Assessment of Pitted Plate Panels, US Ship Structure Committee Report SSC394. Delmar, M. V., and Sorensen, J. D., 1992, Probabilistic Analysis in Management Decision Making, Proceedings of the 1lth International Conference on Offshore Mechanics and Arctic Engineering (OMAE'92), Vol. II, New York, USA, ASME, pp. 273-282. Demsetz L., Cario R. and Schulte-Strathaus, R., 1996, Inspection of Marine Structures. US Ship Structure Committee Report SSC389. Demsetz L. A. and Cabrera J., 1999, Detection Probability Assessment of Visual Inspection of Ships, US Ship Structure Committee Report SSC408. Dharmavasan, S., Peres, S. M. C., Faber, M. H., Dijkstra, O. D., Cervetto, D. C., and Manfredi, E., 1994, The Fatigue Life Improvement of High Strength Steel Welded Joints Using Hammer Peening Techniques, Proceedings of the 14th International Conference on Offshore Mechanics and Arctic Engineering (OMAE'94), Vol. llI, ASME, New York, USA, pp. 227-235. Di Cocco, N. R., Copello, S. and Piva, R., 1997, Improved Processes for Strength Assessment in the Requalification of Offshore Structures, Advances in Safety and Reliability, Vol. 2, pp. 14231435. Di Sciuva M., Icardi U. and Librescu L., 1999, Effects of interfacial damage on the global and local static response of cross-ply laminates, International Journal of Fracture 96, 17-35 Dry M.J., Schulte-Strathaus R. and Bea R.G., 1996, Ship Structural Integrity System (SSIS). Phase 2. US Ship Structure Committee Report SSC388. Duan ML and Liu CT, 1995, Investigation on failure of offshore steel structures under sea ice conditions, Proceedings of the Int. Conference on Technologies for Marine Environment Preservation (MARIENV'95), Tokyo, Japan, Sep.24-29, Vol.1, 71-74 Duan ML, et al., 2003, Finite element method for cracked components and its application to damage assessment of drilling rigs, to be submitted to ISOPE'2003 Duan ML, et al., 1999, Temperature dependent crack propagation theory and its application to offshore structural steels, Proceedings of OMAE'99, St. John, Newfoundland, Canada Duan ML, Li JCM and Li J., 1999, Application of the pivot point on the FCP Diagram to lowtemperature fatigue of materials, International Journal of Offshore and Polar Engineering, Vol.9, No.l, 68-72 Egorov G.V., 2000, Particularities of ensuring strength and reliability of river-sea ship's hulls operated in cold weather regions, ICETECH'2000 Conference. Proceedings, St.-Petersburg, pp. 247-254. Egorov G.V., 2002a, Ship surveys as a tool for ships actual technical condition assessment and maintenance of their safe and efficient exploitation, MARIND 2002 Conference, Varna, Bulgaria. Egorov G.V, 2002b, Study of the hull members relevance factors within the framework of the CAP procedure, ISC'2002 Conference, St. Petersburg.- 8 p. Egorov G.V. and Kozlyakov V.V., 2001a, Investigation of coastal and short sea ship's risk and hull's reliability, Proceedings of the 20th International Conference on offshore Mechanics and Arctic Engineering (OMAE'01), S&R-2109.

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Egorov G.V. and Kozlyakov V.V., 2001b, An assessment of the ultimate plastic strength of the ship's aged hulls, Proceedings of the 8th International Symposium on Practical Design of Ships and Other Floating Structures (PRADS 2001), Shanghai, China, pp. 1013-1019. Ellinas C P, et al., 1995, PARLOC - Pipeline and riser loss of containment north sea experience. Proc. Conf. ISOPE 95, The Hague Emi, H., Arima, T., and Umino, M., 1994, A Study on Developing a Rational Corrosion Protection System of Hull Structures, NK Technical Bulletin, pp. 65-79. Emi, H., Yuasa, M., Kumano, A., Kumamoto, H., Yamamoto, N., and Matsunaga, M., 1993, A Study on Fatigue Strength and Inspection Maintenance of Hull Structures for Planning a Long Life Service, NK Technical Bulletin, pp. 25-51. Facciolli, R., Piva, R., Ferretti, C., and Copello, S., 1995, System Fatigue Reliability Updating for Offshore Structures, Proceedings of the 15th International Conference on Offshore Mechanics and Arctic Engineering (OMAE'95), Vol. III-B, ASME, New York, USA, pp. 235-244. FolsO, R., Torti, F., 2001, Operational control of comfort on HSC, HIPER'01 Conference, Hamburg, Germany, pp. 149-163. Fontana, M. G., 1992, Corrosion Engineering, Third Edition, McGraw-Hill Book Company. Science Limited, pp. 1087-1100 Fujimoto Y. Shintaku E., 1997, Sensitivity Analysis on Fatigue Reliability and Inspection of Ship Structural Members, Journal of the Society of Naval Architects of Japan, Vol.181, pp.367-376. Fujimoto Y., Hamada K., Shintaku, E. and Gernot P., 2000, Development of High-sensitivity Sacrificial Specimen for Long-term Stress Monitoring of Structures, Journal of the Society of Naval Architects of Japan, Vol.187, pp.355-364. Fujimoto, Y., 1994, Planning for Fatigue Deteriorating Structures and Influence of Inevitable Uncertain Parameters, NK Technical Bulletin, pp. 13-45. Fujimoto, Y., and Swilem, A. M., 1992, Inspection Strategy for Deterioration Structures Based on Sequential Cost Minimisation Method, Proceedings of the l lth International Conference on Offshore Mechanics and Arctic Engineering (OMAE'92), Vol. II, New York, USA, ASME, pp. 219-226. Garbatov, Y., and Guedes Soares, C., 1997, Fatigue Reliability of Welded Joints in Tanker Structure, Proceedings of the 16th International Conference on Offshore Mechanics and Arctic Engineering (OMAE'97), Vol. II, New York, USA, ASME, pp. 7219-228. Garbatov, Y., and Guedes Soares, C, 2001a, Bayesian Updating in the Reliability Assessment of Maintained Floating Structures, Proceedings of the 20th International Conference on Offshore Mechanics and Arctic Engineering (OMAE'01), Rio de Janeiro, Brazil. Garbatov, Y. and Guedes Soares, C., 2001b, Cost and Reliability based strategies for Maintenance Planning of Floating structures, Reliability Engineering & System Safety, Vol. 73, pp. 293-301. Goodwin M. J. and Hansen K. A. (1995). Evaluation of Innovative Vessel Inspection Techniques (Phase II), Report for U.S. Coast Guard and U.S. Dept. of Transportation. Grossi, L, Dogliani, M., 2000, Load and Seakeeping Assessment of HSC based on Full Scale Monitoring, NAV2000, Vol. I, Venice, Italy. Guedes Soares, C., and Garbatov, Y., 1996a, Fatigue Reliability of Containership Hull Girders Considering Maintenance Actions, Proceedings of the 1st International Conference on Marine Industry (MARIND'96), Vol. I, pp. 151-166. Guedes Soares, C. and Garbatov, Y., 1996b, Fatigue Reliability of the Ship Hull Girder, Marine Structures, Vol. 9, No. 3, pp. 495-516. Guedes Soares, C. and Garbatov, Y., 1996c, Fatigue Reliability of the Ship Hull Girder Accounting for Inspection and Repair, Reliability Engineering and System Safety, Vol. 51, No. 2, pp. 341-351. Guedes Soares, C. and Garbatov, Y., 1996d, Reliability of Maintained Ship Hulls Subjected to Corrosion, Journal of Ship Research, Vol. 40, No. 3, pp. 235-243. Guedes Soares, C., and Garbatov, Y., 1997, Reliability of Maintained Welded Joints in the Side Shell of Tankers, Proceedings of the 16th International Conference on Offshore Mechanics and Arctic Engineering (OMAE'97), Vol. II, ASME, pp. 13-27.

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Guedes Soares, C., and Garbatov, Y., 1998a, Non-linear Time Dependent Model of Corrosion for the Reliability Assessment of Maintained Structural Components, Safety and Reliability, Vol. 2, Rotterdam, Netherlands, A. A. Balkema Publishers, pp. 929-936. Guedes Soares, C., and Garbatov, Y., 1998b, Reliability of Plate Elements Subjected to Compressive Loads and Accounting for Corrosion and Repair, Structural Safety and Reliability, ICOSSAR'97), Vol. 3, Rotterdam, Balkema, pp. 2013-2020. Guedes Soares, C. and Garbatov, Y., 1999a, Reliability of Maintained Ship Hull Girders of two Bulk Carriers Designs Subjected to Fatigue and Corrosion, Journal of Ship and Ocean Technology, Vol. 3, No. 1, pp. 27-41. Guedes Soares, C., and Garbatov, Y., 2000, Reliability of Maintained Marine Structures, Proceedings of the 6th International Conference Black Sea, pp. 227-237. Hamada K., Fujimoto Y. and Shintaku E., 2000, 2001, Studies on Ship Inspection Supporting System by the use of Product Model (lSt' 2nd and 3rd Report), Journal of The Society of Naval Architects of Japan, Vo1.186, pp.611-619, Vo1.188, pp.409-418, and Vol.190. pp.439-447. (in Japanese). Hamada K., Fujimoto Y. and Shintaku E., 2002, Ship inspection support system using a product model, Journal of Marine Science and Technology, Vol.6. P.205-215. Hansen, K.A., 1995, Evaluating Technology for Marine Inspectors, U.S. Coast Guard R&D Center. Holmes J. M., P. Daniel T., 1998, Human Elements in Bulk Carrier Inspections and Repair, Proceedings of Ship Structure Symposium'96 "Human and Organization Error in Marine Structures". - SNAME.- P. J-1 - J-21. Holzman, R., 1992, Advancements in Tankship Internal Structural Inspection Techniques, Dept. of Naval Architecture & Offshore Engineering, Univ. of California, Berkeley. HSE, 1992, Fatigue Background Document, Her Majesty's Stationary Office, London, Health and Safety Executive Report OTH 92 390. Hu, T. and Bea, R. G., 1997, Transactions of the American Society of Mechanical Engineers (ASME). Huang P, Li Z. and Sun J., 2000, Finite element analysis on evolution process for damage micro crack healing, Acto Mechanica Sinica, Vol.16, No.3,254-263 Huang, R. T., McFarland, B. L., and Hodgman, R. Z., 1997, Microbial Influenced Corrosion in Cargo Oil Tanks of Crude Oil Tankers, Proceedings of Corrosion "97. Iaccarino, R., Monti, S., Sebastiani, L., 2000, 'Evaluation of Hull Loads and Motion of a Fast Vessel Based on Computations and Full Scale Measurements', NAV 2000 - Vol. I, Venice, Italy. IACS, 1998, Repair and Maintenance Quality Standard. IACS, 1999, General Cargo Ships, Guidelines for Survey, Assessment and Repair of Hull Structures. IACS, 2002, Bulk Carriers, Guidelines for Survey, Assessment and Repair of Hull Structures. Second edition. IIMS, 2001, Guidelines for Surveyors Conducting On-Hire Vessel Surveys (Dry Cargo). IIMS, 2001, Guidelines for Surveyors Conducting Pre-Purchase Vessel Condition Surveys IMO, 1993, Resolution A744 (18), Guidelines on the Enhanced Program of Inspections during Survey of Oil Tankers and Bulk Carriers. Ishizuka T. and Takemura M., 1998, Strain Measurement and Crack Detection by Optical Fiber Sensors, Journal of the Society of Naval Architects of Japan, Vol. 183, pp.399-405. Jiao, G. and Moan, T., 1992, Reliability-Based Fatigue and Fracture Design Criteria for Welded Offshore Structure, Engineering Fracture Mechanics, Vol. 41, No. 2, pp. 271-282. Johannesen J.M., Moan T. and Vardal O.T., 2000, Application of Probabilistic Fracture Mechanics for Reassessment of a Floating Production U n i t - Theory and Validation, OMAE 2000, New Orleans, Louisiana. Kageyama K. and Murayama H., 2001, Structural Monitoring by Using Fiber-optic Technology. Journal of the Japanese Society for Non-Destructive Inspection, Vol.50, No.9, pp.595-600. Kageyama K, Kimpara I, Suzuki T, Ohsawa I, Shimamura T, 1995, Fundamental Study on Smart Structure Approach to Marine Structure, (Partl, Novel Methods for Fiber-Optic Measurement of Displacement and Damage), Journal of the Society of Naval Architects of Japan, Vol. 178, pp.583-591. (in Japanese)

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Kai-tung Ma, and Bea, R. G., 1995, Fatigue Life Estimation for Repaired Ship Critical Structural Details, Proceedings of the 15th International Conference on Offshore Mechanics and Arctic Engineering (OMAE'95), Vol. ffl, ASME, New York, USA, pp. 193-202. Kawamura Y., et al., 1998, A Study on the Information System for Damages of Ship Structures. Journal of the Society of Naval Architects of Japan, Vol. 184, pp.523-532.(in Japanese) Kawamura Y., Sakuragi T. and Sumi Y., 2000, A Study of a Ship Inspection Supporting System by Using a Portable Computer Yasumi Journal of the Society of Naval Architects of Japan, Vol. 188, pp.737-743. (in Japanese). Kawamura Y., Sumi Y. and Seki T., 1998, A Study of the Information System for Damages of Ship Structures, Journal of the Society of Naval Architects of Japan, Vol.184, pp.523-532. Kazuro Kageyama, Hideaki Murayama, 2001, Structural Monitoring by Using Fiber-Optic Technology, Journal of the Japanese Society for Non-Destructive Inspection, Vol.50, No.9, pp. 595-600.(in Japanese) Kmiecik, M., Jastrzebsky, T. and Kuzniar, J., 1995, Statistics of Ship Plating Distortions, Marine Structures, Vol. 8, No. 2, pp. 119-132. Kozlyakov V.V. and Stankov B.N., 2000, Assessment of technical condition of steel floating docks hulls. RS Scient.-techn. collect. Issue 23, Vol. 1, pp. 8 4 - 92, St.-Petersburg. Kulesh V.A., et al., 1996, Unified Computer System for Acquisition, Storage, Processing and Analysis of Ship Hulls Fault Detection. DEFHULL. USER'S GUIDE. UG95-5C. NTB-01, Vladivostok. Lassen, T., 1991, Markov Modelling of the Fatigue Damage in Welded Structures under In- service Inspection, International Journal of Fatigue, Vol. 13, No. 5, pp. 417-422. Liu CT, Qin TY and Duan ML, 2000, Finite element analysis of the cracked members of offshore platform structures, Ocean Engineering, Vol. 18, No. 3, pp. 15-19(in Chinese) Liu CT, Qin TY and Duan ML, 2002, Finite element analysis of the deformed legs of offshore platforms, China Ocean Engineering, Vol. 16, No. 3, in press Lotsberg, I., 1997, Evaluation of SN Curves for Fatigue Design, Report No 97-3566, Det Norske Veritas, Oslo. Lu, P., Zhao, B. and Yan, J., 1998, Efficient Algorithm for Fatigue Life Calculations under Broad Band Loading Based on Peak Approximation, Journal of Engineering Mechanics, Vol. 124, No. 2, pp. 233-236. Ma K-T., 1998, Tanker Inspections and a Risk Based Approach, (ABS Americas), ISOPE,. Madsen, H. O., and Sorensen, J. D., 1994, Probability Based Optimisation of Fatigue Design, Inspection and Maintenance, Mobil Offshore Structures, London, UK, Ocean Engineering Research Centre, Department of Civil Engineering, and A. J. Edwards- Ocean Technology (UK) Masuda T., Machida S., Yoshinari H. and Enami K., 1998, Fracture Strength Assessment of Cracked Structures by Ultrasonic Testing, Journal of the Society of Naval Architects of Japan, Vol.184, pp. 443-452. Matoba, M., Kumano, A., and Yamamoto, A., 1994, Based of Design and Planning for Maintenance of Hull Planning of Inspection and Life Design for Preventive, NK Technical Bulletin, pp. 47-63. Melchers, R., and Ahammed, M., 1994, Non-linear Modelling of Corrosion of Steel in Marine Environments, Research Report. 106.09.1994, The University of Newcastle, New South Wales, Australia. Mizukami, T., Ishikawa, I., and Yuasa, M., 1994, Trends of Recent Hull Damage and Countermeasures, NK Technical Bulletin. Moan T., Johannesen J..M. and Vardal O.T., 1999, Probabilistic Inspection Planning of Jacket Structures, OTC 1999, Paper 10848, Houston, Texas. Monti S. and Grossi L., 1999, Definition of a Monitoring System to Collect Structural Data on a New Fast Ship in Service, HSMV'99 Conference. Capri, Italy. Moskvichev VV, Lepikhin AM and Doronin SV, 1999, Simulation in fracture of welded structures with developing damages, International Journal of Fracture 100, pp. 143-153 Murayama H, Kageyama K, Kimpara I, Suzuki T, Ohsawa I and Kanai M, 1997, Fundamental Study on Smart Structure Approach to Marine Structure, Part2, Monitoring Structural Integrity of a Ship

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Using Fiber Optic Strain Sensors in Field Tests), Journal of the Society of Naval Architects of Japan, Vol. 182, pp.579-587. (in Japanese) Murayama H., Kageyama K., Kimpara I., Suzuki T., Ohsawa I. and Kanai M., 1998, Fundamental Study on Smart Structure Approach to Marine Structure, Part 3, Monitoring Structural Integrity of a Ship Using Fiber Optic Strain Sensors in Field Tests, Journal of the Society of Naval Architects of Japan, Vol.184, pp.513-522. Nielsen, K.G., Bryndum, M., Dal, E.G., Pedersen, B., Andreasen, K.K., 1995, Full-Scale Measurements of Wave Load and Wave Kinematics, OMAE 1995, Vol. I-A, pp.503-512. OCIMF & SIGTTO, 1998, Inspection Guidelines for Ships Carrying Liquefied Gases in Bulk 2nd Ed. OCIMF, 2000, Vessel Inspection Questionnaire for Bulk Oil, Chemical Tankers and Gas Carriers Orisamolu I.R and Bea R.G., 1999, Optimal Strategies for Inspection of Ships for Fatigue and Corrosion Damage, US Ship Structure Committee Report SSC407. Ostergaard, C., Dogliani, M., Guedes Soares, C., Parmentier, G. and Petersen, P. T., 1996, Measures of Model Uncertainty in the Assessment of Primary Stresses in Ship Structures, Marine Structures, Vol. 9, No. 3-4, pp. 427-448. Paik J.K., Thayamballi A.K., 1998, The Strength and Reliability of Bulk Carrier Structures Subject to Age and Accidental Flooding, SNAME Transactions, Vol. 106, pp. 1-40. Pirker,G. Fujimoto,Y., Samudro, 2000, A simple Probabilistic Model for Corrosion Evaluation of a Ship's Hull, J. of The Society of Naval Architects of Japan, Vol.188, pp.633-644. Pittaluga, A., Cazzulo, R., Romeo, P. Skjong R. K. and Torhaug, R., 1991, Uncertainties in the Fatigue Design of Offshore Steel Structures, Marine Structures, Vol. 4, pp. 317-332. Reeve H.P. and Bea R.G., 1998, Ship Structural Integrity Information System (SSIS), Phase 3, Ship Quality Information System, US Ship Structure Committee Report SSC404. Schall, G., and Ostergaard, C., 1991, Planning of Inspection and Repair for Ship Operation, Proceedings of the Marine Structural Inspection, Maintenance, and Monitoring Symposium, SNAME, VF1-VF7. Schulte-Strathaus, R., 1991, Department of Naval Architecture and Offshore Engineering, University of California, Berkeley. Shen ZH, Li H. and Xue YN, et al., 1998, The Wll-4 Platform comprehensive strength monitoring system, China Ocean Engineering, Vol.12, No.2, 135-146 Shi, W. B., 1993, In-service Assessment of Ship Structures: Effects of General Corrosion on Ultimate Strength, Transactions Royal Institution of Naval Architects (RINA), Vol. 135, pp. 77-91. Shintaku E., Fujimoto Y., Hamada K. and Takeuchi T., 2000, Study of a Simple Sensor for Stress History Measurements of a Structural Member using Piezoelectric Element, Journal of Marine Science and Technology, Vol.5, Issue 1, pp.40-47. Shintaku E., Fujimoto Y., Hamada K., Takeuchi T.and Takeyabu N., 1999, Study on Simple Sensor for Stress History Measurement of Structural Member using Piezoelectric Element, (Part 2 Investigation on Installation of the Piezoelectric Element to Structures). Journal of the Society of Naval Architects of Japan, Vol.186, pp.401-411. Slaughter S.B., Cheung M.C., Sucharski D. and Cowper B., 1997, State of the Art in Hull Response Monitoring Systems, US Ship Structure Committee Report SSC401 Spencer, J. S., Mansour, A. E., Chen, H. H., and Luckett, M. D., 1996, Reliability-Based Comparative Study of Classification Rules, Proceedings of the 16th International Conference on Offshore Mechanics and Arctic Engineering (OMAE'96), Vol. II, ASME, New York, USA, pp. 207-216. Stacey, A., and Sharp, J. V., 1995, The Revised HSE Fatigue Guidance, Proceeding of 14th OMAE, Vol. Ill, pp. 1-16. Stacey, A., Burdekin, F. M., and Maddox, S. J., 1996, The revised BS PD 6493 Procedure- Application to Offshore Structures, Proceedings of 15th OMAE, Vol. 1II, pp. 13-33. Straumann P., 1997, Ageing vessels- experience from CAP as a condition assessment tool. Proc. of The Ships Ageing Process Conference (IMAS 97), London. Tanker Structure Cooperative Forum, 1992, Condition Evaluation and Maintenance of Tanker Structures, Witherby & Co. Ltd., London.

Inspection and Monitoring

69

Thoft-Christensen, P, and Sorensen, J. D., 1987, Optimal Strategies for Inspection and Repair of Structural Systems, Civil Engineering Systems, Vol. 4, pp. 94-100. Torng, T. Y. and Wirsching, P. H., 1991, Fatigue and Fracture Reliability and Maintainability Process, Journal of the Structural Engineering, Vol. 117, pp. 3805-3821. TSCF & IACS, 1997, Guidance Manual for Tanker Structures. Witherby & Co., London. TSCF, 1992, Condition Evaluation and Maintenance of Tanker Structures. Witherby & Co., London. TSCF, 1995, Guidelines for the Inspection and Maintenance of Double Hull Tanker Structures. Witherby & Co., London. Tseng SS, 2000, Damage assessment of linear structures by a static approach, Numerical simulation studies, Structural Engineering and Mechanics, Vol.9, No.2, 195-208 Tseng SS, 2000, Damage assessment of linear structures by a static approach, Theory and formulation, Structural Engineering and Mechanics, Vol.9, No.2, 181-193 Vardal O.T. and Moan T., 1999, Validation of Inspection Planning Methods, OTH 1999-056, Health & Safety Executive UK, London. Vardal O.T., Moan T. and Johannesen J.M., 2000, Application of Probabilistic Fracture Mechanics for Reassessment of a Floating Production Unit- Philosophy and Target Levels. OMAE 2000, New Orleans, Louisiana. Wang, Z., Folsr R., Bondini, F. and Pedersen, T., 1999, Linear and non-linear Numerical Sea-keeping Evaluation of a Fast Monohull Ferry Compared to Full Scale Measurements, FAST 1999 Conference, Seattle, USA, pp. 443-456. White, G. J., and Ayyub, B. M., 1992, Determining the Effects of Corrosion on Steel Structures: A Probabilistic Approach, Proceedings of the 1lth International Conference on Offshore Mechanics and Arctic Engineering (OMAE'92), Vol. II, New York, USA, ASME, pp. 45-52. Wirsching, P. H., Karsan, D. I., and Hanna, S. Y., 1995, Fatigue Fracture Reliability and Maintainability Analysis of the Heidrun TLP Tether System, Proceedings of the 15th International Conference of Offshore Mechanics and Arctic Engineering (OMAE'95). Wirsching, P. H., Torng, T. Y., Geyer, J. F. and Stahl, B., 1990, Fatigue Reliability and Maintainability of Marine Structures, Marine Structures, Vol. 3, pp. 265-284. Witmer D.J and Lewis J.W., 1995, The BP Oil Tanker Structural Monitoring System, Marine Technology, Vol. 32, No.4, Oct. 1995, pp.277-296. Yamamoto,N., Ikegami, K., 1998, A Study on the Degradation of Coating and Corrosion of Ship's Hull Based on the Probabilistic Approach, J. of Offshore Mechanics and Arctic Engineering, Vol.120, NO.3, pp.121-128. Yamato H., et al., 2001, Application of the wearable system to shipbuilding industrial engineering, Journal of the Society of Naval Architects of Japan, Vol. 190, pp.431-438, (in Japanese) Zhang W., Chen Y. and Jin Y., 2000, A study of dynamic responses of incorporating damaged materials and structures, Structural Engineering and Mechanics, Vol.10, No.2, 139-156 Zhao A. and Yu J., 2000, The overall elastic moduli of orthotropic composite and description of orthotropic damage of materials, International Journal of Solids and Structures 37, pp. 6755-6771 Zhao, W., and Baker M. J., 1995, Probabilistic Sensitivity Analysis of Materials for Structural Fatigue and Fracture, pp. 382-395.

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15tn INTERNATIONAL SHIP AND OFFSHORE STRUCTURES CONGRESS 2003 AUGUST 11-15, 2003 SAN DIEGO, USA

,,q A B ' J / . " l

"~k'~

VOLUME 2

COMMITTEE V.3

COLLISION AND GROUNDING

COMMITTEE MANDATE Concern for structural arrangements on ships and floating structures with regard to their integrity and adequacy in the events of collisions and grounding, taking into account the probabilistic and physical nature of such accidents. Consideration shall be given to the effectiveness of structural arrangements for reducing or avoiding pollution due to leakage, and to residual strength of damaged structures.

COMMITTEE MEMBERS

Chairman:

Prof. Prof. Prof. Dr. Prof. Dr. Mr. Prof. Prof. Dr.

Jeom Kee Paik Jorgen Amdahl Nigel Barltrop Edmond Remy Donner Yonging Gu Hisashi Ito Hans Ludolphy (Deceased) Preben Terndrup Pedersen Udo R~Shr Ge Wang

KEYWORDS

Ship collisions, ship grounding, accidental limit state (ALS) design, collision/grounding scenarios, external mechanics, internal mechanics, post-accident mechanics, oil outflow, residual strength of damaged ships, impact, crushing, tearing, raking, fracture, hull girder collapse.

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CONTENTS

1 INTRODUCTION

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2 C O L L I S I O N A N D G R O U N D I N G I N C I D E N T DATABASES 2.1 Historical Database . . . . . . . . . . . 2.1.1 Ship Collision Database . . . . . . . . . 2.1.2 Ship Grounding Database . . . . . . . . 2.2 Artificial Database . . . . . . . . . . . . 2.2.1 Ship Collision Database . . . . . . . . . 2.2.2 Ship Grounding Database . . . . . . . .

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3 A F R A M E W O R K F O R A C C I D E N T A L LIMIT STATE D E S I G N . . . . . 3.1 Principles o f the Accidental Limit State Design . . . . . . . . . . 3.2 Procedures for the Accidental Limit State Design . . . . . . . . . . 3.3 Accident Scenarios . . . . . . . . . . . . . . . . . . . . . 3.3.1 Ship Collision Scenarios . . . . . . . . . . . . . . 3.3.2 Ship Grounding Scenarios . . . . . . . . . . . . . . . 3.4 External Mechanics . . . . . . . . . . . . . . 3.5 Internal Mechanics . . . . . . . . . . . . . . . . . . . . 3.5.1 Fundamentals . . . . . . . . . . . . . . . . . . . . 3.5.2 Full Non-linear Finite Element Methods . . . . . . . . 3.5.3 Simplified Non-linear Finite Element Methods . . . . . 3.5.4 Simplified Analytical Methods . . . . . . . . . . . 3.5.5 Empirical Formulae . . . . . . . . 3.6 Post-accident Mechanics . . . . . . . . . . . . 3.6.1 Oil Outflow . . . . . . . . . . . . 3.6.2 Hull Girder Collapse . . . . . . . . 3.7 Design Criteria for Accidents . . . . . . . . . .

80 80 80 82 82 83 83 84 84 86 87 87 88 89 89 89 90

4 B E N C H M A R K STUDIES ON N O N L I N E A R FE ANALYSIS OF MECHANICS . . . . . . . . . . . . . 4.1 Tensile Coupon Test o f Mild Steel . . . . . . . 4.2 Collision Test . . . . . . . . . . . . . . 4.3 Grounding Test . . . . . . . . . . . . . .

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INTERNAL . . . . . . . . . . . . . . . . . .

5 ALTERNATIVES TO D O U B L E H U L L D E S I G N . . . . . . . . . . . . 5.1 Varying Structural Arrangements for Double Sides against Collisions . . . . . 5.2 Varying Double Bottom Heights along the Ship Length against Grounding . 5.3 Mid-deck Tankers . . . . . . . . . . 5.4 Soft Bow Design against Collisions . . . . . 73

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Specialist Committee V.3

74

6 PROPOSALS FOR FURTHER RESEARCH AND DEVELOPMENTS 7 CONCLUSIONS REFERENCES

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FURTHER READING ACKNOWLEDGEMENTS

Collision and Grounding

75

1. INTRODUCTION Ship collisions and grounding continue to occur regardless of continuous efforts to prevent such accidents. With the increasing demand for safety at sea and for protection of the environment, it is of crucial importance to be able to reduce the probability of accidents, assess their consequences and ultimately minimize or prevent potential damages to the ships and the marine environment. The Committee report surveys recent research and developments undertaken in the areas of ship collisions and grounding, and addresses the insights developed in the literature. For the sake of completeness in terms of describing accident mechanics, some old publications are also reviewed and included in the report, as necessary. Of primary concern in the report is the research works on structural performance of ships against such accidents. The insights developed from recent studies do indicate that direct Accidental Limit State (ALS) design is a good concept and the structural crashworthiness during collisions and grounding should be taken into consideration in designs. The historical and numerically computed artificial databases of ship collision and grounding damage are surveyed in the report. A framework and procedures for the ALS design are addressed in terms of (1) definition of accident scenarios, (2) analysis of external accident mechanics, (3) analysis of internal accident mechanics, (4) analysis of post-accident mechanics, and (5) design criteria against accidents. While the finite element method (FEM) is a powerful tool to simulate the internal accident mechanics, attention should be paid to finite element size, fracture criteria and material properties so that more reliable simulation results can be obtained. A benchmark study on test structural models is undertaken by the committee members to provide some useful guidelines for the finite element analysis of internal accident mechanics. Design alternatives to the double hull concept are reviewed. Finally, further R & D topics on the ALS design of ships against collisions and grounding are proposed. contact 4%

~

w r e c stranded/ 24%

k

collision ~ 11%

others 19% e d / ~ ~------~\.~" foundered 27% 9, mlssmg fire/explosion 0% 15%

Figure 1: Total loss causes for all types of ships in gross tonnage during the years 1995-1998 (LR 2000)

2. COLLISION AND GROUNDING INCIDENT DATABASES Regulations for damage stability require probability density functions (pdfs) to describe the location and extent of damage. Historical databases of the accidents are very useful to define the probability density distributions of the accident scenarios and to establish accident damage stability standards. However, the size of the historical database is normally not sufficient to precisely define damage characteristics of ship structures. For instance, the probabilistic standards suggested by IMO (1989) use the pdfs derived from limited historical damage statistics, and do not include the influences of ship structural designs or structural crashworthiness on the damage extent. In this regard, mathematical ship models involving structural dimensions, building material and structural crashworthiness are often employed for generating artificial databases for various types of accidents.

76

Specialist Committee V.3

2.1 Historical Database

According to world casualty statistics (LR 2000), the total losses of all types of ships during the years 1995-1998 are 674 in number and 3.26 million in gross tonnage (GT). Figure 1 illustrates breakdown by total loss causes during the same period. Collisions, contact and grounding (wrecked/stranded) account for 39% of total losses. 2.1.1 Ship Collision Database

The damage probability statistics for IMO's 1971 passenger ship regulations A.265 (IMO 1971) were based on reports from member nations of the IMO in the form of "damage cards". Each of these damage cards represented specific technical information about a ship casualty. Most of these data were collected for casualties occurring in the 1950's and 1960's. It consists of 890 records in total and of those 624 records are collisions. Recently the European project HARDER has established a new updated damage database. This database consists of the "old" IMO data, a few updated IMO damage cards and data from the classification societies: DNV, GL, LR of Ship Repair Statistics and the Hellenic Register of Shipping (Tagg et al 2001, Laubenstein et al 2001). This combined HARDER database contains data on 1851 collisions and 930 groundings (Ltitzen 2001). Figure 2 shows the correlations between the striking ship length and the struck ship length or between the striking ship speed and the struck ship speed. It is seen that their correlations are quite scattered, as would be expected.

X

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Figure 2: The striking ship length versus the struck ship length (left) and the striking ship speed versus the struck ship speed (right) based on HARDER database (Ltitzen 2001) Figure 3 shows the probability density distributions for the penetration of the struck ship normalized by the ship breadth and for the collision point along the length of struck ship normalized by ship length. It may be surmised from Figure 3 (left) that the probability density distribution for the collision penetration follows the log-normal or Weibull function. Mean and COV of the probability density function for the normalized collision penetration is given in Figure 3 (left), when a Weibull function is adopted. It is seen from Figure 3 (fight) that the collision location along the struck ship length is more likely to occur at bow and stem. It is noted that the pdfs of the collision damage extent shown in Figure 3 do not identify the structural design characteristics of ships involved.

Collision and Grounding

77

10.0 2.5-

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Figure 3: Probability density distribution: penetration in ship collisions normalized by ship breadth (left) and collision point along the length of the struck ship normalized by ship length (right) 30.0 -

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Figure 4: Probability density distribution: raking damage length normalized by ship length grounding damage height normalized by ship draft (right)

(left)

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Figure 5" Probability density distribution for grounding damage width normalized with ship breadth

Specialist Committee V.3

78

2.1.2 Ship Grounding Database The HARDER database for ship grounding is smaller than the one for collisions. For grounding, the database consists of 930 incidents. Figures 4 and 5 show the probability density distributions for raking damage length normalized by ship length, for grounding damage height normalized by ship draft, and for grounding damage width normalized by ship breadth. The probability density distributions follow the log-normal or Weibull function. Related mean values and COVs are indicated in the figures when the Weibull function is applied. It is again noted that the pdfs of the grounding damage extent shown in Figures 4 and 5 do not account for the structural design characteristics of ships involved.

2.2 Artificial Database 2.2.1 Ship Collision Database Monte Carlo simulation can be applied to develop alternative forms of probability density functions (pdfs), or numerically derived artificial damage databases of the accidents. The process begins with a set of the pdfs defining possible accident scenarios. Using these pdfs, a specific accident scenario is selected in the Monte Carlo simulation and combined with a specific structural design to predict the accidental damage. This process is repeated for a large number of accident scenarios and structural designs until sufficient data are generated (Pedersen et al 1996, Brown & Chen 2002). ...,e- S H 1 5 0 - 2522 cases no

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(m)

Figure 6" Probability density distribution for collision penetration of four struck tankers obtained by the Monte Carlo simulation approach (Brown & Chen 2002) Figure 6 shows the probability density distributions of the maximum collision penetration so obtained by Brown &Chen (2002). Four struck tankers, namely 150k dwt double hull, 150k dwt single hull, 45k dwt double hull and 45k dwt single hull, were analyzed. The damage extent in the pdfs shown in Figure 6 is not normalized by ship breadth so that it would be easier to identify the structural performance of the ships against collisions. In the simulation, the pdfs generated from the historical collision database were used to develop 10,000 collision scenarios that were applied to the four struck tanker designs for a total of 40,000 runs of SIMCOL, which is a computer program for ship collision analysis based on the simplified analytical methodology (Chen 2000). The IMO guidelines (IMO 1995) consider that all tankers have the same level of non-dimensional damage probability distribution, regardless of ship designs. Pedersen & Zhang (2000) extended the IMO guidelines for oil tankers and conventional merchant ships. They include structural dimensions and the building material as parameters of influence. Two groups of ships were considered, namely the small group with the ship length below 100 m and the large ship group with the ship length above 100 m. The average ship length of the large and small ship groups are 135m and 65m, respectively. It was

79

Collision and Grounding

found that the small ship group has a higher probability of larger collision damage, relative to the ship length, than that of the large ship group. 2.2.2 Ship Grounding Database

Pedersen & Zhang (2000) also extended the IMO guidelines to the probabilistic damage distribution of bottom grounding damage for oil tankers and merchant ships, including structural dimensions and building material of the ships. Figure 7 shows the probability density distribution of the bottom grounding damage length for oil tankers and merchant ships that they obtained. Again, two groups of ships were considered, namely the group of small ships and the group of large ships, as defined in Section 2.2.1. It is seen from Figure 7 that the large ship group has a higher probability of larger bottom raking damage extent than that of the small ship group. The probability that the longitudinal grounding damage extent is above 30% of the ship length is about 25% for the large ship group, while it is about 9% for the small ship group. 7....

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Figure 7: Probability density distribution of bottom grounding damage length for oil tanker and merchant ships (Pedersen & Zhang 2000) A similar effort was also made by Zhu et al (2002) to generate a probability density distribution of grounding damage for a single hull VLCC. Figure 8 shows the probability density distribution of the bottom raking damage for the single hull VLCC, obtained by 26,000 Monte Carlo simulations. The historical damage data of 62 cases and the related IMO standard are also plotted in Figure 8. For the present single hull VLCC, it is apparent that the probability density distribution of the IMO standard correlates well with that of both historical and artificial databases. Simonsen (2000) derived simplified analytical expressions to predict the raking damage of fast crafts as a function of vessel size. 4.5 4.0

3.5 3.0 2.5 ~

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i Calculated ~ I'--'-IStatistics ~ I M O requirements

~

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a. 0.5 0.0

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0.3 0.4 0.5 0.6 0.7 0.8 0.9 (Grounding dam age length)/(Ship length)

1.0

Figure 8" Comparison of the probability density distributions of bottom raking damage for a single hull VLCC, as those obtained by historical database, IMO standard and Monte Carlo simulation (Zhu et al 2002)

80

Specialist Committee V.3

3. A F R A M E W O R K FOR ACCIDENTAL LIMIT STATE DESIGN

3.1 Principles of the Accidental Limit State Design A limit state is formally defined as a condition for which a particular structural member (or an entire structure) fails to perform the intended function. From the standpoint of a structural designer, four types of limit states are usually considered, namely the serviceability limit state (SLS), the ultimate limit state (ULS), the fatigue limit state (FLS) and the accidental limit state (ALS) (Paik & Thayamballi 2002). In design, these various types of limit states may be considered against different safety levels: the actual safety level to be attained for a particular type of limit state is a function of the perceived consequences. The primary aim of the ALS design for ship structures may be characterized by the following three broad objectives, namely 9to avoid loss of life, 9to avoid pollution of the environment, and 9to minimize loss of property or financial exposure. The structural design criteria for the ALS are based on limiting accidental consequences such as structural damage, stability, fire, explosion, and environmental pollution. Since the structural damage characteristics depend on the accident, it is not straightforward to establish universally applicable criteria and design accident scenarios. As viable altematives to prescriptive criteria, performance based criteria or risk assessment approaches may be used. The main safety functions of a structure during any accident event or within a certain time period after the accident generally include usability of escape ways, integrity of shelter areas, local/global load-bearing capacity, adequate residual strength, sufficient damage stability, containment of pollutant cargoes, and integrity of environment. Therefore, the ALS based design criteria should be formulated so that the main safety functions mentioned above shall work successfully. Typically either one or any combination of the following factors are considered: 9Energy dissipation related to structural crashworthiness, 9Structural capacity of local strength members or structures, 9Structural capacity of the global structure, 9Allowable tensile strains to avoid tearing or fracture, and 9Endurance for fire protection. While different types of accidental events normally require different analysis methodologies, the resistance evaluation for merchant cargo ship structures against collision and grounding have thus far typically been based on the energy dissipation related to structural crashworthiness.

3.2 Procedures for the Accidental Limit State Design To reduce the probability of outflow of hazardous cargo in ship collisions and grounding, the kinetic energy loss during the accident should with a high probability be entirely absorbed before the cargo hold boundary is punctured. Of crucial importance then is how to design the ship such that the initial kinetic energy is effectively consumed and the structural performance against an accident is maximized.

Collision and Grounding

81

Damage from an accident may be to the vessel or to the environment. The potentially costly consequences to property, material, environment, related industries, and public perceptions have been the driving forces for development of standards for design against accidents. A standard for design against accidents must include the following four items: 9Definition of accident scenarios, 9Methods for assessing structural performance, 9Methods for evaluating post-accident mechanics, and, 9Criteria for approval or acceptance of a design.

Figure 9: A sample of the ALS design procedures for ships against collision and grounding (Amdahl & Kavlie 1995) Figure 9 is a sample design procedure of ship structures against collisions and grounding (Amdahl & Kavlie 1995). In the first step, relevant scenarios of the accident are defined based on statistics of historical data and/or using risk analysis. In the second step, the structural performance (e.g., energy absorption capability) against design accident events is assessed. In the third step, the post-accident effects such as hull girder collapse and oil outflow are evaluated. In the last step, the acceptability of the design is judged based on acceptance criteria for oil outflow, hull girder fracture and so on. Pedersen (1995), Sirkar et al (1997), Brown et al (2000) and Wang et al (2002c) also address the framework for assessing performance of ships in collisions and/or grounding.

82

Specialist Committee V.3

3.3 Accident Scenarios

Accident scenarios define the accident events to be considered in design. They are a description of typical accidents and their associated probability distributions. Accident scenarios may be determined from statistics of past accidents (historical incident databases) and/or using risk analysis (artificial incident databases). Historical collision/grounding databases include Minorsky (1959), ORI (1980, 1981), Engineering Computer Optecnomics (1996), Sandia National Laboratories (1998), Simonsen & Hansen (2000), world casualty statistics (LR 2000) and the European project HARDER (Tagg et al 2001, Laubenstein et a12001). To minimize the danger of misinterpretation of the historical database, it is important to realize that (1) statistics of past accidents may not perfectly reflect the present accident situations, and (2) damage statistics may penalize designs which have been able to resist such damage. In this context, combining historical data, risk analysis and expert opinion may be the best way of defining accident scenarios (Wang et a12002e). 3.3.1 Ship Collision Scenarios

Ship collisions can be classified into two groups: side-collision and head-on collision. The former includes a ship's side struck by another ship or a ship's side bumping a fixed structure such as an offshore platform, while the latter indicates that the ship's bow strikes a large stationary object (e.g., pier, bridge abutment, offshore platform, iceberg or another ship). Parameters needed for defining collision scenarios include both ships' speed, collision angle, collision locations, cargo loading conditions, drafts (laden and ballast usually considered)/trims and striking bow geometry. The flare of a striking ship bow is often modelled as a wedge characterized by spreading angle, wedge length and tip radius. A more refined model for conventional ship bows is a half pyramid defined by stem angle (in vertical plane), spreading angle (in waterline plane) and breadth (Akita et al 1972, McDermott et al 1974, Nagasawa et al 1977, Zhu 1990, Ueda et al 1995, Zhang 1999, Brown et al 2000, Brown 2002). Statistics show that over 40-50% of all types of ships have bulbous bows which are fitted in conjunction with fine hollow waterline entrance to achieve minimum resistance. Bulbous bows may be modelled as conical strikers or elliptic parabola or real shapes (Suhara et al 1970, Ito et al 1984, 1986, Paik et al 1999, Lee et al 1999, Wang et al 2000a). It has been demonstrated that the energy absorption and fracture initiation characteristics of a double hull in collisions are quite different for different bow forms since the geometry of the striking bow has a direct influence on the behaviour of the struck ship: a small spherical bow will cause fracture under the contact area whereas a large spherical bow causes fracture at adjacent hard points (Wang et a12000a). For head-on collision scenarios, it is usually considered that the ship bow strikes rigid objects (Hagiwara et al 1983, Ito et al 1984, Pedersen et al 1993, Kierkegaard 1993, Ohtsubo & Suzuki 1995, Lehmann & Yu 1995, Wang & Ohtsubo 1997, 1999, Paik et al 1999, Kitamura 2000). A risk analysis identifies expected hazards to ships, human beings and the environment. Results of the risk analysis can aid in defining critical accident scenarios that may jeopardize the safety of ships and the environment involved. The risk analysis may include hazard identification, criticality ranking, consequence assessment and risk reducing measures (Fujii et al 1974, Pedersen et al 1993, Gluver & Olsen 1998, Urban et al 1999, Kaneko 2002).

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83

3.3.2 Ship Grounding Scenarios

Parameters for defining grounding scenarios include ship speed, cargo loading condition, draft/trim, sea floor condition, grounding location, water depth and tidal condition. Grounding may be further divided into three groups: raking, stranding and grounding on a sloping sea floor. Stranding is a situation where a ship comes to rest upon a seabed. Seabed reactions are primarily applied in the direction perpendicular to ship bottom plane. Raking is a situation when a ship runs onto protruding rocks for a distance before it stops. Reaction forces from the rocks are in this case along the longitudinal direction of the grounded ship. Grounding on sloping sea floor may be divided into two phases. In the first phase, the ship is subjected to an impulse caused by the sudden contact with the ground, while in the second phase, the ship is sliding with continuous contact with the sloping sea floor. Commercial ships sail mostly at a manoeuvring speed or a cruising speed. A ship's speed in raking is likely between these two speeds (Rawson et al 1998). If a ship runs aground at high tide, the ship may be subjected to larger grounding damage and hull girder loads as the tide drops and buoyancy is lost. The grounding damage may become greater with the increases in longitudinal bending and sheafing forces. For raking type of grounding accident, rocks may be modelled as wedges characterized by spreading angle, inclination and tip radius (Woisin 1979, Jones & Jouri 1987, Lu & Calladine 1990, Wierzbicki & Thomas 1993, Wierzbicki et al 1993, 1993, Astrup 1994, Paik 1994, Ohtsubo & Wang 1995, Simonsen 1997, Wang & Ohtsubo 1997, Zhang 2002, Wang 2002). For stranding, underwater obstructions are often modelled as rigid cones characterized by spreading angle and tip radius of the nose (Kitamura et al 1978, Arita & Aoki 1985, Rodd & Phillipes 1994, Vredvelt & Wevers 1995, Amdahl & Kavlie 1995, Mizukami & Tanigawa 1995, Kuroiwa 1996). Wang et al (2000a) demonstrate that the shape of the obstruction has a significant effect on the damage to the ship. Ravn et al (2001) analyse grounding of high-speed monohulls and catamarans using closedform solutions that include rupture. In grounding on a sloping sea floor, some of the kinetic energy not spent by the end of the first phase noted above becomes potential energy and some is dissipated through friction and deformation if the sea bed is soft. The sloping sea floor may be modelled as an inclined plane. The soil-ship hull interaction can be characterized by a coefficient of friction and a permeability coefficient of soil or rock (Pedersen 1994, Simonsen & Pedersen 1995, Sterndorff & Pedersen 1996, Simonsen 1997).

3.4 External Mechanics

Analysis of the accident mechanics can be classified into two parts, namely the external and internal mechanics. The external accident mechanics deals with the rigid body global motion of the ship(s) under the action of the collision or grounding force and the hydrodynamic pressures acting on the wet surface. The hydrodynamic forces can be split into inertia forces (added mass), the restoring forces (buoyancy-gravity), viscous drag forces and wave forces (Petersen 1982, Pedersen 1995, Suzuki et al 2000, Le Sourne et al 2001). The inertia forces increase the effective kinetic energy of the ships. They are normally calculated using potential flow theory and are generally frequency dependent. In many cases (notably collisions) the motion of the struck ship during the contact phase is small, and the inertia forces are the most important contribution. In simplified analysis of collisions, this is usually represented by a constant added mass term (Minorsky 1959, Pedersen & Zhang 1998). Gu & Wang

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Specialist Committee V.3

(2001) studied the importance of fluid- structure interaction in the case of a bulk carrier rammed by a containership using MSC/DYTRAN. They showed that viscous effects are of minor importance and that an inertia equivalent model (constant added mass) may be used instead of a considerably more computationally demanding integrated model. If the ship is subjected to large motions as a result of the accident, viscous forces and wave forces (which have a "memory" effect) become important. This is the case for ship-submarine impact (Le Sourne et a12001).

3.5 Internal Mechanics 3.5.1 Fundamentals The internal accident mechanics includes evaluation of the structural failure response of the involved ships during the accident. The internal mechanics of collisions and grounding are quite complex. Deformations many times larger than the structural thickness may take place, and the major part of energy dissipation takes place in inelastic straining. The inelastic straining tends to take place in relatively localized regions. Important differences exist between the most common mechanisms. Members subjected to lateral loading (e.g., side plating in collisions, bottom plating in grounding/stranding) will deflect laterally and develop in-plane (membrane) forces during finite deformations. As a result, the energy dissipation capability may become very large. The deformations are, however, limited by finite ductility; once the critical strain level is exceeded, fracture is initiated, resulting in a loss of resistance. Members subjected to axial crushing have a smaller energy absorption capacity (per unit volume). On the other hand, the resistance is not so sensitive to fracture for laterally loaded members. Tearing or cutting are modes associated with relatively little energy dissipation, because it takes place over localized regions. It occurs after initiation of fracture or may be directly enforced on the member, e.g., by a pinnacle ridge in the case of grounding. The initiation and growth of fracture is a local phenomenon, and is influenced by local details such as weld toes, any existing cracks or defects, and constraints to deformation. There are two dynamic aspects to account for, i.e., the effects of inertia and strain rate. Inertia effects may be further subdivided into two classes, namely impact loads and rapidly applied loads. In the former class, stress wave propagation has to be considered, so that at any instant the stress and strain distribution is inhomogeneous (Jones 1997, Paik & Thayamballi 2002). In most collision and grounding events, the load is applied relatively slowly, so that the time to reach the critical load is much larger than the time for a stress wave to travel from one end of the member to the other. Some of the kinetic energy transforms into kinetic energy related to the deformation mechanisms. This effect is normally small, and the energy must, ultimately, transforms into strain energy. For members subjected to axial crushing, inertia forces may affect the failure mode. Considering, however, the fairly low velocity and the gradual increase of forces in bow collision, this effect is often neglected in analysis. Inertia force effects may also reduce the global, horizontal bending resistance of the hull girder because they can deform the shell plating for which a moderate inward displacement may have a detrimental effect on the build-up of membrane tension forces. Kitamura (2002) has shown that this is important when the ramming ship is large compared to the struck ship.

Collision and Grounding 3.5-

85

Fnrmll]a~

3.0- ...... ~ "

t/ C o w p e r & S y m o n d s (1957) for m i l d steel

,/

Paik et al. (1999) for h i g h tensile steel

9 s

2.5~Yvd/cry = 1 + ( ~:/40.4) '/5

2.01.5-

/"

---x- ~

.

-

,~ "" "~"

~

~

~''~

.

1.0XO'vd/Crv = 1 + ( ~ 13200) "5 0.50.0

'

0 .4

' '"'"1

'

' ' .....

10 .3

I

'

10 .2

' ' .....

I

........

10 "1

I

........

1

i

101

........

I

10 2

........

I

10 3

........

I

10 4

k (sec")

Figure 10: Dynamic yield strength, crvd (normalized by the static yield strength, cry ) plotted versus strain rate, ~, for mild and high tensile steels (Paik & Thayamaballi 2002) Another dynamic effect of significance is due to the strain rate. With the increase in the strain rate, the yield stress of material increases, while the ductility is reduced. The most commonly adopted formula to deal with the strain rate sensitivity on yield stress is one proposed by Cowper & Symonds (1957), which is given by crv-----~= 1.0+

(1)

cry

where C,

q = coefficients to be determined based on test data, cry, crvd = static or dynamic yield

stresses, /; - strain rate. Figure 10 plots the Cowper-Symonds formula together with the relevant coefficients for mild or high tensile steels. It is seen from Figure 10 that as the strain rate increases the dynamic yield stress increases. It is also found that for higher tensile steel the percentage increase of the ratio of the dynamic yield stress to the static yield stress is smaller than that for mild steel. It is noted that the Cowper-Symonds formula relates to yield stress; the strain rate dependence is smaller for the flow stress at large strains. The effect of strain rate was also studied by Reckling (1976) and Toni & Funahashi (1978). While both crushing effects and yield strength increase as the loading speed gets faster, any fracture or tearing of steel (and the welded regions) of a structure tends to occur earlier. The following approximate formula, which is the inverse of the Cowper-Symonds formula for the dynamic yield stress, is then useful for estimating the dynamic fracture strain as a function of the strain rate (Jones 1989), namely

[ / lq] -,

~_F~ _ 1.0+

(2)

gF

where eF, eFa = static or dynamic fracture strains. Methodologies for analysis of intemal accident mechanics may be categorized into four groups: full non-linear finite element methods, simplified non-linear finite element methods, simplified analytical methods and empirical formulae. The different types of accident events normally require different methodologies to analyze the crashworthiness of the structure.

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Specialist Committee V.3

3.5.2 Full Non-linear Finite Element Methods

The internal accident mechanics involve yielding, crushing, tearing or fracture. Any non-linear FEM mesh for simulating the internal accident mechanics needs to be fine enough so as to capture such highly non-linear characteristics. Recent advances in computers and calculation algorithms have made nonlinear finite element analysis a viable tool for assessing the internal mechanics of collisions and grounding. Two types of FE methodologies are relevant, namely implicit and explicit techniques. Implicit methodologies require the solution of systems of equations. This places demands on the equation solver and the computer capacity especially in terms of memory resources. Explicit systems do not require equation solving; equilibrium is solved at the element level. However, to comply with stability requirement for equation solving, very small time steps are needed. Explicit methodologies based computer codes include ABAQUS/Explicit, DYTRAN, LS-DYNA, PAM-CRASH and RADIOSS, and implicit methodologies based codes include ABAQUS/Standard, ANSYS, MARC and NASTRAN. There are two important factors to consider in structural modeling of the nonlinear finite element analysis, namely element type and mesh fineness. Higher order elements generally provides better accuracy and allow a less finer mesh, but they require more computational effort. The importance of mesh fineness or element types has been studied by many investigators (Amdal & Kavlie 1992, Kitamura 1997, 2002, Servis et a12002). It is observed that a very large number of elements is required in order to obtain accurate results for components deformed by axial crushing forces. Accounting for realistic size and boundary conditions of FE models has been discussed by Woisin (1999). Analytical formulae derived for evaluating structural damage characteristics (e.g., failure patterns) may be used to determine relevant mesh size. For instance, it is recommended that more than eight (rectangular plate-shell type) finite elements are necessary to capture the structural crushing pattern within a half length of one structural fold, see Figure 11. Available analytical formulations for predicting the length of structural fold are suggested by Amdahl & Kavlie (1992), Wierzbicki & Abramowicz (1983), Wierzbicki et al (1993), Paik & Pedersen (1995), Wang & Ohtsubo (1997) and Suzuki et al (2000). It is cautioned, however, that these formulae were derived for different crushing patterns to different structural geometries.

Figure 11: A thin-walled structure crushed under predominantly axial compressive loads and cut at its mid-section (Paik & Thayamballi 2002) A major challenge in nonlinear finite element analysis is the prediction and simulation of initiation and propagation of fracture. This is essential for in view of the magnitude of energy dissipation, notably for members subjected to extensive membrane stretching, while it is usually less important for axial crushing. The simplest approach to account for the effect of fracture is to remove elements once the critical strain is attained. This is fairly easily done in an explicit code because there is no need to assemble and invert the effective system stiffness matrix. However, deleting elements do not properly

Collision and Grounding

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deal with the aspect that the large stresses can develop around fractured regions, specifically in the direction parallel to the cracks. An improved modeling technique for fracture behavior may be to introduce a double set of nodes such that the elements are allowed to separate once the critical stress is attained (Amdahl & Stornes 2001). The effects of multi-axial state of stresses can be accounted for using approaches based on continuum damage mechanics (Lehmann & Yu 1998a, Zhu & Atkins 1998). The critical strain for fracture depends heavily on the stress-strain measure as well as the mesh size. Various options exist for the stress-strain relationship. Most often engineering stress-strain relationship or true stress-strain relationship is used. The true stress-strain relationship can model the physical process more accurately than the engineering stress-strain relationship, but it is more complex as the change of the element volume needs to be involved. Experience obtained thus far indicates that the difference between these two approaches can be neglected up to ultimate stress as long as the l o a d displacement relation and the associated energy dissipation are concerned. The dependency of fracture strain on element size has been studied by Yu (1996), Lehmann & Yu (1998b) and Simonsen & Lauridsen (2000). By comparing tensile tests with numerical simulations they concluded that a fairly small element mesh is required to capture the features of the tensile test. An empirical relationship between fracture strain and element size has been derived by Lehmann et al (2001), based on thickness measurements of structural elements from actual collisions. 3.5.3 Simplified Non-linear Finite Element Methods

The simplified non-linear finite element methods use a coarse mesh (a super-element approach) or the idealized structural unit method (Paik & Thayamaballi 2002). These approaches often combine the analytical models for crushing, tearing and yielding behaviour of structural components, and the ordinary finite element technique. The computing cost and modelling efforts of analysis are reasonably small, while the accuracy is not lost (Ito et al 1984, 1986, Bockenhauer & Egge 1995, Paik & Pedersen 1996, Paik et al 1999, Paik et al 2002). This group of techniques has the advantage of FEM in modelling of the interactions between structures involved, and also the merit of simplified analytical methods in dealing with complex damage behaviour. 3.5.4 Simplified Analytical Methods

These methods are developed by analytical approach resulting in closed-form analytical formulations which can capture the basic features of structural crashworthiness. Modelling and calculation effort of simplified analytical methods is much lower compared to non-linear FEM simulations. The development of the methods normally consists of four major steps, namely (1) identify primary damage patterns of structural components according to observation of actual damage, (2) develop idealized theoretical models and derive theoretical formulations to capture the main features of the damage patterns, (3) establish global models for the entire damage process of the ship hull, and (4) combine the global damage models with formulations for individual structural components (McDermott et al 1974, Yang & Caldwell 1988, Pedersen et al 1993, Paik & Pedersen 1995, Suzuki & Ohtsubo 1995, 2000, Wang & Ohtsubo 1997, 1999, Wang et a12000a, Simonsen 1997, Zhang 2002). The identified failure modes characterizing the behaviour of ship plates in collisions and grounding include cutting or tearing (Wierzbicki et al 1993, Paik 1994, Ohtsubo & Wang 1995, Paik & Wierzbicki 1997), concertina tearing (Wierzbicki 1995), folding (Wang & Ohtsubo 1997, Simonsen 1997), crushing (Wierzbicki & Abramowicz 1983, Paik & Pedersen 1995, Ohtsubo & Suzuki 1995, Wang & Ohtsubol997), ruptured plate (Wang et al 2000a) and membrane stretching (Yu 1996, Wang et al 1998b, 2002a, Zhang 2002). In addition to these failure modes, relevant topics include initiation of rupture (Zhu & Atkins 1998, Woisin 1998, Wang et al 2000a, Simonsen & Lauridsen 2000), boundary

88

Specialist Committee V.3

condition for the effect of adjacent structures (Woisin 1999, Kitamura 1999) and influences of geometry and size of striking structures (Wang et al 1998a, 2000a). The effect of combined biaxial and shearing crushing forces on the crushing strength of plated structures is dealt with by Paik & Pedersen (1996). The effect of dynamic loading is usually included by considering the material strain rate sensitivity. Benchmark studies on some selected simplified analytical methods for crushing and tearing of steel plated structures are made by Paik & Wierzbicki (1997). Surveys and reviews of research on this behaviour and the fundamental theory can be found in Wierzbiki & Thomas (1993), Zhang (1999) and Wang (2002). The studies of these failure modes contribute to the advance in the knowledge of the plastic behaviour of structures exposed to large impact loads, and also to the advances in the simplified analytical methods. 120[ | 100[

, W-50

,

,

]

c, : TEST --PREDICTION

| l

8~

1

.~ /

401

e

~ .,~~""~~g b

0 0

A' 50

~.~~h

, B , 100 Indentation (mm)

, 150

200

Figure 12: The progressive damage of a double hull penetrated by a cone (Wang et a12000a) The progressive damage process for individual structural members is generally captured in a global damage model, which follows the possible sequences of failure of major structural members including shell plating and main supporting members. Longitudinal support members (stiffeners) are commonly treated as smeared thickness, as they usually deform with the plates they attach to. Under specific circumstances however, supporting members may also be treated as a smeared portion of shell, which simplifies further the structural modelling. Figure 12 shows the load-indentation curve of the progressive damage model for a double hull penetrated by a cone and the comparisons with test observations. Initially the cone pressed the supporting member, which buckled at point a and heavily deformed thereafter. Membrane stresses developed in the shell plate in the stage of ab. The side shell was observed to rupture at point b, resulting in a reduction in resistance. Web girder interaction resulted in buckling at point d, and load raised again. The analytical model, which combines the damage modes of individual structural members, predicts a progressive damage process, which satisfactorily follows the observed phenomenon. Simplified analytical methods have been successfully applied to analysis of a wide spread of accident situations, including head-on collisions to rigid walls (Yang & Caldwell 1988, Pedersen et al 1993, Ohtsubo & Suzuki 1995), ship-ship collisions (Ltitzen et al 2000, Brown 2002), ship platform collisions (Wang & Ohtsubo 1999), ship-bridge collisions (Lehmann & Yu 1998a, Wang et al 1998b), bottom raking accidents (Wierzbicki & Thomas 1993, Wang et al 1997, Simonsen 1997, McGee et al 1999, Zhang 2002), and stranding on sea floor (Wang et a12000a). 3.5.5 Empirical Formulae

Empirical formulae are useful for easy and quick estimation of the energy absorption of a ship in an accident. These formulae correlate empirically the energy absorption with the volume of the damaged

Collision and Grounding

89

ship structures based on the past damage statistics (Minorsky 1959, Akita et al 1972, Hagiwara et al 1983, Kuroiwa 1996, Reardon & Sprung 1996, Paik et al 1999, Suzuki et al 2000). It is noted that empirical formulae are applicable only when the considered ship is similar to the ships used in the derivation of these formulae. To overcome these difficulties, semi-empirical formulae were derived by including structural geometries as parameters of influence (Wang & Ohtsubo 1999, Pedersen & Zhang 2000).

3.6 Post-accident Mechanics

After an accident, the ship may spill hazardous cargoes or lose stability. However, most often the ship will survive the direct consequences of the accident (disregarding the effect of potential fires caused by the accident). Depending on the final state of the accident further damage may result. If the vessel is caught in high sea, further loss of hull girder strength may result, possibly combined with progressive flooding. In case of grounding, the ship may be brought to rest on the obstruction, and further penetration or hull girder damage may result due to tide in addition to waves. Both waves and tidal variations may cause additional outflow of hazardous cargoes. 3.6.1 Oil Outflow

The expected amount of oil outflow is a measure of a ship's performance after an accident. Probabilistic based procedures have been developed for assessing oil outflow. Probability density functions describing the location and extents of side and bottom damages are applied to the vessel's compartment. All cargo or fuel oils are assumed to spill out of tanks that are punctured by collisions, whereas outflow from bottom damage by grounding is based on pressure balance calculations (Michel & Winslow 2000). Studies aiming at improving these procedures include directly considering the influences of the structural designs in terms of damage extents, and deriving probability density functions for damages, which reflect the structural crashworthiness of ship designs. A probabilistic methodology for predicting the oil outflow from a damaged tanker has been established by IMO (1995). Michel et al (1997), Sirkar et al (1997), Samuelides (1999) and Michel & Winslow (2000) also studied methodologies for predicting the amount of oil outflow from ships after collisions or grounding. 3.6.2 Hull Girder Collapse

A ship may collapse after an accident because of the decrease of longitudinal strength. The probability of hull girder collapse can be assessed by comparing the applied hull girder bending moment and the ultimate hull girder moment. The applied bending moment is a sum of still-water and wave-induced bending moments. The wave-induced bending moment calculation should be made by the short-term analysis considering the ship's damaged conditions in the sea states following the accident which involve significant wave height, operating speed and wave persistence (Paik & Thayamballi 2002). Hull girder ultimate moment calculations should account for possible failure modes of local structural components such as local plate buckling, lateral-torsional buckling (tripping), stiffener web buckling and yielding. For ultimate hull girder moment calculation of ships, special purpose computer programs HUH_NT (Yao et al 2000) and ALPS/HULL (Paik & Thayamballi 2002) or closed-form expressions (Paik & Thayamballi 2002, Paik et al 2002) are available. A computer program ALPS/USAS for predicting the safety margin of a ship, as the ratio of the ultimate hull girder moment to the short-term based extreme hull girder loads, can be downloaded from Paik & Thayamballi (2002). Investigations of residual hull girder strength of damaged ships include Zhang et al (1996), Paik & Pedersen (1997), Paik et al (1998), Guo et al (1998), Qi et al (1999), Cui et al (2000) and Wang et al (2000b, 2002a). Hull girder residual strength measures based on the section modulus and the ultimate

90

Specialist Committee V.3

hull girder strength of ships after collisions and grounding were investigated by Wang et al (2002a) for a total of 67 commercial ships including double hull tankers, single hull tankers, bulk carriers and container carriers. As long as the hull girder failure mode is predominantly ductile, it is clear that the section modulus based safety measure does not reveal the realistic safety margin of ships and the ultimate strength based safety measure is much better basis for the safety assessment. However, if brittle failure modes, associated with fracture or unstable buckling, are likely to occur, then the section modulus approach may need to be used.

3.7 Design Criteria for Accidents

Performance of a ship in an accident can normally be measured by the kinetic energy which can be dissipated in the damaged structures before tanks rupture, damage amount (depth, length), and quantity of oil outflow and/or residual hull girder strength. A ship should be designed to minimize the probability that cargo tanks/holds are breached in an accident to prevent spillage of hazardous cargoes, or if the cargo tanks are breached to ensure that the oil outflow following an accident is limited and/or the ship has adequate residual hull girder strength so that the integrity and residual strength of damaged structures is kept at a certain level to continue normal operations for its mission or to minimize further pollution (Wang et al 2002e). Comparative studies of many different designs are useful for establishing relevant design criteria. Recent comparative studies include Kitamura (1997), Michel et al (1997), Ltitzen et al (2000), Brown et al (2000) and Wang et al (2000a, 2002c). The Accidental Limit State (ALS) design format may be deterministic, probabilistic or semiprobabilistic. Design criteria may be a set of deterministic rules representing acceptable safety level or some given limits to the probability of occurrence to adverse events or some specified bounds on the probability of consequences. A practical deterministic form of the ALS design criterion can express a certain energy absorption capability (Paik & Thayamballi 2002):

Er

E k ~/k <-- ~ ')t r

(3)

where E k = characteristic (nominal) value of kinetic energy loss, E r = characteristic value of energy absorption capability until a critical damage occurs, ~'k, ?r = partial factors taking into account the uncertainties related to kinetic energy loss and energy absorption capability, respectively. Risk based (probabilistic) criteria are generally more relevant, while they are sometimes more complex to apply than the deterministic criteria. Shipping - as everything else - may cause risk of loss of lives, economic loss, environmental damage and other unwanted events. Ideally the level of the consequences of any accident should be acceptable to owners, governments and the public. With this in mind, risk acceptance criteria are normally established for two main types of risks, namely (1) the individual fatality risk for seamen shall be approximately the same as those typical for other occupational hazards, and (2) the frequency of accidents with fatalities, or the societal fatality risk, shall not exceed a level defined as unconditionally intolerable. The general ALARP (As Low As Reasonably Practicable) risk management may be applied (Cazzulo 1995). Figure 13 illustrates the principle of this criterion for passenger vessels.

Collision and Grounding

1E-1

,

91

/// ///,////

~1E-3

z

~

1E-5

1E-8

I•lll

1o

II11

Number

100 of

I

I 1 Illll

1000 fatalities

10.oo0

Figure 13" Typical risk acceptance criterion, frequency- number of fatalities diagram (Pedersen 2002) The last-mentioned societal risk acceptance criterion is introduced herein, because a single accident with many fatalities than many accidents with few fatalities per accident is more concerned about. For instance, to have 100 people killed in one accident every 1000 years is considered more serious than to have 1 person killed every 10 years. In the ALARP region, the economic criterion has to be applied to consider the effectiveness of safety measures. That is, the additional cost of risk-reducing measures in the form of construction cost plus present value (PV) of operational costs is evaluated against the effect of the risk. The condition for a decision to introduce risk-reducing measures is given by C< X +I+E+D

(4)

where C = PV of the cost concerned with risk-reducing measure, X = PV of the direct costs associated with the accident, I = accumulated change of PV of all individual user risks, E = quantified change in PV of environmental impact, D - change in PV due to economic loss caused by disruption. For each type of accident, a risk model can be established and the risk contributions are identified and entered into the risk account (Pedersen 2002).

4. B E N C H M A R K STUDIES ON N O N L I N E A R FE ANALYSIS OF I N T E R N A L M E H C A N I C S Full non-linear finite element method is a powerful tool for analysing the structural crashworthiness of ships in collisions and grounding, but to obtain correct results it is necessary to use a relevant combination of element size and fracture strain, as previously addressed in Section 3.5.2. To develop some useful guidelines for full non-linear finite element analyses, benchmark studies on a tensile coupon test of a flat mild steel specimen, a collision test model and a grounding test model were undertaken by the committee members.

4.1 Tensile Coupon Test of Mild Steel The stress-strain curve for a tensile coupon test on flat mild steel specimen (thickness of 2mm) in a quasi-static condition was computed using the non-linear finite element program LS-DYNA. The Belytschko-Tsay type plate-shell elements with a piecewise-linear plasticity material model were employed. The aim of the present benchmark study was to investigate the dependency of the input fracture strain on element size, required to represent the physical test results. Two types of material stress-strain modelling were compared. These accounted for: (1) the strain-hardening effect but neglected the necking effect (denoted by material model I) and (2) both the strain-hardening and necking effects (denoted by material model II).

Specialist Committee V.3

92

In both models, the true stress-strain characteristics of the material are required data input. True stress and strain is related to engineering (or nominal) stress and strain as follows: O'true ~ (Yeng0-t- t~eng))

I~true-

ln0+ 8eng)

where (Ytrue -- true stress (allows approximately for Poisson ratio thinning),

(5) I~true ~-

true strain (integral

of change in length over instantaneous length), ~eng = engineering stress (force / unstressed area),

l~eng

= engineering strain (change in length / initial length). Figure 14 shows the true stress-strain curves of the material transformed by the two approaches noted above. 500

400

300

200

100

: Engineering stress-strain relationship | stress-strain relationship by Material (1) | : True stress-strain relationship by Material (2) I

'

0.2

'

I

'

0.4 Strain

I

0.6

'

I

0.8

Figure 14: Transformation of the engineering stress-strain relationship to the true stress-strain relationship of the material considered

~.....:_~..}..]:::.~..;i.~

~(i~i1

Mesh-type 1 ~i!i!3i!!i!-3!!i~3i~!~-Y~-Ll-L-~-!3!4"----!!:.:!!~:.~Me -t~ sh-type 2

Mesh-type 3 t - ~ ' ~

"~~:-:-::::-~:::_1 Mesh-type 4

rs--.~.~:.:~.~:-~;~.:-~!~N!~~~i-.H

~+-~-++--.-~

~-~:+++-~::-~ Mesh-type 5

Figure 15: Five types of finite element meshes considered on flat mild steel test specimen

Collision and Grounding

93

400400

300-

300-

200

200

Material model I O " Experiment | " Mesh size of 25.5 x 25.0 m m | 9Mesh size of 12.75 x 12.50 m m | " Mesh size of 8.50 x 8.33 m m | " Mesh size of 4.25 x 4.17 m m | 9Mesh size of 2.125 x 2.083 m m

100

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To investigate the effect of element sizes on the material behaviour, five types of element sizes were considered, namely (a) mesh-type 1" 25.5 • 25.0mm, mesh-type 2:12.75 • 12.5mm, mesh-type 3: 8.5• mesh-type 4: 4.25• and mesh type 5" 2.125• as those shown in Figure 15.

Figure 16 compares the finite element solutions (using mesh-type 5) with the test in terms of the engineering stress-strain relationship. It is seen from Figure 16 that both material models I and 11 overestimate the strain-hardening effect. Furthermore, material model I does not reveal the effect of

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necking. It is concluded that material model II is a better approach than material model I in terms of computational accuracy, but it is clear that a more refined modelling technique of material needs to be developed to properly account for both strain-hardening and necking effects. Figure 17 shows the input fracture strains which were obtained by a trial and error technique so that the "critical" fracture strains by FEA corresponded to those by testing. It is interesting to note that this correspondence with material model II was achieved by assuming different level of the fracture strain in the FEA with different meshes, while the fracture strain required with material model I was approximately the same as the fracture strain in the test, regardless of element size. This is because model I does not model the local necking and the narrow section of the whole specimen increases in strain till the specified fracture strain is reached. It is however noted that if the waisted length of the specimen were to be increased the specified fracture strain would have to be reduced because the effect of the local necking would have to be averaged over a longer length. For general use, therefore, the model I would not have the benefit of mesh size independent failure strain that it appears to have in the present benchmark study. On the other hand, in material model U, the local necking and associated strain concentration occurs, and the length of the necking is roughly equal to the element size, because after the stress in one transverse line of elements exceeds the maximum stress, defined by the hump in the stress-strain curve (cYh, eh ), further displacement will concentrate strain in that line of elements. To obtain the measured average strain over the gauge length, the modelled strain to fracture after necking must be inversely proportional to the element size S or symbolically (F~f--F~h ) S -- constant, where [;f is the required fracture strain. This relationship is found to approximately match the findings of the benchmark study, see Figure 17. The coarse mesh finite element analysis together with material model II required the fracture strain to be the similar to the nominal fracture strain in the test. This is because the length of the coarse mesh elements roughly corresponds to the yielding area in the test. From these results and the simplified model it would be expected that the element-size dependent fracture strains together with material model II would be satisfactory for analyses of other structures built from the similar material and thickness. However, a different fracture strain - element size relationship would be expected for different material and thickness (which would result in a different necking length). It is also important to realize that bending may result in a different failure strain element size relationship. Although material model I appears to be mesh size insensitive in the present benchmark study, it is only apply to the particular specimen tested. In practice when setting up different analyses, therefore, the material model II will provide more useful guidance than material model I. More work is needed to provide practical recommendations for the application of material model II.

4.2 Collision Test The Association of Structural Improvement of Shipbuilding Industry of Japan (ASIS 1993) carried out extensive collision and grounding tests. One of the collision test models is a double side structure model made of mild steel. It was impacted from outer side shell by a 82.32 kN weight (striking bow) freely fallen from a height of 4.8m. The weight struck the double hull model at a speed of 9.7 m/s. In FE modelling, the element size should be fine enough so that deformation pattems be properly captured in the analysis. It is desirable that the shape of the element is rectangle and the aspect ratio of

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an element is near 1.0. While the deformation patterns of steel plates under axial compression at the ultimate limit state have a sinusoidal shape, ship collisions and grounding cause more complex deformation patterns involving folding and tearing as well as localized yielding. To investigate the effect of mesh fineness, two meshes are considered: one is a coarse mesh usually applied for an ultimate strength analysis and the other is a fine mesh more suitable for a collision analysis. For an ultimate strength analysis, five elements between a stiffener spacing may be enough to capture the collapse pattern of the plating between stiffeners, i.e., with one element size of 80mm for the collision test model considered. However, for a collision analysis involving structural folding, a finer mesh is required. As previously noted in Section 3.5.2, at least eight elements are needed within a half length of one structural fold, see Figure 11. In the present benchmark study, the formula derived by Wierzbicki & Abramowicz (1983) was employed to predict the fold length, namely H = 0.983b2/3t 1/3

(6)

where b = width of plating between stringers, t = plate thickness, H = a half fold length, see Figure 11. Since b=2000mm and t=7mm for the test model, a half length of one fold predicted by the Wierzbicki Abramowicz formula above is H = 298.5 mm. Therefore, it is recommended that one element size must be smaller than 298.5/8=37.3 mm so that at least 11 elements are necessary between a stiffener spacing (i.e., 400mm). In the present benchmark study, 13 elements were used for modelling of plating between stiffeners, which corresponds to the element size of 400/13=30.77mm. Based on the insights from Figure 17, the same value of the fracture strain obtained by the tensile coupon test was used for both types of element sizes which are relatively large. The strain rate sensitivity effect on the material yield stress of mild steel was accounted for using the CowperSymonds formula (with C = 40.4and q = 5 ) as previously noted in Section 3.5.1, while the effect of strain rate on fracture strain was not considered (since LS-DYNA does not deal with it). Two types of material models as discussed in Section 4.1 are considered: model I for accounting for the strain-hardening effect but neglecting the neck effect and model II accounting for both strain-hardening and necking effects. Figure 18 shows the two types of finite element modelling for the collision test model, namely a coarse mesh for the ultimate strength analysis (left) and a fine mesh for the collision analysis (fight). Figure 19 shows the deformation patterns obtained from the two types of finite element sizes.

Figure 18: Two types of finite element models with meshes: coarse mesh for the ultimate strength analysis (left) and fine mesh for the collision analysis (right)

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Figure 19: Deformation patterns obtained by the mesh type with material model II (a) coarse for the ultimate strength analysis, (b) fine for the collision analysis and (c) by the test

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4.3 Grounding Test A benchmark study on grounding of the double bottom structure model tested by ASIS (1993) in a quasi-static condition was also carried out. The material of the test structure is also mild steel. A sharp wedged rigid body penetrates into the structure quasi-statically. Similar to Section 4.2, two types of finite element sizes are again considered: a coarse mesh generally applied for an ultimate strength

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analysis and a fine mesh more suitable for a cutting analysis. Also, two types of material models are considered.

Figure 21: Two types of finite element models with meshes: coarse mesh for the ultimate strength analysis (left) and fine mesh for the grounding analysis (right)

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Figures 21 and 22 show the finite element models and deformed shapes of the two mesh types, respectively. Figure 23 compares the force-penetration curves and the absorbed energy-penetration curves for the two mesh types and experiment. It is evident that the fine mesh modelling with material model II provides the best solution.

5. ALTERNATIVES TO DOUBLE HULL DESIGN Double hull tankers remain the structural design standard for today, and have indeed been successful in reducing incidents of serious pollution since double hulls were first mandated by the Oil Pollution Act of 1990 (OPA90). Other structural design concepts have been suggested as possible alternatives to the double hull. This committee compares these various structural design alternatives from the standpoint of structural crashworthiness. It is intended to illustrate the technology on structural crashworthiness in collision and grounding accidents, and to propose to introduce the Accidental Limit States. A framework for developing innovative structural design alternatives of ships against collisions and grounding was suggested by Astrup et al (1995) and Michel et al (1997). Paik (2003) surveyed their results.

5.1 Varying Structural Arrangements for Double Sides against Collisions A variety of attempts to vary the structural arrangements and scantlings of double sides have been made to improve the ship structural crashworthiness against collisions, typically in terms of the energy absorption capabilities (Kitamura 1997, Lee et al 1998, Kim & Lee 2001, Naar et al 2001). Figures 24 and 25 compare the energy absorption capabilities for various side structures against side collisions. Some alterative designs improve the structural crashworthiness against ship collisions. However, most of them require additional steel weight and/or upgrade of steel material. When keeping steel weight unchanged, varying structural designs may have limited advantage in improving energy absorption capability. This is specifically true when the penetration depth is less than the double side width. Similar observations also apply to bow penetrations. 600 -

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5.2 Varying Double Bottom Heights along the Ship Length against Grounding The historical databases suggest that in most grounding accidents, breaches take place at the forward tank bottoms. This implies that the probability of oil outflow may be higher for the forward tanks than for the rest of the tank areas. Therefore, the double bottom height does not need to be of the same throughout the cargo area. Higher double bottom can be designed at the forward portion of a tanker, where raking damage is more likely (Amdahl & Kavlie 1995). However, care would have to be taken to design properly for the discontinuities that would result from step changes in the double bottom height. The single bottom may be arranged at the after-body areas where the probability of serious raking damages may be low. The risk of resulting oil outflows is different from a case where an adequate double bottom is employed through the entire vessel, because a low probability of grounding damage is not the same as zero probability of raking damage.

5.3 Mid-deck Tankers The so-called mid-deck tanker design concept has double hulls in sides but a single hull in bottom together with a deck at the mid-height of the ship. Double side widths are typically larger in mid-deck than ordinary double hull tankers because of requirements for ballast water. In the event of side collisions, mid-deck tankers may therefore provide more energy absorption capability than typical double hull tankers. Furthermore, in grounding the cargo oil loaded in the lower tank (i.e., under the mid-deck) may not spill out in large quantity because the pressure of cargo oil is balanced or smaller than the water pressure on the ship bottom. In principle, it is possible to design a mid deck tanker that is superior to existing double hull VLCCs in collisions. The reverse is also likely to be true, that is, for a given premised accident scenario, it may be possible to design a double hull that is of greater or equivalent crashworthiness performance than a given mid-deck vessel.

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5.4 Soft Bow Design against Collisions In a ship-ship collision accident, the striking bow can also be damaged. Given the total loss in kinetic energy, the damage to a struck ship side structure becomes smaller if the striking ship bow structure is 'soft' to deform and absorb more energy. The soft bow design was initiated by Cheung (1969) more than two decades ago. Lee (1983) applied this idea to a bow design against collisions. These studies focused on conventional ramming bows. Recently, Endo et al (2001) revived the soft bow (also called buffer bow) design concept. Their attention is on bulbous bows with ring frames which more efficiently absorb the kinetic energy during a head-on collision. A proper combination of innovative side structure designs with the soft bow design may greatly reduce the probability of oil outflow from tankers. Regarding the soft bow design, however, worldwide consensus is perhaps required. The regulatory bodies or maritime authorities would encourage and set design standards for soft bow structures to all types of ships.

6. PROPOSALS FOR F U R T H E R RESEARCH AND DEVELOPMENTS Although the technologies have advanced remarkably, many areas related to the Accidental Limit State design of ship structures still need to be improved or explored further. Some important R & D topics are proposed as below: 9Collection of traffic data for determining accident probabilities, 9Mathematical models for predicting accidental probabilities in different waterways, 9Standard accident scenarios, 9Acceptance design criteria, 9 Guidelines for non-linear FEM analyses, which include critical element size, dynamic material models and fracture criteria, 9Fracture criteria to be used in crushing and tearing analyses by either full non-linear FEM or simplified analytical methods, which should involve the effects of microscopic/macroscopic fracture characteristics of material and weld toes, 9Benchmark studies on simplified analytical models, numerical methods and experimental results for collisions and grounding.

7. CONCLUSIONS The reduction of spillage of hazardous cargoes from ships and floating structures in the events of collisions and grounding has been a significant issue in the maritime community. This is partly due to the attention of the public who are understandably very concerned with marine pollution. Subsequent to the Exxon Valdez grounding accident, this concern resulted in the enactment of the OPA90 and related IMO requirements for double hulls. It is also of critical issue to avoid loss of life and to minimize loss of property and financial exposure in the events of accidents. Ideally, the vessels should have sufficient structural resistance against reasonably severe potential accidents such as collisions and grounding, and so the structural crashworthiness of vessels should be continually improved. The double hull design is believed to greatly reduce the probability of marine pollution in tanker accidents. Arguably there are design alternatives that have improved the structural performance and reduced oil outflow.

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The majority of rules and regulations are prescriptive, generally conservative and rigid. Performancebased criteria allow for greater design flexibility, and offer more freedom for innovative designs and optimisation of vessels while maintaining safety and pollution performance. IMO has established guidelines for approving alternative tanker designs, which provide a probabilistic method to assess the oil outflow performance of tankers in accidents. Since its publication, some deficiencies have been recognized, with proposals for new or improved standards. While significant advancements in design technologies of ship structures against collisions and grounding have been achieved, the Accidental Limit State design concept has not yet been used in design of ships or floating structures. Such direct design-oriented procedures can be utilized to validate better designs in the future, designs which may prove to be not only efficient for pollution prevention but also perhaps less costly to build and operate. REFERENCES

Akita, Y., Ando, N., Fujita, Y. and Kitamura, K. (1972). Studies on collision-protective structures in nuclear powered ships, Nuclear Engineering and Design, 19, 365-401. Amdahl, J. and Kavlie, D. (1992). Experimental and numerical simulation of double hull stranding, DNV-MIT Workshop on Mechanics of Ship Collision and Grounding, Hcvik, September. Amdahl, J. and Kavlie, D. (1995). Design of tankers for grounding and collision, Proceedings of MARIENV'95, Tokyo, I, 167-174. Amdahl, J. and Stornes, A. (2001). Energy dissipation in aluminium high-speed vessels during collision and grounding, Proceedings of ICCGS'01, Copenhagen, 203-219. Arita, K. and Aoki, G. (1985). Strength of ship bottom in grounding - an investigation into the case of a ship stranded on a rock, Journal of SNAJ, 158,359-367 (in Japanese). ASIS (1993). The Conference on "Prediction Methodology of Tanker Structural Failure & Consequential Oil Spill", Association of Structural Improvement of Shipbuilding Industry of Japan, m.39-111.47. Astrup, O. (1994). Cutting of thick plate with a wedge, Joint MIT-Industry Project on Tanker Safety, Report No.27. Astrup, O., Monnier, I., Sirkar, J. and Cojeen, H.P. (1995). Framework for evaluating alternative designs and configurations for tankers, Proceedings of MARIENV'95, I, 183-190. Bockenhauer, M. and Egge, E.D. (1995). Assessment of the collision resistance of ships for classification purposes, Proceedings of MARIENV'95, Tokyo, I, 44-51. Brown, A. (2002). Collision scenarios and probabilistic collision damage, Proceedings of ICCGS'01, Copenhagen, 259-272. Brown, A. and Chen, D. (2002). Probabilistic method for predicting ship collision damage, Oceanic Engineering International, 6:1, 54-65. Brown, A., Tikka, K., Daidola, J.C., Ltitzen, M. and Choe, I.H. (2000). Structural design and response in collision and grounding, Trans. SNAME, 108,447-473. Cazzulo, R.P. (1995). Maritime safety and risk acceptance criteria, Proceedings of WEGEMT'95, Copenhagen, 1, 1-28. Chen, D. (2000). Simplified ship collision model (SIMCOL), Master Thesis, Department of Aerospace and Ocean Engineering, Virginia Polytechnic Institute and State University, Blacksburg, VA. Cheung, L. (1969). A soft bow for ships, European Shipbuilding, 3, 52-53. Cowper, G.R. and Symonds, P.S. (1957). Strain-hardening and strain rate effects in the impact loading of cantilever beams, Brown University, Technical Report No. 28. Cui, W., Qi, E., Peng, X. and Xu, X. (2000). Asymmetrical bending ultimate strength and reliability analysis of damaged ships, Shipbuilding of China, 42:2, 41-47 (in Chinese). Endo, H., Yamada, Y., Kitamura, O. and Suzuki, K. (2001). Model test on the collapse strength of the buffer bow structures, Proceedings of ICCGS'01, Copenhagen, 145-153.

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Engineering Computer Optecnomics, Inc (1996). World fleet data. Fujii, Y., Yamanouchi, H. and Mizuhi, N. (1974). The effect of darkness on the probability of collision and stranding, Journal of Navigation, 27,243-247. Gluver, H. and Olsen, D. (1998). Current practice in risk analysis of ship collisions to bridges, In: Ship Collision Analysis, Balkema, Rotterdam. Gu, Y. and Wang, Z.L. (2001). An inertia equivalent model for numerical simulation of ship-ship collision, Proceedings of ICCGS '01, Copenhagen, 155-160. Hagiwara, K., Takanabe, H. and Kawano, H. (1983). A proposed method of predicting ship collision damage, International Journal of Impact Engineering, 1:33,257-379. IMO (1971). IMO Resolution A.265 (vm), Regulations on subdivision and stability of passenger ships as an equivalent to Part B of Chapter II of the international convention for the safety of life at sea SOLAS, Chapter II, Part B-1, Subdivision and Damage Stability of Cargo Ships. IMO (1989). Distribution of actual penetrations and damage locations along ship's length for collisions and groundings, IMO Comparative Study on Oil Tanker Design, IMO Paper MEPC 32/7/15, Annex 5. IMO (1995). Interim guidelines for approval of alternative methods of design and construction of oil tankers under regulation 13F(5) of annex I of MARPOL 73/78, Technical Report, Resolution MEPC, 66(37), International Maritime Organization, 1-40. Ito, H., Kondo, K., Yoshimura, N., Kawashima, M. and Yamamoto, S. (1984). A simplified method to analyse the strength of double hulled structures in collision, Journal of SNAJ, 156, 299-312. Ito, H., Kondo, K., Yoshimura, N., Kawashima, M. and Yamamoto, S. (1986). A simplified method to analyse the strength of double hulled structures in collision (3ra Report), Journal of SNAJ, 160, 401-409. Jones, N. (1989). On the dynamic inelastic failure of beams, Chapter 5 in Structural Failure, John Wiley & Sons, New York, 133-159. Jones, N. (1997). Structural impact, Cambridge University Press, Cambridge. Jones, N. and Jouri, W.S. (1987). A study of plate tearing for ship collision and grounding damage, Journal of Ship Research, 31:4, 253-268. Kaneko, F. (2002). A methodology and methods for probabilistic safety assessment of ships, ClassNK Technical Bulletin, 20, 29-53. Kawaichi, K., Kuroiwa, T. and Sueoka, H. (1995). On the structural design of mid-deck tanker, Proceedings of MARIENV'95, Tokyo, I, 213-218. Kierkegaard, H. (1993). Ship collisions with icebergs, Ph.D. Thesis, Department of Mechanical Engineering, Technical University of Denmark, Lyngby, April. Kim, J.Y. and Lee, J.W. (2001). A comparative study of the double hull structures for the collision energy absorption systems, Proceedings of the SNAK Annual Spring Meeting, The Society of the Naval Architects of Korea, April, 303-306 (in Korean). Kitamura, O. (1997). Comparative study on collision resistance of side structure, Marine Technology, 34:4, 292-308. Kitamura, O. (2000). Buffer bow design for the improved safety of ships, Ship Structure Symposium on Ship Structures for the New Millennium: Supporting Quality in Shipbuilding, Arlington, VA, June. Kitamura, O. (2002). FEM Approach to simulation of collision and grounding damage, Marine Structures, 15,403-428. Kitamura, K., Okumoto, Y. and Shibue, T. (1978). On the model tests of double bottom strength for stranding, Journal of SNAJ, 143,346-356 (in Japanese). Kuroiwa, T. (1996). Numerical simulation of actual collision and grounding experiments, Proccedings of International Conference on Design and Methodologies for Collision and Grounding Protection of Ships, San Francisco, 7.1-7.12. Laubenstein, L., Mains, C., Jost, A., Tagg, R. and Bjoerneboe, N. (2001). Updated probabilistic extents of damage based on actual collision data, Proceedings of ICCGS'01, Copenhagen, 93-99.

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Lee, J.W. (1983). On the optimisation design of soft bow structure, Proceedings of PRADS'83, Seoul, 429-435. Lee, T.K., Kim, J.D., Chun, T.B. and Shin, B.C. (1999). Experimental study on the collision strength of VLCC side structures, Proceedings of ISOPE'99, Brest, France, June, 450-455. Lee, J.W., Petershagen, H., Romp, J., Paik, H.Y. and Yoon, J.H. (1998). Collision resistance and fatigue strength of new oil tanker with advanced double hull structure, Proceedings of PRADS'98, The Hague, September, 133-139. Lehmann, E. and Yu, X. (1995). Progressive folding of bulbous bows, Proceedings of PRADS'95, Seoul, 2, 1048-1059. Lehmann, E. and Yu, X. (1998a). Inner dynamics of bow collision to bridge piers, In: Ship Collision Analysis, Balkema, Rotterdam, 61-71. Lehmann, E. and Yu, X. (1998b). On ductile rupture criteria for structural tear in the case of ship collision and grounding, Proceedings of PRADS'98, The Hague, 149-156. Lehmann, E., Egge, E.D., Scharrer, M. and Zhang, L. (2001). Calculation of collisions with the aid of linear FE models, Proceedings of PRADS'01, Shanghai, 1293-1300. Le Sourne, H., Donner, R., Besnier, F. and Ferry, M. (2001). External dynamics of ship-submarine collision, Proceedings of ICCGS'01, Copenhagen, 137-144. LR (2000). Lloyd's Register of Shipping World Casualty Statistics 1996-1999. Lu, G. and Calladine, C.R. (1990). On the Cutting of a Plate by a Wedge, International Journal of Mechanical Sciences, 32, 293-313. Ltitzen, M. (2001). Ship collision damages, Ph.D. Thesis, Department of Mechanical Engineering, Technical University of Denmark, Lyngby, December. Ltizten, M., Simonsen, B.C. and Pedersen, P.T. (2000). Rapid prediction of damage to struck and striking vessels in a collision event, Ship Structure Symposium on Ship Structures for the New Millennium: Supporting Quality in Shipbuilding, Arlington, VA, June. McDermott, J.F., Kline, R.G., Jones, E.L., Maniar, N.M. and Chiang, W.P. (1974). Tanker structural analysis for minor collisions, Trans. SNAME, 82, 382-414. McGee, S.P., Troesch, A. and Vlahopoulos, N. (1999). Damage length predictor for high speed craft, Marine Technology, 36:4, 203-210. Michel, K., Moore, C. and Tagg, R. (1997). A simplified methodology for evaluating alternative tanker configurations, Journal of Marine Science and Technology, 1,209-219. Michel, K. and Winslow, T. (2000). Cargo ship bunker tanks: design to mitigate oil spillage, Marine Technology, 37, 191-199. Minorsky, V.U. (1959). An analysis of ship collision with reference to protection of nuclear power ships, Journal of Ship Research, 3:2, 1-4. Mizukami, M., and Tanigawa, M. (1995). FEM analysis of VLCC bottom raking in grounding, Proceedings of MARIENV'95, Tokyo, I, 86-90. Naar, H., Kujala, P., Simonsen, B.C. and Ludolphy, H. (2001). Comparison of the crashworthiness of various bottom and side structures, Proceedings of ICCGS'01, Copenhagen, 179-188. Nagasawa, H., Arita, K., Tani, M. et al. (1977). A study on the collapse of ship structure in collision with bridge piers, Journal of SNAJ, 142, 323-332 (in Japanese). Ohtsubo, H. and Suzuki, K. (1995). The crashing mechanics of bow structure and its optimal design against head on collision, Proceedings of PRADS'95, 2, 1060-1071. Ohtsubo, H. and Wang, G. (1995). An upper bound solution to the problem of plate tearing, Journal of Marine Science and Technology, 1, 46-51. ORI (1980). Hazardous environment experienced by radioactive material packages transported by water, Silver Spring, MD. ORI (1981). Accident severities experienced by radioactive material packages transported by water, Silver Spring, MD. Paik, J.K. (1994). Cutting of a longitudinally stiffened plate by a wedge, Journal of Ship Research, 38:4, 340-348.

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Paik, J.K. (2003). Innovative structural design of tankers against ship collisions and grounding: a recent state-of-the-art review, Marine Technology, 40:1, 1-9. Paik, J.K., Chung, J.Y., Choe, I.H., Thayamballi, A.K., Pedersen, P.T. and Wang, G. (1999). On rational design of double hull tanker structures against collision, Trans. SNAME, 107, 323-363. Paik, J.K., Choe, I.H. and Thayamballi, A.K. (2002). Predicting resistance of spherical-type LNG carrier structures to ship collisions, Marine Technology, 39:2, 86-94. Paik, J.K. and Pedersen, P.T. (1995). Ultimate and crushing strength of plated structures, Journal of Ship Research, 39:3,250-261. Paik, J.K. and Pedersen, P.T. (1996). Modelling of the internal mechanics in ship collisions, Ocean Engineering, 23, 107-142. Paik, J.K. and Pedersen, P.T. (1997). Simple assessment of post-grounding loads and strength of ships, International Journal of Offshore and Polar Engineering, 7:2, 141-145. Paik, J.K. and Thayamballi, A.K. (2002). Ultimate limit state design of steel-plated structures, John Wiley & Sons, Chichester, U.K. Paik, J.K., Thayamballi, A.K. and Yang, S.H. (1998). Residual strength assessment of ships after collision and grounding, Marine Technology, 35:1, 38-54. Paik, J.K., Wang, G., Kim, B.J. and Thayamballi, A.K. (2002). Ultimate limit state design of ship hulls, SNAME Transactions, 110. Paik, J.K. and Wierzbicki, T. (1997). A benchmark study on crushing and cutting of plated structures, Journal of Ship Research, 41:2, 147-160. Pedersen, P.T. (1994). Ship grounding and hull-girder strength, Marine Structures, 7, 1-29. Pedersen, P.T. (1995). Collision and grounding mechanics, Proceedings of WEGEMT'95, Copenhagen, 1,125-157. Pedersen, P.T. (2002). Collision risk for offshore structure, Journal of Engineering for the Maritime Environment, Proceedings of the Institution of Mechanical Engineers, 216:M1, 29-44. Pedersen, P.T., Hansen, P.F. and Nielsen, L.P. (1996). Collision risk and damage after collisions, Proceedings of the International Conference on the Safety of Passenger RoRo Vessels, RINA, London. Pedersen, P.T, Valsgard, S., Olsen, D. and Spangenberg, S. (1993). Ship impacts: bow collisions, International Journal of Impact Engineering, 13, 163-187. Pedersen, P.T. and Zhang, S. (1998). On impact mechanics in ship collisions, Marine Structures, 11:10, 429-449. Pedersen, P.T. and Zhang, S. (2000). Absorbed energy in ship collisions and grounding - revising Minosky's empirical method, Joumal of Ship Research, 44:2, 40-154. Petersen, M. J. (1982), Dynamics of ship collisions, Ocean Engineering, 9, 259-329. Qi, E., Cui, W., Peng, X. and Xu, X. (1999). The reliability assessment of ship residual strength after collision and grounding, Joumal of Ship Mechanics, 3, 40-46 (in Chinese). Ravn, E.S., Urban, J., Simonsen, B.C. (2001). Damage and loss of stability for HSC in grounding or collision accidents, Proceedings of HIPER, Hamburg. Rawson, C., Crake, K. and Brown, A. (1998). Assessing the environmental performance of tankers in accidental grounding and collision, Trans. SNAME, 106, 41-58. Reardon, P.C. and Sprung, J. (1996). Validation of Minorsky's ship collision model and use of the model to estimate the probability of damaging a radioactive material transportation cask during a ship collision, Proceedings of Intemational Conference on Design and Methodologies for Collision and Grounding Protection of Ships, San Francisco, 5.1-5.7. Reckling, K.A. (1976). Beitrag der elasto - und plastomechanic zur untersuchung von schiffskollisionen, Jahrbuch der Schiffsbautechnischen Gesellscahaft, 70, 443-464 (in German). Rodd, J. and Phillips, M.P. (1994). Stranding experiments on double hull tanker structures, The Advanced (unidirectional) Double-Hull Technical Symposium, Gaitherburg, Maryland, October. Samuelides, M. (1999). Prediction of oil outflow in case of a ship-ship collision based on energy considerations, Joumal of Ship Research, 43:3, 194-200.

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Sandia National Laboratories (1998). Data and methods for the assessment of the risks associated with the maritime transport of radioactive materials results of SeaRAM program studies, SAND 98117/2, Albuquerque, NM. Servis, D., Samuelides, M., Louka, T. and Voudouris, G. (2002). The implementation of finite-element codes for the simulation of ship-ship collisions, Journal of Ship Research, 46:4, 239-247. Simonsen, B.C. (1997). Ship grounding on rock: I & II, Marine Structures, 10, 519-584. Simonsen, B.C. (2000). Bottom raking damage of high speed craft, Trans. RINA, 142, 41-58. Simonsen, B.C. and Hansen, P.F. (2000). Theoretical and statistical analysis of ship grounding accidents, Journal of Offshore Mechanics and Arctic Engineering, 122, 200-207. Simonsen, B.C. and Lauridsen, L.P. (2000). Energy absorption and ductile fracture in metal sheets under lateral indentation by a sphere, International Journal of Impact Engineering, 24, 1017-1039. Simonsen, B.C. and Pedersen, P.T. (1995). Analysis of ship groundings on soft sea beds, Proceedings of PRADS'95, Seoul, 2, 1096-1109. Sirkar, J., Ameer, P., Brown, A., Goss, P., Michel, K., Nicastro, F. and Willis, W. (1997). A framework for assessing the environmental performance of tankers in accidental groundings and collisions, Trans. SNAME, 105,253-295. Sterndorff, M.J. and Pedersen, P.T. (1996). Grounding experiments on soft bottoms, Journal of Marine Science and Technology, 1:3, 174-181. Suhara, T., Shimizu, S., Ando, S., Hiyama, H., Imai, R., Sato, M., Kawano, S. and Maeda, A. (1970). Strength of huge tanker in collision (I), Journal of SNAJ, 128, 281-294 (in Japanese). Suzuki, K., Ohtsubo, H. and Sajit, C. (2000). Evaluation method of absorbed energy in collision of ships with anti-collision structure, Ship Structure Symposium on Ship Structures for the New Millennium: Supporting Quality in Shipbuilding, Arlington, VA, June. Tagg, R., Bartzis, P., Papanikolaou, A., Spyrou, K. and Ltitzen, M. (2001). Updated vertical extent of collision damage, Proceedings of ICCGS'01,101-113. Toni, M. and Funahashi, A. (1978). Energy absorption by plastic deformation of body structural members, SAE Paper, No. 780368, SAE Inc. Ueda, Y., Murakawa, H., Tanigawa, M., Yoneda, N. and Iwata, S. (1995). Strength estimation method of ship side structures subjected to hard bow collision, Journal of SNAJ, 178, 429-437 (in Japanese). Urban, J., Pedersen, P.T. and Simonsen, B.C. (1999). Collision risk analysis for HSC, Proceedings of FAST'99, 181-194. Vredeveldt, A.W. and Wevers, L.T. (1995). Full scale grounding experiments, Proceedings of Conference on Prediction Methodology of Tanker Structural Failure & Consequential Oil Spill, Tokyo, June. Wang, G. (2002). Some recent studies on plastic behaviour of plates subjected to very large load, Journal of Ocean Mechanics and Arctic Engineering, ASME, 124:3, 125-131. Wang, G., Arita, H. and Liu, D. (2000a). Behaviour of a double hull in a variety of stranding or collision scenarios, Marine Structures, 13, 147-187. Wang, G., Chen, Y.J., Zhang, H. and Peng, H. (2002b). Longitudinal strength of ships with accidental damages, Marine Structures, 15, 119-138. Wang, G., Chen, Y., Zhang, H. and Shin, Y. (2000c). Residual strength of damaged ship hull, Ship Structure Symposium on Ship Structures for the New Millennium: Supporting Quality in Shipbuilding, Arlington, VA, June. Wang, G. and Ohtsubo, H. (1997). Deformation of ship plate subjected to very large load, Proceedings of the 16th International Conference on Offshore Mechanics and Arctic Engineering (OMAE'97), Yokohama, 2, 173-180. Wang, G. and Ohtsubo, H. (1999). Impact load of a supply vessel, Proceedings of ISOPE'99, Brest, France, 4, 463-471. Wang, G., Ohtsubo, H. and Arita, K. (1998a). Large deflection of a rigid-plastic circular plate pressed by a rigid sphere, Journal of Applied Mechanics, 65,533-535.

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Wang, G., Ohtsubo, H. and Arita, K. (1998b). Inner dynamics of side collision to bridge piers, In: Ship Collision Analysis, Balkema, Rotterdam, 53-60. Wang, G., Ohtsubo, H. and Liu, D. (1997). A simple method for predicting the grounding strength of ships, Journal of Ship Research, 41:3,241-247. Wang, G., Seah, A.K. and Shin, Y.S. (2002d). Predicting ship structure performance in accidents, Proceedings of MARTECH'2002, Singapore Polytechnic Institute, September, Singapore. Wang G., Spencer J. and Chen, Y.J. (2002e). Assessment of ship's performance in accidents, Marine Structures, 15,313-333. Wierzbicki, T. (1995). Concertina tearing of metal plates, International Journal of Solid Structures, 19, 2923-2943. Wierzbicki, T. and Abramowicz, W. (1983). On the crushing mechanics of thin-walled structures, Journal of Applied Mechanics, ASME, 50, 727-734. Wierzbicki, T., Peer, D.B. and Rady, E. (1993). The anatomy of tanker grounding, Marine Technology, 30:2, 71-78 Wierzbicki, T. and Thomas, P. (1993). Closed-form solution for wedge cutting force through thin metal sheets, International Journal of Mechanical Sciences, 35,209-229. Woisin, G. (1979). Design against collision, Proceeding of International Symposium on Advances in Marine Technology, 2, Trondheim. Woisin, G. (1998). Analysis of the collision between rigid bulb and side shell panel, Proceedings of PRADS'98, The Hague, 165-172. Woisin, G. (1999). Discussion on Kitamura (1997) "Comparative study on collision resistance of side structure", Marine Technology, 36:4, 228-231. Yang, P.D.C. and Caldwell, J.B. (1988). Collision energy absorption of ships, bow structures. International Journal of Impact Engineering, 7:2, 181-196. Yao, T., Astrup, O.C., Caridis, P., Chen, Y.N., Cho, S.R., Dow, R.S., Niho, O. and Rigo, P. (2000). Ultimate hull girder strength, Report of the ISSC Special Task Committee VI.2, Nagasaki, October, 2, 321-391. Yu, X. (1996). Structural analysis with large deformations until fracture and with dynamic failure, Dr. Thesis, Hamburg University (in German). Zhang, S. (1999). The mechanics of ship collision, Ph.D. Thesis, Department of Mechanical Engineering, Technical University of Denmark. Zhang, S. (2002). Plate tearing and bottom damage in ship grounding, Marine Structures, 15, 101-117. Zhang, S., Yu, Q. and Mu, Y. (1996). A semi-analytical method of assessing the residual longitudinal strength of damaged ship hull, Proceedings of ISOPE'96, VI, 510-516. Zhu, L. (1990). Dynamic inelastic behaviour of ship plates in collision, Ph.D. Thesis, Department of Naval Architecture and Ocean Engineering, University of Glasgow, Glasgow. Zhu, L. and Atkins, A.G. (1998). Failure criteria for ship collision and grounding, Proceedings of PRADS'98, The Hague, 141-148. Zhu, L., James, P. and Zhang, S. (2002). Statistics and damage assessment of ship grounding, Marine Structures, 15,515-530.

Abbreviations: ICCGS'01 = Intemational Conference on Collision and Grounding of Ships, Copenhagen, July 1-3, 2001; ISOPE = Intemational Offshore and Polar Engineering Conference; MARIENV'95 = Intemational Conference on Technologies for Marine Environment Preservation, Tokyo, September 24-29, 1995; PRADS =Intemational Symposium on Practical Design of Ships and Mobile Units; SNAJ = The Society of Naval Architects of Japan; WEGEMT = Graduate School organized by a European Association of Universities in Marine Technology.

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FURTHER READING

During the last decade, several international conferences and workshops have been taken place worldwide in the areas of ship collisions and grounding. These include: 9WEGEMT'93: Risk and Reliability in Marine Technology, Instituto Superior Technico, Lisbon, September 6-10, 1993, see the list of Abbreviations, 9WEGEMT'95: Accidental Loadings on Marine Structures: Risk and Response, Technical University of Denmark, April 24-29, 1995, see the list of Abbreviations, 9MARIENV'95, see the list of Abbreviations, 9International Conference of Designs and Methodologies for Collision and Grounding Protection of Ships, San Francisco, California, August 22-23, 1996, 9ICCGS'01, see the list of Abbreviations.

ACKNOWLEDGEMENTS The details of the test results used for the present benchmark study were provided by SSML (Ship Structural Mechanics Laboratory of Pusan National University) for the tensile coupon test of mild steel, and by ASIS (Association of Structural Improvement of Shipbuilding Industry of Japan) for the collision and grounding model tests. The Committee is pleased to acknowledge their support.

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15th INTERNATIONAL SHIP AND OFFSHORE STRUCTURES CONGRESS 2003 AUGUST 11-15, 2003 SAN DIEGO, USA

g _ 4 4 1 "-

VOLUME 2

%IV DIEGO~,~~"

COMMITTEE V.4

STRUCTURAL DESIGN OF HIGH SPEED VESSELS

COMMITTEE MANDATE Concern with structural design methods for high speed vessels, and with uncertainties in these methods. Consideration shall be given to the structural application of advanced composite materials. Attention shall be given to the interaction between loads and structural response, in collaboration with the ITTC.

COMMITTEE MEMBERS

Chairman:

Mr. Mr. Dr. Prof. Mr. Dr. Mr. Prof.

Stefano Ferraris Nuno Fonseca Brian Hayman Owen Hughes Etienne Thiberge Yasumi Toyama Alex Vredeveldt Ping Yang

KEYWORDS

High speed vessels, structural design, classification rules, operational envelope, loads, structural response, strength, fatigue, advanced composite materials, uncertainties, fabrication technologies.

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CONTENTS

1 INTRODUCTION . . . . . . . . . . . . . . . . . . . . . . 1.1 Scope . . . . . . . . . . . . . . . . . . . . . 1.2 Development Trends in Recent Years . . . . . . . . . . . . . .

113 113 113

2 THE UPDATED HSC C O D E A N D C L A S S I F I C A T I O N SOCIETIES R U L E S . . . 2.1 The 2000 HSC Code . . . . . . . . . . . . . . . . . . 2.1.1 The Key Reasons for Updating the HSC Code . . . . . . . 2.1.2 A n Overall View o f the New Issues of the HSC Code (2000) . . . . . 2.1.3 Structural Design . . . . . . . . . . . . . . . . . . 2.2 The Present Situation with Classification Societies . . . . . . . . . . 2.2.1 The Importance o f Class Requirements with Respect to HSC Code Compliance 2.2.2 Comparisons o f Classification Requirements Influencing Structural Design . 2.2.2.1 Operational E n v e l o p e - Characterising Parameters . . . . . . 2.2.2.2 Design Loads and Stress Criteria . . . . . . . . . . . . 2.2.2.3 Service Restriction - Format . . . . . . . . . . . .

113 113

3 LOADS . . . . . . . . . . . . . . . . 3.1 Introduction . . . . . . . . . . . . . . 3.2 Review of Calculation Methods . . . . . . . . . 3.3 Recent Work . . . . . . . . . . . . . . . 3.3.1 Global Loads . . . . . . . . . . . . . . . 3.3.2 Slamming Loads and Structural Responses . . . .

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3.3.3 Hydroelastic Slamming . 2 -rd3 3.3.4 Review o f the LRC Report o f the ITTC

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4 S T R U C T U R A L R E S P O N S E A N D ULTIMATE S T R E N G T H . . . . . . 4.1 Structural Response o f High-Speed Vessels . . . . . . . . 4.2 Ultimate Strength o f High Speed Vessels . . . . . . . . . . . 4.2.1 Overall Strength . . . . . . . . . . . . . . . . . . 4.2.2 Stiffened Panels . . . . . . . . . . . . . . . . .

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5 ADVANCED COMPOSITE MATERIALS . . . . . . . . . . . . . 5.1 Introduction; Definitions . . . . . . . . . . . . . . . . . . . . 5.2 Reasons for Using Advanced Composites; Hindrances to Using Composites 5.2.1 Advantage . . . . . . . . . . . . . . . . . . . . 5.2.2 Special Advantages for Naval Vessels . . . . . . . . . . . . 5.2.3 Hindrances . . . . . . . . . . . . . . . . . . . 5.3 Developments in Materials . . . . . . . . . . . . . . . . . 5.3.1 Carbon Fibre Reinforced Plastics (CFRP) . . . . . . . . . . 5.3.2 Phenolic and other Resins for Improved High-Temperature Behaviour

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5.4 5.5

5.6 5.7 5.8 5.9

Special&t Committee V.4 5.3.3 Thermoplastic Resins . . . . . . . . . . . . . . . . . . . 5.3.4 Sandwich Core Materials . . . . . . . . . . . . . . . . . . Developments in Production Methods . . . . . . . . . . . . . . . Performance o f Structural Elements - Failure Mechanisms and Calculation Methods 5.5.1 Failure under Static Loading . . . . . . . . . . . . . . . . 5.5.2 Fatigue . . . . . . . . . . . . . . . . . . . . . . . . 5.5.3 Vibrations . . . . . . . . . . . . . . . . . . . . . . . 5.5.4 Impact and other D y n a m i c Loading . . . . . . . . . . . . . . Connection Design . . . . . . . . . . . . . . . . . . . . . . Fire Behaviour; Fire Protection . . . . . . . . . . . . . . . . . . Inspection, Repair and D a m a g e Tolerance . . . . . . . . . . . . . . Design Rules for Composites . . . . . . . . . . . . . . . . . .

6 U N C E R T A I N T I E S IN D E S I G N

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7 D E V E L O P M E N T S IN FABRICATION T E C H N O L O G I E S . . . . . . . . . . 7.1 Introduction . . . . . . . . . . . . . . . . . . . . . . . . 7.2 Development in Metallic Materials and Relevant Applications . . . . . . . 7.2.1 Steel . . . . . . . . . . . . . . . . . . . . . . . . 7.2.2 A l u m i n i u m Alloys . . . . . . . . . . . . . . . . . . . . 7.3 Development of N e w Cutting and Joining Techniques . . . . . . . . . . 7.3.1 Influence o f Cutting and Joining Techniques on Design . . . . . . . 7.3.2 Laser B e a m Cutting and Welding . . . . . . . . . . . . . . . 7.3.3 Friction Stir Welding . . . . . . . . . . . . . . . . . . . 7.3.4 Adhesive Bonding . . . . . . . . . . . . . . . . . . . . 8 CONCLUSIONS AND RECOMMENDATIONS REFERENCES

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Structural Design of High Speed Vessels 1.

INTRODUCTION

1.1

Scope

113

Following the report on Structural Design of High Speed Vessels presented at ISSC 2000, this report aims: 9 to give a brief but rather comprehensive view of major research results achieved during the last three years; 9 to focus on the new HSC Code, issued in 2000, and on its impacts on structural design and classification rules; 9 to provide an exhaustive review of the structural application of advanced composite materials, as explicitly required by the mandate. Even though the HSC Code is referred to civilian craft, some military vessels, having HSC characteristics, such as fast patrol and attack craft, have been taken into account.

1.2

Development trends in recent years

The increase in craft size has led to a more extensive use of "first principle design" approaches. The HSC Code and the classification rules allow novel design solutions to be accepted, provided they can be suitably supported and demonstrated by such approaches. New rational ways of predicting motions, loads and structural responses have been developed and included both in the rules and in computational methods, taking into account vessel operational restrictions. The need to ensure high performances, in terms of both speed and payload, has driven designers to the definition of lightweight solutions and the use of lightweight and stronger materials, even though higher design stress levels have led to a need for improved fatigue assessment of such structures, which are often subjected to a large number of cycles (even exceeding 107 to 108) in a rather short period. With regard to civilian vessels, the market has largely stagnated: apart from possible developments in terms of cargo craft, the main progress can be expected in the naval field, where high speed low and medium size vessels are designed and built for fast attack and coast patrolling purposes. This trend is confirmed also by some naval regulations, which make explicit reference to HSC rules, like for instance DNV Rules for Classification of High Speed, Light Craft and Naval Surface Craft and RINA Rules for

the construction and maintenance of the quality standard of Italian Coast Guard Vessels.

2.

THE UPDATED HSC CODE AND CLASSIFICATION SOCIETIES RULES

2.1

The 2000 HSC code

2.1.1 The key reasons for updating the HSC Code The first issue of the HSC Code (resolution MSC.36(63)) was finalised in 1994 and introduced by the 1994 amendments to the SOLAS convention. It came into force for international HSC on January 1st, 1996. Since the very beginning of its implementation, serious deficiencies were faced by designers and yards, in Chapters 2, 4 and 7. In 1996, the Maritime Safety Committee agreed to an urgent revision during its 66 th

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session. The terms of reference for this revision were to address specific known deficiencies rather than to improve the style or presentation. Particularly, a comprehensive revision of Chapter 2 and the associated annexes has been undertaken. Of the changes introduced into the 2000 HSC Code, the most far-reaching is the consideration of extensive bottom raking damage. To ensure that it would indeed be practical to comply with such requirements, an extensive programme of trial application to known designs was undertaken in collaboration with designers and builders of HSC world-wide. Blyth (2001) discussed the background for bottom raking damage, the alternative approaches, areas considered vulnerable, requirements for damage extent and so on. At the final intersectional working group meeting held in April 2000, the bottom raking damage extents as given in Table 1 were agreed upon for all craft for those parts of the hull in contact with the water at service speed, and these were subsequently approved by the SLF Sub-Committee of the IMO Maritime Safety Committee. The new HSC Code (2000) was finalised in November 2000, under resolution MSC.97(73). It came into force on July 1st, 2002.

2.1.2

An overall view of the new issues of the HSC Code (2000)

The main new issues are given in Tables 1, 2 and 3. TABLE 1 - CHAPTER2 ( STABILITYAND BUOYANCY) Topics

Bottom damage (inadequacy of damage extent in 1994 HSC Code)

Bow doors on Ro-Ro Craft Compliance methodology

Improvements

Based on incidents feed-back and theoretical work, new requirements based on raking damage scenario introduced : 9 bottom damage considered only for areas which are underwater at service speed 9 length of the bottom damage depending on craft category and length: (a) 55% L from forward (all craft), (b) 35% L anywhere in length (all craft over 50m length), (c) (L/2+10)% L (all craft up to 50m length), (d) 100% L (Category B craft) Comment: Item (c) is the result of a compromise taking into consideration the validation results from the smaller craft. This compromise can be justified to some extent by the work of Simonsen (2001), which indicates that smaller craft have a lower probability of sustaining full-length raking damage, which is attributable to the scantling being heavier than in direct relation to craft size. On Ro-Ro craft with bow loading openings: - requirements introduced for inner bow door, - relaxation accepted if demonstration that certain conditions are fulfilled starting from worst wave height Alternative methods introduced for demonstrating compliance with criteria (model testing, mathematical simulations, full scale)

Structural Design of High Speed Vessels

Load line

115

Requirements introduced: - watertight doors: door position indicators - integrity of superstructures: weathertight doors and windows - openings in hatchways and machinery spaces ventilators, scuppers, discharge, air pipes freeing ports and guard rails - marking of load-line TABLE 2 - CHAPTER 4 ( ACCOMMODATIONAND ESCAPE )

Topics

Improvements

Empirical formulae modified to avoid arbitrary step changes in case of catamaran: - wetdeck clearance (CH factor simplified) Design head-on - collision speed (reduced down to 2/3 of operational speed) collision acceleration - 6g introduced for hovercraft (in line with British Hovercraft Safety Requirements) gcon For collision affected zone, new energy-based approach introduced Comment: Calibration done such that a SOLAS 22 knots steel ship has a limit similar to that imposed by regulatory location of the collision bulkhead Design accelerations Such accelerations to be considered for: for aft, transverse, positioning and securing of equipment and baggage in public spaces vertical directions deformation or detachment of seats, LSA, heavy items and sub-structures Static tests in 3 directions requested whatever the level of acceleration gcon Testing of seating Dynamic seat testing requirements rationalised TABLE 3 - CHAPTER 7 ( FIRE SAFETY) Topics

Improvements

Classification of hazard areas improved with more accurate definitions Hazard areas Crew accommodation with sleeping berth considered Sales shops classified according to deck area (50m z) Carriage of New part D issued by Fire Protection Sub-committee. dangerous goods Comment: In HSC Code (1994) carriage of dangerous goods was not addressed Ventilation systems Enhanced ventilation systems introduced for galleys, thus not discouraging the fitting of galleys on board HSC Fixed sprinkler system requested in crew accommodations with sleeping berths. Fixed sprinkler systems on cat A HSC no longer requested if simultaneous conditions are verified: smoking not permitted, Sprinkler systems - no sales shops, galleys, service spaces, ro-ro spaces fitted on board, capacity not more than 200 passengers, - voyage duration less than 2 hours at operational speed

2.1.3 Structuraldesign No new requirements have been introduced in Chapter 3 of the HSC Code (2000), which deals with the structural design by means of general requirements and objectives: - Adequate material for the intended use, Under all permitted operating conditions, withstanding of static and dynamic loads without

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Specialist Committee V.4

inadmissible deformation, loss of watertightness or interference with safe operation of the craft, Withstanding of cyclic loads, including vibrations, without impairing the integrity of the structure, without hindering normal functioning of machinery and equipment, and without impairing the ability of the crew to carry out its duties, Design criteria to be accepted by Flag Authorities, Full-scale trials possibly requested by Flag Authorities to crosscheck design assumptions.

For practical application, the above general requirements have to be supplemented by detailed technical guidance and criteria, such as Classification Rules requirements.

2.2

The present situation with Classification Societies

2.2.1

The importance of Class requirements with respect to HSC Code compliance

For practical application, the general requirements of the HSC Code (2000), as listed in 2.1.3, need to be supplemented by detailed technical guidance and criteria. In that respect, dedicated Classification Rule requirements are necessary. However, the lack of detailed information on targeted structural ability makes different interpretations possible, and thus asks for a detailed comparison of dedicated Classification Rules. Applicable Rules of the main Classification Societies are compared hereafter: ABS Rules (1997), BV Rules (2002), DNV Rules (1999), GL Rules (2002), LRS Rules (1996), NKK Rules (2001), RINA Rules (2002), RMS Rules (2000). It is to be noted that BV Rules (2002), GL Rules (2002) and RINA Rules (2002) have been jointly developed in the framework of EEIG UN1TAS, and that their technical content is therefore identical.

2.2.2

Comparisons of Classification requirements influencing structural design

2.2.2.1 Operational envelope - Characterising parameters Comparison of Rules in term of operational envelope is given in Table 4 and Table 5 hereafter. TABLE 4 - DEFINITIONOF HIGH SPEED CRAFT- CRITERIA ABS (1997) BV (2002)- GL (2002) NKK (2001)- RINA (2002) RMS (2000) DNV (1999) LRS (1996)

Minimum Froude number (V / ~/L > 2.36) (ref. Part 1, Sec. 1) Speed versus displacement, similar to HSC Code limit (2000) High Speed Craft: idem BV / GL / NKK / RINA / RMS Light Craft: minimum displacement depending on length and breadth (ref. Part 1, Ch.1, Sec. 2) High Speed Craft: idem BV / GL / NKK / RINA / RMS Light Displacement Craft: minimum displacement depending on length and breadth (ref. Part 1, Ch.2, Sec. 2).

Structural Design of High Speed Vessels

117

TABLE 5 - SERVICELIMITATIONS Service limitations are considered, where applicable (ref. Part 1, Sec. 1): geographical limitation, to be specified in the Class notations, where special modified requirements have been considered, design significant wave height, when less than 4m, to be explicitly mentioned in operating manual for restriction of service. Sea areas (4) are introduced, in terms of maximum significant wave height observed during 90% of the year at least (Ref. Ch.3, C3.3). -

ABS (1997)

-

BV (2002) GL (2002) RINA (2002) DNV (1999)

LRS (1996)

NKK (2001)

RMS (2000)

Service area restrictions (7) are introduced (ref. Pt.1, Ch.1, Sec.2) in terms of maximum distance to the nearest harbour, for winter, summer and tropical conditions as defined in ILLC 1966, Annex II, Appendix A. For light craft, the same service area restrictions apply, reduced by a factor V / 25, where V is the maximum speed. Service area restrictions (6) are introduced (ref. Pt. 1, Ch. 1, Sec.3) in terms of maximum distance to the nearest harbour, or refuge. General statements that <>are introduced (ref. Part 1 - C h 1). Speed versus displacement criteria, similar to the HSC Code (2000). Critical and worst intended conditions are defined identical to HSC Code (2000). Maximum wave height and wind velocity of the worst intended conditions are set out after testing of the first HSC of a series according to a programme approved by RMS.

2.2.2.2 Design loads and stress criteria Comparison of Rules in term of parameters influencing design loads is given in Table 6 hereafter. TABLE 6 - PARAMETERSINFLUENCINGDESIGNLOADS

ABS (1997) BV (2002) GL (2002) RINA (2002)

DNV (1999)

LRS (1996)

NKK (2001)

The 1/100 ~ value of vertical acceleration at LCG. Comment: tabulated from craft parameters, design significant wave height (not to be less than L/12 for unrestricted service) and design speed (Ref. Pt.3, Sec.8) The 1/100 ~ value of vertical acceleration at LCG. Comment: to be given by the designer, without being less than a calculated value, depending on type of service and sea area, for which characterising significant wave heights are explicitly given (Ref. Ch. 3, C3.3) The 1/100 ~ value of vertical acceleration at LCG. Comment: to be given by the designer, without being less than a calculated value, depending on service notation and sea area restriction (Ref. Pt.3, Ch.1, Sec.2) The 1/100 ~ value of vertical acceleration at LCG. Comment: depending on type of vessel, displacement aspect ratio and Froude number, to be tabulated from craft parameters, design significant wave height (not to be less than tabulated value depending on Service Group lower values accepted if properly documented) and design speed ( Ref. Pt.5, Ch.2 ) The significant (1/3 value) vertical acceleration at fore end (ref. Part 5 Ch. 1). Comment: to be specified by the builder, with minimum values given by Rules depending on intended service area (no restriction, coasting service, smooth water service no documentation of corresponding maximum sea state) and type of craft (passenger, cargo

118

Specialist Committee V.4

RMS (2000)

craft) (ref. Pt.5, Ch. 2). Craft of length and speed above given limits must have their primary structure strength and hull girder strength demonstrated by direct analysis accounting for dynamic effect as appropriate, without further guidance. The parameters influencing design loads, the design loads and the criterias are subject to special consideration, except for hydrofoils and ACV' s, where Rules for Dynamically Supported Craft (1990) apply (ref. Part II ). At the same time, the calculation of external forces on hull and appendages is requested for submission to Class (ref. Pt. I - w 5).

A comparison of the Rules longitudinal distribution of the vertical acceleration is given in Fig.1.

3

..........

0 0

2.5

'~

[] ABS

2

---,n--BV, RINA, GL

1.5

•

1

DNV

.......~ .......LRS ( depending also on LCG )

0.5

:

X/L 0

0

0.1

0.2

0.3

0.4

,

,

,

0.5

0.6

0.7

,,

, ...........

0.8

~,,,

0.9

,

,

1

1.1

NKK

1.2

Fig. 1 - Longitudinal distribution of vertical acceleration Comparisons of Rules in terms of design loads, hull girder strength and stress criteria are given in Table 7, Table 8 and Table 9 hereafter. TABLE 7

-

-

BV (2002) GL (2002) RINA (2002)

LOADS

Design loads on bottom and side shell are: impact loads, depending on vertical acceleration at the considered location, sea pressure adjusted by range of design significant wave height, - at fore end, a dynamic pressure on side structure is also considered, depending on design speed of the vessel and adjusted by range of design significant wave height (Ref. Pt.3, Sec.8 ). Specific slamming pressure on cross-deck of multihulls is also given, depending on the design speed and the vertical acceleration at the considered location. Loads on decks are only depending on vertical acceleration at the considered location for concentrated load. Design loads on bottom are : impact loads, depending on vertical acceleration at LCG and longitudinal location of considered section, sea pressure. Sea pressures on bottom and side shell are depending on the vertical craft behaviour, characterised by the vertical acceleration at LCG (Ref. Ch.3, C3.5.5), with minimum value only depending on the craft length. In addition for catamarans, the impact pressure on cross-deck explicitly refers to -

ABS (1997)

- DESIGN

-

Structural Design of High Speed Vessels

significant wave height and craft speed. All the loads on decks (uniform of concentrated) are depending on vertical acceleration at the considered location. Design loads (Ref. Pt.3, Ch.1, Sec.2) on bottom are sea pressure, function of craft parameters, with a minimum depending on sea area notation, and: For all craft, pitching slamming pressure, function of the craft characteristics, applied on a length from fore end function of maximum speed, In addition, for craft with Froude number V / ~/L > 3, slamming impact pressure, depending on vertical acceleration at LCG and longitudinal location of considered section. The sea pressure on side shell is a function of craft parameters, with a minimum depending on sea area notation. The loads on decks are all depending on vertical acceleration at the considered location. Design loads (Ref. Pt.5, Ch.2, Sec.5 & 6) on bottom are hydrodynamic wave pressure, function of craft parameters, type of craft and range of Froude number and volumetric ratio, hydrostatic load pressure and bottom impact pressure. The latest is: - Depending on vertical acceleration at LCG for monohulls or multi-hull craft with V / ~/L _>3 or small volumetric ratio, - Depending on craft speed for craft with V / ~/L < 3 and large volumetric ratio. For all craft, pitching slamming pressure, function of the craft characteristics, applied on a length from fore end function of the maximum speed, Depending on significant wave height and craft speed when craft is supported by hydrodynamic lift. The sea pressure on side shell is a function of craft parameters. The load on weather deck is depending on craft parameters and craft speed (Ref. Pt.5, Ch.2, Sec.7). Design loads are applicable to monohull of less than 50m in length operated in displacement mode, other craft being subject to case by case consideration. The extent of strengthened bottom forward is from the fore end to a point abaft the forward end depending on scantling length, maximum speed and full load displacement (ref. Part 5 - Ch. 1). The design load on bottom is analytically given as a function of : geometrical parameters of the section, - vertical acceleration at fore end, advancing speed, not less than a value derived from a Froude number based on vertical acceleration at fore end, minimum load value function of draft, breadth, vertical acceleration at fore end and type of craft. The design loads on other parts of the structure are all depending on the vertical acceleration at fore end, except load on exposed deck, depending only on craft length and intended service area. -

DNV (1999)

-

LRS (1996)

-

-

NKK (2001)

-

-

-

RMS (2000)

119

See Table 6

Special&t Committee V.4

120

TABLE 8 - HULL GIRDERSTRENGTH For hull girder longitudinal strength (Ref. Pt.3, Sec.6): A minimum hull girder section modulus is given, depending on craft dimensions and maximum speed for the design sea state, - In addition, for craft with maximum speed above 25 knots, a second criterion for hull girder section modulus is given, depending on vertical acceleration at fore end (1.2 times vertical acceleration at LCG for L<6 lm and V<35 knots, determined by model tests otherwise). For hull girder longitudinal strength of monohulls and multi hulls, bending moment and vertical shear force (Ref. Ch.3, C3.4): Analytic still water and wave bending moments are considered for all craft, with wave induced loads depending on main craft parameters and also the sea area, In addition, for craft less than 100m in length, another expression of total bending moment is given, depending on craft parameters and vertical acceleration at LCG, Analytical expression of additional sagging bending moment due to impact at bow is also given, based on actual mass and buoyancy distribution and linear longitudinal distribution of vertical acceleration, - Vertical shear force is derived from total bending moments, Transverse bending moment and shear force on catamarans are given in terms of craft parameters and vertical acceleration at LCG. For hull girder longitudinal strength (Ref. Pt.3, Ch.1, Sec.3), requirements are only given for monohull with length / draft ratio more than 12, or with length more than 50m, and for twin hull. For those craft: Analytic still water and wave longitudinal bending moments are considered, Additionally, total longitudinal bending moment is given, depending on craft parameters and vertical acceleration at LCG, For large bow flare on monohulls, additional sagging bending moment due to impact at bow is taken into account by 20% increase, Transverse bending moment on twin hulls with L<50m and V / ~/L > 3 and shear force on all twin hulls are given in terms of vertical acceleration at LCG. Transverse bending moment on twin hulls with L>50m is given in terms of vertical acceleration at LCG and in terms of maximum speed. For hull girder longitudinal strength (Ref. Pt.3, Ch.5, Sec.l&2), requirements are given as follows: Analytic still water and wave longitudinal bending moments are considered, sum of which is depending on the service area, Dynamic bending moment due to impact is also given, depending on craft parameters and vertical acceleration at LCG for multi-hull vessels or vertical acceleration at LCG, for and aft parts for monohulls. The maximum hull girder bending moment at midship is also given as a function of vertical acceleration at bow and craft dimensions, with additional minimum value similar to conventional vessels for craft with length more than 60m. See Table 6 -

ABS (1997)

-

BV (2002) GL (2002) RINA (2002)

-

-

-

DNV (1999)

-

-

-

-

LRS (1996)

-

-

NKK (2001) RMS (2000)

Structural Design of High Speed Vessels

121

TABLE 9 - STRESS CRITERIA ABS (1997) BV (2002) DNV (1999) GL (2002) LRS (1996) RINA (2002) NKK (2001) RMS (2000)

Rule stress criteria are deterministic and are given for each structural member depending on the nature of the load. Rule stress criteria are deterministic and are given for each structural member depending on the nature of the load and the location of the structural member.

Rule stress criteria are deterministic and not depending on the nature of the load, but only on the nature of the structural element under consideration. See Table 6

2.2.2.3 Service restriction- Format Comparison of Rules in terms of service restriction is given in Table 10 hereafter. TABLE 10 - SERVICERESTR/CTION ABS (1997)

BV (2002) GL (2002) RINA (2002)

DNV (1999)

LRS (1996)

Speed reduction versus wave height is considered and guidance is to be given in the Operating manual (Ref. Pt.3, Sec. 8). Influencing operational parameters for all types of loads are detailed (Ref. C3.3.3). The limit operating conditions format is left to the designer. A format such as a speed reduction versus wave height limitation and indication of maximum allowed wave height is suggested. The limit operating conditions are indicated in the Class Certificate. Installation of a hull monitoring system for vertical acceleration may be required. For monohulls, depending on the volumetric aspect ratio and design Froude number, the Rules state relationships between craft characteristics, actual wave height and actual craft speed for a given vertical acceleration at LCG (Ref. Ch.3, C3.3.3.2). For catamarans, this relationship is to be assessed by tank tests or full-scale tests. Speed reduction versus significant wave height is developed, stated in the <<Appendix to Classification Certificate >>and posted in the wheelhouse (Ref. Pt.3, Ch.1, Sec 2). Notice is given that significant wave height coincides well with visual observation of the <~wave height >>by an experienced person. Fitting of an accelerometer at LCG may be required. The allowable speed corresponding to the design vertical acceleration at LCG may be estimated by proposed formulas between design acceleration, craft characteristics, craft speed and significant wave height, both for V / ~/L > 3 (type of vessel being taken into account by a factor ranging from 0.7 to 1) and for V / ~/L < 3. Operational envelope (allowable speeds, significant wave heights and corresponding displacements) forms an appendix to the Classification certificate and is to be incorporated in the Operational Manual. It is to clearly displayed in the wheelhouse. Fitting of an accelerometer at LCG may be required, with visual display in the wheelhouse. The allowable speed corresponding to the design vertical acceleration at LCG may be estimated by proposed formulas between design acceleration, craft characteristics, craft speed and significant wave height, for monohulls or multi-hull craft with V / ~/L > 3 or small volumetric ratio and for craft with V / ~/L < 3 and large volumetric ratio.

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Specialist C o m m i t t e e V.4

NKK (2001)

RMS (2000)

No information is given regarding an), kind of service restriction. Classification Certificate is to indicate the worst intended conditions, the maximum allowed distance from places of refuges and other limitations, as necessary. In addition, a Seaworthiness Certificate is issued, with indication of particular routes where the craft is allowed to operate with due regard to weather conditions and distance allowed. No mention is made of any speed versus wave height reduction.

3.

LOADS

3.1

Introduction

This chapter addresses the problem of structural loads on high speed vessels. The loads can be grouped in global loads and local loads. Global loads are important from the structural design point of view for larger vessels, with, roughly, L > 50m. Local loads, including slamming loads, determine the structural strength of smaller high speed vessels. For larger vessels slamming type loads are important for the design of local structures, the associated whipping vibration degrades the fatigue life of the structures (Hansen et al., 1995) and in some cases there may be a contribution to the global loads (Jensen et al., 2001). Local loads are external forces and pressures that are applied directly to a given structural element such as a plate or stiffened panel. The report of the ISSC 2000 Specialist Committee V.2 describes the main sources of local loads. Global loads are given by the contributions of static still water loads, inertial loads, hydrodynamic wave induced loads and whipping effects due to impact slamming. Still water loads are usually determined by the difference between the distributions of ship weight and hydrostatic restoring forces, however at high speed the effects of steady sinkage and trim may significantly modify the steady vertical loads compared to the zero speed condition. Inertial loads are easily determined once the wave induced motions and accelerations are known. The greater effort and larger uncertainty is in fact on the determination of the quasi-static wave induced loads and slamming induced loads. This is an area where much research has been carried out in recent years, and it will be the main topic of this chapter. The specific characteristics that distinguish the wave induced loads on high speed monohulls from conventional monohulls are the important higher harmonic content of the time histories even in regular waves (Karppinen et al., 1993, K6hlmoos et al., 2001), and the frequency and severity of slamming occurrences. Due to the high speed of the vessels and the often large flexibility of the structure, the high frequency of encounter in head and bow waves may excite the lower hull girder natural frequency and induce springing vibration (Hansen et al., 1995). Additionally, non-linear effects produce exciting forces with higher harmonics that can also excite one of the natural frequencies of the structure. The effect of springing on the extreme loads is usually accepted to be small, however it may increase drastically the accumulated fatigue damage (Jensen and Dogliani, 1996). Twin hull vessels like catamarans and SWATHs experience critical global loads on the cross deck structure, namely the vertical shear force and bending moment (usually maximum for zero speed and beam waves) and the pitch connecting moment (usually maximum for bow waves). For SWATH ships the structural design of the vertical struts is also critical. Head and following waves may induce large relative motions leading to wet deck impact slamming, which may be more severe than bottom bow slamming since the impact velocity may be nearly perpendicular to the wet deck and the presence of the side hulls restrains the escape of water when the impact with the free surface occurs.

Structural Design o f High Speed Vessels

123

Finally, it should be mentioned that a lower cross deck results in lower vertical bending moments at the cross deck by reducing the moment arm of the horizontal hydrodynamic forces acting on each hull; however the probability of impact of waves on the cross deck increases. In a foil-catamaran the weight of the vessel is supported by submerged foils when it travels at high speed. Compared to catamarans, the foil-catamaran concept has lower resistance and better seakeeping behaviour in small sea states (Faltinsen, 1996). A problem area from the structural point of view is the design of the struts since they will be subjected to large horizontal wave induced loads. Surface Effect Ships (SES) are partly supported by a pressurised cushion of air enclosed between two hulls and flexible seals at the bow and stern. The advantages of the SES concept are lower resistance and better seakeeping characteristics. One problem specific of the SESs is the resonance pressure variations in the air cushion, known as cobblestone effect, that occurs in small wave periods. This effect influences the heave and pitch motions and consequently the wave loads and in particular the relative motions that lead to slamming events. Another phenomena that influences the vessel motions is the air leakage from the cushion, especially in higher sea states. More detailed descriptions of the seakeeping characteristics and problem areas of SES vessels, as well as of methodologies to estimate wave induced motions and loads, can be found at the report of the ISSC 1994 Specialist Committee V.4, and Faltinsen (1996). The pentamaran concept consists of a large and slender central hull and four sponsons connected to the central hull (Gee, 1999, K/3hlmoos et al., 2001, Dudson et al., 2001). The advantages of this concept compared to conventional container ships are lower resistance both in still water and in waves and better seakeeping behaviour concerning vertical motions and added resistance in waves. The characteristics of global load types and slamming loads are basically the same of those of fast monohulls, but additionally it is necessary to consider the shear forces and bending moments between the sponsons and the central hull.

3.2

Review of Calculation Methods

This section reviews the methodologies currently available to calculate the wave induced loads on fast vessels. In general the simpler methods provide more robust and fast solutions; however the range of applicability is more limited and some important physical phenomena may be neglected. More complete methods consider a higher degree of detail, but the solutions are more complex and time consuming and in some cases not practical for engineering applications. The simpler methods to calculate the seakeeping of fast vessels are extensions of the well established strip theories. Strip methods assume that the hull is slender, the frequency is high and the speed of advance is small. The later assumption is obviously a big limitation for fast vessels, however several authors report satisfactory agreement between numerical predictions and experiments of wave induced motions and structural loads, as for example Lee et al. (1973) for a catamaran, Lee and Curphey (1977) for a SWATH, Karppinen et al. (1993) for a fast monohull, and Ge et al. (2002) for a flexible catamaran subjected to wet deck slamming. For twin-hulls the hydrodynamic interaction between the hulls may be considered on the cross flow plane, or neglected if the hull separation and forward speed are sufficiently high. For SWATH vessels it is important to include in the formulation the viscous effects associated with oscillatory motions since for these hulls the potential flow damping is only a part of the total damping. An empirical formulation based on the cross flow approach of aerodynamics may be used (Lee and Curphey, 1977, Chan, 1993, Rathje and Schellin, 1997).

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Specialbst C o m m i t t e e V.4

Another approach is based on a frequency domain panel method applying zero speed Green function. The hull is three dimensional; however the forward speed effects are introduced in a simplistic way similar to strip theories. The numerical solution is robust and relatively simple to obtain. This approach has shown good correlation with experimental data by Chan (1993) for a catamaran and a SWATH, K6hlmoos et al. (2001) for a pentamaran, and Schellin et al. (2002) for a fast monohull. The use of panel methods based on the forward speed Green function to solve the seakeeping problem of fast vessels is more consistent from the theoretical point of view; however the numerical solution becomes much more complex and the results are not always satisfactory. Results for a trimaran are presented by Bingham et al. (2001) and for a fast monohull by Schellin et al. (2002). The high speed strip theory, or 2YzD theory, was developed especially for high speed vessels. The method assumes that the ship is slender and fast so that the generated waves propagate only downstream. This way the hull is divided into a number of cross sections, and the solution is based on the two-dimensional Laplace equation but three-dimensional free surface condition. The problem is solved by starting the solution at the bow and stepping it downstream. The method has been applied to calculate global loads on catamarans by Faltinsen et al. (1992), to investigate the slamming loads on a catamaran by Zhao and Faltinsen (1992) and to predict the slamming occurrence on catamaran cross structures by Grande and Xia (2002). In the former paper (Zhao and Faltinsen, 1992) the authors concluded that the wet deck slamming forces have an important effect on the vertical accelerations and vertical motions, thus slamming loads on the wet deck cannot be estimated by a theory that neglects the effects of slamming loads on the wave induced motions. The high speed strip theory was generalised to represent the flexible hull modes in addition to the rigid body modes. This high speed hydroelastic theory has been applied to investigate the responses of monohulls (Wu and Moan, 1996), catamarans (Hermundstad et al. 1999) and a pentamaran (Dudson et al. 2001). The high speed strip theory may be thought as being between strip methods and fully three-dimensional panel methods, both in terms of detail of the hydrodynamic flow and computational effort. However, while the most important forward speed effects for fast slender hulls are retained in the formulation, the numerical solution is still compatible with practical applications. Important non-linear effects on the global responses have been identified in experimental results obtained with models of conventional ships (ISSC 2000 Specialist Committee VI.1 report) and also with models of fast ships (Karppinen et al., 1993, KOhlmoos et al., 2001). The non-linear effects on the global loads are identified by the large magnitude of higher harmonic components even in regular waves and by the asymmetry of the positive and negative peaks. Additionally other non-linear effects characteristic of high speed vessels, such as wet deck slamming on catamarans, or air leakage on SESs, may need to be coupled to the motion responses to be calculated directly in the time domain. Non-linear solutions of the wave induced motions and loads are almost always obtained in the time domain and there are different levels of complexity, ranging from those methods derived directly from frequency domain strip theories to fully non linear methods. The report of the ISSC 2000 Specialist Committee VI.1 presents a comprehensive review of the existing non-linear procedures, and the ISSC 2003 Loads Committee reports the latest developments in this area. 3.3

Recent Work

This section reviews the research work on wave loads on high speed vessels published since the ISSC congress of 2000.

Structural Design o f High Speed Vessels 3.3.1

125

Global Loads

Heggelund et al. (2001) present and discuss procedures and criteria for determination of operational envelopes and global design loads for non-planing high speed catamarans. The influence of operational restrictions on design loads is discussed, and the authors found that simplified formulas commonly used by classification societies for predicting operational limits significantly overpredict the reduction of motions and wave loads at reduced speed. The operational envelopes and design loads are calculated for a 60m catamaran. Design loads are calculated by short-term statistics based on linear hydrodynamic theory. The use of active foils reduces the global vertical bending moment loads. When informal operational limits are used, the design loads for vertical and transverse bending found by direct calculations are approximately at the same level as design loads given by classification societies. For torsion, however, the design values found by direct calculations are substantially lower than design loads given by the rules. Whereas stresses and deformations due to vertical bending and torsion can be found by simple beam theories, transverse strength analysis has to be performed by a finite element analysis. Jensen et al. (2001) present full-scale measurements of seakeeping trials carried out with a 47m SES fast patrol boat. The authors present a method for estimating global loads based on measurements of strains using networks of fibre optic Bragg strain sensors, together with finite element analysis. The trials were conducted in the North Sea in sea states with significant wave heights up to 6,5m and relatively small wave periods, which may be considered extreme sea states. Depending on the conditions, speeds up to 45 knots were tested. The extreme structural loads obtained from the measurements were compared with the design loads given by Det Norske Veritas - High Speed Light Craft (HSLC) rules (1993, 1996). The measured extreme sagging and hogging bending moments exceed the DNV HSLC (96) design values by a large margin (up to 2 times). The HSLC rules of 1993 are much more conservative, and in this case the measured loads are below the rule values. Horizontal bending, torsion moment and vertical shear are well below the rule values of 1996. An examination of the time series showed that all maximum loads that exceeded the HSLC rules were caused by frontal or bow flare slamming impacts, which shows that impact loads have a significant influence on maximum global loads. Due to the stealth properties required for this particular vessel the front panels are flat and with a relatively steep angle in relation to the horizontal plane. This type of bow section may experience large slamming pressures. The maximum estimated pressure during the trials is around twice as large as the DNV HSLC (1993) requirements. Sebastiani et al. (2001) investigate the slamming pressures acting on the fore body of a large fast monohull vessel. Based on the analysis of the model tests results, a formula for predicting the peak impact pressure as function of the relative vertical velocity is proposed. The formula is validated against full scale results. Thomas et al. (2001) present results of full scale hull stress and motions on a 96m Incat high speed catamaran. The investigation focused on the occurrence and nature of severe slam events. Correlating the strain gauge readings measured during an extreme slam event with results from a refined finite element model, it was possible to develop a realistic load case of an extreme slam event. Detailed analysis of the most severe slam events showed that these are wet deck slams, as opposed to bottom impact slamming or bow flare slamming. It is concluded also that the most severe wet deck slams force the bow to change direction, then influencing the vertical ship motions. Dudson et al. (2001) outline the work undertaken to determine the design global loads on the ADX Express pentamaran, a high-speed trans-Atlantic container ship capable of performing at speeds up to 41 knots. An extensive series of model tests were performed with a segmented model in order to gain insight

126

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into the reliability of the calculated loads. A comparison is made between the loads obtained from direct calculations and those given by the rules. The experience during the project is that both class rules and the L/20 wave underestimate the design vertical bending moment by a considerable amount. The high speed strip theory (Wu et al., 1996) is used to calculate linear and non-linear ship motions and loads and global flexible modal responses. The results from the non-linear version compare very well with the experimental data for the vertical motions and wave loads. The effects of hull flexibility and non-linear springing on the fatigue life of the vessel are discussed as well. In order to investigate to what extent hull flexibility will affect the fatigue life of structures, the authors established two time histories of the nominal stress for each sea state and wave heading. The first is the original time history, the other is the low-pass filtered one. Therefore, the influence of hull flexibility is present in the former time history and absent in the latter. The conclusion is that most of the fatigue damage occurs in head and bow waves, and that the hull flexibility increases the long-term fatigue damage of some structures on the upper deck by around 50%. K6hlmoos et al. (2001) describe the method used by Germanischer Lloyd to determine the global design loads for the same pentamaran. The rules underestimated the design loads, and direct calculations were used to determine them. The wave induced hydrodynamic loads are calculated by using a frequency domain panel method with zero speed Green function, and non-linear corrections are included by integrating pressures up to the wave elevation. The results are pseudo-transfer functions, that are nonlinear with respect to wave amplitude. These are used together with a practical procedure to calculate long term distributions of the wave induced structural loads (Guedes Soares and Schellin, 1995). The design loads were obtained by superimposing still water loads, wave induced loads and slamming loads (estimated according to GL HSC rules). The pressure distribution necessary for FE analysis was calculated for an equivalent regular wave with such an amplitude to generate the corresponding design loads. To validate the hydrodynamic procedure, numerical results of the vertical bending moment are compared to model tests results and a good correlation is reported. FE models gave the opportunity to better evaluate the structural scantling, which had to be increased in certain areas with respect to rule requirements, but could be optimised elsewhere. Reductions in scantling demonstrate the benefit of performing direct calculations, allowing the designer to optimise the vessel structure at an early design stage. Gu et al. (2001) present results of full scale (L=13.65m, Vmax=50kn) and model scale tests for a high speed foil assisted catamaran. The aim of the tests was to obtain slamming pressures, vertical accelerations and local and global stresses in the structures, in order to apply a procedure to calculate design slamming pressures. The measured slamming pressures are reduced in order to transfer the transient peaks into uniformly distributed design pressures that can be used by existing structural rules. It was found that the calculated design slamming pressures are larger than those required by DNV rules in stern and bow regions. Garme (2001) presented a two-dimensional non-linear time-domain simulation model for planing hulls in head seas together with results from model experiments. The local water surface deformation pile-up was added to the wave height. Simulated time histories of heave and pitch motions, accelerations and section loads were compared with measured ones. In irregular severe head seas the hull completely leaves the water and re-enters. In the vicinity of the transom a local model should be superimposed to correctly model the flow and satisfy the Kutta condition in the alongships direction. With a correction of the simulated force at the aftermost sections the agreement between simulations and model tests becomes very good. Coppola and Mandarino (2001) gave a discussion about the preliminary design of trimarans. A method has been developed for the global load evaluation. The numerical example shows that the torsion stresses

Structural Design of High Speed Vessels

127

are negligible for trimarans; the shear and bending stresses are instead very relevant and governing for the transverse strength. Grande and Xia (2002) investigate the statistical distributions of slamming occurrence and slamming pressure magnitudes in random seas for a fast ferry catamaran and a racing sailing catamaran. A partly non-linear high speed strip theory (Wu and Moan, 1996) is used to calculate the ship responses to the waves. The slamming frequency was calculated using the time domain results from the non-linear code and also using the traditional linear frequency domain approach. The latter predicts a drastically higher slamming frequency than the direct approach, which demonstrates the importance of considering nonlinear effects in the calculation of vertical motions. Schellin et al. (2002) present numerical and experimental results of the motions and global loads on a fast ferry in head regular waves. Two sets of numerical results are presented. One is based on the use of an existing frequency domain panel code that formulates the potential flow problem by means of a zero speed Green function. The other set of results is based on a modified version of this code that implements a free surface forward speed Green function using the Fourier-Kochin formulation to fully account for forward speed effects. The latter formulation is theoretically more consistent for high speed vessels; however the results from the zero speed Green function panel method compared better with the experimental data than the results of the more advanced code in almost all cases. This is due to numerical difficulties in solving the forward speed Green function problem, and shows that the three dimensional fully linear problem, accounting completely for the interaction between the steady and unsteady flows in the linear sense, is still a challenge.

3.3.2

Slamming Loads and Structural Responses

Takemoto (2000) proposed a method to calculate hull responses of high-speed vessels in waves taking account of slamming impact loads. He adopted a threshold velocity of water impact and a concept of equi-added-mass lines to compute the relative velocity between hull and wave surface. Model tests of a 40m patrol boat in regular waves were conducted. Measured hull responses showed good agreement with the calculated results. Takemoto et al. (2001, 2002) proposed a simplified method to estimate impact pressures on high-speed vessels in waves. The method is similar to that of Stavovy and Chang (1976), but the impact pressure is calculated by Wagner's formula instead of Chuang's empirical formula. A correction factor for impact pressure can be introduced by using Ferdinande's impact theory when the deadrise angle is large. The measured pressures in Chuang's three dimensional drop tests were compared with the estimated values. Fairly good agreements were observed taking the effects of the elastic deformation of the test device into consideration. Rees et al. (2001) described the development of a numerical modelling technique to generate rough-sea dynamic load data for use in the design of high speed planing craft. The dynamic response of the hull is taken into account when calculating the hydrodynamic behaviour. The code applies the method of Stavovy and Chang (1976) to predict slamming pressure distributions over the hull surface. Faltinsen (2002) proposed a method to analyse the water entry of a wedge with finite deadrise angle by matched asymptotic expansion. The water is incompressible and the flow irrotational. The method is simple and robust. A jet domain, inner domains at the spray roots and an outer domain are defined assuming the deadrise angle of the wedge is finite. Comparisons with the exact similarity solution by Dobrovol'skaya (1969) show that the method is applicable for large deadrise angle. Predictions of jet thickness, kinetic energy and mass flux into the jet agree well up to 45 degrees.

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Sebastiani et al. (2001) have discussed the characterisation of slamming pressures acting on the fore body of a deep-V fast monohull vessel, gathering the results of several years of research in numerical simulation, model-scale and full-scale measurements. In particular a correlation formula for the prediction of the peak impact pressure based on the relative vertical velocity was formulated based on the analysis of a large series of specific model tests. It was found that a 3D forward speed correction to 2D wedge theory provided a fair agreement with the measured data. The validity of the proposed formulation was further checked against available full-scale results on the same vessel. Vredeveldt et al. (2001) presented their work on full-scale slam tests and the results of scaled slam tests with a model-scale transducer. In the adopted approach, a slam is seen as an impact introducing vibrations in the ship structure. The response of a shell subjected to a slam can be described by a set of natural vibration modes. It is shown that the natural frequencies yield a base for damage prediction. It is concluded that relatively simple mechanical models seem adequate for describing hydrodynamic impacts such as slamming, with the aim of predicting damage to the local hull structure. Finite element modelling has become a standard tool for naval architects to conduct structural analyses of ship structures and studies have been conducted seeking to relate the measured stresses on-board vessels with those predicted by finite element analysis. Thomas et al. (2001) reported on an investigation into the nature and effect of severe slam events, whereby extensive full scale hull stress, motion and wave measurements were conducted on a 96m Incat high speed catamaran ferry during regular ferry service. A definition of a slam event for this type of vessel was proposed and used to identify slam events from data records. A slam was defined as having occurred if a peak in the stress record occurred where the rate of change of stress prior to the peak (MPa/s) exceeded 0.1 times the yield stress (MPa). The character and influences of these slamming events were investigated in respect to a number of factors including wave height, vessel speed, relative vertical velocity, location on vessel and time between occurrences. Particular attention is paid to the whipping response of the structure, with the principal structural response frequencies being identified through spectral analysis. A realistic load case for an extreme slam event has been developed and this was achieved by correlating the measured strain gauge with predictions from a refined FE model. Rothe et al. (2001) investigated the use of modem numerical methods to assess slamming induced structural strength of a catamaran wet deck under severe conditions. In their study, they first computed velocities of the wet deck with relation to the simulated regular head waves using a boundary element computer code. Then, for the ship at its operating limit, they applied a short-team stochastic analysis to yield significant amplitudes of relative water entry velocity. Next, they selected three characteristic situations of the wet deck entry into long-crested waves and simulated this process by numerically generating regular wave trains. The CFD code Comet yielded the resulting time series of impact (slamming) pressures acting on the wet deck. Finally, they applied these time dependent pressures as loads on a finite element model of the wet deck structure. The FE code A N S Y S computed deformations and stresses at hot spots in the plating and the longitudinal stiffeners of the wet deck structure. These computations yielded stresses that were higher than allowable limits according to classification society rules. For comparison, a FE strength analysis based on impact loads according to classification society rules was also performed, taking into account the rule based restriction of ship speed. These latter results confirmed that the structure was adequately dimensioned. The partly significant differences between loads predicted by the finite volume method and loads based on the rules and, consequently, the differences in the resulting stresses were due to the 2D application of the directly computed impact loads acting over the entire width of the wet deck. Rosen (2001) discussed a method simply based on linear interpolation and extrapolation to re-construct three-dimensional propagating impact pressure distribution, from recordings with an arbitrary matrix of pressure transducers on a hull. With pressure measurements made with several transducers, on a small

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high-speed craft in both full-scale and model scale, the method is used to formulate realistic load cases for FE-analysis. Resulting structural responses are calculated and compared with full-scale strain measurements, indicating that the re-constructed pressure distribution well resembles the actual load. Further use of the method is discussed, for example investigation of the real pressure distribution influence on panel boundary conditions, and time domain simulation of structural response when the method is used together with methods for seakeeping simulation.

3.3.3

Hydroelastic Slamming

Traditionally the structure used to be modelled as a rigid body for slamming. The water pressure is applied to the structure and then the structural response is determined. This approach often leads to overestimation of the water pressure and hence more severe structural response. If hydroelasticity is considered, it usually leads to reduction of water pressure and more moderate structural response. Hydroelasticity should be taken into account if accurate predictions of loads and structural response are required. Faltinsen (2000) gave an overview of the many water-impact problems in ship and ocean engineering. It is shown that maximum pressure cannot be used to estimate maximum slamming-induced stresses, because dynamic hydroelastic effects become important. It is emphasised that the slamming problem must be hydrodynamically analysed from a structural point of view. Comparisons between theory and full-scale measurements of slamming-induced local strains in the wet-deck of a catamaran are presented. Bereznitski et al. (2001a, 2001b) and Bereznitski (2001) studied extensively the effect of hydroelasticity for a 2D fluid-structure interaction problem of bottom slamming. As a case study a drop test performed by TNO (Netherlands Organisation for Applied Scientific Research) was chosen. In the experiment a steel plate with stiffeners was dropped into a water tank. Several different models and numerical codes such as MSC Dytran and LS-Dyna, were used for the problem of a wedge shaped structure penetrating the water at small deadrise angles. It was found that the ratio between the impact duration and the period of first mode of vibration of dry structure is the key factor in taking the decision when the solution of the structural response should include hydroelastic effects. The effect of hydroelasticity gives very strong reduction of the deflection for impacts with short duration. But when the ratio becomes more than 2.0 the effect of hydroelasticity does not play a significant role and can most likely be neglected. It was also found that the air entrapping is important for deadrise angles between 0 and 5 degrees. Wang (2001) presented a numerical method to calculate the hydrodynamic impulsive pressure acting on side plates and bottom plates at bow by using a mixed Eulerian-Lagrangian method. Non-linear boundary element and finite difference methods were applied to solve the equation of motion of the plate. Water entries of elastic wedges with various deadrise angles and plate thicknesses were solved. The numerical results indicate that impulsive loads vary with the changes of plate thickness. Sano et al. (2001, 2002) developed a computer code to study fluid-structure dynamic interaction combining the non-compressive CFD code with explicit FEM structural elements. Local pressures and stresses on the surface of a stiffened plate in collision with massive water of finite volume were calculated. The effect of stiffness of the stiffened plate and the shape of impinging water on the peak value of pressure was investigated.

3.3.4

Review of the LRC Report of the 23 rd ITTC

The report of the Loads and Responses Committee (LRC) from the 23 rd ITTC (2002) was reviewed in order to identify the aspects relevant to ISSC 2003 committee V.4 that are also of interest to rI'TC. There is not a specific chapter on the LRC report dealing with problems of high speed vessels, but there are a

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few scattered references in the report. Section 5.8 presents a brief review of numerical methods developed to calculate the seakeeping of catamarans and also relevant results from model tests. There are no specific references to structural loads calculations or measurements. A comprehensive review of the numerical and experimental work recently developed on wave impact and slamming is presented in Section 5.9. Although this information is general, it may be also applicable to high speed vessels. Section 8.3, that should cover the hydro-elastic response of high speed marine vehicles, refers to the report of the HighSpeed Marine Vehicles Committee of the 22 nd ITTC (1999) as including a comprehensive survey on this matter. Concluding, the LRC of 23 rd ITI'C only briefly reviews the problems related to the responses and structural loads on high speed vessels.

4.

STRUCTURAL RESPONSE AND ULTIMATE STRENGTH

4.1

Structural response of high-speed vessels

Many papers concerning structural response and ultimate strength have been published over recent years, but only few concern HSC. The approach to structural analysis is largely the same for high-speed vessels and conventional ships. Thus almost all analysis methods developed for conventional ships can be used for HSC, including existing general-purpose FEM programs. However, increased dynamic response (springing, whipping, etc.) due to the use of flexible, lightweight materials may require special attention, as this may increase the effective loading (particularly fatigue loading) on the structure. In general it is necessary to have a better understanding of the loading conditions to be applied in design and structural analyses of HSC, including both local and global loads, together with the need for internationally agreed reliability based ultimate limit state design criteria. There is also a need for a uniform way of treating design for restricted service. Brescia et al. (2001)presented a high speed and low vibration design for a ferry, having a cruise speed of 30 knots, first of a series of four RO/RO passenger vessels to operate in the Mediterranean Sea. The very high power installed was the main new aspect of this fast ferry, leading to doubts concerning the use of standard correction parameters and the extrapolation method from FEM calculation to predict the ship full scale vibration behaviour. A three-dimensional finite element model was made for the main and secondary structure to obtain both global and local vibrations taking into account the damping of the virtual added mass of sea water. The forced vibration response was computed for the whole operational range of the propeller and the vibration amplitude versus frequency response spectra were obtained for all relevant locations on the structure. Together with a low excitation level, the reduced vibration is the result of structural improvements, particularly in aft body areas and in some upper decks. There was no need for additional heavy reinforcement and the resulting structure was light without particular critical areas. Ship performance has been evaluated during sea trials. Comparisons between full-scale results and predictions are given indicating suitable calibrations to be adopted for this kind of high-speed ferry. The structural response of the ADX Express high-speed pentamaran (287m, 40kn) was treated by Dudson et al. (2001) and K6hlmoos et al. (2001), who also gave an overview of the methods used to predict design loads. Refer to Section 3.1.1..

4.2

Ultimate Strength of High Speed Vessels

4.2.1

Overall strength

Boote and Figari (2001) addressed the evaluation of the stress distributions at collapse over midship sections for fast mono-hulls. By using the stress distribution, evaluated from the results of the application of a component approach method to a series of recently built fast ferries, an analytical formulation for the

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prediction of ultimate bending moment has been developed. The assumed stress distribution has been verified by FEM calculations on one ship. It is noted that, in hogging condition, the main contribution to the ultimate strength of the midship section is due to deck and bottom, approximately in the same proportion. The vertical sides give very small influence and the contribution of the other decks is not significant. In the sagging condition the midship section shows a more complicated behaviour, as the contribution of each structural component to the whole capacity of the section greatly differs from one ship to another. As a consequence, the analytical formulation needs further development to improve the results and the influence of non-horizontal bottoms and non-vertical sides should be properly assessed. A large proportion of high speed vessels are multi-hull, mainly catamarans. Multi-hull ships are highly three-dimensional and an important question is whether the traditional simplified methods for global structural analysis (both stress analysis and ultimate strength analysis) are valid for such vessels. For longitudinal (hull girder) analysis, the traditional method is beam theory, while for transverse analysis the most common approach is a two-dimensional analysis of a typical "slice" representing the ship transverse structure. Heggelund and Moan (2001, 2002) investigated the use of beam theory for the longitudinal analysis of catamarans. They modified the theory to account for wide flanges and significant window openings. The benchmark was a global finite element model of a 60 metre catamaran. They reported that for this particular case the effect of warping on shear stress due to torsional loading was moderate and could be neglected. This meant that the modified theory was sufficiently accurate for initial stage design. However, they pointed out that the effect of warping is strongly dependent on vessel geometry and can become important for a vessel with narrower hulls and a superstructure that is lower with respect to total height. This seems to indicate that the only sure way of achieving accuracy is to model the whole ship. For transverse strength, Heggelund et al. (2000) found that the flexibility of the partial bulkheads typical for a twin-hull Ro-Ro vessel is so large that a transverse "slice" approach is not valid any more. They showed that as a minimum the model must include a complete compartment and the interaction with the surrounding structure must be accounted for. If springs are used for this purpose, the results are sensitive to spring stiffness so that the results could still be inaccurate. Here again, the only sure way of avoiding this interaction error is to model the whole vessel.

4.2.2

Stiffened panels

Since weight savings are of great importance in high speed vessels, least weight optimisation has been used to design stiffened panels. In this process the constraints must include the avoidance of all possible modes of buckling collapse: overall panel buckling, plate-induced and flange-induced Euler buckling of stiffeners, stiffener web buckling, and flexural-torsional buckling (tripping). Paik and his colleagues have presented a series of papers setting out a comprehensive set of algorithms for all of these buckling modes for steel panels: Paik and Thayamballi (2000), Paik et al. (2001a, 2001b), Paik and Kim (2002). Besides considering these separate buckling modes, it is also necessary to account for any interaction that may occur between them. This is especially important in doing least weight optimisation because the optimisation process tends to produce a panel for which two or more modes occur simultaneously. In particular, Sheikh et al. (2002) have investigated the simultaneous occurrence of local plate buckling and plate-induced overall buckling and have shown, by a series of non-linear finite element analyses, that this simultaneity causes an interactive buckling in which the post-buckling strength undergoes a sudden and very steep decrease, whereas the separate modes have a gradual post-buckling decrease.

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5.

ADVANCED COMPOSITE MATERIALS

5.1

Introduction; definitions

A historical overview of the use of fibre composites in high speed craft can be gained by from earlier ISSC Reports. In particular, the proceedings of ISSC 1988 and 1991 include reports from Specialist Committees dealing specifically with Composite Structures. In ISSC 1994 there are relevant sections in the reports of no fewer than four committees: II.2 Dynamic Load Effects, 111.3 Material and Fabrication Factors, V.8 Weight Critical Structures, and V.4 Surface-effect Ships. In ISSC 2000 there was some coverage by Committee V.2 Structural Design of High Speed Vessels and Committee V.6 Fabrication Technologies. Traditionally the most commonly used composites for marine applications have consisted of woven or chopped strand glass reinforcements with polyester or epoxy resins. Production has been mainly by hand lay-up or, with epoxies, by use of pre-impregnated reinforcements (pre-pregs). For sandwich cores, polymer foams (mostly PVC) and end-grain balsa have been most commonly used. These technologies have been well established for many years, though subject to constant improvement as attempts have been made to optimise the material utilisation and improve the quality of production. The term "advanced composite" is commonly used but generally undefined. Here it is assumed to include the following materials: 9 Optimised laminates using non-woven glass reinforcement fabrics 9 Composites using high-strength, lightweight reinforcing fibre materials such as aramid (Kevlar| HPPE and carbon, either alone or in hybrid combinations with each other or with glass 9 Lightweight sandwich configurations. The term is also sometimes applied to composites produced by more advanced processes such as vacuum infusion. With such processes it is often possible to achieve much higher fibre content than with more traditional processes such as hand lay-up. Numerous examples of the application of advanced composites in marine craft have been reported, such as patrol boats (Madden, 1998), as well as novel applications of more conventional composites (Anon, 1997).

5.2

Reasons for using advanced composites; hindrances to using composites

5.2.1

Advantage

The following are some of the advantages of using fibre composites in high speed vessel construction: 9 Fibre composites generally offer solutions with low weight. Advanced composites using optimised fabric formats, high-strength fibres and optimised sandwich lay-ups enable even lighter structures to be achieved with the same strength and/or stiffness. This is especially important for high-speed craft, particularly those depending on air-cushion lift systems (ACVs and SESs), and for small craft that may have to be man-handled. 9 Advanced composites offer the potential to tailor and optimise mechanical properties by building in anisotropy. 9 Composites offer good through-life behaviour with little or no maintenance. In particular, provided an appropriate resin and/or finishing gel-coat is used, corrosion is rarely a problem with most reinforcement materials. Also fatigue is rarely a problem with composites in marine applications.

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9 When properly designed, composite ship structures can offer favourable crash performance. In particular, damage tends to be localised rather than extending to areas far removed from the point of impact. Another important feature is the good resistance to tearing as compared to aluminium. 9 It is possible to incorporate multi-functional features (see Section 5.2.2).

5.2.2 Specialadvantages for naval vessels In relation to marine craft composites offer special advantages in addition to those mentioned in Section 5.2.1: 9 Non-magnetic properties that enable the vessel to be used in mine counter-measure vessels and other vessels that might come near to magnetic mines. 9 Potential to build in special structural properties (such as ballistic and shock resistance) as well as non-structural properties such as acoustic damping, fire protection, electromagnetic shielding, radar absorption/reflection/transparency (stealth), and sensors. Examples of such multi-functionality often arise initially from other military applications e.g. armoured fighting vehicles. Examples are described by Fink (2000), Mouring (1998) and Benson (1998). Furthermore, for naval vessels the initial cost penalty is sometimes less problematic than for commercial vessels that may have to show a fast return on investment.

5.2.3 Hindrances The main hindrances to the adoption of advanced composites are to a large extent similar to those for conventional composites: 9 Variability in properties. Even when rather well-controlled, closed production processes are used, some variation of mechanical properties must be expected due to lack of precision in placement of reinforcements, and variations in the basic materials themselves. A point of frequent concern is variations in the quality of bond between fibres and matrix. This is often a function of the sizing used to improve wet-out of the fibres: such sizings are usually a trade secret of the manufacturer. 9 Fire. It is well known that the fire requirements in the 1995 IMO HSC Code virtually stopped the use of composites in passenger craft for several years. The main problem has been the need to use either non-combustible or fire-restricting materials. While it has always been possible to add enough insulation to satisfy these requirements, the added weight made the use of composites unattractive. Only recently have good low-weight fire protection systems become available. 9 Doubts about damage tolerance (see also Section 5.3.1). There are really two separate problems: - Concern about the effects of production defects on the mechanical performance of the structure - Concern about the sensitivity to damage incurred in service. The above topics are discussed in relation to European research on composites for naval applications by Hayman and Echtermeyer (1999) and Hayman et al. (2001a).

5.3

Developments in materials

5.3.1 Carbonfibre reinforced plastics (CFRP) A major development in recent years has been the drop in price of carbon fibres, which has made them much more attractive for use in larger-scale applications such as naval ships and high speed vessels, in both single-skin and sandwich configurations. CFRP is extensively used in the Swedish Navy's corvette

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5.3.2

Phenolic and other resins for improved high-temperature behaviour

Phenolic resins offer good fire performance, but until recently their relatively poor structural performance, combined with the need to use high temperatures for curing (implying the use of an autoclave) has limited their application in large structures such as ships. However, resins combining lower curing temperatures with moderate mechanical strength have been developed. Flexural and interlaminar properties of laminates with phenolic resins are reported by St. John and Brown (1998). Other resin development has also largely focused on providing good high-temperature and fire behaviour. Trimmer et al. (1999) claim good performance at both high and cryogenic temperatures for a new polycrylene resin.

5.3.3

Thermoplastic resins

Most resins in common use in marine applications are thermoset resins. These resins rely on the use of a hardener and, once cured, they cannot be melted down and recycled like steel or aluminium. Thermoplastic resins can be melted by raising the temperature, and can thus (at least in principle) be recycled. The disadvantage is that they require an elevated temperature in the production process, and if this temperature is approached in service they will soften appreciably. Some thermoplastic resins do however offer excellent impact resistance and are thus of interest for use in landing craft.

5.3.4

Sandwich core materials

Developments in sandwich core materials include: 9 Foam materials reinforced by short glass fibres (Farshad and Fernandez, 2000) 9 Materials using 3-D fibre arrangements 9 Foams with improved mechanical properties, in terms of ductility combined with strength.

5.4

Developments in production methods

The main development in production methods for composites used in high-speed and other marine vehicles has been the move towards closed resin-infusion processes. Vacuum bagging has been used for many years in the production of boat hulls and other small to moderately large structures using epoxy prepregs. With pre-pregs the temperature of the entire component must be raised to ensure curing of the resin. Resin transfer moulding has been used extensively in the aerospace and motor-car industries for producing relatively small components, for which the high pressures needed to ensure adequate distribution of the resin before curing are not a great problem. The main breakthrough has been in the development of vacuum-assisted resin infusion processes. This has enabled rather large items such as complete hulls or decks of medium-sized craft to be produced in a single process, without the need for high pressures and temperatures. The motivation for the introduction of such processes has been a mixture of: 9 a wish to provide a combination of high fibre content with good control over the placing of reinforcement by comparison with traditional hand lay-up methods, and 9 a need to use closed processes with little or no release of harmful gases and vapours to the environment (both local and global). The consequences of increased fibre content are not entirely positive, and highlight the need for a slightly different approach to design of composite structures from that used for homogeneous materials like steel and aluminium alloys. While the reduction in the resin content for a given amount of reinforcement

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reduces the weight and increases the tensile strength measured in terms of stress, the effect on other parameters may be negative. For example, compressive strength in terms of compressive force a laminate can withstand for a given amount of reinforcement may be significantly reduced because the thickness, and hence the bending stiffness, of the laminate are reduced. Indeed, the compressive strength measured in terms of stress may also be reduced. Also the sensitivity to production defects or in-service damage may be increased.

5.5

Performance of structural elements- Failure mechanisms and calculation methods

5.5.1 Failure under static loading For FRP laminates, including sandwich skins, the basic failure mechanisms for in-plane stresses are fibre failure and matrix cracking. Usually in a marine environment it is desirable to ensure that the fibre directions are arranged so that fibre failure, rather than matrix cracking, is the critical failure mode. Also there is a trend towards using failure criteria based on strain, rather than stress quantities (DNV, 2002). For out-of-plane loading, it may be necessary to consider delamination between layers of the laminate. For sandwich cores, yielding and ultimate fracture have to be considered; this may be caused by out-ofplane shear forces or by local loading at connections and points of load application. Failure at the bond between the skin and core of a sandwich structure may also occur; this may be associated with crack propagation in the core, the laminate or the adhesive interface. Fatigue may have to be considered; this can be related to any of the failure mechanisms so far mentioned. A number of local and global buckling mechanisms are possible for both single-skin and sandwich structures; excessive deflections may also have to be considered as a failure mechanism. A special failure mechanism for sandwich structures is local buckling, or wrinkling, of the skin laminate. This has rarely been a problem with GRP sandwich structures for marine applications; for these structures wrinkling is only possible with very flexible cores that are rarely used in highly utilised regions. However, for skin laminates with carbon reinforcements wrinkling can occur with somewhat stiffer core materials and this has aroused new interest in this topic. In particular, wrinkling of anisotropic skins with multiaxial loading has been studied by Fagerberg (2000). The more general question of the local compressive strength of CFRP laminates for ships and other large-scale applications is also a topic currently receiving attention. Of particular concern is the influence of production defects (see also Section 5.8). Design of FRP panels is commonly based on linear-elastic analysis using either analytical formulae or finite element analysis. The main exception is the use of a geometrically non-linear formulation in the DNV Rules (DNV, 1991) that includes the development of membrane effects; however, even this formulation is restricted to an imposed limitation on the deflection equal to about the plate thickness. An evaluation of the validity of this formulation, and of the possibility of extending it to cover larger allowed deflections, was made by Hayman et al. (2001c). They showed that the approach was accurate for larger deflections for the case of simply supported edges, but underestimated the combined bending and membrane stresses at the panel edges for the case of rotationally clamped edges. The extent to which ductility of the core material enhances the load-bearing capacity of sandwich panels in relation to the assumption of linear behaviour up to fracture at the ultimate shear strength of the core was investigated by Hayman et al. (2001b, 2002). They showed that giving credit for ductile behaviour might be justified in some cases, but that this was dependent on the provision of in-plane restraint at the edges and/or the panel aspect ratio being not far from 1.0. Also, the post-yield capacity was only brought into play if rather larger deflections were allowed than the currently specified limit of 1% of the shorter panel dimension.

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It has also been suggested previously that significant weight savings could be made if the current deflection limits for sandwich panels were relaxed and a geometrically non-linear formulation, taking account of membrane effects, were used. Previous studies, such as those by Hildebrand and Visuri (1996) and Riihentaus and Hildebrand (1997), have considered cases in which full in-plane restraint is applied either at the edges of the panel in question or at the edges of adjacent panels. Hayman et al. (2002) considered a series of cases where the only in-plane restraint was provided by adjacent panels in the same plane as the loaded panel. They found that even quite narrow adjoining panels provided sufficient inplane restraint to enable membrane effects to develop such that considerable weight savings could be achieved, provided deflections of 3% to 4% of the smaller panel dimension were allowed. Berggreen and Simonsen (2001) investigated whether the membrane effects developed by curved sandwich panels could be exploited in order to save weight in a representative sandwich hull structure. They concluded that considerable savings were possible, but that these were highly dependent on the inplane boundary conditions provided by the adjoining structure.

5.5.2

Fatigue

Fatigue is generally considered to be less of a problem for composite materials used in high speed vessels than for aluminium and steel craft. This may be due, at least in part, to the ability of the fibres in FRP to hinder crack propagation. However, fatigue may have to be considered in sandwich core materials, especially in the presence of initial defects or damage; see, for example, Berman and Zenkert (1995). It is also important to qualify sandwich core materials together with the adhesives used to join the core blocks, as these may be brittle and act as crack-initiators for the core material.

5.5.3

Vibrations

Vibrations can be important in some applications, though they are often designed out by the application of deflection limits. An example of a vibration analysis of composite deck panels is given by Barton and Ratcliffe (1997). A method of transient dynamic analysis of sandwich structures that takes account of the visco-elastic properties has been investigated by Meunier and Shenoi (2001). Optimisation of sandwich panels with regard to their acoustic performance has been studied by Wennhage (2001, 2002).

5.5.4

Impact and other dynamic loading

Impact resistance remains a topic involving some uncertainty for composite structures in the marine environment. A vast amount of research has been performed into the impact resistance of composites. However, the majority has been oriented towards the aerospace industry, where scenarios such as bird strikes and hail storms are of utmost importance. Some of this research may be applicable to composites used in high speed vessels, but the publications in this field are too extensive to review here. Studies on impact focus on the following issues: 9 Modelling and tests to determine the extent of damage (penetration, delamination area and depth, indentation) caused by an impact with an impactor of a given shape, speed and mass 9 Modelling and tests to determine the residual strength properties (particularly tensile and compressive strength) of structures with impact damage (see Section 5.8) 9 The effects of ageing, moisture and other environmental conditions on the impact properties 9 Detection of impact damage by inspection or monitoring (see Section 5.8).

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137

Connection design

A fundamental challenge in using laminated composites is to provide connections that can transfer forces that are out-of-plane in relation to the laminate layers. Theotokoglou and Moan (1996) performed both tests and non-linear finite element analyses of sandwich T-joints subject to tension (pull-out) loading up to ultimate failure. Their paper provides considerable insight into the various modes of failure that are relevant for such joints. Shenoi and Wang (2001) developed a design approach for single-skin composite T-joints that takes account of through-thickness and in-plane stresses in-plane stresses in the curved overlaminates. They studied the effects of various parameters including the stacking sequence, thickness and radius of the overlaminates. Damage modelling in T-joints has been studied by Shenoi (2001). Blake et al. (2001) performed progressive damage analysis of T-joints with visco-elastic inserts; the purpose of these inserts was to increase noise and vibration attenuation across the joints. Adhesive bonding is being used to an increasing extent for both composite-composite and compositemetal joints in ships and high-speed craft. A study of an adhesive joint between a phenolic laminate and a steel plate was carried out by McGeorge et al. (2002).

5.7

Fire behaviour; fire protection

As mentioned in Section 5.2.3, a major hindrance to the use of composites in high speed vessels in recent years has been the need to satisfy the fire requirements of the IMO HSC Code. Relatively little published research addresses directly the fire performance of materials or structures for high speed vessels; however, there are many published papers on fire performance of composites for use in naval surface ships, e.g. Sorathia et al. (2000) and HCyning and Taby (2000), and these address many of the issues that are relevant for high-speed civilian craft. Grenier et al. (2000) proposed criteria to enable small-scale cone calorimeter tests to be used instead of the full-scale room comer test for qualifying materials for use in high speed craft. The recommendations in this first stage of study were confined to fumiture and other room contents, but extension to other items such as compartment linings was planned. Dembsey and Jacoby (2000) performed cone calorimeter tests on ten material systems including some balsa-cored GRP panels in order to test the validity of three simple ignition models, but found that none of the models was able to resolve satisfactorily the effect of skin thickness and core composition. Ohlemiller (2000) addressed the problem of modelling fire growth in a room comer configuration; while they obtained reasonable agreement between predictions and test results, more extensive validation and calibration against tests are needed before the modelling method can be relied on for more than qualitative trend prediction. Mouritz and Mathys (2000) studied the mechanical properties of fire-damaged glass-reinforced phenolic composites of relevance to marine craft and naval ships. Physical damage occurred only when buming began on exposure to a high heat flux, and even this damage caused only a small reduction in the mechanical properties. Dao et al. (1998) performed both experimental and theoretical studies of the degradation of structural single-skin and sandwich panels by fire and proposed a simple design approach.

5.8

Inspection, repah, and damage tolerance

In composite structures it is important to be able to detect and rectify both production defects and inservice damage. It is also important to be able to predict the consequences of such defects and damage in order to decide on the need for corrective measures. For in-service damage it is necessary to decide whether, when, where and how a repair should be carried out. Sumpter et al. (1997) and Elliott and Trask (2001) have addressed several of these aspects in regard to single-skin GRP hulls of naval ships. For sandwich structures in naval ships, this is the subject of an ongoing European project, as mentioned by Hayman et al. (2001a).

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For single-skin laminates and skin laminates of sandwich structures, production defects include dry zones, voids, delaminations, wrinkles, misalignment of fibres, and poor curing (giving reduced physical properties). For sandwich structures, core/skin debonds must be considered, voids and inclusions in the core, and also lack of bond (edge-to-edge and face-to-face) between blocks of core material. In-service damage may include various types of impact damage (ranging from small object impacts at high or low speeds to berthing impacts and collisions with larger objects or other ships), heat damage and numerous types of damage resulting from overloading, such as core fracture or crushing, skin/core debonds, laminate rupture, delamination either within a laminate or at a secondary lamination, and failure at equipment fastenings. Several techniques exist for non-destructive inspection, such as ultrasound, thermography, microwave, shearography and X-ray methods. Some of these were investigated in a project on naval applications of composites; see Artois-Dubois et al. (1999) and Weitzenb6ck et al. (1998). This study showed that there are difficulties in detecting defects or damage in thick sandwich structures if they are not located close to a surface to which access can be gained. Inspection of sandwich panels with end-grain balsa cores is specially difficult because of the large variations in density of the balsa blocks within one panel. There is also a need to be able to scan large areas of composite structures more quickly than it is possible at present. The use of advanced sensor techniques involving piezo-electric or fibre optic devices is becoming established for monitoring response levels in the hulls of composite ships. An application of fibre optic sensors is described by Jensen et al. (2000). The application of these technologies to detect the occurrence of damage in ship hulls is the subject of ongoing research. Established techniques exist for the repair of composite structures; see for example the discussion by Dubois et al. (2001) and Trask et al. (2002). Recently efficient repair methods making use of resin infusion techniques have been developed; see Daniel et al. (2000). 5.9

Design rules f o r composites

Several guides exist to the use of composites in structures, such as that by ASCE (1984). For composite structures in high speed craft and naval vessels the rules of the classification societies are commonly used such as those by DNV (1991 - under revision). A recent development of major importance is a comprehensive standard for the design of composite structures offshore (DNV, 2002); an overview of the requirements for sandwich composites in this draft standard is given by Noury et al. (2000). This document, which is a result of a major industry-wide effort, presents a systematic approach involving the consideration of limit states related to all the relevant failure modes for a given structure and application. It uses a reliability-based load and resistance factor design (LRFD) format, in which partial safety factors are applied to the load effects and resistance variables. In this way the variability and uncertainties in the applied loading, the properties of the composite materials and the modelling of the structure are taken into account in a systematic way.

6.

UNCERTAINTIES IN DESIGN

With respect to the design of high speed ship structures, as with many other types of structure, there are basically three aspects where uncertainties are present: 9 Loads to which the structures are subjected are of a variable and uncertain nature because of the seaway in which the vessel operates. Moreover there are other loads, such as blast, collision and grounding impact which may be of an ill-defined nature.

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9 The load carrying capacity of the structure is not a fully deterministic parameter since material properties and geometry of the structure are not fully deterministic. These may be particularly affected by the quality of production procedures, causing distortions, misalignments, notches, etc., and of maintenance procedures in the operational phase, which influence phenomena such as corrosion. 9 Analysis methods used to predict structural response and strength are always subject to some uncertainty. This is especially true for analysis of fatigue, buckling and hydrodynamic impact response. No specific papers referring to fast ships where found, while several references treat such problems in general terms (see for instance the Proceedings of PRADS 2001): most of the methods outlined by the various authors can be applied to HSC as well as other ship types.

7.

DEVELOPMENTS IN FABRICATION T E C H N O L O G I E S

7.1

Introduction

Practically all basic concepts related to a production friendly design (such as those concerning production engineering and design to cost considerations) have been presented in Chapter 8 - Design considerations for production - of the report of ISSC 2000 Committee V.2. To avoid repetition, the sections below will provide only an update on selected topics. 7.2

Development in metallic materials and relevant applications

A good comparison between the different properties of steel, aluminium alloys and composite materials was presented in Chapter 6 of ISSC 2000 Committee V.2 report. Recent work has been mainly focused on advanced composite materials, which are presented in detail in Chapter 5 of the current report. The following paragraphs briefly outline the latest progress made in the field of metallic materials relevant to HSC applications. 7.2.1

Steel

Used in the late nineties on peculiar portions of some large monohulls, high-strength low-alloy (HSLA) steel plates are now practically limited to naval ship structures. Progress in fabrication technologies, like laser welding and adhesive bonding, gives new chances for interesting applications of special steel hot rolled profiles, as demonstrated by Braidwood and Lofthouse (2000), as well as of steel sandwich panels, generally fabricated using laser welding, which allow a combination of low weight and high structural strength. A presentation of the latter solution for the shipbuilding industry can be found in Roland et al. (2000). 7.2.2

Aluminium alloys

For basic concepts and information about aluminium alloys refer to Section 6.2.2 of the report of ISSC 2000 Committe V.2. Recent research has been mainly focused on static and fatigue properties of the aluminium alloys developed for marine applications in the late nineties, like AA5383 and AA5059. For the former some results can be found in Meynet et al. (2000), who focused on how welding procedure and weld configuration affect the static strength of weldments. For the latter refer to EhrstrOm et al. (2000) and Sampath et al. (2000). Other typical items are as follows: 9 fatigue life assessment of typical details (Benson et al, 2000, Polezhaeva and Malinowski, 2001);

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study of new computational techniques for the evaluation of the fatigue life of aluminium alloy welded joints (Atzori et al., 2001); 9 aluminium profile/honeycomb composite solutions (Dean et al., 2000); 9 use of both open and hollow extruded profiles (G/3nner, 2000); 9 typical damages in operating vessels and relevant repairs (Ferraris and Simpson, 2000, and Wilhelmsen, 2000). Other studies are related to new joining techniques like friction stir welding and adhesive bonding, which are treated respectively in Sections 7.3.3 and 7.3.4.

7.3

Development of new cutting and joining techniques

7.3.1

Influence of cutting and joining techniques on design

Cutting and joining techniques can influence both the geometry and the mechanical properties of structural connections. In the last few years, considerable efforts have been made to industrialise manufacturing techniques that improve the properties in the heat affected zone, reduce distortions and residual stresses, and increase static and/or fatigue performance. 7.3.2

Laser beam cutting and welding

The use of a laser beam is a potentially attractive method for plate cutting and for welding of semimanufactured products that are to be integrated in much larger structural components. However, the application in shipbuilding is limited by the required edge tolerances, the high investment cost, the limited experience of the long term behaviour of laser welded structures and the lack of acceptance rules. In some cases a combination of laser with arc welding techniques (Laser-Hybrid-Welding) may overcome the obstacles, leading to a wide range of possible applications. Research has focused on the development of laser welding techniques and their application to both steel and aluminium alloys (Raspa, 2002), as well as on the use of laser welding to produce lightweight steel sandwich panels. A very interesting presentation of this topic is given by Roland and Reinert (2000), who illustrate not only the fabrication process of such panels but also their operational behaviour and integration with outfitting. 7.3.3

Friction stir welding

Friction stir welding (FSW) is a well established, environmentally friendly and cost-effective technique for the automated fabrication of lightweight panels for HSC applications. As the method is rather recent, a great deal of R&D is still in progress, mainly concerning FSW of aluminium. Many papers are concerned with the properties of welds, essentially fracture toughness (Dawes et al., 2000, and Oosterkamp et al., 2000), effects of imperfections and residual stresses on fatigue crack propagation (Dalle Donne et al., 2000), fatigue behaviour (Ericsson et al,, 2000, and Zhang et al., 2000, who compared FSW with traditional welding). Some papers present the opportunities given by FSW when associated with aluminium extrusions, e.g. Anne de Vries and Backlund (2000), Kallee et al. (2000), Midling et al. (2000). Regulation aspects of FSW and its application are discussed by Przydatek (2000). 7.3.4

Adhesive bonding

The main advantages of adhesive bonding with respect to traditional welding connections are that

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mechanical properties of parent material do not decrease due to the joining technique and that lack of heating implies no distortions. High quality surface finish can be easily achieved, and rework is very limited. The factors which still prevent the wide use of adhesive bonding on vessels are the lack of information about long-term behaviour in the marine environment and strength retention in case of fire, the substantial lack of reference rules, and the need for precise application, inspection and repair procedures. A review of the topic can be found in Weitzenb0ck et al. (2000) and Gambaro et al. (2002).

8.

CONCLUSIONS AND RECOMMENDATIONS

After a decade of major developments, leading to the design and construction of large size monohulls and catamarans, the market for passenger and car ferries is now experiencing a lull. Further progress can be expected in cargo transportation, where structural design would have to cope with significant problems in terms of cargo handling more than in terms of passenger safety. There is nonetheless scope for further R&D work, to reduce uncertainties which still have a strong influence on structural design, and to address certain design aspects, like fatigue assessment of structures subjected to a number of cycles exceeding the traditional limits (107-108). Efficient onboard monitoring systems may help crews to better understand the loading conditions acting on vessels and to operate them accordingly. Progress can be expected in the naval field, where high speed low and medium size vessels are already designed and built, mainly for fast attack and coast patrolling purposes, and there is a growing interest in high speed sea-lift. The need for high performance makes it necessary to extend the use of lightweight solutions and lightweight materials, with new challenges in terms of fire resistance (recent tests have proved that SOLAS requirements for A30 and A60 partitions can be easily achieved by using suitably insulated three and even two millimetres thick steel structures), and resistance to ballistic impacts (mainly due to terrorist attacks), underwater shock and air blast loading.

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Przydatek, J. (2000). Friction stir welding with Lloyd's Register. Proc. Lightweight Construction Latest Development, Royal Institution of Naval Architects, London, U.K. Raspa, A. (2002). Saldatura dell'alluminio con laser Nd:YAG (in Italian), Rivista Italiana della Saldatura, LIV:I, 31-34. Rathje, H. and Schellin, T. (1997). Viscous effects in seakeeping prediction of twin-hull ships. Ship Technology Research, 44, 44-52. Rees, S., Reed, D., Cain, C. and Cripps, B. (2001). Coupled Fluid-structural Modelling to Predict Wave Impact Loads on High-speed Planing Craft. Proc. of FAST 2001, Southampton, U.K. Riihentaus J. and Hildebrand M. (1997). Utilising Non-Linearities in Stiffened Sandwich Panels. Technical Report VALB 180, VTT Manufacturing Technology, Espoo, Finland. RINA Rules (2002). Rules for the Construction and Classification of High Speed Craft. RMS Rules (2000). Rules for the Classification and Construction of High-Speed Craft. Roland, F. and Reinert, T. (2000). Laser welded sandwich panels for the shipbuilding industry. Proc. Lightweight Construction Latest Development, Royal Institution of Naval Architects, London, U.K. Ros6n, A. (2001). Inpact pressure analysis on high-speed craft in waves, through FE-analysis on full-scale and model measurement data. Proc. of PRADS 2001, Shanghai, China. Rothe F., Sames P. C. and Schellin T. E. (2001). Catamaran wet deck structural response to wave impact. Proc. of FAST 2001, Southampton, U.K. Sampath, D., Vloemans, R., Mechsner, K., Ghaziari, H. and Haszler, A. (2000). Alustar alloy AA5059: a better alternative to AA5083 alloy in the marine industry. Proc. 4 th International Forum on Aluminium Ships, New Orleans, USA. St. John, N.A. and Brown, J.R. (1998). Flexural and Interlaminar Shear Properties of Glass-reinforced Phenolic Composites. Composites- Part A: Applied Science and Manufacturing, 29. Sano, A., Ikebuchi, T. and Okada, H. (2001). Water Body Collision Using Fluid-Structure Dynamics Interaction - Development of Simulation Program and Evaluation for Simplified Models -, Journal of the Kansai Society of Naval Architects, Japan, 236, 221-227. Sano, A., Ikebuchi, T. and Okada, H. (2002). A Fundamental Study on Water Body Collision Considering Fluid-Structure Dynamics Interaction ( 2nd Report ) - Evaluation for Stiffened Plates -, Journal of the Kansai Society of Naval Architects, Japan, 237, 209-215. Schellin, T.E, Beiersdorf, C., Chen, X. and Maron, A. (2002). Comparative frequency domain analysis of a fast monohull in regular head waves. Proc. of 21 st International Conference on Offshore Mechanics and Artic Engineering, Oslo, Norway. Sebastiani, L., Valdenazzi, F., Grossi, L. and Kapsenberg, G.K. (2001). A theoretical / experimental investigation of the slamming pressures on fast monohull vessels. Proc. of FAST 2001, Southampton, U.K. Sheikh, Grondin and Elwi (2002). Stiffened Steel Plates Under Uniaxial Compression. J. Constructional Steel Research, 58. Shenoi, R.A. (2001). Damage Modelling in Tee Connections in FRP Ships. Proc. 22nd SAMPE Europe Int. Conf., Paris, France. Shenoi, R.A. and Wang, W. (2001). Design of Tee Connections in FRP Ships Using an Analytical Approach. Proc. of PRADS 2001, Shanghai, China. Simonsen, B.C. (1999). Bottom raking damage to high-speed craft. Trans. RINA. Sorathia, U., Gracik, T., Ness, J., Blum, M., Le, A., Scholl, B. and Long, G. (2000). Fire Safety of Marine Composites. Proc. Int. SAMPE Symposium and Exhibition, 45 (II). Sumpter, J.D.G., Swift, A.C., Elliott, D.M., Faulke, D., Court, R. and Lay, P.W. (1997). Damage Tolerance of GRP Ships. Proc. Int. Conf. Advances in Marine Structures III, Dunfermline, Scotland. Takemoto, H. (2000). On a Method to Calculate Hull Response of High Speed Vessels in Waves Taking Account of Slamming. Journal of the Society of Naval Architects of Japan, 188, 143-154.

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15th INTERNATIONAL SHIP AND OFFSHORE STRUCTURES CONGRESS 2003 AUGUST 11-15, 2003 SAN DIEGO, USA VOLUME 2 ""

O/EGO, u - "

C O M M I T T E E V.5

FLOATING PRODUCTION SYSTEMS

C O M M I T T E E MANDATE Concern for the design of floating production systems. Attention shall be given to the coupling effects between vessel and seabed connections, and to the specific structural behaviour of these systems. Consideration shall be given to identification and quantification of uncertainties for use in reliability methods.

COMMITTEE MEMBERS

Chairman:

Dr D.T. Brown Dr Y. B ai Prof H. Boonstra Dr T.Y. Chung Prof R. Li Dr A. Loeken Prof S. Mavrakos Dr H. Nedergaard Dr T.A. Netto Dr H. Suzuki

KEYWORDS

FPSOs, FPS, floating production, monohull, semi-submersible, spar, tension leg platform, hull, riser, pipe in pipe, steel tube umbilical, steel catenary riser, mooring, anchor, tether, offioading, LNG, vortex induced vibration.

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CONTENTS

1 INTRODUCTION . . . . . . . 1.1 O v e r v i e w o f P r o d u c t i o n C o n c e p t s 1.2 C o n c e p t Selection

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152 6 CONCLUSIONS

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Floating Production Systems 1.

153

INTRODUCTION

This report first provides a brief overview of the Floating Production System (FPS) concept for readers new to the subject, together with a recap for those more familiar with such systems. It then describes the FPS global system lifecycle from project conception through to decommissioning. The present day technical uncertainties associated with FPS schemes are documented, and results of recent investigations associated with the vessel hull and seabed connections presented. The conclusions and recommendations aim to point the way forward for future work.

1.1 Overviewof Production Concepts Over the 1997 to 2001 period there were 85 FPS prospects worldwide with an additional 147 planned for 2002 to 2006 as indicated by Douglas-Westwood (2002). Of the future schemes 100 are monohulls (FPSOs), 18 are semi-submersibles, 15 are tension leg platforms (TLPs), and 14 are spars. A Floating Production System consists of the primary components as listed in Table 1. TABLE 1 MAIN CHOICESFOR UPS COMPONENTS

Component

Main Options

Subsea/Well Completion

A B C A B C D A B C A B C A B C D A B

Riser System

Anchor System

Mooring System

Vessel

Export and Storage

Subsea,on or above seabed Surface,on process support vessel On separate well head platform Straightsteel pipe with motion compensation Flexiblecomposite pipes, geometrically compliant Metalliccatenary shaped pipes (SCRs) Thermo-plasticor steel tube umbilicals Piled anchors Dragembedment anchors Suctionanchors Chainand/or wire catenary Steelcatenary of chain/wire with possibly fibre line Tautfibre line system with steel terminations Monohull(FPSO) Semi-submersible Tensionleg platform (TLP) Sparor deep draught floater Export riser to subsea pipeline to indept storage or CALM buoy Offloadinghose directly to shutter tanker

Monohulls are ship-shaped vessels with lengths and draughts of typically 5 to 6 and 0.3 to 0.4 times the breadth. Vertical motions are of similar magnitude to the wave height, and flexible or other vertically compliant risers are required. Monohulls allow storage and usually offloading in which case they are termed FPSOs. Weather-vaning vessels are advantageous, and numerous turret-moored systems exist. Steel hulls are preferred, but concrete may be competitive because it has small weight sensitivity. Semi-submersible hulls consist of two parallel submerged pontoons, vertical or near vertical columns to support the deck structure that contains production equipment and accommodation. Vessel vertical motions are generally small, allowing rigid risers, though flexible riser connections to the seabed are

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required in rough weather areas for larger vessels to ensure high production uptime. preferred hull material though concrete may be competitive for large platforms.

Steel is the

TLPs have similar geometry to semi-submersibles, with additional transverse submerged pontoons making up a square or rectangular based platform. Vertical tethers are attached to the comers of the pontoons to provide station-keeping. The TLP is weight sensitive, as tethers must be kept in pre-tension, restraining the vertical motions and allowing heave compensated rigid risers and deck based dry wellheads. The main hull of a spar is cylindrical with a central moon pool and tensioned risers, the circular hard tank providing buoyancy. Spars are deep draught floating vessels, and thus have low vertical motions. Additionally compliant towers are possible production alternatives. These are somewhat similar to fixed jacket structures, but achieve compliancy by careful design of the stiffness, particularly tether stiffness, and mass properties. Compliant towers, together with the alternative vessel types, and the various riser, anchor and mooring systems are discussed in more detail in CMPT (1998).

1.2

Concept Selection

Concept selection, described in CMPT (1998), depends on the field requirements, principally: 9 9 9 9 9 9

environment and water depth, oil and gas production volumes, distance to shore or supporting infrastructure, subsea tieback possibilities, required number of drilling centres and wells at each centre, well fluid chemistry and pressure, and intervention frequency for optimum well performance, personnel risk.

The floating structure typically has the following performance requirements: 9 9 9 9 9

appropriate work area, deck load capacity and possibly storage capacity, acceptable stability and motion response to environmental loads, strength to resist extreme conditions, and durability to resist fatigue loading, possible combined capabilities such as drilling and production, may be transportable.

The production concepts identified in Section 1.1 must be matched to the field and performance requirements described above to establish viable field development schemes. A simplified selection procedure has been documented by Inglis (1996) for deep water in Table 2 based on combinations of distance to shore, number of drilling centres and well intervention frequency. The field is assumed to have associated gas that may be exported or re-injected. For short distances to shore, pipeline export is assumed unless an FPSO is used since this already has an integral export system. For long distances, offshore loading is utilised. Where the well entry frequency is low, it is assumed that subsea wells will be the most cost effective, and only concepts with subsea wells are proposed for both single and multiple drilling centres.

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For single drilling centres and high well entry frequency, surface trees are preferred. These require a TLP, compliant tower, spar, or semi-submersible with work-over capability. The combination of multiple drilling centres, with wells that need frequent intervention, is accommodated in shallow water by using a well-head jacket. In deep water this would require multiple TLPs or compliant towers, but this is not economically viable. For deep water fields that combine multiple drilling centres and frequent well intervention a small unmanned TLP, or mini-TLP, which only supports the wells and provides a capability for drilling and work-over is possible. The mini-TLPs processing facilities are provided by another central facility such as a semi-submersible or FPSO moored close by. Fixed structures are primary candidates in water depths less than approximately 150m. However even here an FPSO, in conjunction with subsea wellheads, may become the optimum solution particularly when there are multiple drilling centres, no pipelines, short field life, and infrequent work-over requirements TABLE 2 DEEP WATER FIELD DEVELOPMENTCONCEPT SELECTION GUIDE (INGLIS, 1996) Distance to Shore/ Infrastructure

Number of Drilling Centres

Well Entry Frequency

Development Concept

Semi + Subsea + Pipeline SWP + Subsea + Pipeline FPSO + Subsea One TLP + Pipeline Compliant Tower + Pipeline Spar + Pipeline Semi + Mini-TLP + Pipeline High Semi + Subsea (adjacent)+ Pipeline SWP + Mini-TLP + Pipeline FPSO + Mini-TLP Semi + Subsea + Pipeline SWP + Subsea + Pipeline Low FPSO + Subsea Multiple Semi + Mini-TLP + Pipeline SWP + Mini-TLP + Pipeline High FPSO + Mini-TLP FPSO + Subsea Spar + Subsea + OLS Low Semi + Subsea + FSU or DTL One TLP + FSU or DTL Spar + OLS Compliant Tower + FSU or DTL High FPSO + Mini-TLP Semi + Subsea(adjacent) + FSU or DTL FPSO + Subsea Semi + Subsea + FSU or DTL Low Spar + Subsea + OLS Multiple FPSO + Mini-TLP High Spar + Mini-TLP + OLS OLS=Offshore Loading System, DTL=Direct Tanker Loading FSU=Floating Storage Unit, SWP=Shallow Water Platform Low

Short

Long

Key:

Water depth is also a key driver in concept selection and strongly influences the viable economic and technical solutions. In preparing this report the Committee considered that the term 'shallow' water

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should reflect depths up to 100m (consistent with typical classification rules for first order wave motion analysis). For larger depths the proposed terminology is; 'intermediate' representing depths less than 500m, 'deep' less than 1500m, and 'ultra-deep' greater than 1500m.

2.

G L O B A L SYSTEM LIFE CYCLE

In this section an overview is given of the FPS life cycle from initial project strategy and environmental considerations through to abandonment. This is preceded by a short section on the technical changes to the FPS prescriptive regulations, and guidelines that have been recently implemented. These are instructive as they indicate where previous research investigations have been expended to produce methodology changes that are now considered to be of suitably mature status and thus acceptable for classification purposes.

2.1

Prescriptive Regulations

Competent authorities of the country or state having jurisdiction over the continental shelf in question regulate classification of offshore field developments. Classification is a standardized system for independent verification of quality and safety, which leads to a certificate that declares that the classed vessel complies with the Class Rules covered by the identified class notation. In addition to Class Rules there are National Authorities Standards (eg HSE, NMD), Statutory Regulations (eg MARPOL, SOLAS), and Flag State Requirements. Two classification societies, namely Det Norske Veritas (DNV) and American Bureau of Shipping (ABS) are assessed here. Other classification societies involved in offshore platforms issue similar rules. Recent changes to DNV rules presented in DNV-OS-C101 to 106 (2000 and 2001) associated with classification and structural design of FPS units are as follows: 9 Use of Load and Resistance Factor Design (LRFD) method for offshore structures, where the target safety level of a structure is obtained by applying the loads and resistance factors to the characteristic loads and resistances. Issues such as greenwater and accidental loads as well as combinations of environmental loads are also dealt with. 9 Use of Working Stress Design (WSD) methods that are based on a permissible usage factor. This factor is defined based on the loading condition, failure mode and importance of the strength member. As with the LRFD method, guidelines for greenwater, bow slamming, loss of heading and collision, accidental and other environmental loads are provided. 9 Use of the Partial Safety Factor concept in position mooring design. These are separate safety factors applied to the dynamic and mean line tensions. A 'method factor' is introduced to allow a refinement to the safety factors if more sophisticated numerical models are used in analysis. 9 Introduction of new corrosion allowances for mooring chains. 9 Reduction of scantling sizes allowed for benign environments. 9 Regulations for inspection and corrosion management of permanently moored units. 9 Regulations for tank loads, sea pressures on pontoons and columns for semi-submersibles. 9 Attention is drawn to TLP tendon supporting structures. 9 Consequences of unintended flooding of spar hard tanks.

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ABS classification requirements on floating production systems are provided in the Guide for Building and Classing Floating Production, Storage and Offloading Systems. A revised version was published in June 2000 to replace the 1996 edition. In this revision, guidelines for the following issues are covered: 9 The concept of Environmental Severity Factor (ESF), also known as Fatigue Damage Adjustment Factor, is introduced to account for fatigue damage caused by dynamic environmental loads in different conditions. 9 Reduced scantlings, down to 85%, are allowed for benign service. 9 Equations for greenwater and various slamming loads are provided. Guidelines on planning, design and construction of FPS are covered by the American Petroleum Institute (API) in API RP 2FPS (2001). Key issues in this recommended practice include: 9 Design of FPSOs, spars and semi-submersibles using the Working Stress Design Method (WSD), see API RP 2A-WSD (2000). 9 Class rules for FPSO design and construction. 9 Use of risk assessment approaches in FPS design and construction. Recently operators are actively considering the use of FPSOs in the deep water Gulf of Mexico, and will need regulatory approval from both the Minerals Management Service (MMS) and United States Coast Guard (USCG). To aid the MMS in confirming acceptability of FPSOs a number of studies, as described by Parker and Grove (2001) have been implemented as follows: 9 Environmental impact statement (funded by the DeepStar JIP). 9 Comparative risk assessment covering fatality and oil spill risks for a candidate FPSO, spar, TLP and linked shallow water jacket hub, see Gilbert et al (2001). 9 Assessment of any gaps in existing regulations and development of a regulatory model that the MMS and USCG can use for FPSO project approval. A committee draft report, ISO standard 19904, is presently being reviewed, on materials, equipment and offshore structures for petroleum and natural gas industries - floating structures including stationkeeping. This covers all platform types as well as lifecycle issues. 2.2

ProjectStrategy

At the earliest stages of a possible FPS field development it is essential to define a project strategy. The important components for such a strategy, as given in CMPT (1998), are to: 9 Identify project specific commercial constraints, functional requirements and perform environmental impact assessments (see Section 2.3). 9 Establish applicable regulatory authority requirements (see Section 2.1). 9 Define field requirements and viable development concepts (see Sections 1.1 and 1.2). 9 Develop the various design, engineering, construction, installation, operation and decommissioning phases (see Sections 2.4 to 2.7). 9 Specify realistic project budgets and schedules.

158

2.3

Specialist Committee V.5 Functional Requirements and Environmental Impact Assessment

The key project functional requirements are to establish the environmental criteria, limiting operational conditions, hydrocarbon import, production and export requirements together with manning and replenishment provision, and finally removal requirements. In addition an environmental assessment that considers how the project influences people, nature, physical and ecological resources, is increasingly becoming mandatory. This is established at the earliest opportunity in the planning process, to provide environmental studies, in an open process involving consultation of institutions and the public. Assessments provide for improved environmental decisions, help mitigate environmental incidents, improve project feasibility, and aid future planning decisions at the early design stage.

2.4

Design/Engineering and Construction

Once the field requirements and viable floating development concept is selected the engineering, divided into distinct phases, can commence. Initially front-end engineering and design (FEED) is carried out, involving small multi-disciplinary teams, to establish the main building blocks. At the end of the FEED stage detailed engineering can commence, if the project remains technically and commercially viable, to refine the building blocks into detailed design and working drawings. At this stage the construction yard is selected and work such as fabrication drawings are transferred there. Construction efficiency can be improved by simplifying the global and detailed design, for example a six-column TLP hull structure can be replaced by a four-column structure. TLPs and semi-submersibles with cylindrical columns may be further simplified to box-shaped columns of flat panels. Multi-column braced semi-submersibles can be replaced by four-column non-braced structures. Construction is typically an assembly of several blocks that are fabricated simultaneously. For example, an FPS hull may be divided into several main structure types such as bow, mid-body, stem, accommodation block, turret and process equipment pallet/deck structure. It is also important to recognize the advantages and disadvantages of traditional shipyards and offshore construction yards. A shipyard may efficiently construct plated structures such as hulls, deck structures and accommodation blocks. On the other hand, an offshore construction yard would be well suited to fabricate topsides and specialist structures such as process equipment modules, TLP tether foundations and subsea templates. Hence, both types of yards are usually involved in the construction process. As an example, the hull and deck structures may be built in shipyards in Korea, China or Japan, while topside structures are built and added in Singapore, Dubai or yards near the sites where floating systems will be installed and operated. The construction and installation of mooring systems, pipelines and risers should also be integrated and coordinated with transportation and installation of the FPS. A modem field development project consists of many major contracts and contractors dealing with the hull, pipeline and subsea components, topsides, moorings, risers, umbilicals and transportation/installation phases. A good floating system and facility design will account for integration of all disciplines and contractors.

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Installation

The FPS installation sequence depends on the selected development concept, for example the drilling and well completion can take place from the FPS itself if the well make-up is suitable, or in other cases a dedicated drilling rig is required. The former requires a stable platform such as a semi-submersible, TLP or spar, with considerable load carrying capacity. The installation method used for the stationkeeping system is generally closely connected to the selected hull design. For example TLP tethers can be designed so that they are contained within the hull in sections, and then assembled during installation and lowered, before being stabbed into the anchor blocks. In other projects the tethers are assembled onshore and towed to the field. Some FPS types must be designed to disconnect from the mooring and riser system as a result of hurricanes, storms or icebergs. In such cases this feature can be utilised during installation, where one contractor installs the disconnectable subsea components and another installs the vessel. If the FPS hull and topside is not fabricated and mated at the same yard, as with a spar unit for example, the hull is towed to the field, upended and connected to the mooring system. The topsides are then towed out by barge and lifted in place by a crane vessel.

2.6

Operation and Maintenance

During FPS design and planning it is essential to develop a vessel that has the appropriate adaptability to changing field characteristics, especially for remote locations. For some fields drilling and well intervention are integrated tasks for the FPS, and thus will be a significant part of the vessel operations. The adaptation of the processing units to cater for changes in well fluid composition is important, specific considerations being the associated flow assurance and reservoir development requirements, such as change in water cut, required water injection and possible souring of the well fluids. In conjunction with process equipment maintenance and upgrades, weight sensitive units such as semisubmersibles and TLPs must consider the operational issue of weight management. Older process units may be exchanged with new lighter vessels, thus enabling room for new process optimising equipment. Maintenance is viable for topside equipment and can be planned for in the design, however the subsea elements such as mooring and riser systems are somewhat inaccessible. Operational and maintenance difficulties are exacerbated when such components are exposed to wear and tear. Particularly vulnerable items are fibre rope moorings, flexible risers, and umbilicals together with off-loading hoses to shuttle tankers. FPS schemes that can disconnect from the subsea components can be maintained during the dry-docking phase. However, this has to be considered in conjunction with the requirement that a disconnectable unit requires a complete marine crew full time onboard for manoeuvring the vessel.

2.7

Decommissioning

Offshore installations have to be removed to land, generally within approximately three to five years after operations have finished. The competent national authority at the relevant location publishes rules and guidelines on decommissioning, see for example DTI (1998). For FPS platforms the decommissioning and transfer to land is obviously much easier to accomplish than for fixed structures. An exception to this is the Brent Spar, as described by NERC (1998), the vessel's decommissioning procedures not being fully considered at the design stage.

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3.

PR I M A R Y AREAS OF UNCERTAINTY

3.1

General

Uncertainty can be described as the unknown variation of the basic parameters that govern structure behaviour and durability. For fixed offshore oil installations in the North Sea, the uncertainty is generally considered by industry to be at a comfortable level. This was not the case in the recent past, leading to accidents and unexpected incidents. The industry has since matured further, and due to implementation of the results from experience gained it can be concluded that the uncertainty level is low, that is the variation in the basic parameters is reasonably well known. The FPS industry is not at the same level of maturity, although the first FPS was installed more than 30 years ago. This is because the FPS concept is used in more remote areas at increasing water depths, to produce well fluids of increasingly high pressures, temperatures and having exotic constituents. Furthermore FPS schemes are continuously being developed for diverse but related usage such as LNG offshore processing and offloading plants. The challenges arising from the higher level of uncertainty call for careful FPS design and operation, as often it is not possible to extrapolate from earlier designs. This section highlights the present day areas of high uncertainty related to FPS design and operation.

3.2

Water Depth Related

FPS concepts are utilised in shallow waters as low as 30m and ultra-deep waters exceeding 1500m. In very shallow waters uncertainties exist associated with the wave loading and hull hydrodynamic and structural response, together with its modelling. For example the profile of shallow water steep breaking waves is not well understood, depending on local bathymetry that is not easy to model in tests. There are also uncertainties associated with the green water loading, and non-linear response of shallow water floating structures such as CALM buoys in steep breaking waves, and numerical modelling is not possible at present. The mooring system influence on the vessel response additionally needs careful assessment, as lines cannot maintain catenary shapes in high seas because of insufficient water depth. In water depths less than 300m wave action on the vessel hull is reasonably well understood, though uncertainties again exist in the green water loading and hull structural response. There is limited influence on the global hull motions caused by the hydrodynamic loads on risers and moorings. In deeper waters the wave action will vanish for the lower parts of the risers and moorings and these will act as a significant dampers to the total system. This requires careful modelling and selection of hydrodynamic force coefficients in the commonly used Morison equation. In the upper regions the drag coefficient may be as small as 0.5, whereas in the lower regions, where the riser acts as a damper, the flow conditions may cause drag coefficients of the order of 1.2. However, current and marine growth may result in additional uncertainty here. Furthermore possible vortex shedding induced riser or mooring cable vibrations may cause this already quite uncertain drag force to double. Thus the drag coefficient for deep-water risers and cables with circular cross-section may be in the range of 0.5 to 2.4. A specific uncertainty associated with Gulf of Mexico fields is associated with loop current effects. FPS responses in waves can be studied in model test tanks, however for deep-water locations modelling of the riser and mooring system causes problems. If the total mooring system has to be modelled, the scale becomes very small, or the system has to be truncated. The latter solution is the more common, however to appropriately model a mooring spread by a truncated system introduces uncertainties.

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Combining this with the limited validity of model tests, principally scaling of the different loading mechanisms, caused by wave and viscous forces, the latter leading to an overestimate of damping and drag, results in significant model testing uncertainties in increased water depth. With respect to numerical analysis and simulations, the total system including mooring and risers can be modelled simultaneously to some degree, though it is a complicated and time consuming process. Traditionally in design the separate components are analysed and the mooring forces introduced at the vessel as prescribed characteristics. This split model may be sufficient in shallow water, but for deep water a combined model is required, mainly because the loading, and response, induced by the moorings and risers becomes a relative larger part of the total loading, and behaves in a highly non-linear manner. In deeper water there are also difficulties with flow assurance. Specifically, in risers exposed to seawater cooling and pressure drop due to friction and gravity, wax and hydrate formations may cause flow problems. Riser insulation such as pipe-in-pipe systems are used, and/or heating and injection of inhibitors to help remedy this, but design, installation and operation of such systems and methods in deep water is relatively new to the industry and so increased uncertainty is incurred. New materials and concepts are being introduced to mitigate the effects of high water depth. Because of the steel weight penalty in flexible risers, combined steel and materials such as carbon fibres are used. Steel catenary risers (SCRs) are the most cost-effective deep water solution, but riser structural response must be minimised because fatigue issues are of concern. Deep water umbilicals also utilise steel tubing as thermo-plastics cannot cope with the hydrostatic loading. The steel tubes remove the requirement to have armour wiring and have improved resistance to hydraulic and chemical injection fluids. However there are uncertainties in the tubing fatigue response, and additionally the super-duplex steels used can be prone to heat treatment difficulties. Heavy chain and wire mooring systems are being substituted with lighter materials such as fibre ropes. These require suction anchors that can take vertical uplift. These new concepts and materials inevitably imply increased uncertainty, until sufficient field experience gives improved confidence.

3.3

Environmental Loading Related

Offshore structures are exposed to environmental load effects from wind, waves, current and possibly ice. Determination of the extreme and fatigue related structural response introduces uncertainties, especially with respect to the joint occurrence of the different environmental loading components and their directionality. For turret-moored FPSOs the vessel heading is an additional uncertainty that must also be considered in conjunction with the weather direction, because the extreme load event may not occur during an extreme storm. Current will act as a forcing and/or damping component on parts of an FPS structure including the moorings and risers. Current originates from several physical processes, such as tidal effects, storm surge, surface wind drift, internal waves, inertial and front dependant influences together with circulation. It is very difficult to hindcast and simulate the long term current behaviour without local knowledge and extensive previous measurement records. Furthermore the appropriate repeat periods are of the order of 12 hours for the tidal current component, 30 minutes for the internal wave component and can be weeks for the circulation component. Thus significant uncertainty may be present with respect to the current intensity at remote or poorly unexplored sites, particularly for fast track projects.

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Additionally the spatial and short-term current variation can influence the FPS total hydrodynamic loading. In moderate, but uniform currents, the likelihood of riser lock-in to vortex shedding induced vibrations for example may be far higher than in a more intense current environment. Lock-in induced resonant vibrations lead to increased local fatigue damage, and can result in a doubling of the hydrodynamic loading on the riser element. The local FPS structure and water interaction are difficult to quantify at the design stage, examples being green water on deck and related slamming loads as outlined above, together with wave run up on columns, ringing phenomenon between TLP columns and vortex shedding lock in on a spar hull. More work is required in these areas in order to decrease the uncertainties at the design stage. 3.4

Structure Related

FPS structure related uncertainties include strength, fatigue and corrosion issues. The structural strength uncertainties include that associated with the equations and finite element models that are used, uncertainties associated with the geometric and material properties and errors due to incorrect analysis. Areas that need future research include the development of strength equations for combined loads such as buckling and collapse of plates and shells, the calibration of partial safety factors using risk assessment and structural reliability analysis, the standardization and benchmarking of finite element models, and the development of procedures for the determination of partial safety factors for finite element analysis and strength design based on testing. Fatigue related uncertainties are caused by several factors, such as the selection of appropriate environmental conditions and their combinations, the extrapolation of fatigue stresses at hot spot points, the selection and interpretation of design codes for fatigue assessment, and the calculation of stress concentration factors. Additional uncertainties are associated with fatigue caused jointly by wave loads and vortex induced vibration (VIV), and the selection of safety factors together with inspection/repair methods. A particular fatigue related concern is with FPSO hulls that are converted from existing tankers. Specifically a tanker will be exposed to very different and often unknown loading through its lifetime, whereas an FPSO must remain at location throughout its field life, thus undergoing materially different in-service load histories. Corrosion related defects can significantly reduce ultimate and fatigue strength. A number of mathematical models have been developed to predict corrosion development in structures such as pipelines, risers and pressure vessels, however their field use is uncertain. Various methods have been applied by the industry to measure the magnitude, location and shape of corrosion related defects, as all of these are important for FPSO strength and fatigue assessment. However such methods are costly and it is necessary to develop statistical techniques further, allowing evaluation of the strength and fatigue reliability based on the limited data from corrosion measurement schemes.

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System Related

FPS design and construction involves a significant number of interfaces. A development may consist of separate design and construction contracts for the mooring, risers, hull, turret and topsides. Such a contracting strategy may be beneficial for many reasons, however from a technical point of view additional uncertainties and less optimal solutions will inevitably result. 5000 z

4000

,,_.,

o

3000

. _

o

o

2000 1ooo

0.0

0.5

1.0

1.5

2.0

Current [m/s]

Figure 1" Predictions of maximum hawser tension for a moored tanker (Schellin, 2003).

Understanding and predicting the resonant and chaotic response of a moored vessel in the horizontal plane, such as the fish-tailing behaviour of an FPSO and connected shuttle tanker, are additional uncertainties. Schellin (2002) presents results in Figure 1 of the maximum hawser tension for a tanker subjected to a range of current velocities. Maximum mooring loads vary by a factor of 5, because the results are sensitive to the selected coefficients used in the complex manoeuvring simulation models. As an example, modifying the yaw-damping coefficient by 10% alters the maximum hawser tension by a factor of 4. This work highlights the necessity for detailed modelling of the viscous forces acting on moored vessels, together with the requirement for detailed model tests and full scale feedback. Quantification of the uncertainties is very difficult, but a ranking of the different sources and their consequences is useful. Such ranking allows appropriate effort to be focused on the more critical issues. Such a ranking for FPSOs is found in Noble Denton (2000), see Table 3, covering failure modes that are initiated by environmental overload or fatigue. Note that fire and explosion that may result in a more likely and more onerous hazard are not included.

TABLE 3 FPSO INCIDENT- OCCURRENCE FREQUENCYAND CONSEQUENCE(NOBLE DENTON, 2000) Failure of Limit State

Probability of Occurrence / annum

5 X 1 0 -6 _ 5 X 1 0 -4 Hull Midship Section Strength 1 0 -4 _ l i f e Bow Structure / Slamming . l i f o _ 1 0 -4 Cargo Tank / Sloshing 1 0 .6 _ 1 0 -4 Turret 2X10-4 _ 2X10-2 Station Keeping System 7X10-4_ 2X10-2 Fluid Transfer System 10-4 _ 10-2 Deck & Topside / Green Water

Loss of Life Consequence High Medium Low Medium Low Medium Medium

Loss of Containment Consequence High Medium Medium Low Low Medium Low

Loss of Production Consequence High Medium Medium High High High Medium

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10.5 - 5x 10-6

Intact Stability Structural Fatigue Failure

5x 10-v- 10-4

Damage Stability Ship Impact

5x 10.6 - 5x 10.3 4x 10-~- 4x10-3

High High High High

Medium High High High

High High High High

Although the study notes that the estimated probabilities are unlikely to be better than one order of magnitude, such work is important, and highlights that significant effort should be put into improving the strength of the FPSO hull midship section, fatigue, damage stability and in particular ship impact in order to reduce loss of life.

4.

H U L L STRUCTURES

4.1

FPSOs

The environmental load and motion characteristics of trading ships and FPSO vessels are quite different to each other, as are the hull girder load characteristics. A turret-moored FPSO experiences dynamic loads not only from the environment, but also from the mooring and riser systems. Therefore, an integrated approach to predict the dynamic loads is necessary. Aryawan and Incecik (2001) establish the influence of mooring forces and other non-linear effects on the hull girder loads. Mateus and Witz (2001) investigate the buckling and post-buckling behaviour of imperfect corroded steel plates used in floating marine structures such as FPSOs. The effects of general corrosion are introduced into the finite element models using a quasi-random thickness surface model. The postbuckling strength is influenced by general corrosion and it is shown that the standard deviation of plate thickness correlates well with the plate strength properties. Both maximum and minimum plate thickness do not significantly influence plate post-buckling strength. In recent years, there has been extensive work in practical design of fatigue strength using the SN curve and fracture mechanics approaches. The former is primarily applied for fatigue strength design, and mainly consists of two components, these being determination of hot-spot stress and selection of appropriate SN curves. Recent work on fatigue capacity has focused on the FPSO critical details since fatigue strength has become a key issue. Discrepancies have been observed between the hot-spot stresses predicted by different analyses. It is therefore important to derive an optimum procedure and standardize the analysis as part of the rules/code development. In this aspect, the International Institute of Welding (IIW) has published guidance, given in Hobacher (1996) on the determination of hot-spot stress. As a result of a large joint industry effort, Fricke (2001) recommends hot-spot analysis procedures for structural details of FPSOs and ships based on 'round-robin' finite element (FE) methods. Maddox (2002) presents hot-spot stress design curves for fatigue assessment of welded structures. Relevant design rules have not kept pace with computing developments in design, notably the increasing use of FE based stress analysis. The above work may be applied to develop hot-spot based fatigue assessment procedures that are alternatives to the traditional nominal stress approach in which the selection of SN curve is arbitrary. In parallel to this work, a Marintek led joint industry project (JIP) benchmarks hydrodynamic loads for fatigue analysis, and compares analytical results with full-scale measurement. Boom et al (2000) present the JIP project 'FPSO Integrity' to further understand FPSO fatigue loading, and in particular the

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validity of computational models. Data from the Glas Dowr vessel has been collected from Oct. 1997 to Aug. 1999 and subjected to spectral and statistical analysis as well as rainflow counts. Francois et al (2000) investigate various FPSO fatigue analysis methods. Given the breadth of the methods, reasonably good convergence has been found for certain response parameters. Other values, such as the computed fatigue damage at individual locations, vary substantially. The primary causes of these differences is found to be the way in which hydrodynamic loads are computed, and the method by which these loads are transferred to the structural model, together with the modelling of the local structural detail of stiffeners and brackets. FPSO hull structures are superficially similar to tankers with production facilities mounted on the deck. However, tanker design practices are based largely on simplified methods, such as ship rule-based parametric equations that are calibrated to shipping experience, and not necessarily relevant to FPSOs. Kuo et al (2001) overview a Fatigue Methodology Specification (FMS) for new-build FPSOs. The FMS interprets and extends existing classification society data, tailored to shipyard design capabilities and construction practices. Nordstrom et al (2002) predict mean FPSO headings by considering environmental loads on the hull and topsides throughout the vessel's operational life. A key element of this methodology is a directional representation of met-ocean data for each 3 or 6 hour sea state. Once the time history of the FPSO heading is known, fatigue lives at critical structural connections are predicted using the spectral fatigue method. Comparisons with existing industry practice confirm the need for a first principles based heading methodology for FPSO fatigue design. Zhao et al (2001) present extreme response and fatigue evaluations that are used as a basis for load prediction for FPSO design and conversion. The technical and economic issues key to the selection of FPSOs, TLPs and semi-submersibles are explored by Dorgant et al (2001). Specific projects referred to are Bonga (FPS - West Africa), Brutus (TLP - Gulf of Mexico) and Na Kika (semi-submersible - Gulf of Mexico). Cyranka et al (2001) focus on FPSO design and construction, with particular emphasis on the Petrobras P37 project. A comprehensive review of FPSO hydrodynamic and wind-tunnel model testing requirements, together with identification of key published papers, and discussion of the relative merits of numerical and physical modelling is given in HSE (2000). Other HSE work (2002b) describes 1 to 80 scale FPSO model tests, showing that vessel motions and mooting forces measured in long-crested seas underpredict equivalent short-crested results for 1 year return period steep waves (Hs=10m, Tp=l 1.7s) when vessel headings are up to 30 degrees from the bow. At larger headings pitch is over-predicted leading to conservative mooring loads. Testing with long-crested waves leads to overestimation of the greenwater loading. De Kat and Pauling (2001) discuss extreme motion prediction and capsizing of ships and offshore marine vehicles, overviewing major stability incidents on semi-submersibles, and numerical modelling of intact and damaged vessels. Guedes Soares et al (2001) provide an experimental and numerical study of the motions of a turret-moored FPSO in waves, showing that the panel method and strip theory are valid tools for predicting the surge and heave motion, though strip theory underestimates pitch. Irani et al (2001) present a test program for FPSO responses in non co-linear wave, wind and current conditions and make comparisons with results from co-linear environments. Fan et al (2000) investigate the dynamic performance of a turret-moored tanker. The second-order perturbation method is used to

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predict motions and line tensions using Pierson-Moskowitz and white noise wave spectra. The results are compared with experimental data and good agreement is obtained. Yang et al (2002) model test a 320,000 DWT tanker of 19.5m draught and tower yoke mooring system based on the 100-year environmental conditions with crossed current in the Bohai Bay region of the northern part of China. Water depths from 21-26m are selected. The results indicate that as the water depth decreases, the wave frequency motions in heave, roll and pitch reduce, and the low frequency motions increase. Interestingly the FPSO vessel rarely touches the seabed, even when the water depth is reduced to ultra-shallow. Stansberg et al (2002) develop a numerical design method for analysis of green sea impact loads on FPSO deck structures, validating results against model test data. Steep irregular wave conditions are considered, and numerical time series reconstructions made using the measured wave as input. A second-order numerical random wave description is combined with standard 3-D wave diffraction modelling of the vessel motions to predict the relative wave kinematics. A modified shallow water formulation is applied to predict the water propagation on deck, and resulting local pressures on the deckhouse are estimated by a similarity solution. Comparisons with experiments are made for the relative wave amplitudes, water propagation, and deckhouse loads. Reasonable statistical and individual event agreement is observed. Green sea occurrences are investigated, and characteristics identified. Doyle and Leitch (2000) discuss the Terra Nova FPSO, the first vessel specifically designed for ice infested harsh environments. Hull induced environmental loads have been minimized by the use of a slender shape, restricting the vessel beam, together with a moderate block coefficient. Pitching loads are reduced by avoiding blunt ends and abrupt changes of bow shape. The hull is designed to withstand the loads imposed by icebergs of up to 100,000 tonnes and pack ice of 0.3m thickness with 50% coverage. The mid-body and stern sections have been ice strengthened to satisfy the requirements of Lloyd's Register Baltic Ice Class Notation 1A. Japan National Oil Co (2001) present data on local floating oil storage facilities and the design guideline is now being proposed for ISO/CD 19904 Annex D. Floating oil storage facilities consist of floating oil storage tanks as a main facility, together with mooring equipment such as dolphins, fenders, etc. A significant proportion of large offshore fields have extensive gas reservoirs in locations remote from pipeline infrastructure and gas markets. Recently FPSOs have been proposed to produce liquefied natural gas (LNG) directly at offshore oil and gas fields. Faber et al (2002) summarize FPSO technology aimed at producing oil/condensate and LNG from the same unit. The technology builds on existing production experience with onshore LNG plants, LNG shipping and offshore FPSOs. Key technologies include the combination of oil/condensate and LNG storage in one single vessel, and the marinization of LNG facilities for use in remote offshore locations. Future goals include the complete integration of the oil/condensate production and LNG liquefaction facilities, and design of LNG ship-to-ship dynamic offioading systems suitable for offshore met-ocean conditions. Wadahl and Christiansen (2002) present a floating LNG facility based on the spherical tank system. It is shown that this design can be safely adapted to suit all weather conditions. Alvarado et al (2002) evaluate a number of gas utilization options on a combined oil/gas production FPSO. The paper also investigates project execution issues recognising the increased complexity and timescales associated with large LNG projects.

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Semi-Submersibles

Semi-submersibles have low motions but if dry trees are utilized the vessel requires heave suppression as well as a means to limit pitch motions. Halkyard et al (2002) present a deep draught vessel with a heave plate that is placed in a retracted configuration during fabrication and towing, allowing the deck to be installed inshore. Steen et al (2002) provide two design concepts in which a submerged heave plate is used to trap fluid and dampen heave motions. Additionally air cans, similar to those used to support spar risers, and TLP tensioners, may be used to support semi-submersible risers. The typical draught of a semi-submersible production platform is 20 to 25m. Bindingsbo and Bjorset (2002) demonstrate that increasing the draught from 21 to 40m significantly reduces vessel motions, thus lowering riser fatigue. Steel catenary risers may therefore be considered in harsh North Sea environments at water depths above 1000m. Often (2000) proposes a 4 column ring-pontoon dry tree semi-submersible wellhead platform for West Africa applications, with a lattice structure riser support tower that can be jacked to an elevated position in transit, and lowered to below the wave and current region during operation. Taut leg moorings with combined wire and synthetic rope are used, allowing the platform to be relocated over individual wells for drilling or workover. D'Souza et al (2001) present the design of a deepwater production and drilling semi-submersible that has the potential to significantly reduce capital cost, because a system engineering approach is applied to the design that enables the identification and integration of well construction, completion, drilling, riser, mooring and process technologies. Fachetti et al (2000) address the management challenges of converting the former Spirit of Columbus drilling vessel into the Petrobras production unit P36. Innovative semi-submersible designs are also finding wider applications in offshore exploration. Strong and Michel (2000) describe the design of an unbraced dynamically positioned vessel for deep water well service work, having a column based motion control system that uses low pressure air to rapidly alter draught and control heel and trim.

4.3

TLPs

Many TLP studies focus on the springing and ringing forces, motion response and in particular the nonlinear hull/tendon/riser coupled dynamic behaviour in deep water. Various authors have proposed analysis solutions both by numerical and experimental methods. Teigen et al (2000) investigate the coupled-body time-domain response of a floating mini-TLP and barge. The hydrodynamic interaction effects between the bodies are accounted for to second order. The results show little sensitivity to the grid sizes used for the multi-body diffraction analysis, and the time domain analysis is numerically robust. The system could be feasible for benign environments of up to 8m significant wave height. Li et al (2000) discuss the use of a mini-TLP for marginal oil fields in 500m depth. This utilizes a framed structure supported by a submerged cylindrical buoyancy/ballast tank that is attached by tendons to the seabed. The 100-year surge response is found to be less than 4% of water depth and the heave and roll response is small. Natvig and Johnsen (2000) present a coupled analysis of a mini-TLP in North Sea conditions of 1400m water depth. An inverted pendulum method is used with corrections for the effects of tether dynamics. The results indicate that coupled analyses can be important for tether fatigue. Extreme tether loading is unlikely to be altered however because of the high tether pre-tension.

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Tao and Thiagarajan (2000) investigate the influence of TLP column comer radius on the heave damping. This reduces the form drag exerted on the oscillatory cylinder, the reduction being larger as the comer radius increases. Both linear and non-linear heave-damping regimes exist, and the transition from linear to non-linear behaviour is delayed because of the influence of the column comer radius. Krokstad and Solaas (2000) present a study on the highly non-linear local flow phenomena close to a platform column excited by large steep waves. Experiments reveal that the flow can be considered as a hydraulic jump starting on each side of the column and colliding to give a significant vertical jet. Jumps are observed for all regular wave conditions tested. These are more pronounced with increasing wave steepness and wavelength. The jumps appear unrelated to the tinging force that occurs shortly after the jumps on each side of the cylinder collide downstream. Ma et al (2000) provide a consistent analytical approach for the prediction of non-linear coupled effects between a platform, moorings/tendons and risers. Parametric studies are performed to explore the coupled effects on a Spar and TLP in deep water, revealing that larger differences in mean offset are found in the 100-year loop current condition, because the uncoupled analysis cannot model the system stiffness and current loads on the vessel seabed connections in a consistent manner. Additionally analytical modelling of ringing is difficult and requires an improved 2 nd or higher order wave theory. Pollack et al (2000) present a Tension Leg Deck (TLD) concept, allowing dry trees and full drilling capability. The production risers are pre-tensioned by weights suspended well below the free surface in contrast with the classical buoyancy means. It is found that the measured tension fluctuations are systematically lower than the values predicted using an approximate method. Dynamic amplification factors of less than 6% are established for a West Africa environment. The motions of the suspended weights are well damped and thus of small enough magnitude to adopt a 'safety by distance' approach when it comes to riser/weight clearances. Niedzwecki et al (2001) author a model test study on the directional sea response of a mini-TLP. Various constants and frequency dependent spreading function parameters are used to represent the influence of wave direction on the motions of the vessel. The inplane motions are found to decrease and out-of-plane motions increase. Kim, M e t al (2001) perform a non-linear time-domain hull/tendon/riser coupled TLP dynamic analysis in 3000 feet water depth. The hull first order and second order sum and difference frequency wave loads and other hydrodynamic coefficients are calculated from a second-order diffraction/radiation program, while the tether and riser forces are found using the Morison formula. The calculation is carried out in 100 year hurricane conditions with non-parallel wind, wave and current. Mass-less non-linear springs are used to model the tethers and risers. A sensitivity study indicates that the mooring lines and risers contribute significantly to the surge/sway damping. It is noted that the equivalent static wind modelling may lead to a large underestimation of the slowly varying surge/sway response. The removal of wind or current loading gives respectively a 32% or 8% reduction in maximum tether tension. Teigen et al (2001) study higher order wave diffraction effects for a non-compliant TLP by both experimental and numerical methods to estimate the deck clearance under extreme wave conditions. Results indicate that in regular harmonic waves, in the absence of the structure, there is approximate correspondence between second order Stokes theory and experimental results. However in the presence of the TLP the 2nd order perturbation approach appears inadequate for monochromatic waves, and inconsistent with observations. Additionally higher order terms are more important for the high frequency force contributions. However higher order theory does not improve the correlation between theoretical and experimental estimates of wave height.

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Wang and Kim (2001) present tests in large waves investigating TLP springing and ringing. The tendon length of the TLP was originally 415m but this is extended to 665m and 965m to investigate the low pitching natural frequencies on the hull response. Birkelund and Powers (2001) analyse a mini TLP, investigating the non-linear relationship between random non-Gaussian waves and the vessel surge drift motion using a quadratic Volterra transfer function model. The model reliability is shown to be strongly connected to the properties of the estimators used for the higher order spectral moments. The authors found that it is essential, when interpreting the coherence of the model test data, to minimize spectral leakage. Kim, S and Sclavounos (2001) perform fully coupled large amplitude simulations of various offshore structures in water depths of up to 10,000 feet. A multiple-scale approximation is introduced for the decomposition of the wave and low frequency motions. Efficient solution methods are developed for computation of the frequency domain linear and second-order surface wave problems. Motion simulations are carried out in the time domain by fast Fourier transformation (FFT) summation of the linear and quadratic time series associated with the hydrodynamic coefficients. Non-linear solution algorithms to investigate the static and dynamic response of mooring lines, risers and tethers are developed. Chandrasekaran and Jain (2002) present the coupled dynamic analysis of a triangular model TLP in regular waves. Non-linearities caused by the change in tether tension and hydrodynamic drag force are considered. Results are compared with those for a square TLP. The triangular TLP exhibits lower surge and heave regular wave responses, but higher pitch. Rainey (2002) gives a comparison of similarities and differences between the 3000 foot water depth Brutus platform and other Shell TLPs. The Brutus hull and deck configuration are similar to the Mars and Ram/Powell hulls. The platform is designed to handle tieback to five subsea fields.

4.4

Spars

A Spar unit may be categorized as a deep draught floater with small heave motion. The deck/topside may be modular or integrated. The main hull is cylindrical with a central moon pool for tensioned risers. The circular hard tank provides buoyancy. Below is a circular shell (classic spar), or truss structure (truss spar), connected to a soft tank for fixed ballast, with plates to reduce heave and pitch motion. The spar and TLP are the most commonly selected hull concepts that allow dry tree units. The field development for the world's largest spar is presently under way. This is a truss spar with the first Spar Supported Riser System (SSVR). The platform will be moored by 16 lines and operated in 1400m water depth at the Holstein field in the Gulf of Mexico, producing in 2004. Integrated analysis methods for deep-water platforms such as the spar, as well as procedures to allow for coupling effects from the SSVR, are discussed by Hansen et al (2002). Luo et al (2000) review the selection criteria for implementing dry tree units, and evaluate the advantages of various hull concepts, illustrating recent field development studies. Wang et al (2001) overview truss spar structural design criteria and loading conditions, using random wind and wave time-domain analyses. The comprehensive paper covers the wet tow transportation, upending, truss in-place strength and fatigue analyses, as well as structural design of critical connections and heave plates. Prislin et al (2001) develop a numerical method for time-domain simulation of the spar upending operation, this process introducing motion and structural loading that is substantially different from that occurring during field operations. Up-ending is usually achieved by flooding the

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ballast tanks. As a result, the flooded water effectively becomes part of the spar, making the mass properties time dependent, and thus increasing the complexity of the equations of motion. The authors derive a solution by coupling the equations with the hydraulic equations for the tanks, good agreement being found with model test results. Luo et al (2001) discuss time-domain methods for fatigue analysis of critical connections in truss spars. These include the truss-to-soft tank, hard tank-to-topsides and truss-to-hard tank connections, the latter requiring particular attention since these connections sustain constant significant tensile loads. The procedure requires time-domain wave analysis of the global structural model and local FE analysis of connection details. Forces on connection members are combined with stress concentration factors at selected hot spots. Rainflow counting algorithms are used to calculate fatigue damage at these positions.

4.5

Offloading Systems and Buoys

FPSOs offioad crude oil to a shuttle tanker approximately weekly. For turret-moored FPSOs, the shuttle tanker is separated by a distance of typically 75m. However for spread moored FPSOs, a single point mooring system is used connected by a steel or flexible offloading riser, approximately 2 km away. Steel risers can have fatigue life concems because of the combined motion of the single point mooring and FPSO. Submerging these to approximately 100m below the sea-surface can reduce buoy motions. Alternatively offloading may be conducted using tandem-offloading operations, avoiding the use of buoys. An HSE study (2002a) considers damage resulting from encounters between shuttle tankers and FPSOs, establishing FPSO stem penetrations of up to 5.56m for 100 MJ impact energy. It is additionally noted that heave and pitch induced relative vertical velocities, over which the master has little control, may be up to 3.6 m/s in 6m Hs waves, leading to deformation energies of 270 MJ. Morandini et al (2001) list the specific problems induced by offloading operations including system limits in terms of relative heading and hawser tension, environmental shielding between the vessels, fish-tailing behaviour, and influence of the length and elasticity of the hawser connecting the two hulls. Additionally variations in floater loading conditions, and use of thruster and main engine to reduce instability and large relative headings are also outlined. Tricard et al (2002) describe the development of an offioading riser for the export of oil from a large FPSO to an offloading buoy. The installation of two W-shaped rigid steel pipes is considered appropriate for deep-water fields such as Girassol (1400m). The riser is of only 16 inches diameter and so fatigue strength is not considered to be an issue. Chaudhury (2002) presents an offloading scheme consisting of naturally suspended double catenary, U-shaped, rigid steel risers, and evaluates the strength and fatigue performance, technical/economical merits and safety of the system. Lebon and Remery (2002) present an overview of the design choices for flexible pipe oil offloading systems. The paper considers flow, mechanical and hydrodynamic aspects, and documents the structural integrity and fatigue life of the selected system.

4.6

Risk~ReliabilityBased Design

Continuing work has been performed on risk assessment methods, comparing specific designs with earlier systems that have been accepted to have adequate levels of safety. For example Gilbert et al (2001) present a study to compare FPSO risks with those for existing TLP and spar deepwater floating production systems in the Gulf of Mexico, and a shallow-water jacket serving as a platform hub/host.

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The complete production system is considered, from the wells through to transportation of the product to shore. Three risk measures are assessed and analysed for each system in a 20-year production life. These are the human fatality risk, the total volume of oil spilled as a measure of the chronic environmental risk, and the maximum volume spilled in a single incident. It is concluded that there is no significant difference in the fatality or environmental risks among the four types of systems. The study is useful for the regulatory agency and offshore industry in advance of FPSOs being installed in the Gulf of Mexico. Wolford et al (2001) provide an FPSO integrated risk based design covering the topsides, structural and marine systems, applied to a turret-moored vessel with gas handling. Examples of separator containment loss and mooring line failure are presented. Xu et al (2001) consider risk based inspection methods for FPSO hulls, the work summarising three levels of inspection strategy, namely probability based, risk based and optimum methods. Siddiqui et al (2000) present a reliability analysis of an intact TLP and a vessel with one tether missing for 12 sea states. A Von-Mises failure criterion is adopted to define the tether failure relationship associated with maximum tension. The minimum tension failure occurs when the tethers become slack due to loss of tension. The influence of wind on the progressive failure of the tethers is also studied. The results show that when one tether is failed or missing due to maximum tension the annual and lifetime failure probabilities of the remaining tethers differ in magnitude, but the order of the probability is the same as that of the intact case. However, an immediate replacement or repair of the failed tether is necessary to avoid further progressive failures. The order of probability of failure magnifies in the presence of wind. In the minimum tension case the annual and lifetime probabilities of failure are found to be less compared to those for an intact TLP. Sun and Bai (2001) consider the time-variant reliability and corresponding reliability index for various corrosion rates. Corrosion wastage depends on many factors, including coating properties, cargo composition, inert gas properties, cargo temperature, and maintenance systems and practices. The timevariant reliability, shown in Figure 2, and annual reliability of the hull girder are reduced significantly along with the mean ultimate strength. In the figure the dotted lines account for corrosion effects only, whereas the full lines represent corrosion effects and fatigue cracks. The four cases represent no corrosion (1), half-mean, mean, and double mean corrosion rates (2, 3, and 4). If an 80% reduction of annual reliability index is selected as the reliability threshold in order to maintain the hull reliability level, then hull inspection should be conducted for Cases 2, 3, and 4 at approximately the 20 th, 15 th and 10th service years.

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The degradation effect of fatigue cracks does not appear to be important to the hull girder reliability unless there is unsteady propagation. However, attention to cracks is important during inspections because this may induce a catastrophic event. The work indicates that the reliability index does not decrease if there is no corrosion. On the other hand, severe corrosion rates may significantly reduce the hull girder strength and hence the reliability.

5.

VESSEL

- SEABED

CONNECTIONS

5.1 Riser Systems Riser systems worldwide have increased in both complexity and cost to overcome the technological challenges associated with the growing number of floating production units in deepwater fields such as those offshore Brazil, West Africa and the Gulf of Mexico. They are designed for each particular application depending on the layout of the field, water depth, environmental conditions, vessel, moorings, and produced fluid properties. Different riser systems have been proposed, depending on the application. Various riser materials including flexible, metallic, and composite have been developed to operate, jointly or in isolation, in a large number of configurations such as free hanging or catenary, steep and lazy wave, steep and lazy S, riser tower, etc. A recent survey by Clausen and D'Souza (2001) identifies more than 1550 production risers in operation on floating units. Flexible risers encompass 85% of the total and are installed in water depths up to 1360m. Rigid risers in different configurations, such as top tensioned risers connected to TLPs and spars, together with steel catenary risers (SCRs), are used in water depths up to 1460m. Drilling risers have already been used in water depths of 1853m in a Brazilian early production floating unit. A key issue for deep and ultra-deepwater schemes is continued technological advances to reduce the development cost while improving reliability. The experience gained in designing riser systems in up to 500m water depth is being used to modify the existing technology to work in deeper waters.

5.2

Flexible Risers

Flexible risers have been qualified to 1500m and are expected to be installed in depths of up to 3000m in the next few years. The standard flexible pipe structure, in the market for more than three decades, has

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evolved, as indicated by Rytter et al (2001), to include composite tensile armours, improved insulation properties, and optical fibre monitoring devices. Besides being easy to install and retrieve, flexible pipes are robust solutions particularly in highly dynamic environments such as the shallow waters of the North Sea. In the Janice field on the UK continental shelf for example, a flexible riser system capable of accommodating up to 14 risers supported at mid-depth by two tethered arches has been successfully deployed in only 79m water depth (CSO, 1999). In this so-called pliant-S system, bend limiter devices are installed near the riser base, and the touch down points are attached to the arch anchor bases to avoid over bending. Nevertheless, flexible risers are designed in many configurations, the most common being the lazy wave and the lazy S, as described in Clausen and D'Souza (2001). The modelling of the riser-soil interaction near the touch down point (TDP) has a significant impact on the prediction of extreme response and fatigue damage of the riser system. Both static stiffness and timedomain force-deformation relations between riser and soil foundation are needed to predict the appropriate riser dynamic response. To date riser-soil interaction has been ignored or been approximated as simplistic soil spring models. It is hoped that in the near future the interaction and appropriate seabed configuration will be suitably modelled by 3-D analysis methods, accounting for cyclic pipe-soil behaviour, see for example the FE model discussed in Bai (2001) that has been used in the design of the Aasgard flowlines and export pipelines. Palmer (2000) indicates that observations of the riser-soil interaction have been made by ROVs and it is recognised that trenches can form due to a combination of soil bearing capacity failure and erosion caused by vertical riser motion. The near vertical trench walls may become unstable and collapse in soft soil conditions, and risers ride out of the trench and/or laterally cut the trench sides during storm conditions. Furthermore suction effects may be observed on the underside of the riser during lift-off. Work performed in the STRIDE JIP, reported by Willis (2001), allows improved modelling of the static stiffness and dynamic cyclical behaviour between the risers and soil foundation. However, much of the past research activities have been devoted to shallow rather than deep water experiments. Jacob et al (2002) propose work that captures riser TDP response, including site surveys of existing trenches generated by risers, and the development of detailed analytical riser soil interaction models that may be used to establish the global riser response. Souza (2002) presents results of laboratory tests performed on flexible pipes. A non-linear FE model, calibrated by experimental results, is used to analyse the internal carcass structural behaviour under external pressure. Based on the results, the author proposes an analytical methodology for calculating the pipe collapse pressure. Neto et al (2001) report experimental tests and analyses of armour tendons under compressive loading, that can potentially lead to bird-caging, as part of the flexible pipe development for the Roncador Field. More recently, Troina et al (2002) propose a strategy for design verification of deep water free hanging catenary flexible risers against bird-caging, focusing on tendon buckling. The work includes both global and local analyses of a flexible pipe considering different operational conditions. However, the authors emphasize that the results are sensitive to the conditions considered, and suggest further investigations to account for variable bending stiffness along the pipe length due to temperature variation and pipe curvature. Lemos and Vaz (2002) present studies on the influence of FPSO heading on the flexible riser fatigue performance.

5.3 Steel Risers The steel catenary riser (SCR), made of standard steel pipe, is the simplest and most cost effective riser system. Initially used in conjunction with TLPs, SCRs were first installed on a floating production unit

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in 910m water depth in the Campos Basin. Four years later two SCRs were installed in 1360m on the P36 semi-submersible platform at the Roncador Field, Brazil. Souza et al (2000) report important design, fabrication, and installation issues related to this pioneering system. The study reveals that fatigue of the welded pipe spans is the most critical failure mode. Because of the uncertainties related to the effect of plastic deformations induced during reel laying, it was decided to install the most fatigue critical sections (touch-down and top regions) using the J-lay method. The long-term reliability of the system could not be evaluated due to the unfortunate accident that caused the sinking of the platform in early 2001. The riser tower concept, first employed in deep waters at the 1440m depth Girassol field development, has benefited the use of steel risers. Thus riser-soil interaction problems are eliminated, while flexible jumpers are able to accommodate the large dynamic relative motion between the floating unit and the tower. Additionally, within the tower, production risers, gas lift and optional hot water riser pipes are grouped into insulated riser bundles, to ensure that flow assurance requirements are efficiently met, as described by Dailey et al (2002). Couprie et al (2000) propose a future design of hybrid riser tower that avoids the use of flexible jumpers. The hybrid riser is composed of two sections made of conventional steel pipe, these being a steel vertical tensioned riser anchored to the seabed through a flex-joint and attached to a buoyancy tank at mid-water, and a free hanging SCR between the floating vessel and the buoyancy tank. The authors indicate that the system is less environmentally sensitive than many riser systems, having high fatigue performance, and is applicable to water depths ranging from 500 to 3000m. A number of other alternatives with varying riser material (high strength steel, titanium, composite materials, etc.), mid-water or surface buoyancy devices, and other types of connections are also proposed. Though technically viable, the success of these solutions depends on local cost as well as numerous practical considerations. A number of studies consider the fatigue of SCRs. The main issues affecting the pipe and welded joint fatigue performance are the material itself, the weld procedure, together with the acceptable weld defect sizes and types defined by weld acceptance criteria. Martins et al (2000) present an asymptotic formula to assess the cumulative damage near the SCR top end, using an asymptotic expression for the dynamic curvature. More recently, the fatigue effect of plastic deformations induced during the SCR reel installation process has been studied through several joint industry projects. Hutt and Frazer (1999) show a series of full-scale tests on welded pipes using the Shielded Metal Arc Welding (SMAW) procedure. By comparing the test results with typical design curves, as in Figure 3, the authors report that the fatigue performance of both as-welded and plastically deformed pipes exceed the design requirements for SCRs. Unfortunately, neither the weld acceptance criteria nor the eventual weld discontinuities are reported in the paper, so care should be taken when drawing the same conclusions for different materials, weld procedures, and/or any other weld acceptance criteria.

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5.4

Risers - Alternative Materials

Much work has been carried out recently to qualify titanium and composite materials for riser applications both for drilling and production. In particular, efforts have been put into development of appropriate codes and recommended practices. The lack of guidelines is often an obstacle when utilizing new materials in a reliable and economical way. Recommendations focus on design aspects in terms of criteria to be used, material specifications and testing. General guidelines have been developed for design of load carrying structures and components fabricated from fibre-reinforced plastics and sandwich structures, see Echtermeyer (2000) and Echtermeyer etal (2002). The guidelines are based on a probabilistic foundation using the load and resistance factor design (LRFD) format, developed within an international JIP. A specific guideline for composite risers has been developed along similar lines as an integral part of the DNV LRFD standard for design and analysis of risers, as presented in Mork et al (2000). Materials with lower weight, better corrosion resistance and lower stiffness than steel can be of great advantage as deep-water riser solutions. Titanium and composites meet some of the above criteria. One of the major obstacles in application has been that the solutions can be expensive. One titanium riser was installed in the North Sea in 1995 and further applications are considered. Extensive test experience and theoretical knowledge have been developed over the last ten years. Test programs include evaluation of the mechanical and fracture mechanics as well as fatigue and corrosion testing. Both base material and welding have been tested in realistic environments and temperatures, a good summary paper being given by Torstad and Echtermeyer (1999). The behaviour of composite materials is reasonably well understood, but documentation of long-term performance has been lacking until recently. Composite risers and tendons have been extensively tested in different research programs, see Botker et al (2001). Composite materials have been used offshore for buoyancy modules, subsea protection structures and various other components. The first offshore field installation using a composite riser joint is reported by Salama et al (2002).

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5.5 Pipe-in-Pipe Systems Flow assurance and in particular the prevention of wax and hydrate formation becomes increasingly difficult as hydrocarbon production moves to deeper water, and/or long distances must be covered between wellheads and floating units. To improve flow assurance either active or passive insulation is considered. Active insulation consists of direct electrical resistance heating or re-circulation of hot fluids around the carrier pipe used in flexible bundles, riser tower and pipe-in-pipe systems. Passive insulation involves the use of new efficient materials in different layers, such as pipe-in pipe and sandwich pipe systems. Pipe-in-pipe systems usually consist of two concentric metal pipes in which the annulus is either filled with a non-structural insulating material or used to carry water for well injection, umbilical cables, etc. In this case, internal and external pipes are designed independently against failure under internal collapse and external burst pressure combined with other operational loads such as tension and bending. Sandwich pipes are designed to fulfil both structural and thermal requirements, the core material providing thermal insulation, and also giving resistance against burst or collapse, in conjunction with internal or external metallic pipes. Denniel and Laouir (2001) report the development of an electrically heated pipe-in-pipe system that combines both passive and active insulation properties. Optical fibre temperature sensors are embedded to measure the real time temperature profile. The electrical heating system is activated whenever the passive insulation is not able to keep the temperature above a critical limit. Though effective, active insulation systems using the pipe-in-pipe concept are expensive and more difficult to install. Draaisma and Levit (2002) present a thorough description of installation issues associated with the flowline systems for the King Project connecting three subsea wells in 1650m water depth to a TLP. The system consists of a 219 mm outer diameter (OD) cartier pipe and a 324 mm OD outer pipe with external coating. Active heating is performed by continuously pumping hot water through the annulus. Netto et al (2002) address pipe-in-pipe structures using small scale experiments involving sandwich pipes under external pressure together with a non-linear FE model. Pasqualino et al (2002) study the same structure using 3-D numerical analyses and consider combined bending and external pressure loading. The results indicate that, at any given pressure, the collapse curvature of a 6 inch sandwich pipe is greater than that for an equivalent single pipe. Both papers conclude that sandwich pipes are feasible for ultra-deepwater field applications as alternatives to conventional pipe-in-pipe systems. The integrity of pipe-in-pipe systems in the event of accidental collapse of the carrier pipe is also an issue of concern. The problem has been investigated by Kyriakides (2002) and Kyriakides & Vogler (2002) with an extensive experimental study involving small-scale pipe-in-pipe systems. An empirical equation is proposed to determine the propagation pressure based on the experimental results. In view of the potential danger of initiating a propagating collapse, pipe must be protected with buckle arrestors. All buckle arrestors developed for single pipes are also available for use on the outer pipe in pipe-in-pipe systems. These include slip-on, clamped and integral arrestors. The presence of the annulus in pipe-in-pipe systems allows the development of a new arrestor design consisting of a slip-fit ring in the annulus between the two pipes. The quasi-static effectiveness of this arrestor has been demonstrated in an experimental and analytical study by Oslo and Kyriakides (2002). Kyriakides and Netto (2002) show through pipe-in-pipe experiments in which buckles are propagated dynamically, that this arrestor

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design based on quasi-static calculations and experiments is adequate, as originally concluded by Netto and Kyriakides (2000a,b) for integral arrestors.

5.6 Steel Tube Umbilicals Deep water umbilicals utilize steel tubing as traditional thermo-plastic designs are not adept at coping with the high pressure loads in deep water. Furthermore steel tubes have improved resistance to hydraulic and chemical injection fluids as indicated by Williams and Williams (2001), and the tubes remove the requirement for armour wiring. However the steel components are potentially prone to fatigue failure particularly at butt-welded joints. Additionally the super-duplex steel, used because of its corrosion resistant characteristics, must be subject to rigorous QA/QC and NDT during fabrication to minimise heat treatment problems. Various JIPs, for example BPP (2002), are considering these issues.

5.7 Mooring Lines Publications have primarily dealt with the numerical and experimental evaluation of mooring line dynamic behaviour by accounting for non-linear contributions. In this context, the presentation of efficient numerical and reliable experimental techniques continues to attract particular attention. The developed numerical schemes are primarily concerned with solutions in the time domain. Particular attention has been paid to the examination of line-bottom interaction effects. The cable dynamics in the vicinity of the touchdown point (TDP) are influenced by non-linear phenomena such as snap loading, slack cable phenomena, instantaneous low tensioning, friction effects, and seabed modelling. The latter is represented as a perfectly solid or an elastic foundation. Chatjigeorgiou and Mavrakos (1998) predict the instantaneous TDP from quasi-static considerations and then generate a complete time-domain dynamic solution for the suspended portion of the cable. Figure 4 compares a simple model that represents line portion interacting with the sea bed as a linear spring and a more accurate model that accounts for the time-varying suspended line length obtained through quasi-static considerations. The results show that the simplified model significantly overestimates the tension along the line. Gobat and Grosenbaugh (2001 a) also consider bottom interaction, experimentally verifying the 'shock condition' to capture the tension discontinuity at the TDP of a catenary mooring. 3 x 107Combined

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A number of studies develop efficient numerical and analytical solution techniques for the evaluation of mooring line dynamics. Aranha and Pinto (2001) derive an analytical expression for the dynamic tension variation along the cable's suspended length, whereas Aranha et al (2001) follow the same methodology to obtain an analytical expression for the probability density function of the dynamic tension envelope in risers and mooring lines. Gobat and Grosenbaugh (2001b) propose an empirical model to establish the mooring line dynamic tension together with its standard deviation caused by its upper end vertical motions. The same authors (2001c) introduce a time integration of the cable dynamics equations. Chatjigeorgiou and Mavrakos (2000) present results for the numerical prediction of mooring dynamics, utilizing a pseudo-spectral technique and an implicit finite difference formulation. The effect of submerged buoys on the static and dynamic behaviour of deep water mooring lines is also investigated. Aranha and Pinto (2001) extend previous work by Mavrakos et al (1996), showing that the simplifying assumption of constant dynamic tension along the cable segments is valid for excitation frequencies far below the first elastic natural frequency. The calculation of extreme loads and mooring induced damping for wave and low-frequency vessel motions represent essential aspects of design. Nielsen and Bindingsbo (2000) provide approximate expressions for both the maximum tension and mooring line damping, assuming that the tension in the line is large relative to its submerged weight. Brown and Mavrakos (1999) present comparative results obtained from an ISSC instigated study on the dynamics of suspended chain lines and wire, revealing the sensitivity in line damping levels to the wave and drift frequency amplitude and period. The use of synthetic fibre ropes offshore continues to increase. Lee et al (1999) describe the ABS approach on synthetic ropes, while Stoner et al (1999) present the contents of an engineer' s design guide for fibre moorings, emphasising the limitations in available test data. Stoner et al (2002) outline additional work necessary before fibre moorings can be used at harsh weather locations.

5.8

Tethers

Tethers, used to moor TLPs, are generally air filled to reduce the submerged weight and thus act tp effectively increase the platform payload. Because of the external pressure, a significantly larger pipe wall-thickness is required for water depths beyond 150m. Hanna and Hannus (2000) together with Hanna and Salama (2001) present the alternatives for ultra-deepwater tethers, having stepped diameters and wall thicknesses and composite material characteristics. The upper sections are built from large diameter, thin walled members to provide extra buoyancy. The benefits are illustrated for Gulf of Mexico and West African developments. Botker et al (2001) discuss the feasibility of composite tether technology including the impact on overall TLP design and fabrication. Wybro and Koon (2002) consider the design of deep water TLP tether systems, the prediction of platform/tether performance, platform installation issues and top tension riser design and installation. They view that the API RP 2A LRFD code covering tether collapse in operating conditions is too conservative and could be updated using the more recent design criteria for pipe collapse, as given in Bai (2001). This work continues the development of strength criteria for pipe assessment under combined pressure, axial force and bending moment, thus allowing the required wall-thickness to be significantly optimised. Another solution allowing a reduction in required tether wall thickness is to inject internal pressure along the tether length to balance the increased hydrostatic pressure.

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There are many similarities between analysis and design methodologies applied to tethers and SCRs. In both cases the major design steps are the selection of pipe wall-thickness based on internal and external pressure requirements, the design of components such as flex-joints, the prediction of structural response using frequency and time-domain procedures, and finally extreme and fatigue structural response checks against design codes. It is noted that the prediction of moored system global performance plays an important role in establishing the structural integrity of tethers and risers.

5.9 Vortex-lnduced Vibration Vortex-induced vibration (VIV) of slender marine structures in deep water remains an important research subject. Although improved empirical models continue to be proposed, more work is being carried out using computational fluid dynamics (CFD). Etienne et al (2001) develop a VIV simulation method using strip theory. The code is coupled with FE structural methods. It is noted that strip models do not include any hydrodynamic correlation between the shed vortex cells. Dalheim (2000) uses CFD for the prediction of VIV from SCRs and obtains good agreement with experimental results. Oliveira et al (2000) propose a numerical method for VIV analysis based on the Galerkin finite element method and an explicit finite difference scheme for the time derivative. Oliveira and Sphaier (2001) suggest a hybrid numerical-analytical method to assess the hydrodynamic loads from 3-D flows applied to VIV response of a fixed cylinder and a cylinder supported by springs. Wilden and Graham (2001) present numerical results of the current induced vortex shedding and VIV response of flexible risers using a 2-D Lagrangian vortex method. The results are compared with experimental data for a single riser in both uniform and sheared currents at Reynolds numbers up to the order of 105. Larsen et al (2001) investigate the VIV of slender marine structures in sheared currents. Chang et al (2001) analyse a free-standing drilling riser to determine the fatigue damage contribution from the VIV response, indicating that this riser type can reduce deployment and retrieval time. Experimental VIV data in the critical flow regime rarely exist. Bearman et al (2001) present large-scale laboratory experiments to investigate the effect of Reynolds numbers from 4x104 to 1.2x105 on VIV response. These results can be used in riser full-scale design studies. Wilde and Huijsmans (2001) report VIV experiments with Reynolds number up to 5x105, and observe that the measured in-line drag coefficient of the freely vibrating cylinder in the early critical Reynolds regime is significantly lower than the previously established values obtained for lower Reynolds number. Vikestad and Halse (2000) present experimental work on the combined effect of current and waves using an elastically mounted circular cylinder in a time-varying flow. The results indicate that the cylinder has lower response in combined waves and current than when in current alone. Park et al (2002) report an experimental VIV study on a highly flexible free hanging pipe in water excited by top end oscillations, finding that the structural wave propagation plays an important role in the VIV induced response. Kaasen (2002) propose an estimator to identify the VIV modes and riser lateral motion, making use of pre-calculated mode shapes, and apply this to measured VIV data from drilling risers.

5.10 System Response Extensive effort continues to be put into the evaluation of the non-linear dynamics of moored floating structures. Garza-Rios and Bernitsas (1999) present a mathematical formulation for the horizontal plane motions caused by slowly varying drift forces acting on turret-moored FPSOs, whereas Kim and

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Bernitsas (2001) study coupled spread mooring and riser dynamics, using a design methodology based on non-linear stability analysis and bifurcation theory. Maeda et al (2000) conduct model tests in uniand bi-directional irregular waves on a very large moored floating structure, and compare the measurements with numerical predictions. The results indicate that second order horizontal motions dominate the total mooring line tension, while the first order motions cannot be neglected in irregular waves with smaller significant wave height. Gupta et al (2002) develop a parametric study on spar and riser configurations with associated tensioning systems for operations in water depths greater than 3000m. The authors present analysis results of the coupled hull-riser dynamic behaviour. 6.

CONCLUSIONS AND RECOMMENDATIONS

Floating production systems are extensively used nowadays for both shallow and deep water developments, and are being actively considered in ultra-deep water locations greater than 1500m depth. Monohull FPSOs are the most popular concept, and this is likely to continue with the recent work by the US MMS and USCG enabling consideration of FPSO schemes in the Gulf of Mexico. Additionally TLPs continue to be used and spar platforms are now considered mature development options. Legislation continues to tighten with the increasing requirement for project environmental assessments at the earliest stages of prospective field developments. Extensive and mature rules and guidance, specifically for FPS schemes, are now available from classification societies to guide the designer in the development of fit-for-purpose, efficient systems. An important aspect of recent rule changes is the increasing use of Load and Resistance Factor Design (LRFD) methods, where the target structure safety level is obtained by applying factors to both the characteristic loads and resistances. There continues to be a need for improved quantification of the joint occurrence of wind, waves and current at specific field sites, directional environmental load data being of particular relevance to FPSOs. For deep water developments data on the long-term spatial and temporal current behaviour is not available, though such information is required for design of riser, mooring and tether systems. Additionally in the Gulf of Mexico there remains uncertainty in the prediction of loop current effects. Further work is required in understanding FPS hull water interaction, in particular quantifying the uncertainty in predictions of FPSO green water forcing, slamming loads, wave run up and ringing associated with TLP columns, together with vortex shedding effects on spar hulls. Additionally the fishtailing behaviour of single point moored vessels or an FPSO offloading to a tanker is poorly understood. Areas of future research on FPS hull ultimate response include the development of strength equations for combined loads such as the buckling and collapse of plates and shells, the calibration of partial safety factors using risk assessment and structural reliability methods, the standardization and benchmarking of finite element models, and the development of procedures for determination of partial safety factors for finite element analysis and strength design based on testing. The results of such future investigations need to be implemented in regulatory codes and standards. Work is also required on fatigue related uncertainties, specifically selection of the combination of environmental loads, extrapolation of fatigue stresses at hot spots, selection and interpretation of design codes for fatigue assessment, and calculation of stress concentration factors. Additionally assessment of wave loading in conjunction with VIV, and the selection of safety factors together with inspection/repair methods needs future work.

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The large number of offshore fields with extensive gas reservoirs has led to studies on offshore floating LNG plants. Before these can become viable in deep water, exposed or remote locations further work is necessary to integrate oil/condensate production with LNG liquefaction facilities, and to design suitable dynamic ship-to-ship offloading systems for LNG export. The complexity of flexible pipe products leads to continued uncertainty in its use for deep water production, as designs are limited by pressure and diameter capability. Furthermore temperature limitations exist and sour reservoir service is unproven. Use of steel catenary risers and steel tube umbilicals is attractive from an economic and technical perspective in deep water, though fatigue issues remain to be addressed fully. Work also continues on riser and mooring interaction in the region of the seabed touchdown position as this can strongly influence loading and response. The coupled response between the floating hull and slender seabed connections is of key importance in deep water for the design of the station-keeping system. Efforts should continue here to develop appropriate validated prediction methods, as the validity of small-scale model tests using truncated mooring and riser systems is questionable. Although there are a number of deep water fibre mooring systems installed worldwide there is much work needed to develop an improved understanding of their dynamic response and integrity in field conditions, either through large scale testing or monitoring schemes. Ultimately there is a need for FPS developments that do not require moorings, relying solely on dynamic positioning for station-keeping. As research efforts progress, further emphasis should be placed on the development of results in a form more accessible for use in reliability based methodologies, as these are increasingly being employed both by the classification societies and the offshore industry to quantify and minimize uncertainty.

REFERENCES

The following conference publications are denoted as indicated: OMAE - Offshore Mechanics and Arctic Engineering, OTC - Offshore Technology Conference, ISOPE - International Society of Offshore and Polar Engineers, D O T - Deep Offshore Technology. ABS (2000). Guidefor Building and Classing Floating Production Installations, ABS, USA. Alvarado, C.S., Cone, R.S. and Wagner, J.V. (2002). Next generation FPSO: Combining production and gas utilization. Proc. OTC, 14002. API RP 2FPS (2001): Planning, Designing, and Constructing Floating Production Systems. API RP 2A-WSD (2000): Planning, Designing and Constructing Fixed Offshore Platforms - Working Stress Design. API RP 2T (1997): Planning, Designing and Constructing Tension Leg Platforms. Aranha, J.A.P., Pinto, M.O. (2001) Dynamic tension in risers and mooting lines: An algebraic approximation for harmonic excitation, Applied Ocean Research, 23, 63-81. Aranha, J.A.P., Pinto, M.O., Leite, A.J.P. (2001). Dynamic tension of cables in random sea: Analytical approximation for the envelope probability density function, Applied Ocean Research, 23, 93-101. Aryawan, D. and Incecik, A (2001) Extreme wave-induced loads and ultimate strength assessment of a turret moored FPSO vessel: A case study, Proc. of OMAE, OFT-5104. Bai, Y. (2001). Pipelines and risers, Elsevier Ocean Engineering Series. Bearman, P., Johanning, L. and Owen, J. (2001). Large scale laboratory experiments on vortex-induced vibration. Proc. OMAE, OFT- 1361.

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Bindingsbo, A.U. and Bjorset, A. (2002). Deep draft semi-submersible. Proc. OMAE, 28369. Birkelund, Y., Powers, E..J. (2001). Higher-order spectral estimators and non-linear system identification. Proc. ISOPE, 78-84. Boom, H. V. D., Krekel, M. and Aalberts, P. (2000). FPSO integrity; structural monitoring of Glas Dowr. Proc. OTC, 12143. Botker, S., Storhaug, T., and Salama, M.M. (2001). Composite tethers and risers in deepwater field development: step change technology. Proc. OTC, 13183. BPP Technical Services Ltd. (2002). Steel Tube Umbilicals - Phase 2 Programme (STU2). Brown, D.T., Mavrakos, S.A. (1999). Comparative study on mooting line dynamic loading. Marine Structures, 12, 131-151. Chandrasekaran, S., Jain, A.K. (2002). Dynamic behavior of square and triangular offshore tension leg platform under regular wave loads. Ocean Engineering, 29, 279-313. Chang, S.H.M., Stone, B.A., Vernotzy, R., Fisher, E., and Tompson, H. (2001). Dynamic response and vortex induced vibration of a free-standing drilling riser in deep water. Proc. OMAE, OFF 1140. Chatjigeorgiou I.K., and Mavrakos, S.A. (1998). Assessment of bottom cable interaction effects on mooting line dynamics. Proc. OMAE, 171-176. Chatjigeorgiou, I.K., Mavrakos, S.A (2000). Comparative evaluation of numerical schemes for 2-D mooting dynamics. Int. J. Offshore and Polar Engng., 10(4), 301-309. Chaudhury, G. (2002). Double catenary offioading lines for deepwater fields. Proc. OTC, 14310. Clausen, T. and D'Souza, R. (2001). Dynamic risers key component for deepwater drilling, floating production. Offshore, May, 89-93. CMPT (1998). Floating Structures: A Guide for Design and Analysis. Ed. Barltrop, N. 101/98. CSO (1999). Riser system developed for shallow water development. Offshore, Jun., 68. Couprie, S., Pionetti, F.R., Rocher, R., and Hugnot J. (2000). SHREWD: A single concept of steel hybrid riser for multiple applications. Proc. Rio Oil and Gas and Conference, IBP 212 00. Cyranka, C., Da Silva Pinto, M.C.M. and Dutra, L.A.M. (2001). FPSO Petrobras P-37 project overview. Proc. OMAE, OFT- 1092. Dailey, J.E., Healy, B.E., Zhang, J., and Brown, R.J. (2002). Truss riser tower in deep water. Proc. OTC, 14190. Dalheim, J. (2000). Numerical prediction of vortex-induced vibration on steel catenary risers. Proc. of 10 th ISOPE, Vol. III, 499-503. De Kat, J.O. and Randolph Pauling, J. (2001). Prediction of extreme motions and capsizing of ships and offshore marine vehicles. Proc. OMAE, OF/'- 1280. Denniel, S. and Laouir, N. (2001). Active heating for ultra-deepwater PiP and risers. Proc. OTC, 13138. DNV (2000). Design of Offshore Structures, General (LRFD method)/Structural Design for Offshore Ships/Structural Design of Column-Stabilised Units (LRFD method), DNV-OS-C 101/2/3. DNV (2001). Structural Design of Deep Draught Floating Units (LRFD method), DNV-OSC 106. DNV-OS-C 101 (2001). Position Mooring. Dorgant, P.L., Ballent, S.W., Rodenbusch, G., Lyties, W.H. and Rainey, R.M. (2001). System selection for deepwater production installations. Proc. OTC, 12966. Douglas-Westwood,(2002), The World Floating Production Report 2002-2006. Doyle, T. and Leitch, J. (2000). Terra Nova vessel design and construction. Proc. OTC, 1920. Draaisma, S., and Levit, M. (2002). BP Amoco King development p r o j e c t - Pipeline installation challenges. 25 th IBC Offshore Pipeline Technology Conference.

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D'Souza R., Barton, D. and Hatton S. (2001). Next generation of production drilling semi-submersible based field development system, Proc. OTC, 14260. DTI (1998). Decommissioning of Offshore Installation and Pipelines under the Petroleum Act 1998. Department of Trade and Industry (UK)-Offshore Decommissioning Unit, www.og.dti.gov.uk. Echtermeyer, A.T. (2000). New design guidelines and acceptance criteria for composites offshore. Third International Conference on Composite Materials for Offshore Operations. Echtermeyer, A.T., Osnes, H., Ronold, K.O., and Moe, E. T. (2002). Recommended practice for composite risers. Proc. OTC, 14022. Etienne, S., Biolley, F., Fontaine, E., Cunff, C.L. and Heurtier, J.M. (2001). Numerical simulations of vortex-induced vibrations of slender flexible offshore structures. Proc. ISOPE, Vol. III, 419-425. Faber, F., Bliault, A.E.J., Resweber, L.R. and Jones, P.S. (2002). Floating LNG solutions from the drawing board to reality. Proc. OTC, 14100. Fan, Ju., Chen, Xiaohong., Ji, Chunqun. (2000). The dynamic analysis of a turret-moored tanker. China Ocean Engineering, Vol.14, No.1. Fatchetti, M.B., Valerio, C.G.P., Loureiro, H.E. and Jorge, H.Q. (2000). The conversion of Spirit of Columbus semi-submersible into production vessel P36. Proc. OTC, 12140. Francois, M., Mo, O., Fricke, W., Mitchell, K. and Healy, B. (2000). FPSO integrity: Comparative study of fatigue analysis methods. Proc. OTC, 12148. Fricke, W. (2001). Recommended hot-spot analysis procedures for structural details of FPSOs and ships based on round-robin FE analysis. Proc. ISOPE, Vol. IV, 89-96. Garza-Rios, L.O., Bernitsas, M.M. (1999). Slow motion dynamics of turret mooting and its approximation as single point mooting. Applied Ocean Research, 21, 27-39. Gilbert, R.B., Ward, E.G. and Wolford, A.J. (2001). A comparative risk analysis of FPSOs with other deepwater production systems in the Gulf of Mexico. Proc. OTC, 13173. Gobat, J.I., and Grosenbaugh, M.A. (2001 a). Dynamics in the touchdown region of catenary moorings. Int. J. Offshore and Polar Engng., 11,273-281. Gobat, J.I., Grosenbaugh, M.A. (2001b). A simple model for heave-induced dynamic tension in catenary moorings. Applied Ocean Research, 23, 159-174. Gobat, J.I., Grosenbaugh, M.A. (2001c). Application of the generalized-t~ method to the time integration of the cable dynamics equations. Computer Methods in Applied Mechanics and Engineering, 190, 4817-4829. Soares, G., Fonseca, N. and Pascoal, R. (2001). Experimental and numerical study of the motions of a turret mooting FPSO in waves. Proc. OMAE, OFT-1071. Gupta, H., Finn, L., and Halkyard, J. (2002). Spar riser alternatives for 10.000 ft water depth. Proc. OTC, 14298. Halkyard, J., Chao, J., Abbott, P., Dagleish, J., Banon, H., Thiagarajan, K. (2002). A deep draft semisubmersible with a retractable heave plate. Proc. OTC, 14304. Hanna, S.Y. and Hannus, H. (2000). Ultra-deepwater TLPs - the tendon system challenge and alternative solutions. Proc. OTC. Hanna, S.Y., Salama, M.M. and Hannus, H. (2001). New tendon and riser technologies improve TLP competitiveness in ultra-deepwater. Proc. OTC, 12963. Hansen ,V.,Sodahl, N.,Katla, E.(2002). Coupled analysis of deep water systems. Proc. IBC Conference on Deepwater Risers, Moorings and Anchorings. Hobbacher, A. (1996), Fatigue Design of Welded Joints and Components, International Institute of Welding (IIW), XIII-1539-96/XV-845-96, Abington Pub., Cambridge, UK. HSE. (2000). Review of model testing requirements for FPSOs, Report 2000/123, by BMT Ltd. HSE. (2002a). Damage from shuttle tanker:FPSO encounters, Report 2002/006, by PAFA Cons Engrs. HSE. (2002b). FPSO response in long and short crested seas, Report 2002/018, by HR Wallingford Ltd.

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Huang, K., Bai, Y. (2001). Reliability methods for deepwater position-mooring design and analysis. Proc. OTC, 13269. Hutt, G., Frazer, I. (1999). The effect of plastic deformation on the fatigue performance of metallic risers for floating production systems. Proc. 2 "a Workshop on Subsea Pipelines. Inglis R.B, (1996). Production facilities selection for deep water oil and gas field development. Inst. Engineers and Shipbuilders in Scotland (IESIS), Vol. 139. Irani, M., Johnson, R.P. and Ward, E.G. (2001). FPSO responses to wind, wave and current loading. Proc. OMAE, OFF- 1023. Jacob P., Summers, P.B., Stevenson, P.R. and Audibert, J.M., (2002). Characterization of riser touchdown point behaviour from field investigation and analysis "ChaRT-B". Proc. of IBC Conference on Deepwater Risers, Moorings and Anchorings. Japan National Oil Cooperation (2001). Design Guideline of Floating Oil Storage Facilities. Kaasen, K.E. (2002). On identification of VIV modes from measurements. Proc. 12th ISOPE, 827-833. Kim, Boo-Ki, Bemitsas, M.M. (2001). Non-linear dynamics and stability of spread mooting with riser. Applied Ocean Research, 23, 111-123. Kim, M.H, Tahar, A., Kim, Y.B. (2001). Variability of TLP motion analysis against various design methodologies/parameters. Proc. ISOPE, Vol. III, 467-473. Kim, S., Sclavounos, P. D. (2001). Fully coupled response simulations of theme offshore structures in water depths of up to 10,000 Feet. Proc. ISOP, Vol. III, 457-466. Krokstad, J. R., Solaas, F. (2000). Study of nonlinear local flow. Proc. ISOP, Vol. III, 449-454. Kuo, J,. Lacey, P.B. MacMillan, A. and Zettlemoyer, N. (2001). Fatigue methodology specification for new-build FPSOs. Proc. OMAE, MAT-3016. Kyriakides, S. (2002). Buckle propagation in pipe-in-pipe systems. Part I Experiments. International Journal of Solids and Structures, Vol. 39, 351-366. Kyriakides, S., and Netto, T.A. (2002). Dynamic propagation and arrest of buckles in pipe-in-pipe systems. Proc. OMAE, 28600. Kyriakides, S. and Vogler, T.J. (2002). Buckle propagation in pipe-in-pipe systems. Part II Analysis. International Journal of Solids and Structures, Vol. 39, 367-392. Larsen, C.M., Yttervik, K.V.R. and Passano, E. (2001). Empirical model for analysis of vortex induced vibrations - Theoretical background and case studies. Proc. OMAE, OFT-1203. Lebon, L. and Remery, J. (2002). Bonga: Oil off-loading system using flexible pipe. Proc. OTC, 14307. Lee, M., Flory, J., Yam, R. (1999). ABS guide for synthetic ropes in offshore mooting applications. Proc. OTC, 10910. Lemos, C.A.D. and Vaz, M.A. (2002). Flexibe riser fatigue procedure using a long term distribution of FPSO's heading direction. Proc. OMAE, 28145. Li, R., Gu, Y., Bao, Y. (2000). A study on mini tension leg platform for marginal oil fields in medium water. Proc. 10 th ISOPE, 290-294. Luo, Y., Huang, E., Lai, N., Baudic, S. and Chen, H.T. (2000). Dry completion units for West Africa field development. Proc. OTC, 11927. Luo, Y.H., Lu, R., Wang, J. and Berg, S. (2001).Time-domain analysis for critical connections of a truss spar, Proc. oflSOPE, Vol. I, 362-368. Ma, W., Lee, M., Zou, J., Huang, E. W. (2000). Deepwater non-linear coupled analysis tool. Proc. OTC, 12085. Maddox, S.J. (2001). Hot-spot stress design curves for fatigue assessment of welded structures. Jl of Offshore and Polar Engineering, Vol. 12, 134-141. Maeda, H., Ikoma, T, Masuda K., Rheem, C-K (2000). Time domain analysis of elastic response and 2nd order mooting force on a very large floating structure in irregular waves. Marine Structures, 13, 279 299.

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Martins, C.A., Higashi, E. and Silva, R.M.C. (2000). A parametric analysis of steel catenary risers: Fatigue behavior near the top, Proc. l d h ISOPE, 2, 54-59. Mateus, A.F. and Witz, J.A. (2001). A parametric investigation of thickness variations on post-buckling behavior of corroded steel plates. Proc. OMAE, OFF 1210. Mavrakos, S.A., Papazoglou, V.J., Triantafyllou, M.S., and Chatjigeorgiou, J.K. (1996). Deep water mooting dynamics, Marine Structures, 9, 181-209. Morandini, C., Legerstee, F., Francois, M. and Raposo, C. V. (2001). Numerical analysis of FPSO offioading operations. Proc. OMAE, OFF- 1002. Mork, K.J., Sodahl, N., Kirkemo, F., and Leira, B. (2000). Design and analysis of metallic risers, limit state design. Proc. OMAE, PIPE-5051. Natvig, B.J., Johnsen, ~. (2000). Coupled dynamics of tether and platform for a North Sea wellhead TLP. Proc. l d h ISOPE, 295-301. NERC. (1998). Scientific Group on Decommissioning of Offshore Structures - Second Report (ISBN 1 85531 193 3, see also: http://www.nerc.ac.uk/) Neto, E., Maurfcio, J., and Waclawek, I. (2001). Flexible pipe for ultra-deepwater applications: The Roncador experience. Proc. OT, 13207.. Netto, T.A. and Kyriakides, S. (2000a). Dynamic performance of integral buckle arrestors for offshore pipelines, part I- experiments. International Journal of Mechanical Sciences, Vol.42, 1405-1423. Netto, T.A. and Kyriakides, S. (2000b). Dynamic performance of integral buckle arrestors for offshore pipelines, part II - analysis. International Journal of Mechanical Sciences, Vol.42, 1425-1452. Netto, T.A., Santos, J.M.C., and Estefen, S.F. (2002). Sandwich pipes for ultra-deep waters. Proc. 4 th International Pipeline Conf. Niedzwecki, J.M., Liagre, P.F., Borgman, L.E. (2001). Directional sea response of a mini-TLP. Proc. ISOPE, 447-452 Nielsen, F.G., Bindingsbo, A.U. (2000). Extreme loads in taut mooting lines and mooting line induced damping: an asymptotic approach. Applied Ocean Research, 22, 103 - 118. Noble Denton Europe Ltd. (2000). Rationalisation of FPSO design issues, Relative reliability levels achieved between different FPSO limit stages. HSE Offshore Technology Report 2000/097. Nordstorm, C. D., Grant, B., Lacey, P. B. and Hee, D.D. (2002). Impact of FPSO heading on fatigue design and non-colinear environments. Proc. OMAE, 28133. Often O. (2000). Dry tree semi-reduced costs for dry well completions in deepwater West Africa by application of proven semi-submersible and riser technology. Proc. OTC, 11876. Oliveira, M.C. and Sphaier, S.H. (2001). Numerical simulation of vortex induced vibration in three dimensions using a hybrid method. Proc. OMAE, OFI'-1205. Oliveira, M.C., Sphaier, S.H. and Marros, A.M. (2000). An application of numerical methods to the mechanics of vortex-induced vibrations, lOth ISOPE, Vol. III, 511-518. Oslo, E. and Kyriakides, S. (2002). Internal ring buckle arrestors for pipe-in-pipe systems. International Journal of Nonlinear Mechanics. Palmer, A.C. (2000). Catenary riser interaction with the seabed and the touchdown point. Deepwater Pipeline and Riser Technology Conf. Park, H.I., Hong, Y.P., Nakamura, M. and Koterayama, W. (2002). An experimental study on transverse vibrations of a highly flexible free-hanging pipe in water. 12th ISOPE, 199-204. Parker, W.J., and Grove, T.W., (2001). FPSO standards and rec. practices 2001). Proc. OTC, 13170. Pasqualino, I.P, Pinheiro, B.C., and Estefen, S.F. (2002). Comparative structural analyses between sandwich and steel pipelines for ultra-deep water. Proc. OMAE, 28455. Pasqualino, I.P., Valeriano, I.A., and Alves, T.M.J. (2002). Crack growth prediction in girth welds of steel catenary risers. Proc. 12th ISOPE.

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15th INTERNATIONAL SHIP AND OFFSHORE STRUCTURES CONGRESS 2003 AUGUST 11-15, 2003 SAN DIEGO, USA VOLUME DIEGO,

COMMITTEE V.6

FABRICATION TECHNOLOGIES

COMMITTEE MANDATE Concern for the quantification of general economic and safety criteria for marine structures and for the development of appropriate principles for rational life-cycle design using these criteria. Special attention shall be given to issues affecting code formation and development toward performance design, accounting for all uncertainties affecting actual structural behaviour.

COMMITTEE MEMBERS Chairman: Dipl.-Ing. H. Wilckens Dr. T. Borzecki, Dipl.-Ing. M. Heinemann, Mr. F. Lallart, Professor Nie Wu, Professor D. Olson, Professor J.Y. Park, Professor R.A. Shenoi, Mr. Y. Takeda,

KEYWORDS

High strength steel, Titanium, corrosion resistance, FRP, forming, welding consumables, adhesive bonding, accuracy in production, laser welding, robots, simulation, residual stresses, imperfections

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2 MATERIALS . . . . . . . . . . . . . . . . . . . . . . . . 2.1 Steel . . . . . . . . . . . . . . . . . . . . . . . . . . 2.2 Titanium . . . . . . . . . . . . . . . . . . . . . . . 2.2.1 Characteristics o f Titanium and its Alloys . . . . . . . . . 2.2.2 M a r i n e Applications . . . . . . . . . . . . . . 2.2.3 R e q u i r e m e n t s for the Use o f Titanium in Marine Applications . . . . 2.3 F R P . . . . . . . . . . . . . . . . . 2.4 Welding C o n s u m a b l e s . . . . . . . . . . . . . . . . . . . 2.4.1 Analytical Practice for Weld Metal Property Predictions . . . . . . 2.4.2 Role o f Weld Metal O x y g e n . . . . . . . . . . . . . . 2.4.3 Role o f Solidification on Inclusion Formation . . . . . . . . . . 2.4.4 Influences o f T h e r m a l E x p e r i e n c e . . . . . . . . . . . . 2.4.5 Prediction o f Weld Metal Properties . . . . . . . . . . . . . . 2.4.6 H y d r o g e n M a n a g e m e n t . . . . . . . . . . . . 2.4.7 Flux C o r e d Wires . . . . . . . . . . . . . . . . . . . 2.4.8 A l u m i n i u m Welding . . . . . . . . . . . . . . . 2.4.9 Welding F u m e . . . . . . . . . . . . . . . . . . . . .

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3 TECHNOLOGIES . . . . . 3.1 Cutting . . . . . 3.2 F o r m i n g . . . . . . 3.3 Joining . . . . . . . 3.3.1 Conventional Processes 3.3.2 Laser-Welding . . . 3.3.3 Friction Stir Welding 3.3.4 A d h e s i v e B o n d i n g . . 3.4 Surface Treatment/Coating .

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4 FABRICATION METHODS 4.1 M e c h a n i s a t i o n / R o b o t s . 4.2 A c c u r a c y in Production 4.3 F R P / C o m p o s i t e s . . .

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5 M O D E L L I N G A N D IT A S P E C T S 5.1 Production Aspects in D e s i g n 5.2 Simulation . . . . 5.3 Production Logistics

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6 FABRICATION IMPERFECTIONS 6.1 D i s t o r t i o n a n d R e s i d u a l Stresses 6.2 W e l d M i s m a t c h

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6.3 I n f l u e n c e o n S t r e n g t h 7 CONCLUSIONS REFERENCES

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INTRODUCTION

Fabrication Technologies are no longer seen in isolation; the modern process oriented shipbuilders assess them as one integral part of the business with manifold interrelations ranging from early product definition stages to operational aspects of the ship owner. The strong competition in global shipping markets require novel approaches in product development to arrive at high efficiency solutions. Hence a new quality of co-operation is emerging in the client/customer relationship, which is accentuated by novel possibilities in IT and CT developments and solutions. Aside from this another aspect i.e. the consequence of the strong price competition in commercial shipbuilding over the past ten years needs to be considered. There is a clear distinction between the product lines of the European and the Eastern Shipbuilders which leads to an increasing differentiation in production technologies, the related and interconnected processes and logistics chains. It is obvious, that building vessels with average plate thicknesses between 14 to 25 mm leads to different viewpoints when compared with those using 5 to 8 mm as an average. The consequences are then different directions for research in production technologies like in accurate manufacturing or highly efficient joining technologies. The issue of materials is also seen in new dimensions. The use of high-strength steels (690 MPa) is discussed with steel makers offering novel rolling processes which yield uniform material properties and reduced residual stresses. Filler wires and adequate welding procedures allow to maintain essential properties also in the as welded condition. The work of Specialist Committee V.6 did mainly focus on the fabrication of the ships hull. Design for production however requires a more stringent and detailed view on matters related to component and equipment installation. The modern integrated production process approaches are aiming at concurrent production i.e. installation of equipment at the earliest possible production stage. New fabrication planning tools allow to simulate the production process already at very early design stages and this ability creates new possibilities for integrating hull and outfit production. Additionally IT supported methods for distributed production and resources logistics play an increasing role in modern shipyards. Allan et al. (2000) address these points leading to visions of production technology development on a global environment however with a distinction between naval and commercial shipbuilding

2. MATERIALS For the bulk of ship production - commercial and naval - steel is still the major construction material for the hull. With respect to weight saving specifically with new types of ships (e.g. RoPax, Jumbo Containerships, Chemical Tankers) ultra high strength steels and Duplex Steels are required. The production processes for the new materials frequently lead to residual stresses which later can lead to severe production problems. However one can also observe novel ways of steel usage like in so called sandwich panels. Innovative joining technologies play an important part in new designs and material usage as will be discussed in a later chapter. Also composite structural elements are developed which combine materials of different properties to innovative fit-for-purpose elements. It seems that concepts from other industries (aircraft, car) are introduced into the shipyards. Looking at current research work, like for example in the European Commissions frame work programmes, innovative means of material usage and combination methods with bonding technologies are under development and investigation. This chapter does not deal with Aluminium because no important progress as compared to the last reporting period could be found. The are however some remarks related to welding of aluminium. A separate and in view of so far little information about the use of this material in marine applications extensive subchapter treats Titanium with regarding its properties and potentials in this field.

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A special chapter has been added on welding consumables, because this subject has not yet been dealt with at any depth in earlier reports of this Committee. With respect to new materials and production methods here are potentials for influencing the process and product qualities.. 2.1

Steel

Okano et al (2002) reported the development of high strength steel plates with YP 355 to 460 MPa, up to 65 mm thickness, and corresponding welding materials, in order to respond to the trend towards containership enlargement. These steel plates and welding materials ensure HAZ toughness for highinput welding (for example 50 kJ/mm SEGARC welding). The chemical compositions of these materials are of the low carbon equivalent type, and utilize small amounts of Titanium. Kaiser et al (2001) describes the production techniques and properties of modem high-strength structural steels with up to 690 MPa minimum yield strength. They paid particular attention to obtain the data assessing the fracture toughness and component-like behaviour of these modem steels. The data obtained were used for assessing the failure behaviour of components in terms of Y/T-ratio as well as modem safety concepts like CEGB R6 and Annex C of Eurocode 3. All steels showed high fracture toughness and component safety. And another finding is that the yield to tensile ratio is not a proper value for evaluating the failure behaviour of components made of high strength steels in the case of multi-axial loading. The fracture and deformation behaviour of high strength steels with yield strengths of 690 and 890 MPa was analysed by Langenberg et al. (2000). The ferritic fine grain steels of quenched and tempered type and plate thicknesses between 15 and 70 mm have been investigated with the focus on the clarification of the relevance of the yield to tensile ratio, which in view to the design safety of the construction is used in the design rules especially regarding the true material toughness. The results of the research clearly indicate that the yield to tensile ratio is not a proper value with high strength steels. At the same time it could be proven that the Annex C, Eurocode 3 brittle fracture avoidance safety concept is valid for these steels. With regard to the construction of accurate blocks, a new TMCP steel plate, called as "Residual stress controlled TMCP steel plate", was developed for use in fabricating the cargo tank and hull (Tani et al, 2001). In the investigation, using thermal-elastic-plastic FEM and experiments, the effects of residual stress of TMCP steel plate are quantitatively discussed in slit-slot cutting. As a result, the following advantages are found: 9 Residual stress of TMCP steel plate has great influence on the accuracy of slit-sot cutting as well as on the accuracy of transverse panel after stiffener welding. 9 To achieve high quality of block accuracy, it is effective to build ship blocks using "Residual stress controlled TMCP steel plate". Noelle, P. (2001) presents an analysis of problems associated with welding of Austenitic/Ferritic Duplex steels as being used in transport containments for aggressive liquids. N order to maintain the materials properties in as welded condition special weld procedure and quality assurance specifications need to be applied. 2.2

Titanium

Titanium was discovered at the end of the 19th century. It has been first industrially used in 1950. In the beginning of the application, the main interest focused on its specific strength at high temperatures, which predetermined the wide use of Titanium alloys in aircraft and space industries. Later, attention

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has been paid on its favourable ratio of strength, toughness, and fatigue resistance to weight and its unique corrosion resistance in a wide range of aggressive environments (Peakock, 2000). It resulted in rather a wide application in chemical and shipbuilding industries (Gorynin,1999). At present, the use of Titanium in the offshore industry and also in certain structural parts in ships is one of the growing applications of this special material. TABLE 2.2.1 MECHANICAL PROPERTIES

Titanium (T40) Aluminium Stainless steel (316L) Steel

2.2.1

Density 9 (g/cm3) 4.5 2.7

IL/p (MPa/g/cm 3) 77.8 101.8

Rm/0 (MPa/g/cm 3) 107.8 90.0

7.8

30.1

51.3

El9

UTS*k/p

24.5 25.9 25.3 26.9

64.7 35.6 20.5

Characteristics of Titanium and its alloys

The corrosion resistance of Titanium in marine applications is very good. It is an extremely reactive metal but it seems to be resistant in any environment except to reducing media to any chemical attack due to an immediate passivation, i.e. a thin but tenacious oxide layer of TiO2 is formed, adherent to the metal which protects it from external environment attacks especially in chloride environments such as seawater. Furthermore, when in contact with other metals, Titanium is not subject to galvanic corrosion but may increase corrosion rates of more active metals such as steel, aluminium, copper alloys. Titanium is also insensible to season cracking, to crevice and pitting corrosion, unlike stainless steel. The fire resistance of Titanium is also good. Only Titanium dust is subject to burn or explode due to its high affinity to oxygen. The melting point is above 1650 ~ foreseeing fire resistance. Tests have demonstrated that Titanium pipes have an excellent ability to withstand fire conditions even under shock conditions. In addition, thermal conductivity of Titanium is low, reinforcing its fireproof characteristics. Titanium is demonstrably safe for fire fighting water systems even under shock conditions (Peakock, 1994). Fatigue strength of test specimen of Titanium and its alloys is typically 50 to 60 % of the tensile strength values. Furthermore, fatigue limit and fatigue crack propagation rates are not affected by marine environments. The exceptional properties of Titanium are overshadowed by a high price of raw material, an unsteady supplying in quality and price, additional difficulties for manufacturing. The high price of raw material is partially due to a complex elaborating metallurgy and small scale of the market, mainly linked with aeronautics. But a new process of extraction and processing has been discovered which could open new perspectives for cost reduction. Titanium supplying market is chaotic and easily troubled due to its small scale and newness : prices are linked to the need of aeronautics, the main consumer of Titanium and its alloys. The Titanium market has a volume of about 50 000 tons per year. Ore resources are localized in Australia, South Africa, Ukraine and Canada, Titanium sponge is produced in USA, Russia, Japan and China, ingots and semi-finished products manufacturers are placed in USA, Russia, Japan and U.K. Titanium suffers from the lack of references, experiences and standards in a wide range of applications. Standards are at present not adapted for it, although a design and manufacturing requirements code is in development.

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Titanium machinability is similar to stainless steels: grinding, cutting, planning are realized using tough tools such as carbide tools with a low removing material rate to avoid the heating of the metal. Thermal cutting can be used as a pre-cutting process. Welding is difficult because of the high affinity of Titanium to oxygen above 450~ which leaves primarily inert gas processes with argon or helium. LASER and PLASMA processes could also be used (Osip et al, Inoue, 1995). Although Titanium properties are recognized especially with regards to corrosion resistance and mechanical characteristics, the notch sensibility and Sustained Load Cracking (SLC) effects must be considered. SLC effects have been noticed on several alloys (TA6V ELI, T60) and depend apparently on the strength sustained and the temperature as Millot (2002) explains.

2.2.2

Marine applications

The special characteristics of Titanium favour its use in such applications where extreme resistance to seawater corrosion, high weight saving or high mechanical properties are required. Three examples of marine application are given. Titanium use in heat exchangers is certainly the most widespread and oldest use of Titanium alloys in marine applications. Metals traditionally used were copper-nickel alloys or stainless steels. Those materials have a good thermal conductivity, but their relatively low corrosion and erosion resistance (Erskine, 1997) require to design the exchanger with an extra thickness, which limits the thermal exchange. The service life of those devices is limited to 6 years or less (Gorynin,1999), depending on the environment. Titanium exchangers are thinner because no corrosion/erosion allowance is needed. This permits significantly higher operating velocities, compensating a low thermal conductivity. Furthermore, weight is saved and the service life is at least 40 years, reducing maintenance and life cycle costs and avoiding technical stops. Regarding Offshore applications the 'Heidrun' floating platform may serve as an example (Johnson, 2002). The platform is installed in the North Sea with riser and seawater ballast system made of Titanium. The riser is a combined drilling and production riser with a maximum working pressure of 350 bars. Three reasons in combination led to the choice of Titanium: 9 Fatigue: for the same fatigue life, a steel riser must be considerably ticker than Titanium. 9 Weight: the weight of a Titanium riser will be considerably lower than a steel riser. Therefore the hydraulic positioning system which supports the weight of the riser on the platform structure can also be smaller, saving additional weight of the floating platform. 9 Costs: taking into account the extra thickness of a steel riser plus much larger hydraulic positioning equipment necessary for its support and the considerably stronger structure, a Titanium riser was found to be cheaper. The riser is made of TA6V (ELI). It is an extruded pipe with a flange welded at both ends using a hot wire plasma welding process. Thermal treatment to maximize fracture toughness and shot peening to increase resistance to crack initiation were also in the riser fabrication process. The flange bolts were made of Gr. 5 ELI and the mud line (mounted outside of the riser) was made of Gr. 9. The mast structure for a navy ship may serve as an example for Titanium use in ship design. The High Data Rate mast for the USS VIRGINIA (US Navy web site), was originally designed as a stainless steel and composite structure, but could not meet size and weight specifications. The use of a Titanium alloy allowed to meet these specifications and provided an advantage with respect to reduced life-cycle cost due to unique corrosion resistance properties. A special alloy developed by the Navy in collaboration with TIMET, i.e. Ti-5111, has been used because of its high strength and fracture toughness, its good

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weldability and its stress corrosion resistance. Some of the most common Titanium alloys used for ship and offshore applications are presented in the following table. TABLE 2.2.2 TITANIUM ALLOYS Nominal composition

Unalloyed Titanium (CP)

ASTM grade

Alloy type

(MPa)

1

220

345

2

345

485

345

485

220

345

min. 250~

910

1000

80~

825

895

min. 250~

Rm

80oC

16 17

ot

Ti 0.05 Pd

Ti-6A1-4V

Crevice temp. threshold

Re (MPa)

5 c~-13

Ti-6A1-4V-0.1 Ru (ELI)

29

Timetal 5111

Ti-3A1-8V6Cr-4Zr-4Mo

19

near c~

700

800

[3

1150

1250

min. 200~

Attributes

Application example

Most formable grade Piping, valves, Most formable, pumps, seawater most widely used ballast system grade Added corrosion resistance over Heat exchangers grade 1&2 Basic grade of most widely used alloy Added corrosion resistance over grade 5 High dynamic toughness, seawater stress corrosion cracking resistance. High strength, good corrosion resistance.

High pressure heat exchangers, structural parts Drilling and production riser

Superstructure for navy ships

Alpha alloys or near alpha alloys are widely used because of good corrosion resistance in seawater (crevice corrosion can be improved with palladium or ruthenium addition), their excellent weldability and easy manufacturability in general. Alpha beta or beta alloys are used when a high strength resistance is required, optimum mechanical characteristics are obtained with a heat treatment. weldability is fair, especially for beta alloys.

2.2.3

Requirements for the use of Titanium in Marine applications

Standards exist for usual structural materials such as steel and aluminium. Due to the lack of specific references these standards disfavour Titanium by a high factor of safety. Current welding specifications for Titanium are either specific to the aerospace industry or grouped under the umbrella of the ASME pressure vessel code (Luckowski,2001). Both application areas are not suitable for non aerospace or civilian applications. Therefore a committee (including US Army members) is developing a guide to fusion welding of Titanium and its alloys. This guide contains non-binding practical recommendations and is adaptable to every application. The US Army, needs such a specification and is in the process of generating a large knowledge base in Titanium structural weld performance which would effectively benchmark code requirements. It will include requirements for Titanium in structural application in the following fields:

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9 9 9 9

Design (statically and cyclically loaded structures, welded connections); Process qualifications (preparation, weld profiles, shield); Welder qualification; Inspection (colour inspection, radiography, eddy current, hardness correlation) and acceptance criterion.

Titanium, with its oxide (TiO2), is the fourth metallic element present in the earth crust. The high price of raw material (metallic Titanium) is presently due to the complex Kroll process: TiO2 cannot be reduced by current reducing agents such as carbon and hydrogen which are both inefficient. Metal is first obtained by a chlorination of TiO2, then a reduction with magnesium by electrolysis. The resulting Titanium sponge is then remelted two or three times under vacuum by an electric arc process (Vacuum Arc Remelting). A new process has been developed by the University of Cambridge (Ward-Close et al.) which could lead to considerably cheaper Titanium alloys and thereby enhance its use for marine applications. The industrialization of the new process will however require further research and development efforts. This new process may open wider economic opportunities and could be the trigger for a full size industry. 2.3

FRP

Fibre reinforced plastic (FRP) materials continue to be used for specialist structural applications in marine structures and for vessels such as minehunters, patrol craft, lifeboats, yachts and surveillance boats. The literature in the fields of fibre reinforcements, resin chemistry and core material formulations is extensive. For general marine structural engineers, up to date comprehensive treatments of the fibres can be found in Shindo (2000), Yang (2000) and Dwight (2000) and resin chemistry and properties in Varma et al. (2000) and Muzzy (2000). A study of the challenges for composites into the next millennium from a reinforcement perspective can be found in the work of Bannister (2001). A general treatment of the potential and properties of FRP in the specific context of ships and marine applications can be found in Shenoi et al. (2000). A large proportion of the practical research in the past three years has been devoted to increasing the potential for the use of such materials for high exposure, high performance structures used in America's Cup races. Belgrano (2001) explores the philosophy adopted in exploiting the specific characteristics of toughened resins, high specific strength/stiffness fibres and the compliant core materials for the structural topologies adopted for racing yachts. A major drawback of FRP composite materials however, and one that prevents more extensive usage, is its performance under fire. Because most FRP materials are combustible, there is a need to prove their capacity under such scenarios. Greene (2000) characterises the performance of composites using four criteria: 9 "Flame spread" is defined as the area at which flames travel along the surface of a structure. 9 "Burn through resistance" indicates the ability of a boundary to contain a fire. 9 "Structural integrity" is ability of hull, deck and a bulkhead to support design loads during and after a fire. 9 "Smoke production" indicates the amount of smoke produced by the combustible material Table 2.3.1 below compares, in a qualitative manner, the fire performance of different fibres, resins and core materials.

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TABLE 2.3.1 FIRE PERFORMANCE OF COMPOSITE MATERIALS

Material Polyester resin Vinyl ester resin Epoxy resin Phenolic resin E-Glass fibre Carbon fibre Aramid fibre Balsa core PVC core Phenolic foam core

Flame spread Poor Fair Fair Excellent Good Good Fair Good Fair Good

Burn through resistance Poor Fair Fair Good Excellent Good Fair Good Fair Excellent

Structural integrity Poor Fair Fair Good Good Excellent Good Excellent Good Good

Smoke production Poor Fair Poor Good Excellent Excellent Good Good Poor Excellent

One other aspect of fire resistance is the increasing interest in the use of flame retardant materials within the resin formulations (Murphy 2001). Flame retardants act either by delaying ignition or, after ignition, by depriving the fire of fuel or lowering the temperature to below burning point. This involves technologies of liberating fire-quenching chemicals or the formation of intumescent chars. It is noteworthy that all composite materials considered for marine applications have to conform to SOLAS regulations. In the case of high speed craft, this refers to the IMO HSC code (2000) and the fire testing protocol or FFP document (1998). While fire is an issue that affects operators of FRP craft, the styrene content within certain resin systems affects the processors and manufacturers of FRP structures. Many countries have legislations, or good working practices, that encourage styrene to 50 ppm occupational exposure limit. Much of such emissions arise from open mould constructions practices, principally the spray-up technique used in mass production of leisure craft hulls. Recent advances reported by resin suppliers (Marsh, 2001), (Darby et al. 2001), focus on a multi-pronged effort to combat high styrene levels characterised by: 9 Reducing emissions by altering resin formulation and using low styrene emission (LSE) resins 9 Investigating alternatives to the spray lay-up process 9 Increasing the use of closed mould techniques such as vacuum assisted resin infusion moulding, known variously as VARIM, VARTM, SCIMP, RIFF, etc. 9 End-of-pipe solutions, specific to the reduction of emissions to the environment. Data on the latest range of FRP materials can be found in the standard literature, though increasingly use is being of the world wide web. Table 2.4.2 in the REFERENCES lists some of the principal sites for a range of FRP materials.

2.4

Welding Consumables

The weld deposit joins structural assemblies together. Because the steel weld deposit is a product of solidification followed by solid-state transformation, it produces microstructures and properties that are different from that of the parent metal. Even when weld and parent metal have the same composition, weld metal must be understood and evaluated as its own material in a welded structure. Also with new welding processes like the various options for Laser welding, the consumables have a very important and as yet not fully understood influence on the quality of the joint. In the following, four welding consumable topics are introduced as timely opportunities for research and development. In addition certain special aspects of aluminium welding are discussed.

200 2.4.1

Specialist Committee V.6 Analytical practice for weld metal property predictions

After a half-century in the development of methods to predict weld properties and to select welding parameters, analytical approaches and practices are now achieving more reliable results. The first analytical expressions were based primarily on composition and had a "rule of thumb" usefulness in achieving the selection for the degree of preheat and/or post weld heat treatment that would be required for a specific steel composition (Granjon 1967, Ito et al. 1968). These compositional tools would only correlate to results for the same medium carbon steel, plate thickness and welding parameters, usually for SMA welding. With the introduction of higher strength low carbon steels, having properties based on strengthening mechanisms other than the austenitic decomposition, new predictive expressions were required. As new welding processes increase productivity, it also becomes essential to present the cooling rate into these expressions (Lundin et al. 1989). With new understanding of the interrelationship between welding process parameters, weld metal composition, microstructure and properties as well as the new availability of statistical and neural net analytical approaches as discussed by Metzbower et al. (2001) and Blackburn et al. (1997), the selection of welding consumables can be based on a quantitative footing. It soon should be possible to rapidly select the welding consumable composition for a specific alloy and welding thermal experience. To better understand where this development has come from and where it is going, the necessary features for these analytical methods are described. Fundamentally derived forms of the predictive equations have been suggested by Liu et al. (1986) considering both thermodynamic and kinetic approaches to their formulation. For a thermal transformation, in which phase stability can be expressed by AGy_a, a microstructural sensitive property, such as the carbon equivalent CE, can be given by the following: CE = Ko[C + KM, Mn + KsiSi +... + Kc, CLnC + KM,,MnLnMn + Ksi,SiLnSi +...]

(1)

where Ko is a proportionality constant and Ki are coefficients for the various alloying additions and are subject to fundamental interpretation. Mn, Si, C . . . . are concentrations of the different elements expressed in weight percent. This expression containing both, linear and nonlinear terms is based on the assumption that micro structural sensitive properties are directly related to the amount of alloying elements present in the initial and transformed phases. Omission of the nonlinear terms will simplify the equation to a form similar to the n w carbon equivalent equation. In the case of low alloy steels, omission of the nonlinear terms may not be too distorting and the coefficients could be manipulated to compensate for the omission. However, it becomes apparent that in high alloy systems, some of the information about alloying behaviour may be lost without the nonlinear terms. Assuming that the carbon equivalent can be directly related to the thermodynamic driving force for carbon transport, a new form of the equation for CE can be obtained and expressed as: CE = K:C[1 + K'cC + K'MnMn + K'siSi +...]

(2)

Equation (2) is different from Equation (1) and the form suggests a multiplication relationship to the interaction between carbon and other alloying elements. The presence of interaction terms is reasonable because each alloy addition influences the carbon behaviour and, thus, should be a product term with carbon. This form of the carbon equivalent equation should better fit the low carbon micro alloyed steels in which there is carbon-nitride precipitation. Although the above equations are commonly applied for HAZ in wrought materials, it is questionable whether these expressions have the ability to predict properties and transformation behaviour of weld

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metals. Influences from solidification and second phase particles, such as inclusions, will also affect the weld metal solid state transformation reactions. In addition to the elements considered above, the oxygen content must be included to make these expressions useful in predicting weld metal properties. 2.4.2

Role of Weld Metal Oxygen

Welds typically pick up oxygen to levels of several hundred ppm then deoxidize to oxygen levels of around 200 ppm with the formation of oxide inclusions. Weld metal inclusions, resulting from oxidation, are an important factor in influencing weld metal microstructure in low carbon and high strength steels. The type, size distribution, and density of weld metal inclusions are all important variables to control in order to achieve the desired weld metal microstructures and properties. An oxygen term or terms need to be added to the carbon equivalent expressions, and the magnitude of the coefficient is dependent on the type of inclusions being formed (the type of consumable being used). An adjusted Pc,,, called Pcmo is given by Onsoien et al. (1996): Si Pcmo = C + m + 30 2.4.3

Mn + Cr + Cu 20

V Mo Ni +m+~+m+5B---O 10 15 60

3 4

(3)

Role of Solidification on Inclusion Formation

During cellular or dendritic solidification commonly observed in steel weldments, solute elements segregate to the liquid at the solid/liquid interface, and the liquid concentration of specific solute elements can reach high levels in the interdendritic spaces. As solidification progresses from 0 % to 95 % solid (liquid fraction from 1.0 to 0.05), the interdendritic silicon concentration increases by about 50 %, manganese by about 100 %, oxygen by a factor of eight, and the aluminium concentration also shows a very small increase. The dramatic increase in the interdendritic oxygen concentration as solidification progresses will cause additional oxides to precipitate in the weld metal as discussed by Frost et al. (1992). It is these "last to form" oxides that serve to establish the fine acicular ferrite in the weld metal. This micro segregation must be accounted for in future predictive expression for weld metal and microstructure and properties.

2.4.4

Influences of Thermal Experience

Most carbon equivalent expressions are only a function of composition and do not consider thermal experience and severity of quench. The thickness of plate, edge preparation, and heat input of the process will influence the cooling rate. The time to cool from 800 to 500 ~ Ats/5, is an alternate way to describe the cooling rate of a weldment and has been used to compare the welding thermal cycle to the acceptable thermal experience to achieve properties. Carbon equivalent type expressions that contain a Ats/5, have been introduced (Yurioka 1983). An approximate calculation for Ats/scan be obtained from the Rosenthal solution for welding heat flow behaviour. It is clear that to describe the microstructure and mechanical behaviour of a high strength low alloy steel weld metal, the effects of alloying elements, oxygen, welding parameters, and cooling rate need to be considered. 2.4.5

Prediction of Weld Metal Properties

Empirical expressions have been developed to predict hardness (Hart et al. 1986), yield and ultimate tensile strength (Blackburn et al.1997 and Yurioka 2002) as well as ductility and toughness (Metzbower et al. 1999) for low carbon and low alloy higher strength steels weld metal. Various approaches, including a deterministic scientific approach, statistical approach and neural net analyses, have been attempted and compared.

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Extensive experimental data have been generated by Metzbower et al. (1999) during the welding of high strength steels in naval platforms. These data provide an opportunity for creating a quantitative model for the estimation of weld mechanical properties. A neural network is capable of realizing a great variety of nonlinear relationships. Data are presented to the network in the form of input and output parameters, and the optimum non-linear relationship found by minimizing the difference between the measured value and the predicted value. As in regression analysis, the results then consist of a series of coefficients (called weights) and a specification of the kind of function, which in combination with the weight, relates the inputs to the output. An isoproperty diagram, which correlates the thermal experience and hardenability to weld metal microstructure, can be very useful in the selection of welding consumables for different welding conditions, heat input, and alloy composition. Carbon equivalent diagrams versus Ats/s have been used by Liu et al. (1994) to predict weld properties. The lines on this diagram represent welds with equal property that resulted from different sets of welding conditions and alloy composition. The best carbon equivalent expression to be used with the isoproperty diagrams need to be determined. Acceptable welds can be achieved with a large range of combinations of weld composition and welding parameters, but some combinations are more resilient to process parameter variation. With the evolution of more reliable constitutive equations to predict weld properties, analytical methods are available to select a resilient set of welding parameters and thus reduce the rejection rate. A method based on calculus of variations has been proposed to examine the sensitivity of weld properties to fluctuations in processing and chemical composition (Olson et al. 1991). 2.4.6

Hydrogen Management

Traditionally Hydrogen Assisted Cracking (HAC) is associated with the heat affected zone (HAZ) location of carbon manganese steel welds. The carbon level in the HAZ is fixed at the base material, while low carbon filler metals decrease the occurrence of HAC in the weld metal. With the carbon level of newer high strength steels being much lower than in the conventional steels, the occurrence of HAC in the weld metal is now becoming an issue. Experience would suggest that the weld metal hydrogen problem increases with hardness and that, therefore, a hardness predicator should be useful. However, the situation is complex and there is only a general correlation between weld metal hardness and HAC. In addition to weld metal hardness, a preheat expression based on 10 % cracking in the GBOP test has also been developed by Hart et al. (1986). The primary focus in managing weld metal HAC has been on the development of welding consumables and practice (De Loach et al. 1993, 1999). A crack predication model which is based on tl00 and derived from WIC and modified cruciform tests has been proposed by Wang (1996). It is based on the Yurioka method used to predict HAZ preheat but was adapted to weld metal cracking problems. The transformation temperature of the weld metal can be either higher or lower than that of the base material, depending on the filler and base metal compositions. Wang et al. (1996) report that the location of hydrogen cracks depends to a great extent on the relative martensite start temperatures of the two zones. If the Ms of the weld metal is higher that that of the heat affected zone, hydrogen accumulates in the coarse grained HAZ region. The austenite acts as a barrier to hydrogen movement into the heat affected zone; hydrogen accumulates in the boundary area, causing the heat affected zone problem at this location. However, if the Ms of the weld metal is lower that that of the heat affected zone, there is less hydrogen accumulation in the sensitive boundary region and HAC is more likely in the weld metal. An index based on the weld metal/HAZ difference in martensite start temperature was proposed. Expressions for calculating Ms in both base metal and weld metal are available. The fracture occurs in either the weld metal or HAZ, depending upon which one first reaches the critical combination of sensitive microstructure, residual stress level and hydrogen.

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With the use of higher strength steels, the weld hydrogen content needs to be maintained at even lower content levels and for these new steels the issue is not just the hydrogen content but also hydrogen distribution relative to the weld deposit. New methods to reduce the diffusible hydrogen content include weld metal hydrogen gettering and fluoride additions. Effective control of weld diffusible hydrogen content in higher strength steel weld deposits has been achieved with the use of irreversible hydrogen traps. The benefit of weld metal yttrium as an irreversible hydrogen trap was reported by Maroef et al. (2000) to decrease the diffusible hydrogen content in the weld metal to appreciable levels around 1 to 2 ml of hydrogen per 100 g weld deposit (or 0.89 to 1.8 wt. ppm of hydrogen). The weld metal diffusible hydrogen content is affected by variations in welding parameters. The spray mode, with its fine metal droplets, has the ability to react with oxygen resulting in a large concentration of hydrogen trapping inclusions in the weld metal. The use of selected fluoride additions to welding consumables to promote a plasma chemistry that reduces the weld pool hydrogen content has been demonstrated by Pokhodnya (1996) and Matsunawa et al. (2000). Various fluoride additions have been added to the electrode coveting, resulting in welds in which the diffusible hydrogen contents have been dropped from 5ml/100 gram of Fe to 1.5 ml/100 gram of Fe, a significant reduction. Reductions in diffusible hydrogen levels were measured with additions of A1SiF2, KF, MnF3 and K3A1F6. The hydrogen distribution in a weld becomes significantly more important as the acceptable diffusible hydrogen content decreases. These localized hydrogen contents are probably the cause of the spread in the correlation between the measured diffusible hydrogen contents and the cracking tendencies. A number of methods are being explored to measure hydrogen distributions, including laser induced breakdown spectroscopy (LIBS), hydrogen exposure to silver bromide coatings, laser ablated gas chromatography, and hydrogen changing the electrical conductivity of W30 coatings. A recent optoelectronic technique has been developed to measure the diffusible hydrogen content in less then sixty minutes (Smith et at. 2001) and the hydrogen distribution across the weldment. This W30 oxide coating conductivity device shows the promise of attaching an electrical device to read diffusible hydrogen content directly from the welds of the actual welded structure, rather than just on test coupons. This device provides the potential for alleviating the cost and time expenditure associated with standard diffusible hydrogen testing. Also, the thermoelectric power (Seebeck coefficient) (Olson et al., 2002) has been correlated directly to the diffusible hydrogen content in materials containing transition metals. This correlation supports the development of a simple surface contact (nondestructive) device to measure diffusible hydrogen content in high strength steel weldments. Future work needs to transfer laboratory successes to high strength steel fabricators. 2.4.7

Flux Cored Wires

With the continual efforts to increase productivity, flux cored arc welding has become a major competitive process and these wires are widely used in shipyards especially with automated and robotic processes. The stability of the process and the good adjustment possibilities to specific requirements support this development. Also in Tandem MAG processes with comparably little heat input, metal cored wires offer substantial advantages (Goecke, S. et al. 2001). Leino, K. (1998) analyses the advantages of Flux cored wires in comparison with solid wires with respect to weld seam porosity when welding on zinc silicate primers. Flux cored and metal filled cored welding wires in addition offer the users of specialty alloys the ability to procure acceptable welding consumables for weld repair. With fewer welding consumable manufacturers to produce products for a broader market, flux and metal cored wire products offer an economical way to produce large variation in welding consumable compositions for both large and small specialty needs.

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Recently, flux-cored titanium wires were successfully produced for experimental welding CP-titanium (Liu 2002). The fluoride and chloride-based flux systems resulted in excellent slag coverage and weld pool protection during solidification and cooling from the surrounding atmosphere. The effectiveness of the slag protection resulted in good bead morphologies with no surface or internal defects. This system produced deep penetration, and can be used to achieve full penetration welds. 2.4.8

Aluminium Welding

A general problem with aluminium welding is the degradation of properties in the as welded condition which is also reflected in the Classification Societies rules. Recently attempts are made to minimize this degradation by addition of grain refiners to the weld pool. Grain refinement of aluminium weld metal has been a successful approach to reduce hot cracking and improve mechanical properties (Matsuda, F. et al, 1983, Dvornak, M. J. et al 1989). Grain refinement is achieved by providing heterogeneous nucleation sites in the weld pool. Traditionally, intermetallic aluminide particles, TiA13 and ZrA13, have been used as effective inoculants. Recent studies have shown that both Scandium and TiB TM (5 wt % Ti + 1 wt % B) additions have made significant improvements in weld metal grain refinement. Yushchenko et al. (1995) introduced scandium as a weld metal grain refiner. Weld metal scandium contents greater than 0.55 wt. pct. have achieved grain refinement in commercially pure aluminium with ScA13 serving as the inoculant (Norman, A. F., 1998). Grain refining was experienced in 7108 aluminium welds with concentrations above 0.20 wt. % Sc. With concentrations above 0.25 wt.% Sc, the grain size was smaller than 60 microns and hot cracking was eliminated (Mousavi, 1999). Also Ishchenko et al. (2002) reported increased strength for Sc- and Zr- containing aluminium welds. He draws special attention to the fact that in the as welded condition considerable improvements can be achieved when both, the base metal and the filler wire (MIG - process) are alloyed with Scandium or Zirconium respectively. The resistance against hot cracking increases with Sc-alloyed base and filler material by a factor of up to five. Mousavi (1999) compared especially the effectiveness of Scandium and TiB additions in alloy 7108 welds. 7108 is a common A1-Mg-Zn extrusion alloy often used in welded construction. He reported that TiB addition is a more effective grain refiner for aluminium weld metal than Scandium. It produced finer grains at lower amounts of grain refiner. Hot cracking could be prevented with contents as low as 0.02 wt % Ti with 0.004 wt % B. Only with much larger Scandium contents of greater than 0.20 wt % Sc, one can achieve as fine or finer grain sizes. Scandium additions are also more susceptible to loss of grain refining effectiveness at welding at higher cooling rates. 2.4.9

Welding Fume

Welding consumables, in association with the welding arc, generate welding fumes which may produce an occupational health hazard to the welder and others in the welding workplace. The chemical composition of these particulates depends on the consumable composition and the welding parameters. During welding, these particles are visible because of their quantity, but each particle is in the size range of 0.2 and 1.0 microns. Presently, the primary concern is the metal content of chromium, nickel, manganese and copper in the welding fume. A Danish investigation for lung cancer incidence (Hansen el al. 1996) in over ten thousand metal worker employed during the period of 1964-84 has reported that welders have a significant excess risk. These results are consistent with prior results (Beach et al. 1996, Simonato et al, 1991, Chinn et al. 1995). Although the confounding effects of smoking and exposure to other specific occupational

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carcinogens make it difficult to establish a direct correlation between welding and lung cancer, the ship building industry is experiencing more stringent requirements regarding the management of welding fume. The new limits primarily involve chromium and manganese. The metal content is measured in units of micrograms per cubic meter (AWS 1985). A change is anticipated in a Permissible Exposure Limit (PEL) for hexavalent chromium from the present ceiling level of 100 ~tg/mm3 , as chromates, to an eight-hour time weighted average (TWA) of between 0.5 to 5.0 ~tg/m3. The Threshold Limit Value (TVA) for manganese must be reduced to 200 ~tg/m3, a reduction for 1000 ug/m 3. The Navy Joining Centre (1996) identifies the manufacturing and repair operations, materials and processes that are expected to be impacted by the new limits. The data in current worker exposure levels to nickel and manganese, as well as total and hexavalent chromium are reported. The level of exposure of the shipyard welder depends on the welding process and welding parameters, length of exposure and materials involved with the welding process. The technical and economic impact of the anticipated reduction in hexavalent chromium on U.S. Navy facilities and on public and private shipyards is identified. They also identify future actions that may be required to comply with the recent and anticipated reductions in exposure levels. Future work should require expanded worker experience sampling to produce statistically valid characterization of operations, processes and materials as well as more research and development to minimize hazards during fabrication and repair of ships. These activities should include the development of a long range exposure reduction plan, the evaluation of new, less hazardous base and filler materials, and the evaluation of processes with reduced emissions. In recent investigations (Quimby 1999, Brooks et al. 1997, and Castner et al. 1996, 1998) showed that the rate of fume generation in GMAW increases with evaporation from the molten droplet at the tip of the electrode. The rate of evaporation is controlled largely by the surface temperature of the droplet which in turn, is determined by the heat and pulsed flow in the molten droplet (Mendez et al. 2000). The droplet size increases with the surface temperature. Because the size of droplets is small at a lower temperature, less metal is evaporated, allowing for spray transfer mode or with the application of pulsed current to generate less fume. Methods to reduce fume include the selection of the proper sized weld, the welding process, the welding consumable composition and the welding process parameters (especially the metal transfer mode). With pulsed GMAW less fume is produced than with constant voltage power sources. By pulsing the current at a specific frequency, the total arc energy is reduced decreasing the amount of metal that is vaporized resulting in reduced fume generation. The use of a properly selected wave form control has been demonstrated to make a significant reduction in fume generation. The waveform control with short-circuit transfer has reduced the violent droplet detachment and, thus, spatter and fume. Tightly controlling the current during droplet transfer and allowing the droplet to be pulled into the weld puddle by surface tension force has been demonstrated to result in a significant reduction in spatter (decreased 90 %) and in fume generation (decreased 50 %) compared to conventional short circuit transfer. Harris et al. (2002) gave an excellent summary of the present situation. They describe the reduction of welding fumes through changes in welding equipment, process parameters, consumables and use of local exhaust ventilation (LEV).

3.

TECHNOLOGIES

Coming into the new Millennium, world-class shipbuilders use robotic machines where appropriate and where there is sufficient through put to justify them, such as curved plate forming. They also develop new approaches to ship design and shipbuilding processes as a way to improve productivity and competiveness reported by Lamb (2001). This statement addresses the fact that improving in isolation

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is no longer appropriate and that technologies must be developed and used In the wider perspective of the production chain.

3.1

Cutting

The conventional cutting technologies as used in shipbuilding have reached such a degree of maturity that improvements can only be expected in details. Regarding the increased requirements for parts accuracy when applying accurate manufacturing techniques, one can observe, that the general machine tolerance needs even of very large cutting gantries can be met by the suppliers. It is the detail tolerance which is of interest and which is mainly influenced by the wear and tear of the cutting nozzle during operation. In this area development is going on to receive online information about this influence on the accuracy. Laser cutting has also become a mature technology widely used in the manufacturing industry and it is still expanding in modem IT based production environments. In shipbuilding however it is still limited to smaller yards with relatively thin materials. This opportunity applies in particular if steel is the main building material. It also seems that the advantages of the highly accurate cuts which can be obtained with modem systems are not yet fully appreciated. Olsen (2002) gives a good overview of the present state of the art and a Laser Cutting Working group in IIW has prepared corresponding diagrams. A report about practical applications in a shipyard environment is given by Cahill et al. (2000), which concludes that the high accuracy achieved with laser cutting leads to further downstream cost savings due to increased welding productivity and reduced welding distortions.

3.2

Forming

The term Forming in general describes on the one hand a process adding value in the production chain like shaping profiles and plates, and on the other hand the correcting of unwanted consequences of earlier processes i.e. distortions due to heat input from cutting and welding. For a long period both forming processes have been characterized by manual work based on special experience of workers. Since recently mathematical models of forming and distortion removal processes are being developed in order to be able to rationalize and mechanise these actions. A problem exists however in the relatively small quantities which often make mechanisation uneconomic. As the line heating used for forming curved plates of ship hull requires many man-hours and empirical intuition of skilled shipwrights, therefore the automatic forming method is being widely studied for practical adoption in shipyards. Jang et al. (2000) proposed a new line heating simulator to suggested to acquire the heat information. A thermal elastic-plastic analysis method is employed to predict the heatinduced plate deformation. Heating paths are determined by geometric analysis of a target surface and an initial surface. Also, a surface mesh generation method is developed to link finite element analysis and to determination of heat regions. The simulator includes modules of the preliminary techniques and heat lines can be obtained. Jang's results show the simulator can give heating information with good convergence. Bisgaard Clausen (2000) reported a numerical method for plate forming by line heating. He considers finite element methods to model the behaviour and to predict the heating paths. The finite model is made in ANSYS with linear brick elements. It consists of a fine mesh with six elements through the thickness in the heated region with a transition into a much coarser mesh with only one element in the thickness direction. Considering only plates of identical material and neglecting the edge effects, the problem is reduced only on physical measures as thickness h, torch speed v, and amount of heat Q. To make an equivalent elastic analysis this strain must be applied to an area of the width w p , which is used in the expression to find the average strain contribution:

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1 wP ~"x'Y(Z)--""-~IO cx;y(y'z)dy Two methods for predicting the heating patterns are used, the required bending of the target surface using differential geometry and membrane strains from 'mapping' the target surface onto a flat surface. Liu et al. (2001) studied the process technology for line heat forming with hull fabrication. They proposed a simulation method of line heat forming. It is divided into four stages. The essential local contraction values of the plate deformation, including the contraction area and the maximum contraction length were studied. Based on the processing of data of ship plates, a mathematical model was established, which describes the relation between the local contraction value and the processing parameters of flame bending. According to the technical procedure of line heat forming, the optimal method was applied to determine the parameters in the processing of hull plates of known shapes. The feedback information of ship hull plates was collected to enrich, improve and correct the regression models. Liu also reported that this predicting system was used in building a 61500 t oil tanker where line heat forming was used for all the pillow shaped hull plates. Ishiyama (2001) established an accuracy database for the heating and forming relationship based on parametric experiments and FEM analysis on simple heating lines. It has been confirmed that distribution of inherent strains induced in a plate by flattening the objective curvature, which is defined based on elastic FEM simulation, can be as simulated using the database. This is used for heating process planning for the NC line heating machine with a high frequency induction heater, and facilitates automated thermal forming. In order to clarify the characteristics of line heating process, Tomita et al. (2001) measured the transient 3-dimensional temperature distribution within the gas flame during line heating in detail by a high performance Laser Induced Fluorescence Method measurement system. They found that the relative distribution of gas temperature around the torch is almost the same as that in spot heating. It has also been found that this relative distribution is almost unchanged regardless of the temperature increase in the steel plate. These results show a new hypothesis that the relative distributions of gas temperature and local heat transfer coefficient around the torch remain unchanged and they are almost the same as those in spot heating during line heating process. Regarding the forming of the twisted "T" section longitudinal, Nair et al (2001) and Madhu et al (2001) used the inherent strain method to achieve the desired shape of a longitudinal using elastic FEM analysis. On the basis of this numerical analysis using an inherent strain method, Nair et al (2002) evaluated five probable line heating methods in view of efficiency. They proposed such line heating to be most practical and advantageous where the angle of the moving torch is 45 degree relative to the edge and heating is performed on one side of both web and flange. A simulation of the cold forming process of built-up T sections might be noticed. Mandal (2000) presented that on an average 2.8 % of hull steel weights is accounted for curvilinear welded T sections. The application of cold forming for deep T section is still very much limited. Cold forming of such sections might lead to web/flange weld failure, instability of web, roofing of face plate and reduction in thickness of the web edge. A detailed analysis of all these factors has been carried out to establish a sound basis for this production problem. This simulation approach can be used either to predict the behaviours of welded T sections at the design stage as a design tool so that the required bent shape can be successfully achieved by cold forming. Full scale tests have been carried out and a good agreement between the computed and test data has been observed. Research on laser forming obtained gratifying achievements. Masubuchi et al. (2000) demonstrated the feasibility to predictably laser form a variety of ferrous and non-ferrous metals of different thickness.

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Laser forming provides a method of producing complex shapes in sheet, plate, and tubing without the use of tooling, moulds, or dies. By heating a localized area with a laser beam, it is possible to create stress states that result in predictable deformation. This research program has developed, refined and demonstrated constitutive and empirical, and neural network models to predict deformation as a function of critical parametric variables and established an understanding of the effect of laser forming on some metallurgical properties of materials, the first task of this program involved forming flat plates to one-dimensional shapes, such as, hinge bends in low-carbon steel, high-strength steels and aluminium alloys. The second task of the program expanded the work to investigate three-dimensional configurations. The models were updated, 3-D specimen fabricated and evaluated.

3.3

Joining

Welding and allied processes have for many decades served as the major joining technologies. The room for improvement in the basically traditional arc processes is seen primarily in process control and in sensor developments for online control of automated procedures. The certainly considerable improvements obtained with respect to power sources, consumables and process control systems will nevertheless not alleviate this situation due to inherent limitations caused by the physics of the processes. Low heat input methods like Laser- and Friction Stir welding are promising and cold rolling techniques or even adhesive bonding seem to show significant potentials for shipbuilding applications. In consequence new approaches in research and development must to be of a holistic nature as illustrated below in order to find really new solutions demanded by society and the product users.

environment J _

Figure 3.3.1 Holistic approach to modem welding process development Klaestrup Christensen (2002) in his comprehensive overview about the present state of the art and future development requirements refers in this context to the concept of sustainability which nowadays has obtained a widespread understanding and acceptance. One can observe that the focus of attention has moved from the process itself, over the process as a means for making the product further to satisfying the client and recently to fulfilling a market demand also taking sustainability and consequently society demands into consideration.

3.3.1

Conventional processes

The conventional welding technologies have reached a degree of maturity which leads to improvements in details and in new attempts in process control which is enhanced by advances in power electronics and sensor technology. New developments for specific applications can be developed on the basis of results of recent extensive research work.

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209

Laser-welding

The application of Laser-Welding in shipbuilding has been a matter for research and process development over more than ten years. After initial attempts to use indigenous processes (without additional filler material) it became soon apparent that the tight tolerances required by the process (zero gap situations) could hardly be met in practical applications on large structures. The introduction of cold filler wire did only partly alleviate the situation because much of the expensive Laser energy was then used to melt the filler wire. A combination of MSG and Laser Process showed promising perspectives. Klaestrup Christensen (2001) presented an overview of the various options and their problems. Dilthey et al. (2001) presented an overview of the process variants under consideration. With respect to the problem of deformation due to the thermal effects it is noteworthy that according to their findings hybrid welding does not lead to an increased heat input. Matsunawa et al. (2002) carried out fundamental research with CW CO2-Lasers related the understanding of the physical phenomena of keyhole laser welding. They report about the reasons of keyhole instabilities being associated with localized evaporation and resulting in fluctuating dynamic pressure and complex liquid metal flows. In order to improve the understanding of these basic phenomena additional mathematical modelling and innovative process monitoring seems to be necessary. Kujanp~i~i et al. (2002) report about result of innovative investigations of CO2 and Nd:YAG-Laser welding with filler wire. They demonstrate that with cold wire feeding the geometry of the feed position is an important element in the set of weld parameters. This is due to laser beam reflection on the feed wire. The problem is reduced with hybrid welding where the GMAW heat input melts the feed wire. Hybrid welding is more tolerant to misalignment and gab variation. Roland et al. (2002) gives a comprehensive overview about practical applications in a shipyard, ranging from production of specialized sandwich structures to hybrid welding of butt welds and T-joints over seam length of up to 20 m. The repeat shows especially for large deck assemblies in cruise ship production the considerable increase in productivity due to reduces thermal distortions and high welding speed (3 - 4,5 m/rain with 5 mm plate and 12 KW CO2 - Laser). The paper illustrates again the advantages of hybrid welding due to the reduced accuracy requirements in part fit-up. Further investigations are planned with respect to partial penetration welds for T-joints leading to additional reductions in heat input while again increasing the process speed. Especially the beam transmitting problems (mirror optics) with C02 -Lasers over wider application areas or remote construction sites led to investigations for the use of Nd:YAG lasers in combination (GMAW) processes. Jokinen et al. (2001) report about their investigations and point out that the advantages of the process can be seen in the reduced heat input leading to less deformation whilst the combination reduces the fit up tolerance and increases productivity. The development of YAG Laser applications is also the subject of a number of research projects under the funding of the EU Commission (www.cordis.lu). Walz (2001) applied the hybrid welding process to offshore structures. Compared to laser welding, the combination of the laser and one conventional welding process (e.g. MIG) offers many advantages like wider gap bridge ability, enhanced process stability and higher welding speed. Hybrid welding allows the geometrical design of weld seams on demand. He shows the results of hybrid welding experiments with austenitic stainless steel and super-martensitic 13 wt. % Cr steel. Tensile tests and hardness measurements were carried out. All tests have been performed in the as-welded and in the post weld

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heat treated conditions. The results show a very good behaviour of the hybrid welded samples with seam properties close to the properties of the base material. 3.3.3

Friction Stir Welding

Friction Stir Welding has been a new development which initially aimed at aluminium alloys. Here certain applications are in use for commercial high speed craft building. A good summary is given by Kallee (1999). The main advantage over traditional welding technologies is that joining is achieved below the melting temperature prohibiting some of the deterioration to the material microstructure and joint mechanical properties often seen in traditional welding and thus adding new dimensions to design and component optimisation. A very important aspect is the absence of problems related to the working environment as being present with traditional arc welding processes, namely air pollution (in particular Ozone) and ultraviolet light. Tsujino (2001) investigated optimum condition for the friction welding of type 2017 aluminium similar alloy by statistic analysis. The results are as follows. (1) In general, selection of the optimum friction welding condition for similar materials are easy. However, it was not always the case for type 2017 aluminium alloys. For the optimum friction welding condition of this material, it is necessary to add relatively a larger upset pressure after obtaining a certain friction heat. (2) Joint efficiencies obtained under the optimum friction welding condition show near normal distribution with a small dispersion, and the large shape parameter (m value) of a Weibull. (3) The m value calculated on the small number of date can be substituted for m value on the 30 data. Therefore, the m value is useful for the practical use in the factory for simply assuming the propriety of the friction welding condition. Meyer et al. (2000) presented results of a joint Industry project on Friction Hydro Pillar Processing (FHPP) which has been developed for repairing thick steels under water. This application aims at maintenance problems in the offshore industry. It is especially noteworthy to observe the very low hardness values across a repair weld, which can be attributed to the fact that friction welding is a solid phase process with relatively small temperature gradients. 3.3.4

Adhesive Bonding

The use of adhesive bonding as a means of assembling and joining structural components is now well established in the marine industry, particularly with regard to FRP and hybrid metal-composite structures. Some of the knowledge can be transferred from one sector to another; for example, the manner of joining components in buses and aircraft would have lessons for the marine industry and vice-versa. Burchardt et al. (2000) have published a set of simple, practical guidelines coveting a series of issues including simple design formulations, selection of adhesive type, disassembly and repairs, quality assurance, long term serviceability, environmental factors etc. There are several types of adhesives suitable for different adherents and different applications. When dealing with FRP materials when needing to be bonded to metallic materials, the experts suggest three principal types, namely polyurethanes (PUs), acrylics and epoxys. Table 3.3.1 summarises the principal characteristics of structural adhesives. The marine applications of adhesives have led to a re-appraisal of the production technology and led to the concepts of modular construction in boats (Strand 2002, Prince 2002). The concept, in one instance, involves building the boat hull within a traditional mould, with cross-stiffening members being fabricated separately. After both the parts have cured, the stiffening structure is then placed in the hull plating and joined together by adhesives. This, it is argued saves production time and hence leads to lower costs.

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The concept of modularisation can be extended to kit construction, where a boat is idealised into a series of standardised structural parts. These structural components are fabricated in factory conditions and shipped in containers to boatyards, where they are assembled using both adhesive bonding and, in limited locations, mechanical fastenings using bolted connections. In such cases, it is imperative that the adhesive joints are most carefully designed and sized. The assessment of such joints needs to be done using theoretical approaches and validated by selective experimental modelling (Cossich 2000).

TABLE 3.3.1 PRINCIPAL CHARACTERISTICSOF STRUCTURALADHESIVES

Type , Main characteristics ] Principal advantages PU Low modulus Very simple to use. Hot One part Very low strength metal variants very convenient on suitably sized components. Fills large gaps. No mixing. PU Very low to medium Fast curing possible. Very Two modulus good application part Very low to medium characteristics. Fills large strength gaps.

Acrylic Pseudo One part

Acr3 Acrylic Two part

Epoxy One part

Epoxy Two part

, Principal disadvantages Sensitive to moisture. Not true structural adhesives. Slow curing. Must be applied to non-metal surface . for long term durability Sensitive to moisture. Often requires heating to achieve acceptable production times. Must be applied on a nonmetallic surface for long term durability. Must be mixed. Medium modulus Very fast curing. Easy to Needs good fit and narrow Medium strength apply. Extremely durable, gaps to function effectively. Bonds metals particularly Best below 2-3 mm. well. A true structural adhesive. No mixing. Medium modulus Fast curing. Easy to apply. Must be mixed. Needs good Medium strength Benefit of delayed action fit and narrow gaps. Best cure. Extremely durable, below 2-3 mm. Copes well with light contamination. A true structural adhesive. High modulus Fast curing. Easy to apply. Needs to be heat cured. Very high strength Extremely durable with Needs good fit and narrow robust all round gaps. Best below 2-3 mm. performance. No mixing. Medium to high Easy to apply. Durable. Can Must be mixed. Needs good modulus be speeded by fit and narrow gaps. Best Medium to high warming/heating. True below 2-3 mm. Slow curing. strength . structural adhesive. .

Kim et al. (2001) describe a recent novel application of adhesive bonding has been to the ensuring joints within composite propeller shafts, see This form of connections is argued to be advantageous in some circumstances from the viewpoints of fatigue and maintenance, especially with the adherents being composite materials where the matching of adhesive-adherent properties is better achieved. For larger, naval ships there is active consideration in the use of FRP composite materials for superstructure construction. Joints between the FRP superstructure and the metallic hull have been studied by Wright et al. (2000) and Clifford et al. (2002). Both studies looked at different aspects of the mechanical response of a prototype joint between an FRP superstructure and a metallic hull produced using the resin infusion process. The FRP-steel interface was shown to be critical, indicating the importance of

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the correct surface preparation of the two adherents. The modelling also identified the phenomenological causes for damage inception and propagation. There were some indicators for improved designs based on the test and analysis results. 3.4

Surface Treatment / Coating

Ultra high pressurized water blasting, pressure of up to 196 MPa, is introduced by IHI to surface treatment of hull-blocks before painting (SNAJ, 2001). Traditionally cleaning has been conducted by sand blasting or power tool, however, these treatments produced dust such as sand and rust, which required much man-hour to remove them. By using water blasting, rust is washed away so easily that cleaning is rationalized. In addition to that, water blasting can prevent from rusting of surface because shop-primer is not damaged due to water blasting. TABLE 3.4.1 SURFACETREATMENTOPTIONSFOR ADHESIVEBONDING

Material Anodised aluminium Bare aluminium Steel Wood FRP composites

Abrading/cleaning

Degreasing/activating

Priming

X

X

X

X

X

X

X

X

X

X

X X

Surface preparation is also important in adhesive bonding, see Burchardt et.al. (2000). A critical factor in this regard is the condition of the substrate - the surface of the material being joined. Since adhesion takes place only at the interface between the structural plate/stiffener part and the adhesive, it is evident that surface preparation has a crucial bearing on the quality of the structural bond. The options for surface preparation and treatment are many and varied. They include simple cleaning of the surface, mechanical abrasion, the chemical alteration of the surface by pickling or phosphatising, thermal process such as flame treatment, as well as specialised physical-chemical techniques. In addition, there are various kinds of paint systems and coatings, including lacquers and primers, which can be used to provide a satisfactory substrate for adhesives. The table below lists typical surface treatment options for a range of substrates/adherents.

4.

FABRICATION METHODS

The term Fabrication Methods describes the use of Technologies in the actual enterprise environment. The advancements of IT/IC technologies play the important part in combining technologies to methods. Some of the more recent technology developments like Laser or Friction Stir Welding could only be applied to shop floor level together with mechanised systems. An interesting vision of future yard technology combinations has been presented by Wilckens (2001) where especially the multidisciplinary approach was addressed. One can also observe that the use of robots is increasingly connected to workspace improvements and to environmental aspects rather than only to productivity gains. 4.1

Mechanisation/Robots

A good analysis of problems related to the introduction of robots in smaller shipyards is given by Hengst et al. (2000). An initial two phase research period resulted in a definition project for mobile portable welding robot application based on off-line programming. As an interesting side effect the research work lead to a far more detailed and process oriented production planning concept.

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One of the most common operations in ship building is the alignment of a tool with respect to an assembly part. If the tool is mounted onto a robot manipulator then the problem is to properly control the latter. Among various sensors to be used in conjunction with robot control, vision has a number of advantages: it is low cost, fast and covers a wide field of view. Lamiroy et al. describe the development and prototype tests. Ames et al. (2001) introduced fully automated planning and generation of control programs (AUTOGEN) for robotic welding of the ship structure. The software automatically derives torch and robot motion from the geometry of parts, rather than attempting to fit pre-designed trajectories to part geometry. The software is designed with a constant and mathematically coherent core with well-defined and open portals for separable code modules that tailor the system to the requirements and constraints of various shipyards. The openness of the system should make it attractive in other related industries, such as railcar building and bridge fabrication. The system may also be applicable to other processes that require similar tool motion, such as laser welding, manipulating routers and deburring tools over large complex parts in the aircraft manufacturing industry, or delivering adhesives and similar fluids. Moon (2002) describes a system where conventional SAW units with five torches and FCAW units using six torches are combined with various automatic features such as vision sensors and adaptive control systems. The combinations developed are applied to the shipyard and give the outstandingly high productivity and quality. Shin et a/.(2000) have conducted extensive research and experimental work related to the automation of the plate bending process by line heating. The system finally installed consists of hardware- (automated machine) and software- (control system) components. They point out that only by concurrently developing hard and software one can expect the desired progress with respect to quality and productivity gains and including user friendliness. Another application of PC system to shipbuilding is conducted by Yamato et al. (2001). They investigated the application of a wearable PC system to the work measurement in shipbuilding. They made the work content input system that the worker input by using the voice recognition, the speech synthesis and the positional measurement system of the worker that used a magnetic frequency. They were processed with the computer that the worker wore. The work analysis of the current state could be done at once for the acquired data, and effectiveness was found as a prototype system of the work measurement in shipbuilding. The laser hybrid welding process supported by various sensors and elaborate control systems, which involves CO2 laser and MIG, is applied to the welding of the panels of a cruise ship (Meyer Shipyard, 2002). This innovative technology is more precise, faster and more efficient and achieves a level of automation never before imagined. It makes it possible to process very large panels and to reduce distortion thanks to its relatively low level of heat transfer. A new type painting robot which operates in a complex system, such as a double hull block has been developed by Miyawaki et al. (2001) with special attention to avoid collisions between the paint supply hose and the structure. An in-line supply system has been used to address this problem. Of special importance is the sealing performance of the swivel joint utilized by the in-line mechanism. This type of in-line robot was applied to trial testing for painting of a ship hull-block; the results of which showed the effectiveness and durability of this mechanism. 4.2

Accuracy in Production

Accuracy control in ship building has become a well proven standard in production processes worldwide. In this context Heinemann et al (1998) point out that accuracy is not meant as a specific product performance quality. The aim is to produce parts and modules with a predefined precision

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for minimising the assembling efforts. The importance of accuracy control is seen in raising the productivity in a defined production environment. Therefore the aim is to be as accurate as necessary instead of being as accurate as possible. Thus one can state that accuracy control can not be generalised, but is depending upon the production processes at the specific shipyard. The potential for productivity gains by accuracy control are seen in different areas: 9 Avoidance of unnecessary processes, e.g. accurate assembly cut or transfer of section contours to the connecting modules 9 Optimisation of processes by using all possibilities of more efficient machinery, e.g. for edge preparation of plates, profile cutting machines, welding of profiles and beams to the end on welding portals Zorn et al. (1999) report that it is possible to increase the productivity by up to 30% if the production logistics is taken into account. Even larger economic benefits can be achieved by stabilising of the production process. This results in enhanced production planning security or quality and thus can lead to an increased operational capacity. For example the reduction of the production time on the slipway or in the building dock increases the possible number of deliveries per year thereby increasing the productivity of the whole shipyard. The major challenge in production accuracy is the control of deformations of components resulting from welding stresses. The goal is the prediction of these deformation processes which makes it necessary to understand the different parameters leading to shrinkage due to the welding energy. Equally important is to quantify these influences of the different parameters. This will allow to take appropriate advance measures to reach a higher accuracy in the production of modules. The major parameters are: 9 Structure of the section 9 Connecting seams between sections 9 Building position Structural influences are taken into account by two approaches: 9 The Shrinkage Manager (Heinemann et al. 1998, Nikolay et al. 1999) is based on differentiating between intermediate production steps and assembling methods of the different modules. 9 Miebach et al. (2000) describes an FE-based method breaking down the process into basic elementary cases. In these the specific boundary conditions and process parameters and the specific welding energy at the seams are used. The resulting stresses in the FE-model are used for further processing steps. Several investigations on the impact of welding stresses under well defined conditions are available today and using the results of such methods the designer is able to take predicted deformations into account. The use of these results is restricted to cases where the actual conditions in the production process are comparable to the base cases, e.g. during the welding process. With the different production processes and their individual possibilities and limitations in mind the ship's body can be divided into three main areas (Nikolay et al., 2000): 9 Panels and volume sections with mainly 2D-boundaries 9 Panels and volume sections with mainly curved 3D-boundaries 9 Complex volume sections of the fore and aft body of a ship

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A high rate of repetitions for the accuracy of the joining geometry is possible for sections consisting of flat panels. This is possible throughout all production levels. The reason for this is mainly seen in the increased rate of mechanised welding leading to a reproducible impact of welding energy. Zorn et al. (2000) investigated the possibilities of transferring experiences gained from the development of the Shrinkage Manager with 2D-panels onto the more complex sections of the second main area. Possibilities for the transfer of the methodology to more complex blocks are seen on the basis of good results with intermediate production steps. Difficulties exist in transferring the required geometry information from CAD-systems into the production process for these more complex blocks and in addition the lesser rate of reproducible detail joint geometries. Wanner (2002) points out the possibilities of accuracy control as they can be seen today for the very complex blocks in the aft and fore body of a ship. For these areas accuracy control using the Shrinkage Manager concept is not possible. The main reasons are: 9 9 9 9

The variation of the weld gaps width is much larger High internal material stresses of the complex warped plates Highly different basic geometries The high percentage of manual welding in these areas results in a considerable influence of the individual welder's skills and to varying heat induced stresses.

One solution for overcoming these drawbacks is the breakdown of complex sections into simpler modules with a high degree of stiffness which again allow the application of the Shrinkage Manager. Nevertheless increased needs for further R&D are clearly visible for increasing the knowledge and prediction of the deformation processes in the structures. The final tolerances defined for the complete assembly and the nested structure defines the required tolerances for the intermediate modules. Production methods and the quality of the machines used govern the tolerances in the production steps. Different alternatives of the detailed breakdown of the main structure and the resultant tolerance requirements suggest potentials for alternative and possibly more economic production methods. This question is investigated systematically in an Office of Naval Research (ONR) funded Small Business Innovation Research (SBIR) project from an American engineering and consulting company in close cooperation with the University of Michigan and several American shipyards. One of the important problems of the shipbuilding industry, which is not yet solved completely, is the prediction of welding deformations of hull structures during the manufacturing processes. The accurate prediction of post-welding distortions is of primary importance for those actual hull assembly methods which do not rely on excessive, redundant material left on edges to be welded and which has to be removed prior to welding during final assembly. Jang et al. (2002) present an efficient method for predicting weld deformation of complicated structures based on the inherent strain theory combined with the finite element method. The inherent strain is defined as the residual plastic strain after the welding heat cycle, and is determined by the highest temperature and the degree of restraint. In order to calculate precisely the inherent strain in real structures, it should be considered that the degree of restraint changes differently according to different fabrication stages. A simulation of a stiffened plate confirmed the applicability of this method to simple ship hull units. As the welding distortion causes re-work, adds cost and may affect strength, Bruce et al. (2001) address the complexity of the problem using a neural network model as a predictor and compare the results with experimental distortion data. The sensitivity of results to different variables is reported. The economical potential in detailed process oriented production methods innovation is substantial as De Payrebrune et al. (20029 prove in the description of a novel panel production line arrangement. This new production line produces panels based on edge-prepared and accurately cut single steel

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sheets in predefined minimal tolerances. This led to a considerable reduction of the installed equipment by omitting one complete intermediate process step. The design and the process structure have been optimised based on in-depth numerical simulation. The gain in productivity was achieved with comparably conventional technology by addressing specifically and detailed accuracy controlled processes in a very short production line. This led to a reduction of the required personnel and a doubling of the production rate. The increasing application of thin steel sheets in shipbuilding raised the requirements for accuracy control. Thin steel sheets can not compensate the welding stresses which results in buckling due to tension. Due to unknown inherent plate stresses it is impossible to predict the necessary energy for reducing the buckling after welding. Thus the heat induced stresses in such components and the resulting deformations become equally unpredictable. Some shipyards tend to solve the problem by leaving the bottom-up approach of defining low tolerances. They apply direct correction methods, e.g. fitting cuts in the final section. The fits are transferred from one section to the next using computer based measurement methods. This methodology requires some extra work, but by decoupling the assembly process and the fitting cut, most of the productivity potentials are still available. New methods for the treatment of buckling and similar deformations can reduce the impact on the accuracy control. The availability of high focus plasma, defocused lasers or even pointed light can reduce the heat input required for post weld straightening and thereby reduce the fabrication induced stresses in the panels. Such new concepts could b less critical to the bottom up approach for final tolerances of blocks and subassemblies. Goldan et al. (2002) describe a "New Ship Repair Process" developed in the framework of the research project AMORES. The developed methodology is based on entering the ship hull geometry into the CAD-system using photogrammetry. The accuracy reaches deviations from the true geometry as low as 1 mm. The system has been verified for different CAD-systems with different ships. The main area of application is the repair and reconstruction of ships if no or a very limited amount of data is available. The necessary time for conventional methods is reduced quite dramatically. Additionally the conditions for performing a correct damage assessment become more flexible.

4.3

FRP! Composites

A thorough, state of the art, generic study outlining various FRP fabrication processes can be found in the work of M~nson et.al. (2000), which outlines both the fundamental issues coveting different processing techniques and the overall modelling methodologies that could be employed. Progress in FRP fabrication could be considered under three broad categories, namely practical production research, studies of theoretical parameters and modelling of the infusion process and process control features. On the practical front, most ship and boat building continues to rely principally on spray and hand layup techniques though closed mould techniques are finding increasing usage due to emergence of stricter environmental pollution controls and to meet demands of better mechanical properties in the resulting structure. Spray lay-up techniques, often thought of as the 'dirty' end of the reinforced plastics industry, are benefiting from the new LSE resin formulations and better spray up equipment, leading to more 'greener' boats, see Jacob (2002). The environmental styrene emission limits are probably the most pressing reason for manufacturers and fabricators turning towards closed mould processing techniques. Marsh (2002) discusses several variations of this process variously termed as Vacuum Assisted Resin Moulding (VARIM), Vacuum Assisted Resin Transfer Moulding (VARTM), Seaman Composites Resin Infusion Moulding Process (SCRIMP) or Resin Infusion Flexible Tool (RIFF), etc. Stewart

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(2002) reviews the benefits of the US-derived SCRIMP process while Swiatecki (2002) investigates the techniques used to build Danish Coast Rescue Service craft. A comprehensive discussion of the infusion process, of a practical nature, can be found in Cocquyt (2001). The paper includes a coverage of the process parameters affecting the fabrication process as well as the background sources for information such as the original patents that affect the process development, tracing the chronology from the 1940's to the present. The topics covered in the report include the basic principles, resin delivery systems, equipment and training issues and surface coatings (or cosmetics). Theoretical modelling of the resin transfer or infusion moulding process continues to be the subject of several academic research studies. The basic formulation is still based on the 19th century Darcy's Law, derived from ground water seepage principles (D'Arcy, 1856). One of the key features in this model, for both RTM and VARIM, is the permeability of the reinforcement fabric, which governs resin flow rate. Xing (2001) proposed a generalised analytical model for the estimation of permeabilities under generalised conditions of variable flow rates and pressures. A further unique feature of this work was to include in his model the effects of gravity on the resin flow through the fabric preform. Nedonov and Advani (2000) reported on the theoretical modelling of the mould filling in sandwich construction. They used finite element techniques to simulate sandwich structure manufacture. Dong et al. (2001) recognised the need to consider the draping characteristics of different fabrics and their effect on the resin flow during the infusion process and set out a numerical procedure for assessing this feature. Kang et al. (2000) developed a control volume finite element model to study the RTM process with multiple gate injection. They confirmed the validity of their model through a series of experiments with good correlation being achieved. Kang et al. (2001) developed an analytical model to assess the resin flow through deformable fabric performs. Interestingly their application is the simulation of the moulding of the section of a three-dimensional ship hull form. Lin et al. (2000) recognised the multivariant nature of the process modelling and the consequential need for developing an optimisation tool for identifying the best combination of process and product variables to yield optimum performance. The tool incorporates discrete search, random search and genetic algorithm strategies for searching for the optimum. In a ship or boat building yard, process control is essential to ensure consistent delivery of quality. Gibson (2002) discusses several practical approaches and systematic recording and assurance procedures that enable this objective. The discussion covers bought-in materials, yard storage controls, optimality of checklists, pressure points for control, inspection strategies, etc. Barry (1999) argues the emergence of accuracy as an ever important feature of FRP production and the use of robots in ensuring of tight dimensional tolerances in commercial production. Robotic routers are now routine equipment in vast segments of the boatbuilding industry and they are leading to savings in fabrication costs and increases in production quality.

5.

MODELLING AND IT ASPECTS

5.1

Production aspects in design

An integrated software system was developed for the manufacturing information of the roll bending process for ship hull pieces (Shin et al, 2002). To this end, the information flow in the process was studied and designed using the object-oriented method. Separate program modules, obtained by introducing new approaches or modifying existing methods, were integrated to yield a software system that can provide all the manufacturing information required for the roll bending of hull pieces of general shape. The information includes roll orientations, roll lines, roll region, and the centre roller movement. Two examples are presented to demonstrate the performance of the developed software.

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Farkas et al. (1998) present an illustrative example of the economic design dealing with the minimum cost design of uniaxially compressed, longitudinally stiffened plates. The fiat or L-shaped stiffeners are welded to the base plate by two longitudinal fillet welds using SMAW (shielded metal arc welding) or SAW ( submerged arc welding) technology. In the cost function the material and welding costs are formulated using the COSTCOMP software for welding times. The stability constraints are as follows: (1) global buckling constraint according to the design rules of the American Petroleum Institute (API) considering the effect of initial imperfections and residual welding stresses; (2) local buckling constraints using the limiting plate slendemess concept for flat stiffeners according to Eurocode 3 and for L-shaped stiffeners according to German DASt Richtlinie 016; (3) limiting slenderness for the whole plate according to API rules which makes it superfluous to consider the effective width; (4) the required moment of inertia of stiffeners is prescribed also according to API rules. In numerical examples the thickness of the base plate, the number and dimensions of stiffeners are sought which minimise the cost function and fulfil the design constrains. For the constrained function minimisation the Rosenbrock's mathematical programming method is applied. The optimum continuous values are rounded using a complementary computer method. The comparisons of the optimum solutions show the following: (1) when the plates with flat stiffeners are designed without the distortion constraint, the initial imperfections can be too large. To fulfil the distortion constrain the flat stiffeners should be larger, but this increases the cost; (2) the plates with L-stiffeners are more economic than the plates with fiat stiffeners. The distortion constraint does not determine the number and dimensions of Lstiffeners; (3) the cost of both types of plates is less for SAW than that for SMAW technology. Lim (2002) introduced the production planning system for shipbuilding. His system is integrated into the CAD system and gives the production information such as optimal work amount in each stage of fabrication, at the design stage. By comparing design alternatives optimum solutions can be identified. Related to production aspects in design for FRP Composites, Pfund (2001) states that FRP ship and boat builders change laminate schedules for a variety of reasons: 9 to minimise the number of fabric reinforcement types in stock; 9 to simplify worker skills; 9 to comply with environmental regulations; and 9 to make a lighter, stronger and better product. Pfund (2001) argues that a further, important objective is minimising labour costs or production time. This can be achieved by using non-linear programming techniques, for instance. Walker et al. (2002) adapted the traditional optimisation techniques by incorporating a sequential solution procedure to account simultaneously for a number of continuous design variables such as ply angles and for a range of discrete design variables such as material combinations and numbers of plies. They demonstrated their methodology for optimal design of sandwich panels. Bader (2002) conducted a study to estimate the costs of manufacture of a simple component in a number of different composite materials and made by different manufacturing techniques. The materials and routes selected span the range of composites from those appropriate for general engineering applications to aerospace structures. A simple methodology is introduced for a comparison on the basis of cost-performance efficiency. An interesting demonstration, with possible lessons for marine applications, concluded that more economic solutions may often be realised by choice of 'expensive' carbon rather than the 'cheaper' glass as the reinforcing fibre. Traditional analytical tools used to increase process efficiency and reduce costs have tended to focus on the physical processes involved in each stage of the FRP component manufacture. Kendall et.al. (1998) have argued for an alternative approach for increasing efficiency by viewing the component manufacturing process as a whole and optimising the integration of each processing stage, such that throughput is optimised for a given manufacturing cost. They describe the application of discrete event

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simulation in combination with technical cost modelling to the production of a resin infusion moulded structural component. 5.2

Simulation

Virtual or digital shipbuilding has become an important issue and is regarded as a next-generation key technology in shipbuilding. The virtual manufacturing using the simulation technology was adopted earlier by other key industries such as automotive, aerospace, nuclear industries. Many companies leading these industries are realizing tremendous cost savings and increased quality. The application of Simulation Technologies covers a wide field like total factory planning, detail process analysis, ergonomic aspects or robot programming optimisation. The latter has been described in an extensive paper by Fridenfalk et al. (2002). The work deals with planning of Robot welding processes in the dock assembly area and it concludes that major parts of sensor guided control algorithms should be performed by simulation. It is demonstrated that time and cost can be saved when simulation is included in the methodology. The digital shipbuilding is the new concept combining the traditional shipbuilding with IT and an implementation of the shipbuilding process including design and manufacturing as a computer model using VR technology. It consists of three components, PPR model (Product for 3D-design, Process for manufacturing process and Resources for facilities) supported by UML (Unified Modelling Language) information model. On the basis of these three components, it builds a virtual shipyard and simulates the entire shipbuilding process in the virtual shipyard. This technology is being actively researched by several research institutes and companies. Shin et al (2001) presented the concept of a digital manufacturing and its application to shipbuilding and some digital models for forming-shop of shipyard. Park (2002) performed the research on modelling and simulation of assembly in shipbuilding. He analyzed the detail assembly process on the basis of UML model and implemented three dimensional simulation. Scotton (1999) introduced virtual manufacturing for shipbuilding and several examples which are being carried out in the US companies. Recently Sasaki et al. (2002) presented a 3-D digital mock-up system for work strategy planning. The system enables the visualisation and simulation of process planning for hull blocks as well as interference checks I the assembly stage and an evaluation of production stage workability. Okumoto (2002) introduced the concept of "Ergo factory" to assess the workability and safety of welding works by means of bio-mechanics and ergonomics, together with computer simulation technique, which combines productivity and humanity well. He shows an example of human task simulation by the virtual human model on a computer; the static strength prediction on human body and low back compression analysis to prevent low back pain, industrial disease of welders. He summarized that the torque and/or forces on each segment and joint in a human body were calculated in accordance with the various working postures of welding, and that the welder would suffer from severe torque on the knee and also heavy compression force on low back at the usual squat posture (flat position), and the standing position of the welder would be better for the health comparing with the squat posture. 5.3

Production Logistics

The requirements for short delivery time and reduced cycle are highly increased. To respond to these requirements, it is inevitable to build several ships with the limited facilities in the limited space of shipyard at the same time. This limitation needs more precise production planning and logistics of shipbuilding. Several meaningful research projects were carried out recently. One of them is TOLOS (Total Logistics System) which was developed by Hwang (2000). This system monitors and controls deployment and flow of material, and finds optimal logistics planning and scheduling by simulation. Hultin et al. (1999) and Hwang (2001) defined the requirements of production planning and

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management and suggested an integrated IT system and simulation technology for supporting the advanced shipbuilding. Shin et al. (2002) showed the possibilities for improving the utilization of the facilities and the output of the forming-shop of a shipyard by means of logistics simulation using the discrete system-based simulation software, QUEST. A special case arises when the requirement exists to optimize cooperation processes between a number of yards working on common larger projects. Bentin et al. address this aspect and underline the necessity for the implementation of a suitable supply chain management system. The paper describes a solution based on a master planning system with high requirements for user-friendly system interfaces.

6.

FABRICATION IMPERFECTIONS

6.1

Distortion and Residual Stresses

Takeda (2002) proposed an estimation method for welding deformation of curved shell plates using the inherent strain method. In the analysis, contact condition between the plate and the positioning jig is properly modelled and the rigid body motion of the block being welded is considered. The proposed prediction method including the above effects is confirmed to be successful by showing good agreement between the calculated deformations and the measured ones of large size specimen. Another prediction method of welding deformation for curved shell plates is proposed by Matsuoka et al (1999). They showed the results of experiments, in which weld deformations were measured in two models and then compared with the calculated values using the proposed method. Both experimental and calculated values are in good agreement, and the availability of the methods is confirmed. Zhu et al. (2002) carried out detailed three-dimensional nonlinear thermal and thermo-mechanical analysis using the finite element welding simulation code WELDSIM. The objective was to investigate the effect of each temperature-dependent material property on the transient temperature, residual stress and distortion in a computational simulation of the welding process. Welding of an aluminium plate using three sets of material properties, namely, properties that are functions of temperature, room temperature values, and average values over the entire temperature history in welding, are considered in the simulation. Results show that the thermal conductivity has certain effects on the distribution of transient temperature fields during welding. Also the yield stress and Young's Modulus have significant and small effects, respectively, on the residual stress and the distortion, after welding. Finally it was shown that except for the yield stress, the use of material properties at room temperature gives reasonable predictions for the transient temperature fields, residual stress and distortion. Since high temperature material properties are either difficult to obtain or do not exist for many materials, an engineering approach is proposed based on the results in this study. This approach suggests to use simplified properties constituted by a piecewise linear function with temperature for the yield stress and constant room-temperature values of all other properties for the computational weld simulation. Procedures for the structural integrity assessment of welded components have provided limited guidance on the treatment of residual stresses due to insufficient information on the residual stress distributions in welded joints and uncertainties in the behaviour of residual stress distributions under applied loading. However, the assumptions made about the residual stresses can have a very significant effect on the structural integrity assessment and improved guidance on this subject is required. The EC funded project Structural Integrity Assessment Procedures for European Industry, SINTAP, has provided the opportunity to perform an extensive investigation on residual stresses and develop further the BS 7910 and Nuclear Electric R6 procedures (Stacey et a/.,1999). It has entailed an extensive literature review of residual stresses in the principal weld geometries (including plate butt, pipe butt, pipe to plate, T-butt and tubular welded joints), experimental and numerical investigations and the development and validation of procedures.

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Measurement of residual stresses is a major problem when destructive methods are not applicable. The problem arises already when rolled plates especially from TMCP processes or in quenched and tempered conditions are to be used as in light weight high speed craft or in modem RO/RO vessels. Kandil et al. (2001) published a very comprehensive state of the art report about the methods and possibilities to measure residual stresses in materials and in fabricated components. Dimensional differences caused by weld deformation often occur when multipass welding is used to join thick plates. The root openings must be controlled to below 6 mm for butt-joint welds in plates under 20 mm, but a root opening of 30 mm can develop in reality. In that case, the gap 24 mm out of tolerance is generally closed by built-up welding. However, no accumulated data and standards have been developed regarding these built-up welded parts. Jang et al. (2001) performed a study to accumulate data on the behaviour of built-up parts and to verify the effects of root opening on the mechanical properties of the welded parts. The finite element common code (MARC) was used to study the effects of 6 and 30 mm root openings on residual stress and weld deformation in multipass welding. The experimental and analytical results show good correlation. Weld induced residual stresses and the deformation distribution of the specimen with the 30-mm root opening appeared to be asymmetric and the magnitude was larger than those of root opening specimen within the tolerance. Many authors have discussed the effect of residual stresses due to welding on the life properties of welds. As it has been demonstrated that usual welds have residual stresses at the level of yield strength of the base material, many codes or standard assume this point for the design of welded structures. Because residual stresses are linked to thermal cycles and metallurgical transformation in the HAZ as well as in the weld metal, certain authors have demonstrated that residual stresses can be lower than expected. Mabelly et al. (2001) present investigations carried out to understand the effect of the combination of different filler materials and weld metal ranging from HY80 to HY130 with matching and undermatching combinations. A combination with martensitic stainless steels has also been performed. According to these results, it can be shown that the chemical composition of the weld metal can lead to a complete change of the restrain cycles. When the weld was sufficiently alloyed to have martensite or lower bainite in weld metal, residual stresses have been lower. As the tests were performed on steels already used for the fabrication of JACK-Ups, a comparison of the lab results to industrial ones was made. For this purpose, a rack delivered for the fabrication of Jack-Up legs has been assessed for residual stresses. As expected, the residual stresses measured with drilling hole method were far lower than the yield strengths of weld metal or base material (E 690 type / SUPERELSO 702). These results confirmed the tendencies established in laboratory tests. It was demonstrated that use properties of high-strength steel, especially when sufficiently alloyed, are to be higher than expected because the residual stresses are much lower than the assumed level. Wen et al. (2001) dealt with a multi-wire submerged arc welding (SAW) process which was modelled using a general purpose finite element package ABAQUS. The paper explains the welding process and its application in thick wall line pipe manufacturing. Corresponding 2D and 3D finite element (FE) models of the SAW process are presented. FE analyses were carried out to investigate the heat transfer characteristics in the fusion zone and the HAZ during welding. The effect of process parameters and weldment geometry were evaluated with and without considering residual stresses and strains induced from the forming process prior to welding. Comparisons of FE predictions with experimental results were presented. It was shown that the geometrical distortion and residual stresses and strains caused by element analysis can be applied to better understand the SAW process and hence be a useful tool for future process development and control with the view of optimising product properties. Distortion is a potential problem with all welded fabrications. To a large extent, industrial control of weld induced distortion is achieved by reliance on past experience, simple empirical formulae or rectification procedures. Rectification can be costly, whilst in large complex structures, empirical

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formulae are rarely applicable. Classical approaches to the modelling of welding distortion and residual stress, even when accurate, have not been readily adapted within the industry. The time and costs associated with running such models appear to be the main reasons. Nevertheless, the use of computer simulation techniques has the potential to significantly reduce the cost of welded fabrications by allowing for predictions to be made long before actual production on shop floor level. Therefore, computer models aiming at predicting such welding effects not only need to be accurate, but must also be affordable and capable of making predictions within industrial time frames. Bachorski et al. (1999) present one such strategy. The Shrinkage Volume Method is a linear elastic finite-element modelling technique that has been developed to predict post-weld distortion. By assuming that the linear thermal contraction of a nominal shrinkage volume is the main driving force for distortion, the need to determine the transient temperature field and microstructural changes is eliminated. In so doing, the computation times are reduced significantly and the use of linear elastic finite-element methods permits large, highly complex welded structures to be modelled within a reasonable time frame. Verification of the modelled results was carried out by an experimental program that investigated the distortion of plain carbon steel plates having differing vee-butt preparations. The initial models assuming the edge preparation to be representative for the overall shrinkage volume, were in reasonable agreement with the experimentally determined distortion values. Further improvements were made by using a thermal model for the definition of the effective weld shrinkage volume. The increasing use of thin steel plates has given rise to massive deformation problems and the necessity for straightening in on-line processes. Kuo et a/.(2001) explain the results of experiments and predict techniques for the control of deformation in thin panels. For the prediction of deformation, a method to estimate input heating of laser and torch is introduced. The proposed prediction method can be used during the forming process. Laser beam (LB) welding is increasingly being used in welding of structural steels. The associated thermal cycles are generally much faster than those involved in conventional arc welding processes, leading to a rather small weld zone, usually exhibiting a high hardness for C-Mn structural steels due to martensite formation. It is rather difficult to determine the tensile properties of a laser weld joint area due to the small size of the fusion zone. Complete information on the tensile and fracture toughness properties of the fusion zone is essential for prequalification and a complete understanding of the joint performance in service, as well as for conducting the defect assessment procedure for such weld joints. Cam et al. (1999) carried out an experimental investigation on the mechanical properties of laser welded joints using flat microtensile specimen (0,5 mm thick, 2 mm wide) to establish a testing procedure to determine the tensile properties of the weld metal and heat-affected zone (HAZ). Two similar joints, namely, ferritic-ferritic and austenitic-austenitic and one dissimilar ferritic-austenitic joint were made with a CO2 laser of 6 mm plates. The results of the microtensile specimen were compared with standard tensile specimen. This showed clearly the suitability of the microtensile specimen technique for such joints. The crack tip opening displacement (CTOD) tests were also performed to determine the fracture toughness of the LB welds using three-point bend specimen. The effect of strength heterogeneity (mismatching) across the weld joint and at the vicinity of the crack tip on the CTOD fracture toughness values was also discussed. 6.2

Weld mismatch

Welds are often made with substantial strength mismatch between base material and weld metal. However, when applying traditional defect assessment practices to crack-like defects in welds, it is normally assumed that defects are located in material of uniform mechanical and microstructural properties. This strength mismatch issue associated with weld performance has been investigated by Schwalbe et al. (1993). Petrovski et al. (1990) evaluated the crack driving force in the HAZ of mismatched weldments using direct J-integral measurements in tensile panels. Assessment expressions

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and models have been modified by using a mismatch factor to evaluate the fracture tendency for a defect in mismatched welds (Schwalbe, 1991). Kocak et al. (1993) summarized the fracture aspects of under-and over-matched weld joints. They conclude that in highly over-matched weld panels, the deformation behaviour is controlled by the crack size and low toughness level of the weld metal. Extremely over-matched weld metal could not provide any protection for large cracks of a/W= 0.3 and 0.5. Slightly over-matched ferritic weld metal with good toughness exhibited the best deformation pattern by shielding for the whole loading range compared to the highly under- and over-matched weld metals. The fracture behaviour of the tension loaded mismatched weld joints with large defect sizes was found to be mainly dependant on the fracture toughness of the defective zone. Even an extremely over-matched weld metal cannot provide a protection against a crack tip from applied strain due to significantly reduced toughness.

6.3

Influence on strength

Park et al. (2001) report about the effects of alloying elements on the thermal properties of HSLA (High Strength Low Alloy) cast steels investigated in view of thermal fatigue, thermal shock, and tensile tests. According to their investigations, the optimum composition of HSLA cast steels in order to obtain the highest thermal fatigue resistance was 0.1% C, 1.2 % Mn, 0.05 % Nb, 0.05 %V steel, resulting in polygonal ferrite plus small amounts of bainite microstructure. Thermal shock resistance of HSLA cast steels was also superior to that of SC42 cast steels. However, the difference between the HSLA cast steels with both niobium and vanadium, and those with niobium or vanadium individually has not been found. Fatigue strength of laser beam-welded lap joints is usually assessed on basis of structural stresses in the joints. Stresses are nevertheless difficult to determine due to overlap of welded sheets. Zhang (2002) describes a strain gauge method that can determine inaccessible stresses by attaching two or three strain gauges to the outer surface of the overlap sheets near the joint. The method was validated by finite element simulations. Results obtained from the current method are compared with those from the literature.

7.

CONCLUSIONS

Regarding materials it is interesting to note that in the search for lightweight designs new Aluminium and Titanium alloys seem to offer interesting options, whilst at the same time also high strength steels will have their place provided the matter of residual stresses in the rolling process and subsequent non destructive detection possibilities are properly solved. Titanium suffers from its relatively new appearance as a structural material and bears an image of an exotic and hence expensive material, and its use had been restricted to aeronautical and biomedical applications. Consequently, it's difficult for this recent material to compete without references with steel or aluminium, strongly established in the structural material market. Nevertheless, titanium must be considered wherever corrosion is a problem or weight a factor, appearing as an elegant solution. Furthermore, the new elaboration process may open wider economic opportunities and could be the trigger for a full size industry. FRP/Composite materials gain added attention in connection with adhesive bonding technologies both, with respect to light weight designs and to cost reductions. The problem of fire resistance needs particular attention in this matter.

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Regarding welding consumables there is still need for research conceming new analytical practices to select more resilient welding consumable compositions and process parameters to reduce weld rework for welding steels of a large compositional range. Also the problem of new low hydrogen welding consumables based on the use of irreversible hydrogen trapping and fluoride additions needs to be revisited. It is also necessary to develop, evaluate and demonstrate advanced electronic and magnetic sensors for rapid diffusible hydrogen content determination on welded structures, not just on qualification coupons. The problems associated with environmental and health conditions related to welding need further observation. The basic technologies are seen in the context of the total production chain and further reduction of rework attributed to preceding thermal processes. The increasing penetration of shop floor activities with integrated IT solutions lead to innovations in technology combinations and attempts to substitute manual work by mechanised and robotic set ups. New technologies like Lasers or Stir-Welding lead to new structural solutions like steel and aluminium sandwich panel elements for major parts of ship hull construction. Adhesive bonding becomes increasingly important in the context of distortion free joining. Coating processes are also integrated into process chain. The need for further cost reductions has triggered systematic and scientific approaches to accuracy in production especially with thin plate fabrication. Substantial progress in this matter in hull fabrication leads to innovative aspects in early outfitting, a subject that certainly needs further research efforts. The 'digital' shipyard seems to be closer with improving methods for integrating product and production information. New embedded production logistic systems reduce costs and error potentials. The 'total' solution approach can result in production process and production plant improvements. Substantial further research seems to be necessary in order to harvest the possibilities principally available. Fabrication imperfections are dealt with under many aspects. The matter of accurate manufacturing is one obvious at the fabrication level. The remaining stresses in finished structures are a matter of life cycle suitability. Because of the complexity of large welded structures a complete numerical analysis of the total structure is not yet feasible. It seems necessary to undertake further research on this subject in order to arrive at a better and rational assessment of this problem.

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Noelle, P., (2001). Schweissen und Nachbehandeln Austenitisch-Ferritischer GefUgestrukturen, Schiff & Hafen, Vol 8, 2001, 77-80 (in German). Norman A.F., Prangnell, P. B., and McEwen, R. S., (1998). Acta Materialia,, vol. 46, 5715-5732. Okano, S., Kobayashi, Y., Shibata, M., Shiwaku, T., Nagatani, K. and Sasakura, S., (2002). 355-460 MPa yield point steel plates and welding consumables for large heat-input welding for giant container ships, R&D Kobe Steel Engineering Reports, 52:1, 2-5 (in Japanese). Okumoto, Y., (2002). Study of workability and safety of welding works using virtual human model, Journal of Ship Production, 18:1, 47-53. Olsen,F., (2002). 30 years with laser cutting - from laboratory to shipyard, Proceedings IIW Conference on Advanced Processes and Technologies in Welding and Allied Processes, Copenhagen, Denmark June 2002. Olson, D.L., Ibarra, S. and Liu, S., (1991). Selection of Resilient Welding Parameters for Arctic Fabrications and Repair, Proceedings 10th Intl. Conference OMAE, Vol m-A, 125-130, Olson, D.L., Park, Y.D., Kaydanov, V., Smith, R.D., Gavra, Z., (2002). Advances in Hydrogen Sensors for Welds, Proceedings 6th Intl. Conference On Trends in Welding Research, Pine Mountain, Georgia, OH, 1-6. Onsoien, M.I., Liu, S. and Olson, D.L., (1996). Shielding Gas Oxygen Equivalent in Weld Metal Microstructure Optimization, Welding Journal, vol. 75, (7), 216-224. Osip, H., Wojcik, J., (1995). The use of titanium alloys in shipbuilding, Welding International, 1995. Park, J.H., Kim, H.J., Kim, I.B., (2001). The effects of alloying elements on thermal fatigue and thermal shock resistance of the HSLA cast steel, Proceedings ISOPE 2001, IV, 250-255. Park, J.Y., Park, H.C., Cha, T.I. (2002). Modelling and simulation of assembly process in shipbuilding, Journal of KWS, Vol.20, No.l, 18-22 (in Korean). Payrebmne, de, J., Dietze, J.U., Heinemann, M., Lengen, H., Werftentwicklung unter dem Gesichtspunkt der Genaufertigung, SchweiJ3en und Schneiden 2002, DVS - Berichte Band 220, 282- 286 Peacock, D.K., Skauge, J.,(1994) Titanium fights fire. Stainless steel Europe, Feb 1994. Peacock, D.K., (2000). Effective design of high performance corrosion resistant systems for oceanic environments using titanium. Titanium Information Group, vol 18, Nos 4-5. Petrovski, B. and Sedmak, S., (1990). Evaluation of crack driving force for HAZ of mismatched weldments using direct J-integral measurements in tensile panels, Proceedings of the Intl. Conf. Welding, Geesthacht, FRG, 341-354. Pfund B., (2001). Re-engineering Production Boat Laminates, Professional Boatbuilder, 71, 38-51. Pokhodnya, I.K., (1996). Hydrogen Behaviour in Welded Joints, E.O.Paton Electric Welding Institute, National Academy of Science of Ukraine, Kiev. Prince K., (2002). Boats Benefit from Adhesive Bonding, Reinforced Plastics, 45:3, 46-48. Quimby, B.J. and Ulrich, G.D., (1999). Fume formation rates in gas metal arc welding, Welding Journal, vol. 78, (4), 142-149. Roland, F., Reinert, T., Pethan, G., (2002). Laser Welding in Shipbuilding - an Overview, Proceedings IIW Conference on Advanced Processes and Technologies in Welding and Allied Processes, Copenhagen, Denmark, Paper B II. Sasaki, Y., Sonda, M., Ito, K., (2002). A Study on 3-D Digital Mockup Systems for Work Strategy Planning, Proceedings ICCAS 2002 Malm6, Sweden. Schwalbe, K. H. and Cornec, A., (1991) The Engineering Treatment (ETM) and its Practical Application, Fatigue and Fracture of Eng. Mat. and Stmct., Vol. 14, 405-412. Schwalbe, K. H., Kocak, M., (1993). Fracture Mechanics of Weldments: Properties and Application to Components, Proceedings 3rd Intl. Trends in Welding Science and Technology, 479- 494, ASM, Materials Park, OH.

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Scotton, T.W., (1999). Virtual Manufacturing for Shipbuilding in a Globally Collaborative Environment, Proceedings ICCAS 99, 205-216. Shenoi R.A. and Dodkins A.R., (2000). Design of Ships and Marine Structures Made from FRP Composite Materials, Comprehensive Composite Materials Encyclopedia, 6, Elsevier, 429449. Shin J., (2002). A Digital Manufacturing Model of Shipyard Forming-Shop, Journal of Korean Welding Society, 20:1, 10-17 (in Korean). Shin, J., G., Won, S., Ryu, C.H., Lee, J.H., Kim, W.D., (2000). User-Friendly, Advanced Line Heating Automation for Accurate Plate Forming, Proceedings 2000 Ship production Symposium, Williamsburg, Va. 33-43. Shin, J., G., Won, S., Ryu, C.H., Yim, H., Lee, J.H., (2002). Object-oriented development of an integrated system for manufacturing information of roll bending process, Journal of Ship Production, 18:2, 86-91. Shin,J., (2001). Digital Shipbuilding, Proceedings of the Annual Spring Meeting, Society of Naval Architects of Korea, 8-13 (in Korean). Shindo A., (2000). Polyacrylonitrile (PAN)-based Carbon Fibres, Comprehensive Composite Materials Encyclopaedia, 1, Elsevier, 1-34. Simonato, L., Fletcher, A.C., Andersen, A., Anderson, K., Becker, N., Changclaude, J., Ferro, G., Gerin, M., Gray, C.N., Hansen, KS., Kalliomaki, P.L., Kurppa, K., Langard, S., Merlo, F., Moulin, J.J., Newhouse, M.L., Peto, J., Pukkala, E., Sjogren, B., Wild, P., Winkelmann, R. Saracci, R., (1991). A Historical Prospective Study of European Stainless Steel, Mild Steel and Shipyard Welders, British Journal oflndustrial Medicine, vol. 48, (3), 145-154. Smith, R.D., Benson, D.K., Maroef, I., Olson, D.L. and Wildeman, T.R., (2001). The Determination of Hydrogen Distribution in High Strength Steel Weldments, Part II: Optoelectric Diffusible Hydrogen Sensor, Welding Journal, vol. 80, 122-126. Smith, R.D., Plandis, G., Maroef, I., Olson, D.L. and Wildeman, T.R., (2001). The Determination of Hydrogen Distribution in High Strength Steel Weldments, Part I: Laser Ablation Methods, Welding Journal, vol. 80, 115-121. SNAJ (2001). TECHNO MARINE, Society of Naval Architects of Japan, 862, 42 (in Japanese). Stacey, A., Barthelemy, J.Y., Ainsworth, R.A., Leggatt, R.H., Bate, S.K., (1999). Quantification of residual stress effects in the SINTAP defect assessment procedure for welded components. Proceedings OMAE 99,18 th International Conference on Offshore Mechanics and Arctic Engineering July 11-16,1999, St. Johns, Newfoundland, Canada. Stewart R., (2002). SCRIMP Offers a Cleaner Alternative, Reinforced Plastics, 46:5, 26-29. Strand R., (2002). Adhesive Bonding Meets Production Boatbuilding, Professional Boatbuilder, 75, 58-71. Swiatecki, S. (2001), Vacuum moulding route chosen for rescue vessels, Reinforced Plastics, 45 (5), 30-33. Takeda, Y., (2002). Prediction of welding deformation in full scale curved shell, Journal of Ship Production, 18:2, 99-104. Tani, T., Okada, N., Ohe, K., Miyazaki, T., (2001). Effect of Residual stress controlled TMCP steel plate on accuracy ship blocks, (lSt report), Influence of residual stress of TMCP steel plate on precision of slit-slot cutting, Journal of Society of Naval Architect of Japan, 189, 299-307 (in Japanese). Tani, T., Ueda, T., Ohe, K., Miyazaki, T., Nakashima, Y., (2001). Effect of Residual stress controlled TMCP steel plate on accuracy ship blocks, (2nd report), Influence of residual stress of TMCP steel plate on precision of fabricating straight block, Journal of Society of Naval Architect of Japan, 190, 599-606 (in Japanese).

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Tomita, Y., Osawa, N., Hashimoto, K., Shinkai, N., Sawamura, J., Sugiyama, K., Tsubouchi, D., Deguchi, Y., Yamaura, T., (2001). Study on heat transfer phenomena between the combustion flow field and the heating plate during line heating process, (3rd report), Journal of Society of Naval Architect of Japan, 190, 479-488 (in Japanese). Tsujino, R., Morikawa, K., Yamaguchi, H., Ogawa, K., Ochi, H., (2001). Statistical investigation for optimum welding condition of 2017 aluminium similar alloy friction welded joints, Proceedings ISOPE 2001, IV, 304-308. US Navy web site : www.dt.navy.mil/div/capabilities/accomplishments/D-14.html Varma I.K. and Gupta V.B., (2000). Thermosetting Resin- Properties, in Comprehensive Composite Materials Encyclopaedia, 2, Elsevier, 1-56. Walker M. and Smith R., (2002). A Computational Methodology to Select the Best Material Combination and Optimally Design Composite Sandwich Panels for Minimum Cost, Computers and Structures, 80, 1457-1460. Walz, C., (2001). Hybrid welding of steel for offshore applications, Proceedings ISOPE 2001, IV, 263-266. Wang, W.W., Wong, R., Liu, S. and Olson, D.L., (1996). Use of Martensite Start Temperature for Hydrogen Control, Minerals, Metals and Materials Society/AIME, 17-31. Wanner, M.C., Kothe, U., Nikolay, P., Schneidenbach, R., Wiebeck, E. (2002). Grundlagen der Genaufertigung komplizierter Sektionen des Vor- und Hinterschiffes (Accurate Manufacturing of forebody and stern blocks), Institut for Fertigungstechnik und Logistik, Rostock, (in German). Ward-Close, M., Godfrey, A., (2002). New low cost titanium extraction and processing. Proceedings Journ6es technologiques titane, 4 & 5 April, Nantes (France). Wen, S.W., Hilton, P., Farrugia, D.C., (2001). Finite element modelling of submerged arc welding process. Journal of Materials Processing Technology, VoL 119, issues 1-3, 203-209 Wilckens, H. (2001). Innovation in Ship Production, What can we expect?, Proceedings PRADS 2001, Shanghai, 359-365. Wong, R.J., (1996). Hydrogen Induced Cracking in High Strength Steel Weldments, Proceedings of Intlernational Conference on Advances in Welding Technology, EWI, 347-357. Wright P.N.H., Wu Y., and Gibson A.G., (2000). Fatigue Behaviour of Steel-Composite Joints for Ships Structures, Plastics, Rubber and Composites - Processing and Applications, 29, 549-557. Xing J., (2001). Permeability Identification of Porous Media in Resin Transfer Moulding, PhD Thesis, University of Southampton. Yamato, H., Kobayashi, I., Shirayama, S., Masuda, H., Enomoto, M. (2001). Application of the wearable system to shipbuilding industrial engineering, Journal of Society of Naval Architect of Japan, 190, 431-438 (in Japanese). Yang H.H. (2000). Aramid Fibres, in Comprehensive Composite Materials Encyclopaedia, 1, Elsevier, 199-230. Yurioka, N., (2002) Prediction of Strength of Weld Metal (Report II), IIW Doc. IX-2026-02, American Council, AWS, Miami, FL. Yurioka, N., Suzuki, H., Ohshita, S. and Saito, S., (1983), Determination of Necessary Preheating Temperature in Steel Welding, Welding Journal, vol. 62, (6), 147-153. Yushchenko, K. A., Ishchenko, A. Y., Tretyak, N. G., Lozovskaya, A. V., and Sklabinskaya, I. E., (1995). Welding in the World, vol. 35, pp. 48-51. Zhang, S., (2002). Stresses in Laser-Beam-Welded Lap Joints Determined by Outer Surface Strains. Welding Research, supplement to Welding Journal, January 2002, 14- 18. Zhu, X. K., Chao, Y.J., (2002). Effects of temperature-dependent material properties on welding simulation, Computers and Structures, Vol. 80 967-976.

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Zorn, H., Henkel, K.-M., Kothe, U., Kunkel,J., (2000). Genaufertigung von zweiachsig gekriimmten Flachen- und Volumenbaugruppen aus Stahl (curved panel), Report of Institut far Fertigungstechnik und Logistik, Universit~it Rostock, (in German).

TABLE 2.3.1 WEB BASEDSOURCESFORINFORMATIONONFRP MATERIALS Raw Material Type

Fibre reinforcements

Resin systems

Core materials

Adhesive materials

Web Address www.ahlstrom papergroup.com www.chomarat.com www.reichold.com www.sgva.com www.advanced-composites.com www.dupont.com www.aoc-resins.com www.ashpec.com www.dow.com www.dsm.com www.scottbader.com www.spsystems.com www.crayvalley.com www.alcanairex.com www.diabgroup.com www.lantor.nl www.rhoem.de www.itwplexus.com www.ipscorp.com www.asgcem.com www.lordadhesives.com www.permabond.com www.sika.com

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15th INTERNATIONAL SHIP AND OFFSHORE STRUCTURES CONGRESS 2003 AUGUST 11-15, 2003 SAN DIEGO, USA VOLUME 1 fv DIEGO, US

COMMITTEE VI.1

FATIGUE LOADING

C O M M I T T E E MANDATE Evaluate and develop direct calculation procedures for fatigue loads on displacement ships. Due consideration shall be given to the combination of the global and local responses induced by external wave pressure and internal cargo loads. Simple and more refined procedures should be compared with each other and with the statistical distributions obtained from in-service experience and/or experimental tests. Recommendations of standard design analysis shall be given.

COMMITTEE MEMBERS Chairman:

Dr. Dr. Mr. Prof. Dr. Mr. Dr.

Iwao Watanabe Kim Branner Alain Cariou Toichi Fukasawa Xue Kang Gu Geert Kapsenberg Enrico Rizzuto

KEYWORDS Fatigue load, Wave-induced hull girder loads, Wave pressure, strip theory, validation Comparative calculation, stochastic load description, design wave, stress transfer function.

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CONTENTS

1 INTRODUCTION

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2 SIMPLIFIED PROCEDURES BY CLASSIFICATION SOCIETIES . . . . . 2.1 Introduction . . . . . . . . . . . . . . . . . . . . . 2.2 General F r a m e w o r k o f Simplified Checks . . . . . . . . . . . 2.3 N o m i n a l P r i m a r y L o a d s . . . . . . . . . . . . . . . . . . 2.3.1 F o r m u l a t i o n s o f Wave Loads . . . . . . . . . . . . . . . 2.3.2 R e f e r e n c e E x c e e d i n g Probability . . . . . . . . . . . . . 2.3.3 C o m b i n a t i o n o f P r i m a r y Stress C o m p o n e n t s . . . . . . . . . . . 2.4 Local Stresses . . . . . . . . . . . . . . . . . . . . 2.4.1 Ship M o t i o n s . . . . . . . . . . . . . . . . . . . . 2.4.2 Inertial Forces and Stresses . . . . . . . . . . . . . 2.4.3 External Pressures and Inherent Stresses . . . . . . . . . 2.4.4 C o m b i n a t i o n o f Local Stresses . . . . . . . . . . . . . 2.5 C o m b i n a t i o n o f P r i m a r y and Local Stresses . . . . . . . . 2.6 Stress C o n c e n t r a t i o n Factors . . . . . . . . . . . . . . . 2.7 L o n g Term Stress R a n g e Distribution . . . . . . . . . . . 2.8 Material Fatigue Strength Curve ( S - N ) . . . . . . . . . . . . . . 2.9 Fatigue Checks . . . . . . . . . . . . . . . . . . . . . . 2.10 C o m m e n t s . . . . . . . . . . . . . . . . . . . . . . .

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3 HYDRODYNAMIC LOADS FOR FATIGUE . . . . . . . . . . . . . 3.1 General R e m a r k s . . . . . . . . . . . . . . . . . . . . . 3.2 Linear Estimation . . . . . . . . . . . . . . . . . . . . 3.2.1 Strip M e t h o d . . . . . . . . . . . . . . . . . . . . 3.2.2 Three D i m e n s i o n a l M e t h o d . . . . . . . . . . . . . . . . 3.2.3 Time Simulation Technique . . . . . . . . . . . . . . 3.3 N o n - L i n e a r i t y . . . . . . . . . . . . . . . . . . . 3.3.1 N o n - L i n e a r i t y in Global L o a d . . . . . . . . . . . . . 3.3.2 N o n - L i n e a r B e h a v i o u r o f Wave Pressure Distribution . . . . . . 3.4 H y d r o e l a s t i c i t y . . . . . . . . . . . . . . . . . . . . . . . 3.4.1 Global Effects . . . . . . . . . . . . . . . . . . . 3.4.2 Local Effects . . . . . . . . . . . . . . . . . . . . . 3.5 Internal L o a d s . . . . . . . . . . . . . . . . . . . . . . 3.6 Other L o a d s . . . . . . . . . . . . . . . . . . . . . . .

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4 STRESS TRANSFER FUNCTION . . . . . . . . . . . . . . . . . . 4.1 Structural M o d e l . . . . . . . . . . . . . . . . . . . . 4.2 L o a d Application . . . . . . . . . . . . . . . . . . . . . . 4.3 Stress Analysis . . . . . . . . . . . . . . . . . . . . .

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238 4.4 Stress Combination

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5 SHORT A N D L O N G T E R M ESTIMATION . . . . . . . . . . . . . 5.1 Short Term Predictions . . . . . . . . . . . . . . . . . 5.1.1 Sea Spectral Shapes . . . . . . . . . . . . . . . . . . 5.1.2 Effect o f Short-Long Crest Description . . . . . . . . . . . 5.1.3 Effect o f the Shape o f Stress Range Distributions . . . . . . . . 5.2 Long Term Characteristics . . . . . . . . . . . . . . . . . . 5.2.1 Environmental Conditions . . . . . . . . . . . . . . . . 5.2.2 Operational Conditions . . . . . . . . . . . . . . . . . 5.3 Remarks . . . . . . . . . . . . . . . . . . . . . . . . 6 D E S I G N WAVE M E T H O D . . . . . . . . . . . 6.1 Selection o f Design Wave Condition . . . . . . 6.1.1 Considered Design Responses . . . . . . 6.1.2 Selection o f Wave Length and Height . . . . 6.1.3 Selection o f Heading and Speed . . . . . . 6.2 Long Term Stress Distribution . . . . . . . . 6.2.1 Stress Calculation . . . . . . . . . . . 6.2.2 Long Term Load Distribution Model . . . .

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7 BENCHMARK CALCULATIONS . . . . . . . . . . . . . . . . . . 7.1 Benchmark by Direct Method . . . . . . . . . . . . . . . . . . 7.1.1 Vessel and Calculation Condition . . . . . . . . . . . . . . 7.1.2 Calculation Method . . . . . . . . . . . . . . . . . . . 7.1.3 Load Outputs . . . . . . . . . . . . . . . . . . . . . 7.1.4 Stress Outputs . . . . . . . . . . . . . . . . . . . . . 7.1.5 Long Term Prediction . . . . . . . . . . . . . . . . . 7.1.6 Loads Results and Comments . . . . . . . . . . . . . . . . 7.1.7 Stress Results . . . . . . . . . . . . . . . . . . . . . 7.1.8 Damage Results . . . . . . . . . . . . . . . . . . . . 7.2 Design Wave Method . . . . . . . . . . . . . . . . . . . . 7.2.1 Considered Design Responses . . . . . . . . . . . . . . . 7.2.2 Selection o f Design Wave . . . . . . . . . . . . . . . . . 7.2.3 Long Term Stress Distribution . . . . . . . . . . . . . . . 7.2.4 Results and Considerations . . . . . . . . . . . . . . . . 7.3 Consideration o f Calculated Results . . . . . . . . . . . . . . . . 7.3.1 Comparison between Results from Methods o f Different Level o f Complexity 7.3.2 Comparison between Results o f Total and Component Stress Ranges . .

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8 CONCLUSIONS AND RECOMMENDATIONS

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REFERENCES

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Fatigue Loading 1.

239

INTRODUCTION

The present ISSC has set up two Special Task Committees to deal with fatigue analysis: STC VI.1 for the fatigue loading and STC VI.2 for the fatigue strength assessment respectively. The present committee tries to develop procedure to estimate nominal stress and items to be considered in the procedure. Fatigue failure is caused by cyclic repetition of loads due to encountering wave and resulting ship motions in her lifetime. There are various ways of estimation method for fatigue loads depending on simplification introduced in the computation model. The most commonly used methods for fatigue analysis are rule-based methods by classification societies. These methods are well calibrated to the actual damage cases and easy to use. However, it is known that they may give wide deviations in the prediction of ship motions, wave loads and in the combination of the resulting stresses. This reflects in a wide scatter in the resulting fatigue damage estimation. It is difficult to find out what the causes of the deviation are since the load estimation is made implicitly. We need to establish more consistent method based on state of the art of wave load estimation and structural analysis. The most detailed method would be to simulate the time history of structural responses in the computer taking whole hydrodynamic and structural characteristics into account. The method could give load characteristics caused by ship response necessary for fatigue strength assessment when applied to all the ship conditions and all the cruising conditions. However, it is not feasible for practical purpose at present. The simplified but most detailed way for the today practice is to use approaches based on linear superposition. These approaches are sometimes called direct methods or spectral methods. They calculate frequency response of the load components first by linear hydrodynamic tools taking into account non-linear corrections and then calculate the frequency response of stress at local points of interest by using FEM models. This solves the problem of a better prediction of motions and loads and also tackles the problem of the stress combination. The committee intends to discuss this approach as the direct procedure for wave loading as requested in the mandate as schematically described in Figure 1. There may be more simplified methods than the approach mentioned above, depending on simplifications in the treatment of structural response estimation. The committee discusses design wave method as a simplified method. The design wave method is an effort to decrease laborious works of the structural analysis substantially by selecting a few design wave cases. The report consists in the following steps. First we review procedures adopted by class societies and give some general comparison to point out scatter in the estimation given by them. Next we proceed to discuss the direct method. We review first the state of the art knowledge on hydrodynamic aspect for estimating wave load in linear approaches. Due consideration is given to significant non-linearities, which may arise in relation to ship types, ship conditions and wave conditions. Next discussion will be given on the conversion of the hydrodynamic loads into the stress at a local point in the ship structure. In the discussion, we limit ourselves to discuss only nominal stresses to avoid over- lapping to the works by STC VI.2. The issues about short-term and long-term predictions of load and stress are discussed next so that the derived stress amplitude could be related to the cumulative distribution for further analysis developed by the other committee. Then the design wave method is discussed. This approach is based on the concept of selecting a few design waves replacing a reference situation obtained with a long-term prediction (Figure 1.1). All general discussions have been illustrated quantitatively by an example, which was selected for a benchmark calculation. In a specific chapter

Special Task Committee VI.1

240

results are presented in order to demonstrate how the different wave load estimation methods give load deviations and how their difference will affect stress response. As a conclusion, recommendations on the various aspects on procedures for wave loading are given. One remark is due. In seeking data for the benchmark calculation, it became clear that no ship was available of which data set can be open to the comparison study because of confidentiality of hull form or structural design. Therefore, the benchmark the report has to deal only with numerical computations of a hypothetical VLCC hull form.

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Figure 1.1: Calculation flow of fatigue damage factor

2.

S I M P L I F I E D P R O C E D U R E S BY CLASSIFICATION SOCIETIES

2.1

Introduction

Simplified procedures for fatigue checks are issued by all Classification Societies for specific locations on board, selected as those most prone to the phenomenon. The IACS unified recommendation for the fatigue assessment of ship structures (IACS Recomm. no. 56, 1999) provides a common framework for all procedures, but each Society developed its own verification, giving rise to some differences in the methodology adopted and to quite significant discrepancies in the results of the checks. This was pointed out in ISSC 2000 Committee 11].2, where a comparative study of a detail of the bulwark of a container ship is presented. That test case was selected in order to have a comparatively limited number of stress components, but nevertheless a wide dispersion in results was found in terms of predicted fatigue life.

Fatigue Loading

241

The mandate of our committee deals specifically with load effects in fatigue predictions and, accordingly, a first step has been to compare simplified procedures formulated by Class Societies specifically as regards load definition, stress derivation and stress component combination. To this aim, a comparison based on a detail including more stress components was selected within the double hull tanker ship which was made available for evaluation. The ship is presented in chapter 7, where results are presented of load predictions coming from different procedures of increasing accurateness, as described in the chapters 3-6. The detail selected for comparison of simplified checks in the present chapter is the connection between longitudinal No.5 of the outer side and the supporting primary structure at mid-ship. The vertical position of the longitudinal corresponds to about 3.9 m below the full load waterline, (while being about 7.8 m above the ballast waterline) as better described in Chapter 7. With reference to this test case, formulations by Class Societies have been qualitatively compared in the present chapter, with the aim of outlining differences and similarities in the approaches. A few quantitative results about load components estimation are also presented here, while long term stress range Weibull distributions on which checks are based are compared in Chapter 7 with distributions obtained with direct methods and other procedures. The incomplete number of computed results was due in part to the extensive work required to pass through the whole evaluation process step by step (automated procedures are often available, but usually they do not provide intermediate results) in part to the fact that in some cases complete procedures were available only as computer programs and it was difficult even to find out documentation of the algorithms embedded in the procedure. Numerical results given under Register names in this chapter were not checked by Registers and represent an independent application of Rules."

2.2

General framework of simplified checks

All procedures are based on the identification of one or more load cases for each of the two loading conditions (ballast and full load, in the examined case). In each load case an evaluation is performed of the stresses induced by primary loads of the hull girder in the check position. A separate evaluation is performed of stresses due to local loads (external wave pressure and internal inertial pressure exerted in the liquid cargo by ship motions). Primary and local stresses are then combined to define a reference stress range. Stress concentration factors are applied to account for the geometry of joint and weld. From the reference stress range and its associated probability a stress range distribution is derived. This distribution, coupled with the Miner's hypothesis of linear accumulated damage and to S-N curves describing the material fatigue strength, provides the background for all verifications, even if the way such verification are formulated and the type of final output can differ. The various steps of this general framework are analysed in more details in the following.

2.3

Nominal primary loads

2.3.1 Formulations of wave loads All Societies consider contributions coming from bending of the hull girder in the vertical and the horizontal plane. They all refer to Unified Rule S 11 for quantification of the wave induced vertical bending moment. The horizontal bending component is provided, on the contrary, with different formulations, all based in general on ship geometrical characteristics. Explicit formulations for torque

Special Task Committee VI. 1

242

moment are provided by some Societies, but this component was disregarded in numerical values in the present case (ship with a closed section). All checks adopt the scheme of a simple beam when deriving stress components from vertical and horizontal bending moments.

2.3.2

Reference exceeding probability

In IACS Unified Rule S 11, reference is made to an exceeding probability of 10 -8. The same applies in general to reference values of horizontal bending. This poses the problem of scaling such an extreme value (likely to occur once in the ship's lifetime) to a more representative value of the ship fatigue behaviour. In IACS Recomm.no.56, an exceeding probability of 10-3 to 10 .5 is suggested as representative for fatigue loading and most of the Societies mention a reference probability complying with the indication, but with values different from each other. Other sources of differences in the procedures are represented by the scaling algorithm, which is always based (explicitly or implicitly) on Weibull distributions, but with different shape parameters (in some cases related to the ship's length, in other considered as a constant (=1: straight line in a semilogarithmic plot). TABLE 2.1 GLOBALLOADS GLOBAL LOADS Class society Reference probability .~ Weibull shape for scaling probability of global loads ~ Factor correcting probability from 10.8 to reference one

BV&RINA 10-5

CCS 10-8

DNV 10-4

GL 10-6

KR 10-4

(1.000)

NA

0.857

(1.000)

0.843

0.625

NA

0.445

0.750

0.440

Simple beam

Structural scheme for derivation of stresses Extreme Vertical Wave Ind. BM (hog) [MN m]

10102

Extreme Vertical Wave Ind. BM (sag) [MN m]

-10934 5020 28

3159 45

6041 20

5463 34

6375 20

Nominal stress from VWIBM (sag) [MPa]

-30

-49

-22

-36

-21

Nominal stress from HWIBM (amplitude) [MPa]

23

24

20

31

21

2781

"~'Extreme Horizontal Wave Ind. BM [MN m] Nominal stress from VWIBM (hog) [MPa]

.i

9437

Extreme Vertical Wave Ind. BM (hog) [MN m]

10549

Extreme Vertical Wave Ind. BM (sag) [MN m] ,,,a

2120

1984

3965

3673

~ Nominal stress from VWIBM (hog) [MPa]

26

42

19

31

18

Nominal stress from VWIBM (sag) [MPa]

-29

-47

-21

-35

-21

Nominal stress from HWIBM (amplitude) [MPa]

10

15

13

21

9

,.,.,

Extreme Horizontal Wave Ind. BM [MN m]

2.3.3

Combination of primary stress components

The combination of the two primary components is performed in different ways: generally speaking, combination criteria can be grouped according to the following categories: (i) based on stresses or stress ranges (ii) expressed by linear or non-linear laws, (iii) formulated in a unique way or for different load cases (which, in turn, can be separately accounted for, or among which the worst case is selected). These alternative choices reflect different logical schemes for describing the long-term statistical correlation between the two primary components. In particular, point (i) reflects a different way of treating mean stress values and the superposition of the various components: in one case ranges and mean values are derived after combination of all

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243

components, in the other case the single ranges are computed and combined, while mean values corrections, if any, are evaluated separately. Point (ii) refers to the statistic relationship between components. In the time domain, there is a phase lag between the components to be combined (as they will not have the reference values at the same instant). The question is how this phase lag (which is a stochastic variable) can influence the statistic of the combined load. One strategy (linear combination) is to consider the combined load in phase with of one component and to add a certain percentage of the other one (hypothesising a representative phase angle). Another strategy (quadratic combination) is to compute the sum of the two component vectors, again with an hypothesis on the phase lag between them. The third approach is to cumulate independently the fatigue damage of the two fluctuations deriving an equivalent stress range (cubic average, if a slope o f - 3 is selected in the S-N curves). Point (iii) relates to the interpretation of the various load cases: sometimes they are considered as contributing separately to the total fatigue damage, sometimes they are used to define a unique reference situation (corresponding to the worst case, or to a weighed sum of the various ones). In Table 2.2, a comparison is presented of the various formulations and also of numerical values (the latter comparison performed at a nominal stress range value, excluding all stress concentration factors). TABLE 2.2 COMBINATION OF GLOBAL LOADS Class

BV

Formul a, b AV c, d 0.4AV+AH

CCS

DNV

(AV2+AH2 +.2AVAH) ~

(AV2+z~--I2 +2pAVAH) ~ p= correl, coeff.

GL

KR

max~0.5AV + AH

LAY

RINA AV 0.4AV+AH 1/2 {[max(A~a;ACrb)]3 +[rnax(Acra;A~b)]3}0.33 a, b c, d

Full load

a,b: 67 c,d: 81

109

61

128

41

97

Ballast

a,b: 29 c,d: 48

96

50

99

39

53

2.4

Local stresses

Local stresses are generated on the side by external hydrodynamic pressures on the wetted surface and inertial pressure exerted in the internal liquid by ship's local accelerations. Motions of the ship are therefore necessary to define both the internal and external components. A stress component corresponds in all cases to a beam uniformly loaded by the pressure directly acting on either side of the plate. There are differences however among the way different Class Societies treat stresses at ends. This reflects in different values of stress per unit pressure at the verification point. To these components are in some Rules added those deriving from the relative deflection of the different primary supporting members at ends of the stiffener. These effects are represented by a stress increase at the stiffer support (bulkhead) and by a decrease at the other end (due to the deflection of the other support: transverse). For a proper evaluation of these components, the response of the ship side as a whole is to be investigated, which requires a more complex structural model. As the beam selected for investigation has both ends on transverses, no relative deflection was accounted for in numerical evaluations. Other components modelled in some cases are related to the deflection of the double side: to obtain such components a further complicated structural model is needed. In this case additional stresses can be generated in the stiffener also by inertial pressure not directly acting on the stiffener plate. This applies in the considered case to the inertial forces acting on the crude oil contained in the inner tank:

244

Special Task Committee VI. 1

they induce stress components also on the outer side as a part of the double hull structure. Due to the considerable rigidity of the double hull structure, these components are neglected in the numerical values presented. 2.4.1

Ship motions

Formulations are provided for ship linear and angular motions and their second time derivatives. Adopted reference values are likely to be at the same reference exceeding probability of primary loads, even if this is not always clearly stated in the Rules. Sometimes they are calculated directly at this reference probability, sometimes they are obtained by correction of (extreme) values at a lower (10 -8) probability (same procedure as for primary loads). In one case reference value of combined motions are provided for the whole ship instead of single components. TABLE 2.3 SHIP MOTIONS

Quantity (full load / ballast) reference probability heave acceleration [m s"2] surge acceleration [m s"2] extreme pitch amplitude [rad] extreme pitch period [s] extreme pitch angular accel [rad s "2] sway period [s] sway acceleration [m s "z] extreme roll amplitude [rad] extreme roll period [s] extreme roll angular accel [rad s z] Yaw acceleration [rad s "z] Longitudinal acceleration (surge + pitch) [m s"2] Transversal acceleration (sway + yaw + roll) [m s "2] Vertical acceleration (heave + pitch) [m s"2]

2.4.2

BV&RINA (10 .5)

CCS (10 .8)

DNV 10 -8

1.531 / 1.531

1.920 / 1.987

1.884 / 1.949

0.500 / 0.500

0.442 / 0.427

0.433 / 0.419

0.053 / 0.056

0.076 / 0.082

0.076 / 0.082

10.286 / 10.286

10.182 / 10.182

10.280 / 10.280

0.020 / 0.021

0.029 / 0.031

0.029 / 0.031 0.724 / 0.724

GL 10-6

6.793 / 6.793 1.187 / 1.187

0.738 / 0.738

0.182/0.182

0.301 / 0.318

0.301 / 0.366

17.218 / 17.218

17.441 / 15.652

17.441 / 10.518

0.024 / 0.024

0.039 / 0.051

0.039 / 0.130

=

=

=

0.471 / 0.690

=

=

=

6.072 / 4.535

=

=

=

1.849 / 1.853

0.00810.008

Inertial forces and stresses

Local accelerations at the centre of gravity of the tanks are derived to define inertial loads on the tank's walls in contact with the fluid. The definition of reference values of inertial pressures is done with reference to local accelerations at the tank barycentre occurring in the three directions: longitudinal, transversal and vertical. If a single load case is defined, it corresponds to the worst situation, if more load cases are accounted for, they are all brought to the following phases of the evaluation (Table 2.4). 2.4.3

External pressures and inherent stresses

The external hydrodynamic pressure field is modelled with simplified patterns, referred to ship motions of two main types: in the transverse plane ("roll dominated motions") or in the longitudinal one ("pitch dominated").

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245

T A B L E 2 . 4 L O C A L LOADS Class society

Structural scheme Bending stress per unit pressure [Mpa/kPa] longitudinal acceleration at barycentre of ballast tank [1TI]S2] transversal acceleration at barycentre of ballast tank [m]s 2] vertical acceleration at barycentre of ballast tank [ITI/S2]

BV

CCS

beam +

beam +

rel.defl.supp

rel.defl.supp

1.561 a,b: 0.813 c,d: 0.000 a,b: 0.000 c,d: 2.263 a,b: 1.771 c,d: 0.698 a: O, b:76 c: 8, d: 11

DNV beam +

GL

RINA

beam +

beam +

rel.defl.supp +DH deft. rel.defl.supp

1.098

1.426 1.08

1.36

0.690

3.49

4.58

4.535

2.51

4.23

1.853

57.4

44.2

15.9

a:O, b: 120 c: 13, d: 18

82

49

19

a, b: 6.56

12.44

7.84

External pressure range at the stiffener [kPa]

O, b:O c:a:10, d: 5

107

1

28

Stress range due to external pressure [MPa]

c: a:15,O,d:b:O 7.5

152.0

1.1

33.4

Amplitude of relative motion [m]

a, b: 6.80 c, d: 8.86

7.22

7.14

External pressure range at the stiffener [kPa]

a: 100, b: 50 c: 124, d: 62

354

115

86

Stress range due to external pressure [MPa]

a:155, b: 78 c." 194, d.' 97

505

127

103

Internal pressure range at the stiffener [kPa] Stress range due to internal pressure [MPa] Amplitude of relative motion [m]

c, d: 8.74

rel.defl.supp

1.207

1.561 a,b: 0.813 c,d: 0.000 a,b: 0.000 c,d: 2.263 a,b: 1.771 c,d: 0.698 a: O, b:76 c: 8, d: 11 a:O, b:120 c:13, d: 18 a, b: 6.56 c, d: 8.74

60 (equiv.pr. range) a: O,

b:O

c: 15, d: 7.5 6.80 8.86 143 (eq. pr. range ) 223 (eq. str. range) a, b: c, d:

The ship side, which is the ship's area under consideration in the present study, is generally subdivided into three regions: (a) one across the still water line, which represents the side area which is alternatively wet and dry when the ship is subjected to reference motions relative to the free surface (b) a second zone above the previous one, which is "always" dry and, therefore, is not at all exposed to wave pressures (or it is exposed to a minimum threshold value), and (c) a third zone "always" immersed. A key point is the evaluation of the amplitude of the relative motion. The position that has been chosen for the present evaluation is about 3.9 m below the still water plane in full load and about 7.8 m above the one in ballast (see Table 2.4 for values of relative motions and external pressures).

2.4.4 Combinationof local stresses As for primary loads, also for local loads different strategies for combination of internal and external pressures are followed, including linear and non linear formulations. Load cases, when modelled, are either considered separately or averaged in different ways (Table 2.5). TABLE 2.5

C O M B I N A T I O N OF PRIMARY AND LOCAL STRESS RANGES Combination of local effects (E=external; I= internal)

BV Load condition

Full load Ballast

a,b,c,d: A~ = A~ext+A(~in t a: 187, b: 93 c: 232, d: 116 a: O, b: 143 c: 33, d: 30

CCS (A(IE2+A(~I2 +2pAcrEA~I)05 19= correl, coeff 505 19=0.237 186 9=0.200

DNV (A~E2+A(~I 2 +2pAcrzAch)~ 19= correl, coeff 127 19=0.587 49 19=0.550

GL

RINA

m(y = m(Yext+m(yhlt

1/2 {[max(A~a;A(yb)]3 +[max(Acr~;Acra)13}0.33

103

177

53

106

Special Task Committee VI. 1

246

2.5

Combination of primary and local stresses

The final step of reference stress range(s) definition is generally the combination of the various components coming from primary and local loads. This is not always the case, as some checks regard separately hull girder stresses and local component. Table 2.6 provides a comparison on the subject (numerical values presented are at notch stress level, when applicable: see the following paragraph). As apparent, linear combinations are preferred for the purpose. TABLE 2.6 COMBINATION

Class Soc.

OF PRIMARY AND LOCAL STRESS RANGES

(AP=primary; AL= local) CCS DNV

BV

a,b,c,d: 0.9x 0.5AP Max{ (0.6AP+AL); Max{ (0.6AP+AL); Criterion + 0.75AL (AP+0.6 AL) } (AP+0.6 AL) } + Adefl (ign.) a,b: 486 Full 513 202 c,d: 496 Range load

values

2.6

Ballast

a,b: 404 c,d: 134

220

132

GL AP+AL

RINA {[max(p.aACYa;~b A(Yb)]3 ~.[mRx(j.tcAoc;j.taACYd)]3}0.33

231

479

152

341

Stress concentration factors

When a notch stress approach is followed, nominal stresses components (primary and local) are multiplied by stress concentration factors (SCF) to get hot spot stresses and notch stresses. These corrections are sometimes applied during the evaluation of single components (particularly if different SCF are to be applied to specific components), sometimes after the combination. Table 2.7 presents some numerical values. The asymmetry in the stiffener section is often accounted for in SCF for local loads. This effect has not been accounted for in numerical results (symmetrical profile). When different approaches are followed, based on nominal stresses, concentration factors are not provided (and the same effect is accounted for with the selection of different S-N curves) TABLE 2.7 STRESSCONCENTRATION FACTORS

Hot spot

BV- RINA Hull gird I Local

1.30 Notch

2. 7

] 1.65 2.63

CCS NA NA

DNV 2.10

GL NA NA

Long term stress range distribution

All procedures are more or less explicitly based on a long term stress range probability distribution, which is used in combination with S-N curves to compute the damage accumulated according to the Miner's Rule. In most cases, the long term distribution is a Weibull, which is identified by a shape parameter and by a 'calibration point' belonging to the curve itself. The shape parameter is related to ship main

Fatigue Loading

247

dimensions as in the reference(IACS Recomm. no. 56), while point coordinates correspond basically to the reference stress range SR described in the previous sections and to its inherent exceeding probability PR. Table 2.8 presents a comparison of some numerical values. The equation for the Weibull distribution of stress range S can be expressed as a function of these three elements as: (IACS Recomm. no. 56)

f(S) = ~ / S / ~ - ' e x p l - (S/~ ]

TABLE 2.8

2.8

where: k =

SR [ln(1/PR)] ~/~

P A R A M E T E R S OF THE W E I B U L L LONG TERM STRESS RANGE DISTRIBUTIONS

Class society Reference probability PR

BV

CCS

DNV

GL

10.5

10.8

10-4

10-6

shape param. ~ Full load reference stress range Sa shape param. ~ Ballast reference stress range SR

0.813 a,b: 486 c,d: 496 0.829

0.848

0.878

513

202

0.836

0.887

220

132

a,b:404 c,d: 134

RINA 10.5 0.850

NA

479 0.850

NA

341

Materialfatigue strength curve (S-N)

S-N curves characterise the fatigue strength of the material. Such curves are represented by one or two strait lines in a semi-logarithmic plot. Accordingly 2 or 4 parameters identify the curves. Slope values are generally quite homogenous (3 for single slope and 3 and 5 for double slope, even if different values are possible) and also the change of slope, if present, is often at a probability 10-7. On the contrary, different values of the constant identifying the position of the first strait line are used in various Rules, particularly in those which adopt a nominal stress approach, which implies to adapt the S-N curve to the specific detail (Table 2.9) TABLE 2.9 PARAMETERSFOR S-N CURVES

S-N curve parameters

2.9

BV- RINA

CCS

DNV

GL

Constant

5.09E+12

4.33E+11

5.75E+12

NA

Slope(s)

3 and 5

3

3 (3 and 5)

NA

Prob. at sl. change

107

=

(107)

=

Fatigue Checks

As a final step of all procedures, a check is performed on the detail to assess if it satisfies a minimum fatigue strength requirement. While the general background of checks is the same, the way they are formulated is quite different and varies from a comparison against limit values of the predicted damage over the ship life, to a comparison of the reference stress range with an allowable value, to prescriptions on the minimum scantlings of the stiffener. Results of some checks are presented in Table 2.10

Special Task Committee VI. 1

248

T A B L E 2 . 1 0 D A M A G E CHECK RESULTS BV

CCS

DNV

type

Damage

Damage

Damage

limit

<0.98

<1

<1

<225

< 350

> 4138

Full

1.06

3.73

1.15

231

NA

NA

Ballast

0.30

0.22

0.28

152

NA

NA

to be checked

1.36

3.95

1.44

231

421

1285

Verified?

NO

NO

NO

NO

NO

NO

Check

Result

GL

RINA

Allowable Allowable

Stiff.

modulus

As it can be seen, in the specific case all checks agree on the inadequacy of the scantling of the stiffener, but margins appear to be quite different.

2.10

Comments

On the basis of the case analysed, even for the limited number of numerical data compared, a considerable dispersion is shown among the results obtained from simplified procedures. Analysing in more details the various steps, main sources of uncertainties are represented, as regards loads, by the formulation of ship motions and local loads and of their combination, while stress derivation from local loads and stress concentration factors are the main problems in the area of structural response. More detailed models able to deal better with these aspects are needed to increase the quality of predictions.

3.

H Y D R O D Y N A M I C LOADS F O R F A T I G U E

3.1

General Remarks

A ship is subject to wave loads due to wave encounter and subsequent ship motion and acceleration while navigating in the ocean. These loads dominate the fatigue life of the ship structure. The direct alculation of fatigue load is to estimate these load as accurately as possible. The estimation of the fatigue loads is not far different from the estimation of ultimate loads which was described in the report of ISSC 2000 Comm.VI However there are several features in estimating the fatigue load to be taken special care of compared to the ultimate strength load. One of them is that since the latigue loads affect in cumulative way throughout her navigation history the moderate sea condition which the ship is likely to encounter most frequently has relatively important influence rather than very rarely occuring severe sea states. And in the same context another point is that not only head sea condition but also other encounetring condition such as beam sea and following sea conditions becomes equally important. Another point is that the shorter wave which causes no significant ship motion may have important influence to the fatigue load through pressure application on the hull surface. The last point is that the fatigue stress is result of the combined application of various load components the phase relation among the load components. Therefore The estimation method has to be able to give accurately the phase relation among the load components. In summary the estimation method has to be able to predict wave loads with equal accuracy irrespective of wave heading and wave length. This is contrast to the ultimate load where the head sea case is the dominant concern.

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249

The load components are categorized into Global and local load components. The Global load components may be defined as the ones governing the response of the primary structure of the hull or entire hull. They are the vertical and horizontal bending moments, the longitudinal torsion moment, the vertical and horizontal shear forces generated by waves and by consequent motions. In that respect they might also be called hull girder wave loads. They are considered as global entities since they result from an integration of both local pressure and local density of inertia (and gravity) loads over the whole part of the hull located on one side of the section at which the hull girder loads are considered. The local load components may be defined as the forces and moments governing the response (stress and deflexion) of the local structure such as single panel of stiffened plating and of the tertiary structure such as elementary panel of plating bounded by adjacent stiffeners of the hull. The pressure distribution over the hull surface is typical example. From different point of view the wave load is categorized into hydrostatic, hydrodynamic and inertial load components. The hydrodynamic components are proportional to moiton amplitudes. And it is known for conventionall ships the hydrostatic component plays substantial role. But this is not to say hydrodynamic components is nogligible. Hydrodynamic component is not oly itself a dominating part of wave load but also also affects hydrostatic components through ship motions. Therefore it is important to have good estimation of hydrodynamic forces around the ship. So in following the discussion will be given to the estimation of hydrodynamic components.

3.2

Linear Estimation

The basic level of approximation for the estimation of the fatigue load is the linear potential flow theory. The estimation method is grouped into the well-known 2D strip theory approximations and 3D methods like Green's function method and Panel methods (Salvesen, Tuck and Faltinsen (1970)). This linearity hypothesis is very convenient for the calculation and use of transfer functions (amplitude and phase) for the estimation of responses in frequency domain analysis. This is the most often performed analysis.

3.2.1 Strip Method It is thought the 2D flow assumption gives good estimation of the wave loads in general for the practical purpose with less computing time and less labour than those of 3D methods. It is especially so for the fatigue load estimation when considering the direct calculation of fatigue load based on spectral method requires vast amount of wave length and heading combination cases. There are many computing codes with variation in details within framework of strip theory. ITTC Sea Keeping committee(15 th 1TTC (1978) and 16th ITI'C(1981)) had given the comparative study of the codes from different organizations on the ship motion, acceleration and bending moments and concluded that the codes give fairly good agreements at least on ship motion and vertical bending moments. However there are problems that remain to be solved when applying to the fatigue load. Major ones are as follows. (1) The wave load in following sea The following sea condition is not considered crucial when ultimate load is discussed. However since in some wave length the yawing and rolling can become large in the following wave encounter condition, the lateral load such as horizontal bending moment and torsion moment tends to become

Special Task Committee 111.1

250

significantly large and may be crucial influence to the fatigue load. The accurate estimation of the load in this region is necessary when the direct calculation is made. The estimation of rolling is very important when the transverse load is discussed. As is well known, the roll damping provided by the linear potential theory, which is due to wave radiation, is usually negligible compared to the viscous roll damping. Thus it is essential to take viscous effect in to account of the roll damping. The methods for taking into account of this effect may range from empirical formula based on past experimental data to the most sophisticated CFD methods (Salui, Sarker and Vassalos(2000)) The viscous roll damping is non-linear in nature and generally proportional to the square of the roll velocity. This means that the pure sinusoidal roll response to a sinusoidal wave input given by the linear theory is no longer valid. In order to circumvent these unwelcome problem equivalent linear damping coefficient is usually used, which approximately accounts for the effects of the non-linearity. This allows us to continue using the linear equations of motion and the spectral techniques for irregular waves calculations. (2) Estimation of the load in short wave length range As the ship size increases the length of the ocean wave becomes shorter compared to the ship length. These shorter waves are not likely to excite significant ship motions but the pressure caused by these waves on the ship hull may become a source for the fatigue load. There has fatigue load for the side shell structures. The important part of the estimation is how to estimate diffraction potential. The conventional strip method, which employs approximate the diffraction potential may give erroneous pressure distribution. (Watanabe(1994))

3.2.2

Three Dimensional Method

3D method is thought to be the most rigorous way of solving the hydrodynamic problem around the ship hull in treating the forward speed effects and longitudinal variation of the load distribution along the ship hull. Typical methods are the 3D Green's function method and the Rankine source method. The Green' s function method is the method to solve linear boundary value problem using the Green's function which possesses source singularity and satisfies all the boundary conditions except the body boundary condition. The body condition is satisfied numerically by distributing the source singularity on the body boundary only. Hence, the matrix can be relatively smaller than other methods. The difficulty of this method lies in finding efficient and reliable computation scheme of Green's function. The straight forward computation requires lots of computer resources resulting numerically unreliable results. As for the numerical scheme and related problems of the Green's function refer to Inglis and Price(1982), Ohkusu and Iwashita(1989) and Takaki, Iwashita and Lin(1992) The most of the application of this method is limited to estimation of hydrodynamic forces and ship motions. The example to the estimation of the wave load is very few. Rankine source method is the method to solve the boundary value problem by distributing simple source over not only the body boundary but also the free surface boundary. It can take into account nonlinear boundary condition into account easily. This method originally was applied successfully to the wave-ship interaction problem by Nakos(1989) Sclavounos and Nakos(1990). The problem of this method lies in how to satisfy the radiation condition at the far away from the body. It could be said that the 3D method is close to the practical method for fatigue load estimation but it is requested to prove itself better than the strip method considering the computing hours required.

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251

Within this first order level of approximation, the formulation for taking into account the forward speed can be approximated when a forward speed is very low. The use of a zero speed Green function at encounter frequency is often preferred to the option of an exact Green function with forward speed. As can be seen from the following benchmark calculations and also from the study of the influence of modelling in ISSC 2000 Loads committee report discrepancies of results may occur between 2D and 3D methods and also between the two different 3D Green functions. Recent progress of CFD technique (Direct solution of Navier-Stokes equation) has widened possibility of its application to the ship motion problem. However practically feasible code is not at hand to date especially to the satisfaction of fatigue loads estimation.

3.2.3 Time Simulation Technique Time histories may be helpful when the ranges and cycles of responses are not or cannot be derived from the frequency domain or spectral analysis. This is the case when the Rayleigh distribution cannot be used because the stress response is a non-narrow band process and the cycle counting method such as the Rain-flow method accounts for the intermediate maxima within a zero-crossing period. Time histories can be derived either through convolution techniques applied to frequency domain results or direct integration of equation of motion of motions. They may also be issued from direct calculations solving the linear potential flow around the ship formulated in time domain it requires computer resources though. Rizzuto and Hansen (2000) concluded that as far as only the highest maxima within a zero crossing period are counted the Rayleigh distribution could still apply (they also outlined the differences in damage results introduced by the various cycles counting methods), and, as presented by Wirshing and Light (1980) and by Jiao and Moan (1990), theoretical or semi-empirical approaches are able to deal with cycle counting of wide band processes provided that they are Gaussian. The non-linearity which will be mentioned in the later section generally makes the loads and structure responses processes to be non-Gaussian. While the determination of ranges and number of cycles is discussed in chapter 5, these preliminary comments enlighten the possible importance of non-linear loads and usefulness of time domain simulations for damage assessment.

3.3

Non-linearity

3.3.1 Non-linearity in Global load As already mentioned, the other most important non linear behaviour of wave loads comes from the variations of the wetted area of the hull due to the combined actions of the ship motions and of the deformation of the free surface by waves. Two aspects have to be distinguished, the impulsive loads such as slams and green-water which are highly non-linear local loads with possible global consequences and the non-impulsive cyclic wave loads. It is often considered that fatigue life is mainly affected by loads in moderate seas which the ship encounter most frequently. In such moderate sea states non-impulsive type of non-linearity occurs when the hull walls are not vertical in the vicinity of the waterline. This is generally the case at both ends of the hull with the bow flare and with the stem shape (and its counter slope), and these non-linearities are responsible, among other things, for the well known hog sag asymmetry. This type of non-linearity was approximated by a quadratic theory by Jensen and Pedersen (1979) and by a third order theory by Adegeest (1996). It was also experimentally investigated for bow flare effects by Watanabe, Ueno and Sawada (1989). Such non linearity introduce higher and also lower harmonics in loads response spectra, in addition to the linear one, which are able to change significantly the fatigue damage estimation by modification of the ranges and of the number

Special Task Committee VI.1

252

of cycles. Jha and Winterstein (1998) concluded that the fatigue damage suffered by ship structures with large bow flares under non linear wave loads are 2.2 times larger than that under linear wave loads when the slope of the SN curve m=3 and 3.4 times larger when m=5. Cariou and Jancart (2003) obtained the same orders of magnitudes with m=3 and by varying the geometry (counter slope) of simplified hulls.

3.3.2

Non-linear behaviour of wave pressure distribution

The local aspect of non linearity due to the cyclic variations of the wetted hull area lies in the fact that, points of the hull surface located in the vicinity of the waterline in calm water, are alternately wet and dry in waves, which makes the pressure time histories at such locations to be non linear. And this effect should naturally be taken into account when estimating the damage to the local structure in the vicinity of the waterline. Cramer, Loseth and Bitner-Gregersen (1993) treated this problem with a simplified procedure for evaluating the long-term fatigue damage on side shell longitudinals. This last kind of non-linearity becomes worse when impacts such as slams occurs. Then they cannot be tackled by the linear flow approximation applied to the whole hull and a local fully non linear approach has better to be applied to approximate the consequent impact loads and pressure time histories. The available calculation methods and levels of approximation range, from the 2D methods initiated by Von Karman (1929) and by Wagner (1932) (improved and extended to 3D problems for instance by Watanabe (1986, 1987) or more recently by Scolan and Korobkin (2000)) to the use of emerging CFD techniques.

3.4 3.4.1

Hydroelasticity Global Effects

From view point of the fatigue load hydroelasticity may have significant meaning when vibration or wave-induced vibration are expected. There are two types of hull girder hydroelastic responses to wave loads; whipping which is the response to impact loads such as slams and springing which is the response to the non impulsive periodic wave loads. These responses are characterised by a dynamic amplification due to the excitation of deformation of the hull girder. This occurs in the whipping case provided that the intensity of the impulse is sufficient to excite significantly the hull girder vibration, and in the springing case provided that the loads spectrum density of energy at frequencies of wetted modes of resonant vibration of the hull girder is sufficient to significantly excite at least one among them. So, in the springing case, as the wave frequencies are always quite lower than the modes frequencies, the linear part of the loads spectrum will possibly contribute to this excitation only for fast ships with high frequencies of encountering waves or for very flexible structures with modes at low frequencies, this leads to suspect the non linear high frequency part of the loads spectrum to be the main cause of excitation of the hull girder modes of most conventional ships in the above defined moderate sea states. When the condition of sufficient excitation of one or several wetted modes of deformation is fulfilled, then the structural problem of calculating stresses and deformations and the hydrodynamic problem of calculating flow and hydrodynamic pressures are dependant each other and have to be coupled and solved as so-called hydroelastic problem. The basic principles of hydroelasticity of ships can be read, for instance, from Bishop and Price (1979). When its additional contribution to damage has to be considered, due to the time scale difference between impulsive and cyclic responses and to the non Gaussian resulting process, the usual frequency domain analytical formulations for the determination of ranges and number of cycles are no more well

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adapted and often replaced by time domain simulations and counting methods such as the Rainflow method needs to be adopted.

3.4.2 Local Effects Whipping and springing are the possible global hydroelastic response of the primary structure to such impulsive local loads but the local secondary and primary structures may also respond hydroelastically. An overview of slamming and ensuing hydroelastic response is given by Faltinsen (2000). The use of CFD methods to approach these local responses is emerging today (a recent illustration is given by Couty (2002)). Model tests techniques as presented by Kapsenberg and Brizzolara (1999) are also available. As for global loads, due to the time scale difference between impulsive and cyclic responses and to the non Gaussian resulting process, the determination of the resulting damage contribution is generally not practicable by an analytical frequency domain formulation.

3.5

Internal Loads

The internal loads include internal pressures on ship structures induced by solid cargos, liquid cargos and ballast water. Because of the six degrees-of-freedom motions of ships in a seaway, especially rolling of mono-hulls, these pressures are fluctuating with the accelerations and resultant stress ranges in structures will contribute to the structural fatigue damage. For partially filled tanks, the ship motions may excite liquid sloshing and induce additional dynamic internal pressures, or high slamming pressures when severe liquid motions happen, on ship structures. The tank walls could be damaged because of an instantaneous high pressure or long term accumulative fatigue. The asymmetric behaviours of external dynamic sea pressure and internal liquid cargo pressure acting on the side shell and bottom of the hull are important fatigue load of ships, Zhang and Hu (2001) discussed on this problem. For rationality sake it has to be considered that liquid sloshing motions and ship motions are mutually dependent and should be modelled as coupled mechanical systems in the calculation methods. This is particularly true for large tanks and for the roll motion of mono-hulls, one have to bear in mind that such ships can be stabilised by passive anti-roll flume tanks. Therefore neglect of such a coupling may lead to significant errors in the determination of the motions of both ship and internal fluid and consequently to erroneous magnitudes of internal and external loads. Finally that may cause significant errors in stresses and damage estimations. When the ship and internal fluid motions are moderate they can be rationally predicted by 2D or 3D potential theories as those mentioned in previous sections. Furthermore, they can detect the risk of violent motions of the liquid cargo, especially when resonance of the liquid and the ship motions occurs. But, although several linear or nonlinear potential theories and their codes were developed, the simulation of these violent sloshing motions and the determination of the induced impact loads and pressures by them is still not totally reliable. This is due to the strong non-linearity of liquid motions, and so the use of CFD methods to tackle this problem is now arising. A combination of sea-keeping, CFD and fatigue cumulative damage calculations was applied to a typical moored FSO tanker by Casella and Dogliani (1996). Finally, it is generally thought that the hydroelastic interaction of the impacting liquid cargo and its boundary structures should be taken into account.

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3.6

Other Loads

Fatigue life of ship structures are influenced not only by stress ranges due to dynamic wave loads but also by mean stresses due to welding load conditions, still water static loads, and asymmetry of nonlinear wave loads. These mean stresses affect the damages estimations by varying the characteristics of the S-N curves. This subject was addressed for instance by Marquis and Mikkola (2001). The residual stresses are usually very high but they relax rapidly with the time going on, they might have remarkable effects on low-stress high-cycle structural fatigue. Cargo ships are usually subject to different load conditions like fully loaded and ballast, for instance. The mean stresses caused by these still water loads are characterized by block-wise tensile and compressive variations, they cause the so-called low cycle fatigue and might have significant effects on the structural fatigue behaviour. The mean stresses induced by non-linear wave loads may be high especially for slamming loads but they have not yet received significant attention in the open literature. Finally we have also to mention the fatigue loads due to the vibrations of the structures excited by the propulsion plant and by the pressure fluctuations due to the rotation of propellers in the vicinity of the hull.

4.

STRESS TRANSFER FUNCTION

Stress transfer function method is based on the spectral analysis. Long-term prediction can be carried out with the use of wave spectrum and wave scattering diagram together with the stress transfer function. It is generally inevitable to conduct structural analysis using FEM to obtain the stress transfer function. 4.1

Structural Model

Structural models are chosen according to the purpose of stress analysis. Ship strength is traditionally categorized into 3 types; that is, longitudinal strength, transverse strength and local strength. FEM has been mainly adopted to investigate the transverse and local strengths so far. However, the total strength of whole ship is also tend to be evaluated using FEM recently. The modelling range and the mesh size in FEM are determined according to the stress under investigation. Structural models of short length (1/2+1/2 hold, say) are often adopted for the stress analysis in a transverse section, so that the longitudinal bending is to be excluded maintaining the structural continuity in ship longitudinal direction. On the other hand, the longer structural model (1 hold, 1/2+1+1/2 hold) or a complete hull model is used so as to consider the altemate loading condition or to investigate the stress combination due to longitudinal and transverse strength of ship. Fine mesh model or zooming technique is necessary for the local strength analysis taking the stress concentration into account, while coarse mesh is used for the nominal stress analysis. Boundary condition is very important and should be chosen properly, particularly for a partial model of ship. Forces or moments caused by longitudinal bending and local pressures are generally not balanced in such models. This fact is caused by the inconsistency of calculation method of global and local loads which mainly arises in strip theory, the neglect of the inertia force due to hull weight, the pressure nonlinearity due to the variation of wetted area along side shell, the numerical inconsistency between load analysis and structural analysis, and so on. In these cases, artificial counterbalance force is necessary to apply to the model so that the stress concentration due to the force unbalance does not affect on the stress under investigation. The location of constraint point is also important, because the actual ship is

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not supported by a point but by the buoyancy as a whole. Special attentions should be paid to the forces or moments applied at the model end of partial model particularly in the case where the model end is not rigid.

4.2

Load Application

Calculation of stress transfer function in its strict definition requires enormous number of FEM analysis because of the treatment of load non-linearity mentioned in Chapter 3, and the load combinations involved. There are two ways roughly speaking to derive stress due to hull girder bending depending on the degree of the simplification. One way is to take the simple beam theory. In this approach the bending stress can be calculated by dividing the bending moment of hull girder by the section modulus of the ship section. The other one is to use FE analysis with a longer structural model or a more detailed hull model. Stress obtained in the former case is superposed on the local stress calculated by FE analysis without the longitudinal bending effects. This approach has such a merit that the hull girder bending moment can be calculated independently with the use of various prediction methods. On the other hand, it can take the interaction of longitudinal and transverse or local stress and the stress concentration into account in the latter case although it requires a number of FE analysis. In the case of internal pressure of liquid cargo or ballast water, one of issues to be discussed is the way of defining the pressure distribution within the liquid mass. The problem is very complicated in case of partially filled tank because of the sloshing phenomenon mentioned in Chapter 3. But even fully laden tank case problem arises on how to distribute pressure on the tank wall. In the tank without free surface, zero-pressure point cannot be determined theoretically under the incompressible fluid assumption. Practical solution is to take that the pressure increases linearly from the centre of the tank toward tank sides in the case of horizontal acceleration and increases from the upper boundary downwards in the case of vertical acceleration. Because of this assumption of linear pressure distribution and the linear load-stress relation, one FE analysis for each pattern of pressure distribution due to horizontal and vertical acceleration of unit amplitude is sufficient to obtain the stress transfer function caused by internal liquid pressure of motion by superposition principle. In the case of bulk cargo, similar methodology is usually adopted assuming the bulk cargo behaves like liquid. However, further investigation may be necessary on the pressure caused by the bulk cargo since its behaviour is more complicated than the liquid cargo because of the shear stress effect and other characteristics as solid body. Derivation of transfer function for extemal wave pressure requires FE analysis applying pressure over the hull surface for one cycle. Generally speaking the transfer function of linear response can be defined as the difference between maxima and minima in sinusoidal time response. It would be very time saving if the pressure distribution to give maxima and minima of stress is known before FE analysis. However because of the complexity of the pressure contribution to the stress field there is no means to predict it. Therefore only way to derive the maxima and minima of stress is to do FE analysis for sets of pressure distribution taken for one cycle. The problem is made much harder when taking into consideration "pressure non-linearity" caused by the intermittent wetting, as was mentioned in Chapter 3. Since only the external sea pressure below still water surface can be predicted in frequency domain by conventional linear theories, it is necessary to apply fully non-linear approach. However from practical point of view, several approximate methods have been proposed to estimate the sea pressure on the hull surface above the still water surface.

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4.3

Stress Analysis

Stress analysis is usually carried out with the use of FEM. Although a time domain analysis is preferable if pressure non-linearity is to be treated in rigorous way, some simplified methods may be adopted to reduce the cost and effort of calculation. Most strict method may be a quasi-static FE analysis in time domain. Stress analyses are conducted applying instantaneous loads or pressure at several different instances during one wave encounter cycle. Obtained stresses of these moments are interpolated as a sinusoidal function in time, and the maxima and minima of stress are calculated. This method requires enormous number of calculation in general to obtain the stress transfer function. On the other hand, assuming the linearity of structural response, a linear superposition method is available to calculate the stress transfer function. Dividing the surface of ship's hull into a number of panels and applying a unit load on each panel one by one, the load-stress influence matrix, or the stress coefficient, can be calculated by FE analysis. Pressure distributing on the hull can be decomposed into discrete loads corresponding to the abovementioned unit load, and superposing the load-stress influence matrix according to the pressure distribution pattern, the stress can be obtained. Calculating the stress in this way from the instantaneous pressure distribution, the stress amplitude can be obtained by interpolating the instantaneous stress as a sinusoidal function in time. This method is called as DISAM, the DIScrete Analysis Method. This method is more useful than that of the quasi-static FE analysis in time domain if the number of panel on which unit load is applied is not so large. In the case of the load in which the linearity can be assumed, such as the intemal pressure due to liquid cargo or ballast water, stress amplitude can be calculated by applying the real and imaginary parts of the load. As for the external sea pressure, if only hull surface elements below or at the still-water plane is loaded with extemal pressure, this method can be applicable. In order to take into account of the pressure non-linearity, some techniques are necessary. Since the dynamic pressure is calculated in the frequency domain, each loaded element will experience a sinusoidal loading around a mean level of zero. Each element is therefore loaded with a unique pressure amplitude and phase, which can be expressed by a real and imaginary part using complex numbers. The fatigue load cases will therefore consist of a real and imaginary load case, which actually corresponds to the dynamic loading at the instant 0 s and at a quarter wave-encounter cycle before. The results from the real and imaginary load case will then be the real and imaginary deflections and stresses, which then again can be used to calculate the stress amplitude in any structural part of the vessel. To reduce the number of calculation in FE analysis, several simplified methods can be adopted. One is the approximate method to assume the load or the pressure distribution which corresponds to the maxima or minima of stress. The load used in the rule of classification society is categorized into this method. More accurately, it is also possible to estimate the instant when the stress will be maximum in time from the phase angle of the load or the pressure of a certain point. Another approximate method is to estimate the stress transfer function itself. For example, stress transfer function due to external pressure can be estimated by considering the weighted sum of pressure transfer functions of several points. FE analysis is only carried out at typical wave conditions in this case.

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257

StressCombination

It is possible to estimate the combined stress due to various loads by conducting a time domain stress analysis with the use of a complete hull structural model. However, this method is considered to be unrealistic in the present stage from the viewpoint of calculation cost and accuracy. Applying the method called DISAM, or the method using real and imaginary part of load in linear case, it is possible to calculate the stress take into account of the combination of load. Stress transfer function is sometimes calculated for the case of single load and by a simplified way. A problem of stress combination arises in this case. Load components have phase differences each other, and this causes the phase difference between resulting stresses. Therefore, it is necessary to take into account of the phase difference to combine these stresses. Amplitude of combined stress, O',to,a,,caused by two kinds of stresses, 00, and o 2 , are given by 2

2

O'total -- ~/ 001 nt- O'2 + 2cos(~bl - ~b2)00i 002

(4.1)

where ~1 and ~2 are the phase angles of these stresses. This means that the complex stress transfer function due to single load is necessary to obtain the combined stress, however, it may be impossible for the simplified prediction method of stress transfer function. To approximate Eqn.(4.1), a method using the phase angle of load or pressure of a certain point instead of the stress phase angle has been proposed. On the other hand, instead of using Eqn.(4.1), Eqn.(4.2) is sometimes used, where a correlation factor, or a combination coefficient, of two kind of stress components, P,2, is introduced instead of the phase angle of each stress. O'tota I =

3~

2 2 O"1 or- 0~ "~" 2/912 001 002

(4.2)

As the correlation factor can be estimated from the past calculation results or the experience of type ships, this method is practically useful if the stress transfer function is obtained by a simplified method.

5.

S H O R T AND L O N G T E R M E S T I M A T I O N

In the context of ship structural analysis by short term estimation it is meant in general the prediction of the probability distribution of one or more loads or structural responses in a period of stationariety of the process under consideration. The stationariety hypothesis is needed to enforce stable probabilistic characteristics in the variables to be studied, thus enabling their description in terms of probability distributions. This situation is realised by a stationary input to the ship system and by time-constant characteristics of the system itself. In the case of wave loads, the stationary condition implies that the ship is sailing with time-invariant operational parameters (weight distribution, speed, heading .... ) in a sea with statistical characteristics not dependent on time (generally represented by significant height, mean period, main wave direction, spreading of wave energy around the main direction .... ). Short-term predictions are therefore conditioned to the occurrence of a specific sea state and to specific values of the operational parameters accounted for in predictions themselves.

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To obtain a distribution for a time period covering different ship or sea conditions (in fatigue loading prediction usually the entire ship's life is of interest), a probabilistic characterisation of all such conditions is needed. This is used to de-condition the short-term distributions, this way obtaining a probability distribution of the variable(s) referred to the whole ship life. PLT[Y _< y]= ~P[(Y _< y)l(s,c)i]'P(s,c)i

(5.1)

i=l

PeT [Y <- Y] = probability for variable Y to be less than value y in the long term

P[~Y _
sea and ship conditions (short term prediction) = probability associated to the ith combination of sea and ship conditions

n

= total number of possible combinations of sea states and ship conditions

5.1

Short Term Predictions

Different means are available to obtain short term predictions, including direct derivation of statistics from experimental surveys in full or model scale or from time domain numerical simulations. However, frequency domain simulations represent the most common methodology for the purpose: the following paragraph will focus on them, with the aim of discussing some practical aspects of their application. The general framework of frequency domain analysis is based on (a) a characterisation of the input sea by a power spectral density distribution (spectrum), (b) the computation of a response amplitude operator (RAO) and of a response spectrum, (c) the derivation of a response probability distribution from response spectral characteristics. This well assessed procedure is often called 'spectral method'. Within the common framework of such method, different choices regarding the three steps above mentioned reflect into different results: in the following, the influence of some of the main parameters will be discussed, with the aim of giving a picture of the impact of different models.

5.1.1

Sea Spectral Shapes

The subject of sea spectral characterisation has been widely covered by all ISSC committees on environment since 1964, when the first formulation of the ISSC spectrum was presented. During the years different formulations have been proposed, based on meteorological parameters connected to the mechanism of wave generation (wind speed, fetch extension) or directly to the characteristics of the wave process itself (significant height, mean period, energy concentration at peak frequency for one or more sea components). The debate about wave spectral formulations is continuously stimulated by the opposite requirements of an accurate representation of actual sea states and the need for simple models for which a limited number of parameters can be made available from meteorological surveys. The subject is important for load predictions as different energy distributions in input to the ship system imply both different amounts and distributions of energy in the load response, with different consequences in terms of structural effects (fatigue, in particular). Without further discussion, (for more information see ISSC 1997 Committee 1.1. (1997), ISSC 2000 Committee I. 1. (2000) and previous reports of the same ISSC committee), the subject is here illustrated with an example regarding the comparison between two classical bi-parametric formulations of wave spectrum, originally proposed by Bretschneider and by the JONSWAP project. They have been recommended in the last decades by both ISSC and the ITTC (in particular 15th ITTC(1978), 17th ITrc(1984)) and are quite widely used in common practice to model wave conditions respectively in

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259

open oceans and in seas with a limited fetch area. The difference between the two formulations is in the different spreading in frequency of the same wave energy content (more concentrated in the JONSWAP spectral shape, see Figure 5.2) Different formulations for sea spectra (significant height = 5 m , zerocrossing period = 7 , 9 , 1 1 , 1 3 14

s)

...........................................................................................................................................................................................................................

12

i,

; ,'

~la 8 ",5 o ~6

~Bretschneider

: ', , ,

,

:':=.~,

.

.

.

.

JONSWAP

I.

o~

2

0

0.0

0.2

04

0.6

08

angular frequency

1.0 [rad

1.2

1.4

s "1]

Figure 5.2

TABLE 5.1 Ratios between IIl0y values obtained with JONSWAP and Bretschneider wave spectra Wave/ship encount, angle 0 30 60 90 120 150 180 210 240 270 300 330 average Zerocrossing period Tz

7 9 11 13

0.84 1.11 1.12 1.02

0.90 1.26 1.00 0.91

1.18 1.08 0.97 0.88

0.81 1.11 1.13 1.02

0.86 1.03 1.20 1.01

0.87 0.99 1.30 1.00

1.01 0.97 1.31 0.99

0.89 1.08 1.03 1.08

1.14 0.87 1.09 1.21

0.82 0.99 1.13 1.19

0.91 1.18 1.06 0.94

0.73 1.16 1.10 0.99

0.91 1.07 1.12 1.02

For the comparison of short term predictions, reference is made for the present example to the stress Response Amplitude Operators (RAOs) calculated at various heading angles for stiffener No.5, placed on the side of the ship adopted for the benchmark test described in Chapter 7. Results are reported in Table 5.1 in terms of the ratio between the spectral moment of order 0 (m0v values) of the stress response obtained with the two wave spectral formulations. Values relate to four different sea states, each identified by a mean zero-crossing period Tz and to 12 ship/wave encounter angles. The ratio depends on the relative position of the modal frequencies in the input spectra and in the RAOs. The former quantity is controlled by the mean zero-crossing period of the sea, the latter one is influenced by the ship/wave encounter angle. Values reported in this exercise refer to the specific case and depend also on the particular frequency resolution adopted in the computation; however, some general trends can be found: (a) significant variations in the energy content of the response can be obtained, for specific combinations of parameters, depending on the input spectral shape (b) such variations tend in part to 'cancel out' if equiprobable encounter angles are considered in computing an average response energy.

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5.1.2 Effectof Short- Long Crest Description Another aspect which may have an influence in short term load predictions is the directionality of the sea spectrum. The same wave energy can be modelled as concentrated in a single direction ('long crested sea') or spread over a number of wave systems proceeding at different angles with a principal direction ('short crested sea'). Real sea states do show an angular distribution of energy and the characterisation of such distribution through suitable 'spreading functions' has been debated since a long time (see ISSC 1997 Committee 1.1. (1997), ISSC 2000 Committee 1.1. (2000) and the previous reports of ISSC Committees on 'Environment'). The practical implication of introducing a short crested model is to reduce differences in the ship responses obtained with different encounter angles. Each response spectrum is in fact obtained as a weighed average of responses that the ship has to single wave systems coming from various directions distributed around the main one. TABLE 5.2 Comparison between adimensional m0v values with and without a cos 2 spreading function Wave/ship encounter angle

-~'

7 9

~ 11

13

long cr cos2 longcr cos2 longcr cos2 longcr cos2

0

30

60

90

120

150

180

210

240

270

300

330 average

0.15 0.37 0.14 0.29 0.14 0.25 0.15 0.22

0.75 0.60 0.57 0.45 0.47 0.38 0.41 0.34

0.86 0.81 0.51 0.66 0.38 0.59 0.31 0.55

1.00 0.86 1.00 0.78 1.00 0.76 0.98 0.73

0.87 0.73 0.95 0.74 0.99 0.75 1.00 0.74

0.41 0.49 0.45 0.52 0.49 0.55 0.50 0.55

0.18 0.28 0.18 0.29 0.18 0.31 0.18 0.32

0.11 0.18 0.10 0.17 0.12 0.20 0.13 0.21

0.19 0.17 0.17 0.16 0.21 0.19 0.25 0.21

0.16 0.19 0.18 0.18 0.23 0.20 0.26 0.22

0.28 0.19 0.23 0.18 0.20 0.19 0.19 0.19

0.14 0.23 0.13 0.20 0.13 0.19 0.13 0.18

0.42 0.42 0.39 0.39 0.38 0.38 0.37 0.37

i

From the point of view of wave load predictions, an important point to quantify the influence of a spreading function seems to be the dependency of the spreading function on frequency. If such dependency holds, a different energy distribution in frequency is generated for each wave direction and this implies an effect on the total amount of energy in the response spectrum inherent to the total sea state. This occurs at each single encounter angle, but also, in principle, in the average value (equiprobable angles). On the contrary, if the frequency dependence is de-coupled from angular dependence, like in the classical 'cos 2' spreading function, the introduction of a directional spectrum affects the response energy at single encounter angles, but not in the average value. Such effect is illustrated by the following example, where are compared m0v values derived from input sea spectra having a Bretschneider shape with and without a 'cos 2' spreading function. The case examined is the same of the previous paragraph. Values are adimensionalised using the maximum m0 value computed with long crested sea.

5.1.3 Effectof the Shape of Stress Range Distributions Once a stress spectrum is obtained, the next step for fatigue loading definition is the derivation of a distribution for stress ranges. To this aim, additional hypotheses are to be set: the most common choice is to postulate a Rayleigh distribution (which corresponds to the model of a Gaussian process with zero

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mean and narrow band characteristics). In this case, the distribution for stress range S is readily obtained as a function of m0u only (Equation 2) _S 2

f(S)=

S e ~8moy 0<S<~ (5.2) m0v If, in addition, the Miner's rule for fatigue damage accumulation is followed, coupled with a single slope S-N curve, the short term fatigue damage Dsv accumulated in a single stationary period of exposure to sea can also be expressed in terms of m0v.

DST = --~-C 1

2mov

tm

(5.3)

where: K,m parameters of the S-N curve for the material nc number of cycles in the period F ( ): gamma function This simplifies greatly fatigue prediction procedures, as the long term damage can be expressed as a weighed sum of elementary contributions coming from the exposure of the structure to single sea states. Equation 3 above was generalised by S.R.Winterstein. (1988) to include slightly non-Gaussian stress responses due to vibration. The effect is modelled through higher order indexes of the probability distribution (skewness and kurtosis). The same model was applied in (Jensen, J. J., Dogliani, M (1996)) to non linear wave loads induced on a containership hull both in the low and high frequency range (springing): effects of these non-linearities on fatigue damage were evaluated, showing to be very small as regards springing and of the order of 10% for low frequency components (this for a ship with considerable bow flare). Other non-linearities that play a role in fatigue predictions are those recalled in chapter 3.1.3, also representing non-Gaussian processes. In the report by Rizzuto, E., Hansen, P.F. (2000) a wide band model was used to study the influence of the bandwidth parameter e (Equation 4) on the statistics of ranges in a Gaussian process with zero mean. Signals corresponding to spectra with different ~ values were generated and analysed in the time domain to compute histograms of ranges. In the time domain, a large band stress signal can show several relative maxima and minima in a zerocrossing period: different definitions for 'range' can be given: Figure 5.2 shows results derived according to some of them. As apparent, not only the distribution but also the number of fluctuations counted may be very different, even though differences regard smaller ranges, which, in a stress signal, are less effective in fatigue damage accumulation, particularly if a double-slope S-N curve is adopted in Rizzuto, E., Hansen, P.F. (2000).

Bandwidth parameter ~ = I1

m2 m0m4

(5.4)

oo

where: m n = ~cons(co)do = spectral moment of order n 0

Large band effects are accounted for in practical cases by means of corrections like those proposed in Wirsching et.al(1980) (based on the Rainflow counting) and applied to results derived with the narrow band scheme.

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Fig. 5.2: Histograms of Ranges according to Different Definitions in Wide Band Signals (Rizzuto, E., Hansen, P.F. (2000))

5.2

Long Term Characteristics

In order to de-condition short-term predictions, a joint probability of all parameters identifying ship and environmental conditions is needed (Equation 5.1).

5.2.1

Environmental Conditions

Joint probability histograms of significant wave height and mean zero-crossing period (scatter diagrams) are usually adopted to characterise statistically the various sea conditions the ship can encounter in her life. Figure 5.3 reports the damage caused, for the test case, by sea states of different characteristics, as percentage of maximum value. Values are derived on the hypothesis of equi-probable encounter angles, of constant ship speed and of a single slope S-N curve. This figure establishes a ranking between sea state characteristics as regards fatigue damage generation. Damage is always increasing with significant height at fixed Tz (equation 3), but the dependence on zero-crossing period is more complex and the relative weight of the two variable is shown by the figure.

Figure 5.3

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263

The actual contribution to damage of the various sea states is to be obtained coupling information like those from Figure 5.3 to a scatter diagram representative for the ship's life. This rises the problem of choosing 'a priori' such representative scatter diagram. Usually data from North Atlantic areas are used as standard values, but this is recognised to be conservative for most applications: decreasing correction coefficients are often used in computations. Another option could be to associate typical routes (and inherent scatter diagrams) to particular ship typologies, in order to match better the future operational life of the ship. However, this conflicts with the concept of 'unrestricted navigation' reported in class notations, as verifications based on specific geographical areas would imply in principle to limit the ship navigation only in those areas (or in other ones less severe). Numerical results are presented in Figure 5.4 for the test case, using a North Atlantic average scatter diagram. Values are presented as percentage of the total damage for the whole scatter diagram on the hypotheses of: (a) equal number of cycles in each sea state, (b) equi-probable encounter angles, (c) constant ship speed and (d) single slope S-N curve. The last hypothesis in particular is conservative, as double slope curves imply a lower contribution from smaller amplitude stress cycles and, accordingly, less fatigue damage accumulation. The figure provides a picture of the relative contribution to damage of the various sea states contained in the scatter diagram, accounting for their probability of occurrence.

Figure 5.4

5.2.2 OperationalConditions Operational parameters are often regarded as independent from environmental conditions, but this may not always be the case, particularly in extreme bad weather. In these situations, the speed of the ship, the encounter angle formed with the main wave direction and even the distribution of ballast on board can be strongly correlated with the sea conditions. This makes complex the stochastic characterisation of these variables, which cannot be treated as independent from the characterisation of the environment. The subject is further complicated by the large degree of subjectivity present in the master's decisions about speed reduction, weather routing or change in ballast conditions. This matter has been addressed by various authors for the evaluation of extreme wave loads. In this context, periods of exposure to very

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severe (and rare) seas give the major contribution to the probability distribution of extreme responses: a proper model of these situations also in terms of operational parameters is therefore very important. For fatigue damage evaluation, as shown by Figure 5.4, the importance of moderate, more frequent, seas is higher. Accordingly, voluntary or involuntary changes of the operational conditions in heavy seas have less pronounced effects, as mentioned in ISSC 2000 Committee 111.2. (2000), where the adoption of a speed profile (velocity reduction as a function of significant height) was found to have a modest influence. In the present case, Table 5.3 below has been extracted from Figure 5.4, reporting the cumulative percentage of damage as a function of significant height only. TABLE 5.3 CUMULATIVE DAMAGE DISTRIBUTION VS SIGNIFICANT HEIGHT OF SEA (NORTH ATL. SCATTER DIAGRAM)

I-Is [m]

0.5

1.5

2.5

3.5

4.5

5.5

6.5

7.5

8.5

9.5

10.5 11.5 12.5 13.5 14.5 15.5 16.5

Dam. % 0.05 3.47 13.78 29.95 48.24 64.99 78.16 87.33 93.13 96.50 98.32 99.23 99.67 99.87 99.95 99.99 100.0 It can be seen from the table that sea states with a significant height over 10.5 m are responsible of less than 2% of the total fatigue damage experienced by the ship, while sea states between 8.5 and 10.5 m generate about 5% of the damage. Any modification in the operation profile for these environmental conditions can only have the effect of decreasing to some extent this portion of the total damage. 5.3

Remarks

Most of practical applications of the spectral method for long term predictions aimed at fatigue damage evaluations are based on simple models, represented by standard input wave spectra, linear ship response, Rayleigh short term probability distributions and long term predictions based on standard scatter diagrams and constant or almost constant ship operational parameters. More sophisticated models are available to improve several steps of this procedure, increasing the adherence to the reality. However, uncertainties affecting the whole procedure are such that it is to be evaluated if improvements on single aspects are more cost effective than a proper calibration of the base procedure (in particular as regards the derivation of stress range response operators, discussed in the previous and following chapters). The aim of this chapter has been to quantify some effects in order to make it possible to compare with other ones and concentrate efforts in reducing the largest uncertainties.

6.

DESIGN WAVE M E T H O D

The general idea with using a design wave method is to make the structural analysis simpler as the number of load cases is reduced to a minimum compared with the massive number of load cases needed for the stress transfer function method. The design wave method is in general 1. Select a representative regular wave condition (load condition, heading, speed, wavelength, wave height) 2. Calculate the stresses in the structure for that wave condition 3. Assume a representative long-term stress distribution 4. Calculate the cumulative fatigue damage The method can be made more complicated (and adequate) by including several design wave conditions and/or several loading conditions.

Fatigue Loading 6.1

265

Selectionof Design Wave Condition

The selection of a representative design waves should be based on some kind of long-term analysis. From the operational profile of the ship and the wave scatter diagrams from the sea-areas where the ship will operate long-term predictions can be made. The question is which kind of long-term predictions is needed in order to make the selection of a representative design wave condition.

6.1.1 ConsideredDesign Responses One way of selecting the design wave conditions is to look at the design responses that will cause fatigue damage. The design responses to be considered in the fatigue analysis can be 9 Vertical bending moment 9 Horizontal bending moment 9 Torsional moment 9 Transverse acceleration (racking) 9 Vertical acceleration 9 Dynamic sea pressure 9 Etc... For different ship types and different structural details, different of these design loads are important for the fatigue damage. For ships with large deck openings like Container ships the torsional moment will probably be important. For ships with large transverse accelerations and/or structure with few or open transverse bulkheads like RoRo ships the racking will probably be important.

6.1.2 Selectionof Wave Length and Height An important step in the selection of the design wave condition is the determination of wavelength. The wave amplitude will be calibrated in order to reach the target design response. Therefore, the choice of wavelength, speed and heading will determine how the target design response is combined with other load components. The method should be to choose these parameters such that the load combination becomes as realistic as possible. How this is done is not obvious, but there exist several suggestions for such a procedure. DNV suggest that the maximum value of the transfer function is chosen for the worst heading and frequency for the wave load of interest (see Det Norske Veritas (2001)). In other words the wave frequency is chosen where the RAO of the wave load of interest (say VBM) reach its maximum. ClassNK suggest a similar procedure and state that the wavelength of the regular design wave is taken as the wavelength when the response of the dominant load becomes maximum in regular waves (see ClassNK (2002)). In the BV Rules different headings and formulas for wavelength are suggested for each of the wave loads of interest (dominant load effects). For VBM head sea is selected and the wavelength is selected at peak value of vertical wave bending moment RAO without being less than 90% of ship length (see BV Rules Pt.B Ch.7 App.3 Table 5). It can be argued that the load combination must be dependant on the sea-states in which the ship operates. Therefore, it is suggested that the probability of wavelength is derived from the wave scatter diagram and that the wavelength with highest probability is chosen. This can be done in the following manner: 9 Make a probability distribution of Tz based on the wave scatter diagram. 9 Select Tz with the highest probability.

Special Task Committee 111.1

266

9 Convert Tz to the peak period Tp in the corresponding sea spectrum. Tp = 1.408 Tz. 9 Use the found Tp to calculate the corresponding wavelength.

6.1.3

Selection of Heading and Speed

Given the wavelength the next step is to select the speed and heading. It is recommended to use the ship speed equal to the most probably speed at the given probability level (say 10-4). This will in most cases be the service speed. However, for the heading the choice is often not so obvious. Take for example the VBM design response. It is often the VBM RAO in head sea (180 degrees) that is highest for a given wavelength. On the other hand the VBM RAO in 150 degrees will often be very close to the head sea RAO and the probability of heading 150 degrees is double the probability of being in head sea (assuming equal probability for all headings). So which heading should be chosen? The ultimate design load will in principle only occur once in the lifetime of the ship and therefore only for one heading. However, for the fatigue target response the situation is quite different. The fatigue target response will occur 104 times (if probability level 10-4 is used) in the lifetime of the ship and will occur for several different headings. For some design responses it may even occur for all headings. The method should again be to choose the heading such that the load combination becomes as realistic as possible. Two methods are suggested. The first is based on the transfer functions. 9 Calculate the RAO value for each heading. 9 Multiply with the probability of that heading 9 Choose the heading for which the value is highest. The second method is based on short-term prediction. 9 Calculate the RMS value for each heading. 9 Multiply with In N, where N is the numbers of wave encounters in some period (say 3 hours) 9 Multiply with the probability of that heading 9 Choose the heading for which the value is highest.

6.2

Long Term Stress Distribution

6.2.1 Stress Calculation Normally, only the time-varying loads needs to be applied in fatigue load cases. The reason is that when performing fatigue assessment calculations the need is to calculate the stress amplitude and not the mean stress level. This means that the hydrostatic and steady pressures and the gravity forces can be neglected. However, for structural details where high mean stresses are expected, the mean stress effect should be taken into account as the stress amplitude required to cause fatigue decreases at high mean stress levels. The used load application procedure is linear. This means that only hull surface elements below or at the still-water plane is loading by external dynamic sea pressure. Since the dynamic pressure is calculated in the frequency domain each loaded element will experience a sinusoidal loading around a mean level of zero. Each element is therefore loaded with a unique pressure amplitude and phase, which can be expressed by a real and imaginary part using complex numbers. The fatigue load cases will therefore consist of a real and imaginary load case, which actually corresponds to the dynamic loading at the instant 0 s and at a quarter wave-encounter cycle before. The results from the real and

Fatigue Loading

267

imaginary load case will then be the real and imaginary deflections and stresses, which then again can be used to calculate the stress amplitude in any structural part of the vessel. In the waterline region the real non-linear pressure effect should be taken into account. The negative dynamic sea-pressure load should not be taken larger than the static pressure at the considered position.

6.2.2 Long Term Load Distribution Model The long-term stress distribution is approximated with a Weibull distribution. The long-term stress distribution is assumed to be proportional with the long-term response distribution such that the same shape parameter is found for the two curves.

7.

B E N C H M A R K CALCULATION

7.1

Benchmark by Direct Method

Taking into consideration the situation mentioned in the previous section, a benchmark calculation by direct method was performed according to the calculation flow shown in Section 4. The transfer functions of nominal stresses in selected position of the ship structure are final targets of the calculation, as well as the long-term distributions of loads and of nominal stress.

7.1.1 Vesseland Calculation Condition The vessel adopted in the benchmark calculation is a Japanese VLCC, whose main characteristics are shown in TABLE 7.1. Full load and ballast conditions are taken into account, whose weight distributions are shown in Figures 7.1 and 7.2. The ship speed is assumed to be 13 knots in both loading cases. TABLE 7.1 MAIN CHARACTERISTICSOF VLCC Description Length over all Length between perpendiculars Breadth moulded Depth Draft at APP Draft at FPP Center of floatation fwd of APP Height center of buoyancy Height metacenter above keel Metacentric height Waterplane area Wetted surface Volume Displacement in seawater Block coefficient (excl skeg) Prismatic coefficient (excl skeg) Midship section coefficient Water plane coefficient

units [m] [m] [m] [m] [m] [m] [m] [m] [m] [m] [m2] [m2] [m3] [ton] [-] [-] [-] [-]

Abbr. Loa Lpp B D TA TF LcF KB KM GM Awp S V A cB Cp CM Cwp

Full Load

Ballast

320.00 60.00 28.60 21.431 21.325 159.861 11.076 24.775 8.225 17101 27910 331775 340069 0.805 0.807 0.998 0.872

320.00 60.00 28.60 10.898 8.432 171.837 4.947 34.064 21.484 15588 19806 140154 143658 0.752 0.756 0.995 0.829

268

Special Task Committee II1.1

Height center of gravity above keel Long. position CoG fwd of APP Roll radius of gyration Pitch radius of gyration Yaw radius of gyration

KG LcG kxx kyy kzz

[m] [m] [m] [m] [m]

12.580 170.594 26.400 80.574 83.253

16.55 170.388 19.800 77.220 78.802

In the benchmark calculations, ship motions and wave loads are firstly calculated in the following wave conditions. 9 Wave Headings: 7 cases from 0 (following sea) to 180 (head sea) step 30 degrees 9 Wavelengths" 14 cases (0.2, 0.4, 0.6, 0.8, 1.0, 1.2, 1.4, 1.6, 1.8, 2.0, 2.5, 3.0, 3.5, 4.0 L) 9 Wave Height: 5m (1 case)

[ton/m]

[ton/m ]

F j~

0

L,

F

'_.__1--~ ~ ~

Full Load Condition

,

,

,

,

,

,

,

i

32

64

96

128

160

192

224

256

i

288 ~ ] 320

Figure 7.1: Weight distribution (Full load condition)

7.1.2

L'

Ballast Concition

0

32

64

96

128

160

192

224

256 ~ ]288

320

Figure 7.2: Weight distribution (Ballast condition)

Calculation Method

Fatigue loads can be estimated by using a strip method or 3D panel method. Four institutions participated in the benchmark calculation, and five different calculation methods shown in TABLE 7.2 were adopted, which are indicated as A, B, C, D, and E, hereafter. TABLE 7.2 CALCULATIONMETHODUSEDIN BENCHMARK S~,mbol A

E

Calculation Method 3D panel method with zero forward speed 3D panel method with forward speed 3D panel method with zero forward speed Strip method Strip method

Fatigue Loading

269

7.1.3 Load Outputs The following outputs were provided by the mentioned computation methods. 9 Ship Motions at CG ( Surge, Sway, Heave, Roll, Pitch, Yaw) 9 Longitudinal, Horizontal and vertical Accelerations at Tank Centre in midship Center Cargo Tank: CL & 16.450m above keel Wing Cargo Tanks: 19.780m aside from CL & 16.134m above keel Ballast Tanks: 7.428m above keel & 21.025m aside from CL 9 Longitudinal Horizontal & Vertical Bending Moment at midship 9 Outer Sea Pressure Distribution at midship

7.1.4 StressOutput Two types of structural analysis are developed to obtain the nominal stress transfer functions. The first analysis is performed on the transverse reinforced rings near midship and is based on the FEM model shown in Figure 7.3 The model is loaded by internal and extemal pressures and nominal stresses are derived in 10 positions, shown in Figure 7.4. For the sake of simplicity, in the present computation pressures acting on the model were assumed as constant and corresponding to the value at station 5 (mid-length of the model and of the ship)

Figure 7.3: FEM model for trans ring stress analysis

270

Special Task Committee VI. 1

Figure 7.4: Stress investigation point in trans ring near midship

The other positions under consideration corresponded to the connection of side longitudinal stiffeners with the supporting structure. Unfortunately, the FEM model was not fine enough to describe the geometry of stiffeners in details, so, while able to capture the contribution of stiffeners to the whole hull girder bending rigidity, it was unable to model the flexural behaviour of the single longitudinals under local transverse load. The total stress in the selected locations (end of stiffeners) was therefore derived as a sum of the normal stresses obtained from the longitudinal bending moments of the hull girder plus an additional term due to the bending of the stiffener under the local action of the external pressure (and the internal one, if applicable). Stiffeners are considered as clamped to supporting structures. This model based on the superposition of hull girder and local stresses is pretty much similar to the one of simplified procedures from Class Societies, with the difference that in this case the combination is performed on each frequency with the proper phases between components. The properties of the ship side structure are assumed as follows: Outer Side Shell Plating: 18.5 mm Side Longitudinal Spacing: 880 mm Side Longitudinal: 400 x 11 (W) 150 x 14 (F)T.BAR Trans Ring Interval: 5m Midship section characteristics adopted to derive stresses from vertical and horizontal components of hull girder bending moment are: lzz = 1180760952 mm*mm*m*m Iyy = 4028990065 mm*mm*m*m Neutral Axis: 12.208m from Base Line

Fatigue Loading

271

Long Term Prediction

7.1.5

Long term distributions of nominal stresses were derived in the hypothesis of an ISSC (Bretschneider) wave spectrum, a cos 2 spreading function and a IACS North Atlantic wave scatter diagram (TABLE 7.3) with equi-probable heading angles. TABLE 7.3 WAVE SCATTER DIAGRAM Hs\Tz 0.5 1.5 2.5 3.5 4.5 5.5 6.5 7.5 8.5 9.5 10.5 11.5 12.5 13.5 14.5 SUM

7.1.6

4.5 1350 293 22 2 0 0 0 0 0 0 0 0 0 0 0 1667

5.5 8656 9860 1975 349 60 10 2 0 0 0 0 0 0 0 0 20912

6.5 11860 49760 21588 6955 1961 510 126 30 7 2 0 0 0 0 0 92799

7.5 6342 77380 62300 32265 13543 4984 1670 521 154 43 12 3 1 0 0 199218

8.5 1863 55697 74495 56750 32885 16029 6903 2701 979 332 107 33 10 3 1 248788

9.5 369 23757 48604 50991 38575 23727 12579 5944 2559 1019 379 133 44 14 5 208699

10.5 56 7035 20660 28380 26855 20083 12686 7032 3506 1599 675 266 99 35 17 128984

11.5 7 1607 6445 11141 12752 11260 8259 5249 2969 1522 717 314 128 50 26 62446

12.5 1 305 1602 3377 4551 4636 3868 2767 1746 992 515 247 110 46 27 24790

13.5 0 51 337 843 1309 1509 1408 1117 776 483 273 142 68 31 20 8367

14.5 0 9 76 224 403 533 562 501 389 270 170 98 51 26 18 3330

SUM 30504 225754 238104 191277 132894 83281 48063 25862 13085 6262 2848 1236 511 205 114 1000000

Loads Results and Comments

Calculated results are shown in Figures 7.5 to 7.14. A few remarks on each figure are summarized in the following. Ship motions Not so much differences can be seen between the results of five calculation methods in ship motion in general. Some differences are seen at the resonant frequency of heave and roll. As regards roll, it is noted that the amplitude in the response largely depends on damping characteristics. Figure 7.5 reports a selection of Response Amplitude Operators for ship motions.

2.0

[m/m]

Heave (Ful~ 180 deg. 1

[deg./m ]

_

P itch (FuR 180 deg. )1

1.5 l.O

3.5 3.0 0.0

0.5

1.5

1.0

2.0

2.5

3.0

3.5

4.0

0.5

0.0

~A

1.0

1.5

--lU- B

2.0

2.5

C

::

3.0

3.5

D

~E

4.0

4.0 Roll (Full, 60 deg. 3.0

.

.

.

.

Sway (Full, 60 deg. )

Era/m]

.

2.0 1.0 WL/S L

~.0 0.0

0.5

1.0

1.5

Figure 7.5: Examples of ship motions RAOs

2.0

2.5

3.0

3.5

4.0

Special Task Committee VI. 1

272

Accelerations (Figure 7.6) Results in terms of accelerations, being derived by superposition of several motions with phase difference, show slightly larger discrepancies.

A

+

---I--B

.....~,:+D

C

[1/s/s] [ - - N ~ .

~

A

E

HorizontalA c c e l e r a t b n P ortW ng Tank (Full, 90 deg.)

-"all-"B

C

......!~:~iD

~

E

VerttJcalA cce bratJon .] Port Wng Tank (FuR 180 deg.~

El/s/s]

.,.,.r~...~,,

0.0

0.5

1.0

1.5

2.0

2.5

3.0

3.5

0.0

4.0

0.5

1.0

1.5

2.0

2.5

3.0

3.5

4.0

Figure 7.6: Accelerations RAOs (examples)

Vertical bending moment (Figure 7.7) Differences among the five sets of results are very small in the vertical bending moment except in one case (method B). Horizontal bending moment (Figure 7.8) Discrepancies among results of horizontal bending moment are a little larger than in the case of vertical bending moment: this is considered to be caused by differences in rolling characteristics.

~

A

~

B

500

500f)00 [[MN-m/m] ~

C

.......i~i. 9D

@

E

VerticalBendhgMoment (FuR 180 d e g

I iii':~.................. ...

000 500 0 0.0

0.5

1.0

1.5

2.0

2.5

3.0

3.5

4.0

0.0

0.5

1.0

1.5

2.0

2.5

3.0

3.5

4.0

Figure 7.7: Vertical bending moment

--0-- A

~

MN-m/m]

B

C

D

~

A

E

--m-- B

C

..... !~,,,:,iD..

HorizontalBendhg Mom ent (Fun, 120 d e g

0.5

1.0

1.5

2.0

2.5

3.0

3.5

4.0

E

~WLISL~

=WL/SL~ 0.0

+

ment (Fuji, 60 d e g

0.0

0.5

1.0

1.5

Figure 7.8: Horizontal bending moment

2.0

2.5

3.0

3.5

4.0

Fatigue Loading

273

Outer pressure (Figure 7.9) Roll behaviour influences greatly the outer pressure amplitude. This explains a certain spreading in results (in Figure 7.9 examples are presented of pressure RAOs in full load at two locations on weather side: at waterline and at bilge). Results at centre line are in better agreement.

-.-__~

-.-~_

~_

__~ ....

!00+--/~ +--/~

+__~

~

--

10.0

--

'o:2

t . /---~1"--"+-- ~ lt'''~'~-- . . . .<................. , ,

0.0

0.5

1.0

1.5

2.0

,

1

" i~

................... .......

2.5

3.0

3.5

+--El

"2

-W ea~her-ff~e (Fu~ 12--O--~eg.

ivt

- -

0

4.0

Figure 7.9: External pressure RAOs (examples)

7.1.7 Stressresults Stress transfer functions in the various locations reflect differences among the various methods as regards the various load components above mentioned. For example, in the position corresponding to side longitudinal stiffener No.5 in full load (Figure 7.10), contributions from vertical and horizontal bending moments are added to those coming from direct action by external pressure in the specific location to generate stress amplitude. In this case, the dispersion in results increases in general as a consequence of the complexity of the phenomenon described. Differences can be enhanced for particular values of heading angle and/or main frequency of the encountered sea. In Figure 7.10 examples of stress range Transfer Functions for two positions on board are presented. Ii

80

--e--A

C

.-t-- B

:~, D

--gK---E

---___B.

_.___Z~ 240

150

1210

No.12 (Full, Heading=120 deg.

I

I I 3

~

/

~......

.....

0.5

1.0

1.5

2.0

2.5

3.0

3.5

--~-__A~

~ d e Lonqitudinal No.5 (Full, Headin.q=120 deg ) /\

WLISL

.,, 300011

0.0

~_

4.C

0

0.5

I 1.0

, .....

I 1.5

....".

2.0

.............' - - - r ~

2.5

3.0

3.5

4.(

Figure 7.10

When passing to a long term prediction, an averaging process between different angles and different sea frequencies takes place, sometime leading to a smoothing of discrepancies among the various methods, at least in some locations. Examples of long term results are presented in Figure 7.11, which represent the final outcome of prediction procedures.

Special Task Committee VI. 1

274

--h

B

C

................ D

--^

--E

--B

C

...............D

--E

700

Trans R lag No.12 (Fu11Load)

[ldPa] 600

500

8

7-Log @ ) 6

5

4

3

2

1

0

!8

7-Log~)6

5

4

3

2

1

0

Figure 7.11

7.1.8 Damage results A comparison among damage predictions coming from different methods was also made. The calculation was performed according to equation 3 (chapter5) for the location corresponding to the side stiffener No.5 in full load. Results are presented in the following table in terms of damages normalised with the average of values provided by the different procedures. The results show a certain dispersion: the relative values reflect the same order as Figure 7.12, but differences among the procedures are affected by the non-linearity contained in the definition of damage (in particular due to the slope of the S-N curve [value:-3]). TABLE 7.4 DAMAGE(RELATIVEVALUES) A B C D E Aver.of procedures A-E

7.2

117% 103% 101% 65% 114% 100%

Design Wave Method

In this section the results from benchmark calculations using the more simplified design wave method will be reviewed. Again the benchmark calculations are performed for the midship part of the VLCC described in section 7.1.1.

7.2.1 Considered Design Responses The first step is to select the design responses that will be important for fatigue damage of the structure. In the results presented here only the vertical bending moment is considered as an important design response to simplify matters. Horizontal bending and torsion will most probably not be important as the

Fatigue Loading

275

hull girder box is a closed and quit wide compared to its height. Racking might be important for some parts of the transverse web structure. Vertical acceleration is most probably not important for the midship part. However, dynamic sea pressure at the water level will probably be important for the side longitudinal. Based on the agreed operational profile and the IACS North Atlantic wave scatter diagram a 20 years long-term prediction of the vertical bending moment is performed for both the ballast and full-load conditions using both linear and 2 nd order strip theory. The FLS (Fatigue Limit State) target responses for the two load conditions are listed in Table 7.5 below.

TABLE 7.5 TARGETDESIGNRESPONSES(10 "4 PROBABILITY LEVEL)

Linear strip theory Full-load Ballast Target VBM Longitudinal position Number of peaks in 20 years Weibull shape parameter

5.69 GNm 160 m from AP 7.12E+07 0.92

5.64 GNm 160 m from AP 7.12E+07 0.92

2 nd order strip theory Full-load Ballast

5.40 GNm 160 m from AP 6.76E+07 0.79

5.66 GNm 160 m from AP 7.95E+07 0.84

7.2.2 Selectionof Design Wave The next step is then to select a representative regular sinusoidal design wave for each of the considered target design responses (here VBM full-load and ballast using linear and 2nd order theory). This is done using two different approaches. The first approach is the one adopted by many classification societies where the wavelength and heading are chosen where the RAO of VBM reaches its maximum. This approach is called the max RAO method in the following. The second approach is the one suggested in chapter 6 where the wavelength is derived from the period with highest probability in the wave scatter diagram. For this benchmark case the same heading is found using the two methods suggested in chapter 6. This approach is called the RMS method in the

Special Task Committee VI. 1

276

Figure 7.12: Wave-induced vertical bending moment RAOs for the full load condition

The RAO curves for the full-load condition are shown in Figure 7.12 where the design wave condition found using the 'max RAO method' also is indicated. The design wave conditions for the two load conditions are listed in Table 7.6 below. The range in wave height is the minimum and maximum wave height resulting in the target VBM response from the 5 considered hydrodynamic procedures.

TABLE 7.6 DESIGN WAVECONDITIONS

Wave length Wave length/Lpp Wave height Speed

7.2.3

Max RAO method Full-load Ballast 398.4 m 294.4 m 0.92 1.25 2.7 6.2m 4.0 5.1m 13 knots 13 knots

RMS method Full-load Ballast 223.6 m 223.6 m 0.70 0.70 7.3 - 9.6 m 6.0 7.8 m 13 knots 13 knots

Long Term Stress Distribution

The long-term nominal stress range (double amplitude) distributions are calculated using the regular design waves listed in Table 7.6 with the target responses listed in Table 7.5. Some results are compared in the next paragraph with those derived according to other methods. Therefore, the Weibull shape parameter is found by fitting a Weibull curve to the long-term response prediction around the 10-4 probability level. The results for the full-load condition are shown in Figure 7.13 below.

Fatigue Loading

1.6E+10

.

.

.

.

.

.

.

.

.

.

.

.

.

.

.

.

.

.

.

.

.

277

.

.

.

.

.

.

.

.

.

1.4E+10,

1.2E+10

-

1.0E+10 8.0E+09

-

6.0E+09

-

4.0E+09

-

2.0E+09

-

O.OE+O0 . . . . . . . 1.0E-01

~. . . . . . . .

I .........

I. . . . . . . . . .

1.0E-02 1.0E-03

I

..............

1.0E-04 1.0E-05

t.

. . . . . . .

I . . . . . . . .

1.0E-06 1.0E-07

I. . . . . . . .

1.0E-08 1.0E-Os

Probability of exceedance for individual peaks, log O

Figure 7.13" Wave-induced vertical bending moment RAO for the full load condition

7.2.4 Resultsand Considerations The long-term nominal stress range (double amplitude) distributions are calculated using the regular design waves listed in Table 7.4 with the target responses listed in Table 1. Some results are compared in the next paragraph with those derived according to other methods. In Figure 7.14 a comparison is presented between linear and 2 nd order results in the two application of the equivalent wave method to the prediction of the stress range distribution of side longitudinal 5 in full load. It is noted that the two different definitions of equivalent wave provide different results in absolute values and also as regards the relative position between curves. Definitely lower are the values from equivalent wave method (version maximum RAO value): this is probably due to the fact that the wave selected has a direction 180deg (head sea) which, for a long crested wave, means no horizontal bending and relatively small value of local pressure fluctuations. Further, by selecting the max RAO point, automatically the amplitude of the equivalent wave is minimised, being obtained as a ratio between the long-term target response and the selected RAO value. This implies small amplitude for the equivalent wave (Table 7.6) and, also for this reason, small values of stress ranges. ....

A-Lin

~B-Lin

C-Lin

D-Lin

A-NL

....

C-NL

D-NL

Side

B-NL

Longitudinal

max.

No. 5

response

(Full

E-Lin ....

Load)

E-NL

A-Lin ....

A-NL

-.

[MPal

method

Side

~B-Lin .... B-NL

C-Lin C-NL

Longitudinal

.......................D - L i n ...... D-NL

No. 5

Load)[Mea

(Full

hod

9

~

E-Lin E-NL

....

~

~

-

~

100

0 8

7-Log (~)6

5

4

3

(a) max response

2

1

0

?-Log (Q)6

5

4

3

2

1

(b) average wave

Figure 7.14 Comparison between linear and non-linear long term evaluations by equivalent waves

0

Special Task Committee VI. 1

278

A comparison among damage predictions coming from the different methods is also made. The calculation is performed for the location at side longitudinal no.5 in full load based on a one-slope S-N curve and a Weibull distribution. Results are presented in the following Table 7.7 in terms of damages normalised with the average damage values provided by the different hydrodynamic procedures using the RMS method with linear long-term evaluation. The results show some dispersion, especially between the RMS and Max RAO methods. Interestingly the results show that using a non-linear longterm evaluation instead of a linear long-term evaluation, the predicted damage is approximately halved even for this VLCC where non-linear wave load prediction usually are considered unnecessary. TABLE 7.7 FATIGUEDAMAGE(RELATIVEVALUES)

B C D E Aver.of procedures A-E

Z3

RMS method Linear analysis 120% 130% 81% 17% 152% 100%

Max RAO method Linear analysis 3% 4% 29% 4% 5% 9%

RMS method Non-linear analysis 64% 69% 43% 9% 80% 53%

Consideration of Calculated Results

In this chapter some comments on comparisons between the various results obtained on the benchmark case will be given.

7.3.1 Comparisonbetween Results from Methods of Different Level of Complexity A first comparison regards data obtained from the direct method described in chapter 4, from the design wave method described in chapter 6 (two sets of results, depending on the definition of equivalent wave) and from simplified method by class societies (chapter 2). As regards the design wave method, only linear results are taken into consideration as more homogeneous with the direct method. These three methods are of descending level of complexity. A further note regarding the design wave method is that the application described in 7.2 is to be regarded as a pilot application, in which the target response was assumed to be Vertical Bending Moment. Other choices are actually possible and should be considered, with the aim of a proper selection. The comparison is performed in the following figures on longitudinal No.5 (the only position for which all the three classes of predictions were available). Figure 7.15a (full load) shows that stress range predictions with the direct method provides long term stress range distributions quite similar to each other and almost always similar to the design wave method with average wave (RMS version). Values from simplified formulas have a larger scatter. They however embrace the values of the other two methods. In Ballast (Figure 7.15b) the agreement between the direct and design wave methods is less good (the latter giving lower values).

Fatigue Loading

.%:::~ ~:~Side

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279

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-- E-lin Average Wave

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........................ .....

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9

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I

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(b) Side longitudinal No.5 in ballast

a) Side longitudinal No.5 in full load

Figure 7.15 Comparison between long term predictions according to different classes of procedures

7.3.2 Comparisonbetween Results of Total and Component Stress Ranges A long term prediction of the stress coming from two loads A and B (VBM and HBM, or Total Bending and local pressure) can be obtained and compared with the same prediction of the two components (obtained separately). This way an empirical 'combination coefficient' CAB between the two loads can be obtained at a fixed probability level (e.g. 10 -4) as: (A+B)p=10-4 = (A)p=10-4 + CAB (B)p=10-4 It is to be noted that, in this definition, the combination coefficient CAB is actually applied to load B and load A is considered in its entire value (independently from the fact it is smaller or larger). This coefficient has not a precise significance in statistic terms, as the fact that the probability level of the single loads or of their combination is the same does not mean at all that they are contemporary. It represents only an engineering way of obtaining a total load from components analysed separately. This coefficient can be compared with the 'correlation' hypothesised in simplified formulas (table 2.2 and 2.5 of Chapter 2) This procedure has been applied to the long term stress ranges (at probability 10 -4) coming from the three loads insisting on the position of longitudinal stiffener No.5 (rows 1 to 3 in the table 7.8) and also to the two long term transversal stress range components at position 12 on the transversal ring (rows 1 and 2 in table 7.9). TABLE 7.8 LOADS AND LOAD COMBINATIONCOEFFICIENTSAT LONGITUDINALNO.5 (PROBABILITY=10 -4)

V [MPa] H [MPa] P=L [MPa] G = V&H [MPa] T = V&H&P [MPa] CGL = (T G) / L CVH = ( G - V ) / H

A 43 67 324 82 366 0.88 0.58

B 72 76 288 104 346 0.84 0.42

C 38 67 289 81 328 0.86 0.63

D 53 88 327 107 308 0.61 0.61

E 51 100 291 113 354 0.83 0.62

Special Task Committee VI. 1

280

Note: symbol '&' indicates combination, symbol '-' indicates arithmetic subtraction V = vertical bending moment; H = horizontal B.M., P = external pressure; G = global stress range coming from direct combination of V and H; T = total stress range obtained combining V, H and P; CCL and CVH = respectively combination coefficient of global and local loads and of vertical and horizontal hull girder loads as above defined. TABLE 7.9 LOADS AND LOAD COMBINATION COEFFICIENTSAT TRANSV. RING No.12 (PROBABILITY =10 -4)

I&O [MPa] I [MPa] O [MPa] Cio=(O&I- I) [ O

A 279 120 193 0.82

B 219 106 156 0.72

C 229 100 164 0.78

D 237 92 195 0.74

E 222 79 170 0.84

I = stress range coming from internal pressure; O = stress range coming from wave outer pressure C~o = combination coefficient of internal and outer pressures as above defined. All results presented above are related to the single positions examined, and cannot be in principle generalised to other locations on board. Values reported show comparatively high values of combination coefficients for the three types of load combination examined.

8.

CONCLUSIONS AND RECOMMENDATIONS

The review of the current state of the art regarding fatigue loading and the benchmark calculation presented in this report have identified points and areas that should be dealt with care in the procedure for fatigue loading. The follwing conclusions and recommendations result from the study. DIRECT METHOD 1.

A significant improvement in prediction procedures can be achieved by means of direct methods based on linear methods in the frequency domain for the computation of hydrodynamic forces and ship motions. Even though such procedures are not the most complete methodologies available, they are considered accurate enough to capture the main components of fatigue loading for conventional types of displacement ships. A reference model of this kind has been developed on a test case represented by a tanker ship. The prediction method needs to be capable to calculate with satisfactory accuracy hydrodynamic forces and wave loads for wide range of wave heading and wave length. Methods based on potential flow are available tools for the practical purpose. Among them it is thought the strip method is still the most robust and reliable tool for fatigue loading. Three dimensional method and CFD technique has a room for improvement for easy use in the design and practical stage. The above statement may be not applicable to specific ship types, for which the impact on fatigue life of nonlinear phenomena become dominant. Phenomena of this kind may be represented by wave loads in the presence of particular hull geometries, by specific types of loads (sloshing, external or internal water impacts) or by hydroelastic coupling between hull and sea water

Fatigue Loading

281

(springing, slamming, whipping). These cases have been reviewed, but not directly included in the reference model. 4.

It was felt necessary to include in the reference model non linear phenomena connected with the wet to dry transition in hull portion close to waterline. Even in an approximate way, the introduction of corrections to account for this effect is found essential for the quantification of fatigue damage in this part of the structure.

5.

Mean stresses due to cargo gravity loads (still water loads) or to mean values of non linear hydrodynamic loads have a direct influence on the fatigue behaviour, but this influence is usually accounted for through strength characteristics (S-N curves) of the material. A Finite Element structural response model coupled to the load model allows to evaluate the influence on the stress response in any location on board resulting from different inputs from the load model. In particular it allows to compare different ways of characterising the input sea and the ship operational conditions and to study the combination of load components. The base of such procedure is the evaluation of stress transfer functions in the frequency domain for each wave component.

INFLUENCING PARAMETERS 7.

Different procedures for load determination provide results with differences that increase according to the number of effects involved: this refers to short term predictions, while, on the long term, discrepancies tend to smooth away.

8.

Long term fatigue damage results to be not very affected by the spectral shape adopted for the input sea states: large influences are found on single sea states, but the long term averaging process once again tends to limit the impact.

9.

A description of sea states with a cos 2 spreading functions tends to limit the differences between responses to different heading angles, but does not affect the long term result in terms of damage, as computed according to the Miner's Rule and a single slope S-N curve (for equi-probable heading angles).

10. Comparatively rare modifications of ship operational parameters due to bad weather (speed, in first instance) have limited impact on long term damage. SIMPLIFIED METHODS 11. The adopted test case shows considerable dispersion among the results obtained from simple procedures adopted by Classification societies. Even though the final result was coherent for the test case examined, intermediate results were not. Main sources of scatter are represented by the formulations of ship motions and local loads and by their combination. More detailed models capable of dealing with these problems are needed to increase the quality of predictions. 12. Results from the design wave method proved to be very sensitive to the choice of the target response and to the selection of the design wave. 13. A general actually, a damage in load cases

problem of simplified methods for fatigue evaluation is represented by the fact that, large number of different load situations contribute, to a different extent, to the fatigue a given location. This is partially accounted for in those Rules which consider multiple and could be implemented in the design wave approach with a set of different design

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282

waves (possibly coming from different target responses). This trend tends to reduce the difference between simplified and directmethods, both in theory and in practical terms (computational effort).

REFERENCES

Adegeest, L. J. M. (1996). Third order Volterra modelling of ship responses based on regular wave results.21st Symposium on naval Hydrodydnamics, Trondheim, 189-203. Bishop, R. E. D. and Price, W, G, (1979). Hydroelasticity of ships.Cambridge University Press. Bureau Veritas (2000), Rules for Steel Ships. Cariou, A. and Jancart, F. (2003). Influence of modelling on calculated sea loads on ships. Marine Structures, 16:2. Casella, G. and Dogliani, M. (1996). Evaluation of sloshing-induced fatigue damage on a FSO tanker. Proceedings of the 6th International Offshore and Polar Engineering Conference (ISOPE), Los Angeles, CA, USA, 1, 1,304-313, China Classification Society (2001). Guidance Notes GD01-2001, Guidance Notes for Fatigue Strength of ship Structure. Clarke, J. D. (1986). Wave loading in warships. Advances in Marine Structures., Elsevier Clarke, J. D. (1987). Prediction of fatigue cracking in warship hulls. Proceedings of the third PRADS, Trondheim. ClassNK (2002). Guidelines for Bulk Carrier Structures, Nippon Kaiji Kyokai, Japan Couty, N. (2002).Three-dimensional modelization of slamming impacts with an explicit finite element software. Application to fast ships. ICCAS 2002. Cramer, E. H., Loseth, R. and Bitner-Gregersen, E. Fatigue in side shell longitudinals due to external wave pressure. Proceedings of the 12th International Conference on Offshore Mechanics and Arctic Engineering (OMAE), Glasgow, Scotland, England, 2, 267-272. Det Norske Veritas (2001). Fatigue Assessment of Ship Structures, Classification Notes No. 30.7, Det Norske Veritas, Hcvik, Norway Det NorskeVefitas (2001). Classification Notes 30.7, Fatigue Assessment of Ship Structures. Faltinsr O. M. (2000). Hydroelastic slamming. Journal of Marine Science and Technology, 5, 49-65. Fukasawa T., Mofikawa M., Fukuoka T., Shibazaki K. and Ito A. and Tomita Y. (2000). On the Simplified Method to Predict the Long-term Probability of Occurrence of Stress caused by Wave Pressure acting on a Ship. Int. Conf. on Ship and Shipping Research (NAV2000) 2, 7.4.1-7.4.10. Fukasawa, T. (2000). Simplified prediction method of long-term probability of occurrence of stress fluctuation caused by wave pressure. Trans. Korea Ship and Offshore Structures Congress 14:2. 13-20. Germanischer Lloyds (2000). Rules and Guidelines 2000. Gu, X. K. and Moan, T. (2002). Long term fatigue damage of ship structures under non linear wave loads. Marine Technology, 39:2, 95-104 Gu, X. K., Shen, J. W. and Moan, T. (2001). Time domain simulation of non linear responses of ships in waves. Submitted to Journal of Ship Research. Guedes Soares, C. and Moan, T. (1991). Model uncertainty in the long term distribution of wave induced bending moments for fatigue design of ship structures. Marine Structures, 4:4. Hansen, P. F. and Thayamballi, A. K. (1995). Fatigue damage considering whipping arising from slamming. Fourteenth OMAE, Copenhagen. Huss, M. (1987). Combined wave induced stresses in a Lo/Lo Container Ship. Rep. TRITA-SKP 1059. IACS. (1999). Recommendation No. 56, Fatigue Assessment of Ship Structures. IACS. (2001). Unified Rule S 11. Longitudinal Strength Standards, Rev.2. Inglis,R.B. and Price, W.G. (1981). A Three Dimensional Ship Motion Theory; Comparison between Theoretical Predictions and Experimental Data of the Hydrodynamic Coefficients with Forward Speed. Trans. RINA, 124

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283

ISSC 1997 Committee 1.1 (1997). Report on "Environment". ISSC 1997 Committee 1.2 (1997). Report on "Loads". ISSC 1997 Committee 111.2 (1997). Report on "Fatigue and Fracture". ISSC 2000 Committee 1.1 (2000). Report on "Environment". ISSC 2000 Committee 1.2 (2000). Report on "Loads". ISSC 2000 Committee 111.2 (2000). Report on "Fatigue and Fracture". ISSC 2000 Committee 111.2(2000). Report on Fatigue and Fracture. 15tlaITTC (1978). Seakeeping Committee Report Proceedings of15 th ITTC. 16th lq~C (1981). Seakeeping Committee Report Proceedings of 16th ITTC. 17th 1T/'C (1984). Seakeeping Committee Report Proceedings of 17th ITI'C. Iwashita, H. and Ohkusu, M. (1992) Green Function Method for Ship Motions at Free Speed, Ship Technology Research, 32, 2 Jensen, J. J. and Pedersen, P. T. (1979). Wave induced bending moments in ships. A quadratic theory. Transactions RINA, 121, 121-165. Jensen, J. J., Dogliani, M (1996). Wave-induced hull vibrations in stochastic seaways. Marine Structures 9 353-387. Jha, A. K. and Winterstein, S. R. (1998). Stochastic fatigue damage accumulation due to non linear ship loads. Proceedings of the seventeenth International Conference on Offshore Mechanics and Arctic Engineering, Lisbon. Jiao, G. Y. and Moan, T. (1990). Probabilistic analysis of fatigue due to gaussian loads processes. Probabilistic Engineering Mechanics,5:2. Kapsenberg, G., K. and Brizzolara, S. (1999). Hydroelastic effects of bow ~are slamming on a fast monohull. Fast 1999, 699-708. Kuramoto, Y., Hashimoto, K., Sato, K., Hiraki, T., Sueoka, H. and Hokazono, K. (1995). Study on Strength Analysis Method for Transverse Member of Double Hulls VLSS. Trans. West-Japan Soc. Naval Arch. 91.147-157. Kuramoto, Y., Kawamura, A., Hashimoto, K., Inoue, S. and Toki, N. (1991). Study on the LoadStress Simulation for Ship Structure in Wave. J. Soc. Naval Arch. Japan 170. 425-437. Marquis, G. B. and Mikkola, T.P.J. 2001). Effect of mean stress changes on the fatigue strength of spectrum loaded welds. Proceedings of the 8th PRADS, Shanghai, China Nagamoto, R., Takahashi, T., Kawamura, A. and Toshio Yamashita. (1980). On the design pressure for transverse members by wave induced external pressure and internal liquid pressure due to ship's motion. Trans. West-Japan Soc. Naval Arch. 59. (in Japanese) Nakos, D. (1989). Free surface mehods for unsteady forward speed flows, Proc. 4 th Int. workshop on water wave and floating bodies, Norway Registro Italiano Navale (2001). Rules for the Classification of Ships. Rizzuto, E., Hansen, P.Friis. (2000). Stress range distributions for the computation of fatigue cumulated damage in marine structures, Proc. NAV 2000, Venice (I). Salui, K., B., Sarkar, T. and Vassalos, D. (2000). An improved method for determining hydrodydnamic coefficients in roll motion using CFD techniques. Ship Technology Research, 47. Salvesen, N., Tuck, E. O. and Faltinsen, O. M. (1970). Ship motions and sea loads. Transactions SNAME, 78, 250-287. Sclavounos,P. and Nakos, D.(1990). Ship motions by Three-dimensional Rankine panel method. 18th Symposium on Naval Hydrdynamics. Scolan, Y., M. and Korobkin, A., A. (2000). Design of three-dimensional bodies subject to water impact. 15th International workshop on water waves and floating bodies,141-144. Shirakihara, H., Sato, K., Hashimoto, K. and Sueoka, H. (2001). Study on Strength Analysis Method for Transverse Member of Double Hulls VLCC. Trans. West-Japan Soc. Naval Arch. 101. 145153. Takai, M., Iwashita,H. and Xin Lin. (1992). Forces on a ship with forward speed in waves, Proc. Int. Ocean Space and resource utilization seminar and 29 th ocean engineering research workshop, Ulsan, Korea

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Toki, N., Fukushima, Y., Tozawa, S. and Wada, Y. (1994). On the characteristics and long-term prediction procedure of wave-induced pressure fluctuation on a VLCC hull. J. Soc. Naval Arch. Japan 176. Toyama, Y. and Akashi, T. (1998). A method to estimate internal pressure of fluid cargo in ship tanks. J. Kansai Soc. Naval Arch. Japan 229. Von Karman, T. H. (1929). The impact on seaplane floats during landing.NACA TN, 321. Wagner, V. H. (1932).Uber stotz und gleitvorgange an der oberflache von flussigkeiten. Zeitchrififur Angewandte Mathematik und Mechanic,12, 4, 193-215. Wang, D. H. (1998). Fatigue strength analysis of ship structures under vertical wave loads. Ph. D.

Thesis, Harbin Engineering University, China. Watanabe, E., Inoue, S., Hashimoto, K., Sato, K. and Sueoka, H. (1995). Proposal of Simplified Fatigue Design Method for Side Longitudinals. J. Soc. Naval Arch. Japan 177.391-398. Watanabe, I. and Guedes Soares, C. (1999). Comparative study on the time domain analysis of non linear ship motions and loads. Marine Structures, 12: 3, 153-170. Watanabe, I.(1994). Discussion to the Report of the ISSC 1994 Committee 1.2 on Load. Proceedings of

ISSC 1994 Watanabe, I., Ueno, M. and Sawada, H. (1989). Effect of bow flare shape on wave loads of a container ship. Journal of the Society of Naval Architects of Japan, 166, 259-266. Winterstein, S.R. (1988). Non linear vibration models for extremes and fatigue, ASCE, J.Engng.Mech. Div. 114 1772-1790. Wirshing, P. H. and Light,M. (1980). Fatigue under wide band random stresses. Journal of Structure Engineering, ASCE Structural Division, 106: 7. Zhang, A.-N. and Hu, Y.-R. (2001). Discussion on the asymmetric behaviours of external dynamic sea pressure and internal cargo pressure (In Chinese). Journal of Ship Mechanics, 5, 5, 36-42

15th INTERNATIONAL SHIP AND OFFSHORE STRUCTURES CONGRESS 2003 AUGUST 11-15, 2003 SAN DIEGO, USA

,R~sv~oF~E Sm~TRuc~

,o,

,q ~ % ~

'q .4.7 e~tvOtEGO~5~

VOLUME 1

SPECIAL TASK C O M M I T T E E VI.2

FATIGUE STRENGTH ASSESSMENT

C O M M I T T E E MANDATE Evaluate and develop procedures for estimating the fatigue strength of ship structures. Due consideration shall be given to the effect of stress combinations and to the evaluation of stress concentrations for ship hull details. The influence of fabrication imperfections, workmanship, and environmental conditions on the fatigue strength shall be discussed. The procedures shall be assessed by comparison with model tests and service experience. Recommendations for fatigue strength assessment methods shall be given.

COMMITTEE MEMBERS Chairman:

Prof. S. Berge Dr. D. Kihl Dr. I. Lotsberg Dr. S. Maherault Dr. T.P.J. Mikkola Dr. L.P. Nielsen Dr. H. Paetzold Dr.-Ing. C.-H. Shin Dr. H.-H. Sun Prof. Y. Tomita

KEYWORDS

Fatigue, fatigue strength, fatigue design, S-N curves, welded structures, ship structures, hot spot stress, structural stress, cumulative damage, finite element analysis, stress concentration factors, corrosive environment, mean stress, improvement techniques.

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CONTENTS

1 INTRODUCTION

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2 FATIGUE S T R E N G T H OF SHIP S T R U C T U R E S - G E N E R A L O V E R V IE W 2.1 Fatigue Design o f Ships . . . . . . . . . . . . . . . . . . . 2.2 Characteristics o f Ship Structural Details . . . . . . . . . . . . 2.3 Fatigue Strength Assessment o f Ship Structures . . . . . . . . . . .

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3 E L E M E N T S OF FATIGUE S T R E N G T H A S S E S S M E N T . . . . . . . . . . 3.1 Definition o f Design Stress . . . . . . . . . . . . . . . . . . . 3.2 Finite Element Modelling and Hot Spot Stress Calculation . . . . . . . . 3.2.1 B a c k g r o u n d . . . . . . . . . . . . . . . . . . . . . 3.2.2 Guidelines . . . . . . . . . . . . . . . . . . . . . . 3.3 S - N Curve Formulation . . . . . . . . . . . . . . . . . . . . 3.4 Combination o f Stresses and Link to S - N Curves . . . . . . . . . . . 3.5 Effect of Tolerances . . . . . . . . . . . . . . . . . . . . . 3.5.1 Introduction . . . . . . . . . . . . . . . . . . . . . . 3.5.2 Eccentric Butt Welds in Plates . . . . . . . . . . . . . . . . 3.5.3 Fabrication Tolerances in Plated Structures . . . . . . . . . . . . 3.6 Effect o f Corrosive Environment . . . . . . . . . . . . . . 3.7 Residual Stress and Effect o f Mean Stress . . . . . . . . . . . . 3.8 Plate Thickness Effects . . . . . . . . . . . . . . . . . . . . 3.9 Improvement o f Fatigue Life by Fabrication . . . . . . . . . . . . .

291 291 293 293 294 298 299 301 301 302 302 304 305 307 308

4 FRACTURE MECHANICS . . . . . . . . . . . . . . . . . . . . 4.1 Introduction . . . . . . . . . . . . . . . . . . . . . . . . 4.2 Fatigue Design Based on Fracture Mechanics Approach . . . . . . . . . 4.3 Working Stress and Sequence Effects . . . . . . . . . . . . . . 4.4 Crack Growth Law and Stress Intensity Factor . . . . . . . . . . . . 4.5 Initial Conditions o f Surface Crack Appearance along Weld Toe . . . . . . 4.6 Application of Fatigue Crack Growth Analysis to Fatigue Strength o f Ship Hull 4.7 Fracture Mechanics A n a l y s i s - Conclusions . . . . . . . . . . . . .

309 309 309 309 310 310 310 312

5 CUMULATIVE DAMAGE . . . . . . . . 5.1 Constant Amplitude S - N Curve . . . . 5.2 Stress Histories and Cycle Count . . . . . 5.3 Fatigue D a m a g e Accumulation . . . . .

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6 C O M P AR A TI VE STUDY . . . . . . . . . . . . . . . . . . 6.1 Bilge KnuckleJoint o f V L C C . . . . . . . . . . . . . . . . . . 6.1.1 Description o f Test and Experimental Results . . . . . . .

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Special Task Committee VI.2

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6.1.2 Fatigue Strength Assessment according to various Codes - Discussion 6.2 Side Longitudinal/Frame Connection o f an F P S O . . . . . . . . . . 6.2.1 Description o f Test and Experimental Results . . . . . . . . . . 6.2.2 Fatigue Strength Assessment according to various Codes - Discussion 6.3 Conclusions from the Comparative Study

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Fatigue Strength Assessment 1.

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INTRODUCTION

Committee VI.2 Fatigue Strength Assessment is mandated to "evaluate and develop procedures for estimating the fatigue strength of ship structures" and to give "recommendations". Key words of the mandate are: stress combinations, stress concentrations, fabrication imperfections, workmanship and environmental conditions. Considering the large number of existing design codes, recommended practices etc. on fatigue design of welded structures, and the amount of time and effort that has been input to this work, the mandate may be considered a very ambitious one. To put this statement in perspective, the latest n w recommendations for fatigue strength assessment (IIW, 2002) is a document of 150 pages. The main objective of the committee report is to contribute towards more commonality between existing codes and practices. The work has partly been undertaken as an interpretative review of published literature, where recent research activities have been perused in particular. A comparative study has also been carried out, validating fatigue strength assessment procedures of different classifying societies against relevant fatigue test data. The study was based on data from two large scale tests of rather different ship structural details; bilge knuckle joint in a VLCC (1/3 scale) and connection between a longitudinal stiffener and a transverse frame for an FPSO (1/1 scale).

Envronmen, I and function

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Loads

-~ [

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Environment and loads

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Load effect analysis

stress

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S-N curve and fract, mech.

<:=

Material properties, technol, factors, environment

Comm Vl.2

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I

Strength

Figure 1. Schematic view of a fatigue design procedure for ships, and the borderline between the two committees. There is considerable interaction and possible overlap between Comm. VI.2 and other committees, in particular Comm. 111.2 Fatigue and Fracture and Comm. VI.1 Fatigue Loading. Overlap with Comm. 111.2 seems to be unavoidable, and has been resolved by emphasising recommendations. With regard to Comm. VI.1 a borderline between the two committees was agreed, as shown by the schematic "design chain" in Figure 1.

2.

F A T I G U E S T R E N G T H OF SHIP STRUCTURES - G E N E R A L O V E R V I E W

2.1

Fatigue Design of Ships

Until approximately ten years ago, there was no explicit fatigue design of cargo ships. Based on service experience, ships were designed against ultimate loads with the implicit assumption that fatigue would - at w o r s t - be a maintenance problem. During the 1980's new grades of steel with improved weldability and increased tensile strength were developed for ship structures. Thus, the allowable stress level increased and scantlings were decreased to minimise fabrication costs and weight of structures. As is well known, fatigue strength of welded joints is relatively insensitive to the tensile strength of the steel, and it turned out that implicit fatigue design based on static strength criteria was not sufficient to cater for the increased stress level.

290

Special Task Committee VI.2

Another development was the advent of floating production, storage and offloading (FPSO) ships for oil and gas. These structures are moored offshore and operated in a stationary mode for the production life of an oil and gas field, typically up to 20 years. Docking for inspection and repair is very costly for an FPSO, and design is thus aimed at uninterrupted operation. Fatigue design should thus be based on a safe life approach. For these reasons, during the latter half of the 1990's several major classification societies have issued new or revised rules and recommendations for fatigue design of ships, ISSC Comm. 111.2 Report (1997). Various methods are prescribed for assessment of load and load effects, from simplified design wave approaches to full spectral analysis. For strength assessment, a range of schemes are applied, based on different formulations and families of S-N curves for welded joints.

2.2

Characteristicsof Ship Structural Details

For several decades, fatigue design procedures have been established for bridges, pressure vessels, etc., with fatigue strength criteria - S-N curves - based on small scale testing of planar welded joints, Gurney and Maddox (1972). During the development of oil and gas fields in the North Sea in the early 1970's it became evident that offshore structures in this harsh environment had to be designed against fatigue, using explicit procedures. The existing S-N curves covered simple joints with uniaxial loading only. For tubular joints, featuring three-dimensional geometry, combined loading and large stress concentrations and gradients, a fatigue design procedure based on hot spot stress and full scale testing was developed, Radenkoviq (1981). Ship structures pose new challenges with respect to fatigue design. Welded details in ships differ from those of bridges and offshore structures in several respects. Firstly the loading is complex, with a combination of global loads (hull bending), local loads (wave action against side shell), and service loads (changing load/ballast conditions). The combined effect of these load components for fatigue strength assessment is not clear. Secondly the geometry of the welded connections can be very complex, e.g. stiffener to transverse frame connections, reinforced cut-outs in plates, etc. with large stress gradients and no clear definition of a nominal design stress. A third factor is the threedimensional nature of the connections, leading to load shedding and in most cases of practical design, a large crack tolerance. The last statement may not be true for cases of cracking in low toughness Grade A steel. Fatigue design based on crack initiation or growth to short crack lengths may be overly conservative. At the same time, prediction methods for crack growth are uncertain, due to the load shedding effects. The last, but not least factor is the scarcity of full scale fatigue test data to back up any procedure.

Fatigue Strength Assessment

2.3

291

FatigueStrength Assessment of Ship Structures

Fatigue strength assessment procedures used by different classifying societies are summarised in Table 1, which is an update of a similar table in ISSC Comm. 111.2 Report (1997). TABLE 1 SUMMARY Class soc. ABS I~

Stress analysis guide nominal SCF simple/FE yes B V 1) simple/FE yes DNV j~ simple/FE yes GL simple yes LR simple/FE yes NK FE yes RINA 4) rsimple no IKRl~ simple/FE yes l) C h a n g e d since 1997

Nom. DoE DoE DNV IIW no BS IIW no

OF FATIGUE STRENGTH ASSESSMENT PROCEDURES

Local DoE DoE no IIW no BS IlIW DoE

Fatigue strength mean no yes yes yes no yes yes yes

thick Spec. case 16 mm 25 mm Spec. case 22 mm no no no

Corrosion method

no

Safety Program factor2~ name no SafeHull yes VeriStar yes 3~ Nauticus yes Poseidon no ShipRight yes PrimeShip

no

no

no

true

no

Sea Trust

net

time net

no net

Guidance on details yes yes yes yes (incl. rules) yes (in program) yes yes yes

2) S a f e t y f a c t o r ( t y p i c a l l y 3 or 10) on M i n e r s u m d e p e n d i n g on criticality o f detail 3) F o r s o m e w e l d c l a s s e s 4) S a m e as B V till 2 0 0 0 .

A comparative study has shown that for a rather simple weld detail, subjected to nominal ship girder stress only, the variability in fatigue life prediction using different rules and recommendations, was a factor of ten, ISSC Comm. I11.2 Report (2000). The variability in the design approaches as displayed in Table 1 is a clear indication of the need for a more unified approach.

3.

ELEMENTS OF FATIGUE STRENGTH ASSESSMENT

3.1

Definitionof Design Stress

Very important for a reliable fatigue analysis is a clear definition of the type of stress at the respective detail, which will also be the basis of the associated S-N curve. Normally, the following types of stress are applied:

Nominal stress (Snore) The nominal stress is defined according to elementary beam theory for idealised geometry, or by global FE analysis. It is calculated accordingly as force divided by the cross sectional area, bending moment or torsional moment divided by the section modulus and their superposition. For this the location and the measurement of the critical cross section and the location of crack initiation sites have to be known or assumed. The local stress concentrations created by the joint themselves and by the weld profile are not considered in the calculation of nominal stress. For a ship hull structure the effects of macrogeometric features must be included in the nominal stress, like larger cut-outs and notches, as well as reduced effective width effects, Figure 2. Structural stress (hot spot stress, geometric stress, Shot spot) Structural stress is the stress that occurs at weld connections right at the weld toe. Structural stress is a fictitious value that includes, in addition to the nominal stress, the increase in stress due to macrogeometric effects, cf. Figure 2. The stress concentration (non-linear peak stress) due to the weld geometry is not explicitly considered. This is considered in the structural S-N curve. The latter is explained in Figure 3. At a greater distance from the weld connection the elementary nominal stress

Special Task Committee 111.2

292

takes effect, then follows an area of nearly linearly increasing stress influenced by the structural geometry. Finally, the weld geometry causes another extremely high increase in stress in the immediate vicinity of the weld toe.

nom

CYnom

'

~om

(e)~~~:i-~:~ i

~/_ ~.~_______~___j:.

(f)

~

.~

o

m

Figure 2. Examples of macro-geometric effects, Niemi (2001).

Bracket

~."~

Bracket Sk I / ~./t ~'"-k AxialStressat

~ Snom Figure 3: Explanation of structural stress. For fatigue assessment, the structural stress has to be determined in the critical direction at the weld toe (hot spot), where fatigue crack initiation is expected. The hot spot stress is obtained by linear or quadratic extrapolation to the weld toe over 2 or 3 points in front of the weld toe. In general, maximum principal stress or normal stress perpendicular to the weld are used. It is not possible to determine the structural stress in structural components using analytical methods. Finite element (FEM) analysis is mostly applied. Basic guidance for application of FEM analysis is given by Niemi (1997). The structural stress acting at the weld toe is also known as hot spot stress or geometric stress.

Notch stress (Sk) The nominal stress is exceeded at all notches. Considered as notches in unwelded structures are structural discontinuities, variation in width, drillings, chases, grooves etc.. The local peak s t r e s s notch stress Sk - is calculated for the failure location of the structural detail on the basis of the linearelastic theory. In the welded structures the notch stress Sk is located at the weld toe and caused by the global geometry and the local geometry of the weld transition. Linear-elastic material behaviour is assumed. The notch

Fatigue Strength Assessment

293

stress is the total stress at the root of the notch comprising membrane stress, shell bending stress and non-linear stress peak caused by the weld. The increase of the effective stress caused by the weld with respect to the structural stress is sometimes described by a weld shape factor Kw. The fatigue effective notch factor Kf is then obtained as the product of the structural stress concentration factor Ks and the weld shape factor Kw: Sk = Kf" Snom S k : K s " Kw" Snom

3.2

Finite Element Modelling and Hot Spot Stress Calculation

3.2.1 Background Due to the nature of the stress field at a hot spot region there are questions on how to establish the hot spot stress, see Figure 4. For tubular joints the notch effect due to the weld is included in the S-N curve and the hot spot stress is derived by some extrapolation of the geometric stress to the weld toe as indicated in Figure 4. It is observed from this figure that the stress used as basis for such an extrapolation should be outside that affected by the weld notch, but close enough to pick up the geometric stress. The link between the hot spot stress and a corresponding design S-N curve was investigated in a joint industry project, Lotsberg (2001 a). Three different methods for derivation of hot spot stress (geometric stress) were investigated, Fricke (2001): 9 Linear extrapolation of stresses to the weld toe from stress at distances 0.5t and 1.5t from the toe (t = plate thickness). This method is used by some of the Classification societies, Figure 4. 9 Stress at a distance 0.5t from the weld toe, no extrapolation. 9 Linear extrapolation of stresses to the weld toe from stress at distances 0.4t and t from the toe This method is recommended by the International Institute of Welding (I/W, 1996) For analysis by shell elements the distance to the stress read out points is in most cases measured from the intersection lines as the weld is normally not included in the finite element model. For analysis by solid elements the distance to the stress read out points is measured from the weld toe. It should be noted that the finite element modelling might influence the calculated stress at the hot spot region. Parameters affecting this are: 9 type of element used 9 size of elements at the hot spot region 9 how the stresses are derived from the analysis (Gaussian stress, nodal stress etc.). The IIW method with derivation of hot spot stress with linear extrapolation from 0.4t and 1.0t was initially developed for assessment of measured stresses in fatigue testing. It requires FE analysis with a fine element mesh. For this reason the n w method for hot spot stress derivation is not found very attractive by the ship industry.

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294

Stress [

<

Notch stress Extrapolation of geometric stress to / d e r i v e the hot spot stress

Hot spot stress i

G't/2

:

t/2 Ii

l

G'3t/2

3t/2

Distance from hot spot

~i y

Region effected by the notch stress

Figure 4. Stress distribution in hot spot region. A number of classification societies have recommended derivation of hot spot stress from linear extrapolation using the stress values calculated at points 0.5t and 1.5t from the weld toe (or intersection line) from FE analysis with a well defined mesh size at the hot spot region. This procedure has been found acceptable with respect to accuracy and robustness. However, it is time-consuming in terms of engineering time. This is due to manual work involved in the derivation of the hot spot stress. It is also observed that engineers may perform the analysis differently if it is not properly described in the guideline how to establish the stress at 0.5t and 1.5t. Some derive these stresses from the Gaussian stresses and others directly from the nodal stresses. The procedure for calculation of element nodal stresses are not clearly described in the computer programs and this introduces some uncertainty when unproven programs are used for hot spot stress derivation. An example of hot spot stress derivation is shown in Figure 5. First the Gaussian stresses at the top surface are extrapolated to the section line for hot spot stress analysis. Then the stresses at 0.5t and 1.5t are calculated and finally the hot spot stress is derived by linear extrapolation. An alternative is to use a number of stress points along line A-B and fit some function for derivation of stress at 0.5t and 1.5t. Another alternative is to use the element stresses at the nodal points along line A-B as basis for such a calculation. From an engineering point of view a simple calculation of the hot spot stress at a position 0.5t from the weld toe is very attractive. It also shows small scatter in the calculated results, Fricke (2001). Normally this procedure provides somewhat smaller hot spot stress values than the method of extrapolation from 0.5t and 1.5t. However, this difference can be compensated for by using a lower hot spot S-N curve or simply by increasing the 0.5t hot spot stress by a factor 1.10 which is approximately the difference between the FAT90 curve (applicable for 0.5t and 1.5t extrapolation) and the FAT 80 curve (applicable for the 0.5t hot spot value).

3.2.2

Guidelines

The following guidance is made to the computation of hot spot stresses with local models using the finite element method: Hot spot stresses are calculated assuming linear material behaviour and using an idealised structural model with no fabrication-related misalignment. The extent of the local model has to be chosen such that effects due to the boundaries on the structural detail considered are sufficiently small and reasonable boundary conditions can be formulated.

Fatigue Strength Assessment

295

In plate structures, three types of hot spots at weld toes can be identified as exemplified in Figure 6, Fricke (2001): a) at the weld toe on the plate surface at an ending attachment b) at the weld toe around the plate edge of an ending attachment c) along the weld of an attached plate (weld toes on both the plate and attachment surface).

~ Intersection line

Extrapolated ot spot stress

-->

A

Gaussian integration point

Figure 5. Example of derivation of hot spot stress.

Figure 6. Different hot spot positions, Fricke (2001).

Models with thin plate or shell elements or alternatively with solid elements are normally used. It should be noted that on the one hand the arrangement and type of elements have to allow for steep stress gradients as well as for the formation of plate bending, and on the other hand, only the linear stress distribution in the plate thickness direction needs to be evaluated with respect to the definition of structural stress. The following methods of modelling are recommended:

296

Special Task Committee 111.2

9 The simplest way of modelling is offered by thin plate and shell elements which have to be arranged in the mid-plane of the structural components, see also Figure 7.8-node elements are recommended particularly in case of steep stress gradients. Care should be given to possible stress underestimation especially at weld toes of type b) in connection with 4-node elements, which should contain improved in-plane bending modes. The welds are usually not modelled except for special cases where the results are affected by high local bending, e.g. due to an offset between plates or due to a small free plate length between adjacent welds such as at lug (or collar) plates. Here, the weld may be included by vertical or inclined plate elements having appropriate stiffness or by introducing constrained equations for coupled node displacements. 9 An alternative particularly for complex cases is offered by solid elements which need to have a displacement function allowing steep stress gradients as well as plate bending with linear stress distribution in the plate thickness direction. This is offered, e.g., by isoparametric 20-node elements (with mid-side nodes at the edges) which means that only one element in plate thickness direction is required. An easy evaluation of the membrane and bending stress components is then possible if a reduced integration order with only two integration points in the thickness direction is chosen. A finer mesh sub-division is necessary particularly if 8-node solid elements are selected. Here, at least four elements are recommended in thickness direction. Modelling of the welds is generally recommended and easily possible as shown in Figure 8. 9 For both types of modelling, the dimensions of the first two or three elements in front of the weld toe should be chosen as follows. The element length should correspond to the plate thickness. In the transverse direction, the plate thickness may be chosen again for the breadth of the plate elements. However, the breadth over the first two elements should not exceed the "attachment width", i.e. the thickness of the attached plate plus 2 times the weld leg length (in case of type c: the thickness of the web plate behind plus 2 times weld leg length). This attachment width may also be taken for the width of solid elements in front of the weld toe, see Figure 8. The structural stress components on the plate surface should be evaluated along the paths shown in Figure 7 and Figure 8 and extrapolated to the hot spot. The average stress components between adjacent elements are used for the extrapolation. Recommended stress evaluation points are located at distances 0.5t and 1.5t away from the hot spot modelled, where t is the plate thickness at the weld toe. If the weld is not modelled, the hot spot is the structural intersection point modelled. The principal stress at the hot spot is calculated from the extrapolated component values. If the element sizes mentioned above are chosen, the stresses may be evaluated as follows: 9 In case of plate or shell elements the surface stress may be evaluated at the corresponding mid-side points. 9 In case of solid elements the stress may be extrapolated linearly to the surface centre. Normally, a linear extrapolation of the stresses to the hot spot modelled is performed. Altematively, a simplified approach without stress extrapolation is reasonable where the stress is taken at the location 0.5t away from the hot spot modelled and assessed with a reduced design S-N curve as described below. Much effort has been made during the last years to define S-N design curves for ship details, consistent with the assessment of hot spot stress. S-N data for weld toe cracking of tanker- and FPSO-specific details were developed based on fatigue tests, Kim (1997, 2001). This work included small-scale testing of 5 different types of typical FPSO details (75 specimens of each geometry) and 5 full-scale specimens of FPSO longitudinal to transverse frame connections for verification of the proposed procedure. The approach of using one hot spot S-N curve was supported by full scale tests of side longitudinals in ships, Lotsberg et al. (200 lb).

Fatigue Strength Assessment

297

Figure 7. Stress extrapolation in a three-dimensional FE model with shell elements.

Figure 8. Stress extrapolation in a three-dimensional FE model with solid elements. A design hot spot S-N curve was recommended based on a literature survey, The Welding Institute (TWI) in-house database and fatigue test data, Maddox (2001). The recommended S-N curve is linked to the finite element modelling and the method used for derivation of hot spot stress. For the linear extrapolation methods based on the FE modelling described above it is recommended to link the derived hot spot stress to the FAT90 curve IIW (1996). For stress calculated at 0.5t it is recommended to link the derived hot spot stress to that of FAT80 curve IIW (1996). The FAT90 curve may also be used together with this hot spot stress assessment. Then the hot spot stress can be calculated as 1.10 times the stress from 0.5t outside the weld toe or the intersection line. The nodal stress at the mid side node along line A-B in Figure 5 may be used for this stress using 8-node shell elements of size t by t at the hot spot region. As an alternative to this, Niemi (2001) proposes FAT90 for load-carrying fillet welds, while FAT100 is recommended for all other cases. The described procedure is considered to be conservative for details with significant stress gradients at the hot spots such as at plate bending at a hopper corner detail. Then a reduced hot spot stress may be

Special Task Committee VI.2

298

calculated taking this stress gradient into account, Kang et al. (2002). Reference is also made to the comparative study of the analysed bilge knuckle.

3.3

S-N Curve Formulation

Fatigue design criteria for structural components are based on a statistical analysis of fatigue test data obtained from constant amplitude tests. A linear relationship between log AS and log N is assumed (Figure 9), where AS is the stress range and N is the number of cycles for the fatigue life: (AS) m" N = C

or, on logarithmic form: m. log (AS) + log N = log C The exponent m is the slope of the S-N curve, taken with reference to the vertical axis. The value of the exponent m is generally between 3 and 7, i.e. a smaller slope exponent for a notched or welded component in relation to a smooth component. The S-N curve is the basis of the national and international codes for the calculation of fatigue design life. Thus for standard welded joints m = 3.0, consistent with fracture mechanics theory. The definition of a fatigue limit differs among design codes, for example N = 5" 106 (IIW, 1996) or 1.107 (NORSOK, 1998). In the I1W (1996) recommendations fatigue design curves are catalogued according to the characteristic stress range value ASR at N = 2. 106 (FAT classes). Only little is said about the scatter band of the individual joint types in the codes. In general the design S-N curves are calculated from a regression analysis giving a mean life curve with a two-sided confidence band. Design curves are assessed on the basis of a notional probability of survival of at least 95%. In most codes a "mean life curve minus two standard deviations" is used for design. To comply with the 95% probability of survival criterion the analysis must be based on a sample with at least 50 test results. In constant amplitude loading of steel components there will in general be a fatigue limit below which the fatigue life is "infinite." In design, the fatigue limit or the concept of a non-growing crack may be used only when the following conditions are met: - No cycles in the load history are above the fatigue limit No deterioration mechanism that could lead to an increase in the size of the initial defect is acting (corrosion, wear, etc.) -

Fatigue Strength Assessment

"9- .

Scatterband ~';'~'"..

2 Stdv S-N desig n~ c u r v e ' ' ~

R = Sm,n = const.

Smax

299

Mean curve

""'.. ""

x~"x

""-. """

"~'x'x """ x """ .................. ~]

k Number of cycles N [log]

Figure 9. S-N curve formulation in the codes.

A common case in fatigue design is to have a design load history with some cycles above, and some cycles below the fatigue limit. During the load history the cycles above the fatigue limit will contribute to crack initiation and growth, and the fatigue limit will be gradually lowered. For ships and offshore structures a major contribution to the fatigue limit comes from the small cycles in the spectrum, and the lowering of the fatigue limit may have a significant effect on cumulative damage. The lowering of the fatigue limit may be taken into account by a fictitious extrapolation of the S-N curve with a slope m2, Haibach (1970): m2 = 2 m l - 1 For welded joints m~ = 3.0 and m2 = 5.0. An alternative method is to apply a linear extrapolation of the S-N curve with slope ml to a fictitious cut-off level at N = 2- 107, Gurney (1976). Both these models were developed on the basis of fracture mechanics analysis. The Haibach model has been implemented in most design codes.

3.4

Combination of Stresses and Link to S-N Curves

In some design codes it is normal practice to use the principal stress within + 45 ~ normal to the weld toe together with an S-N curve for the weld for fatigue assessment as shown in Figure 11. Then the principal stress is linked to the FAT90 following the IIW(1996) notation.

Special Task Committee VI.2

300

& r-

Co._.nst_ant..am.,plitude .fati_gue l i m i t

03

R =

s

rain = const. Smax

z I

I

{

/ @1

z

1

Variable amplitude ] cut off life, / e'g" N = 10'

Number of cycles N [log]

Figure 10.

Modification of S-N curve for cumulative damage calculations.

If the angle between the principal stress is larger than 45 ~ the fatigue life with a stress component parallel with the weld together with an appropriate S-N curve that depends on the welding process should be assessed, IIW (1996). Some actual fatigue cracks can only be explained if it assumed that it is a principal stress at an angle larger than 45 ~ that is initiating and driving the crack growth. Thus IIW has suggested increasing this angle to 60 ~ Niemi (2001). In fatigue assessment of ship structures it is in most cases appropriate to assume that the calculated principal stress acts either normal to or parallel with the weld. However, there are some cases at hatches and circumferential welds around penetrations in deck structures where the new definition of principal stress and link to S-N curve may be of importance. Designers should be aware that fatigue cracking may occur at several locations around a circumferential weld. What region is most critical depends on geometrical properties as indicated in Figure 12. Fatigue cracking around a circumferential weld may occur at several locations on reinforced rings in plates depending on geometry of ring and weld size, ref. Figure 12. Fatigue cracking transverse to the weld toe in a region with a large stress concentration giving large stress parallel to the weld (Flexible reinforcement). See Figure 12a. Fatigue cracking parallel to the weld toe (Stiff reinforcement with large weld size). See Figure 12b. Fatigue cracking from the weld root (Stiff reinforcement with small fillet weld size). See Figure 12c. All these potential regions for fatigue cracking should be assessed in a design with use of appropriate stress concentration factors for holes with reinforcement, DNV-RP-C203 (2001).

Fatigue Strength Assessment

301

(3"2

ltlllItllllllllllll

0"1

45~

lllllllllllllllllll 0"2 Figure 11. Definition of stress in relation to S-N curve.

Position of fatigue crack

Comment

Fillet weld Fatigue crack growing normal to the weld toe due to large stress concentration when insert tubular is thin.

a)

1

Fatigue crack initiating from the weld toe for thicker insert tubular. The principal stress o 1 is the crack driving stress.

b) (~n

Fatigue crack in the fillet weld (initiating from the weld root) at region with large normal stress and shear stress (Small fillet weld size in relation to thickness of insert tubular or stiffening ring).

Figure 12. Fatigue cracks at reinforced cut-outs.

3.5

Effect of Tolerances

3.5.1 Introduction During the last years the reliability of equations for stress concentration factors for butt welds in design rules for floating production vessels (FPSOs) have become an important issue for fatigue design. Block sections welded from one side only and buttwelds at plates going from a thin to a thicker plate are

302

S p e c i a l Task C o m m i t t e e VI.2

found to be critical areas in terms of calculated fatigue life. A representative stress distribution at butt welds is required in order to perform a reliable fatigue design of these areas and to establish a sound basis for planning in-service inspection. A stress concentration factor can be defined as a stress magnification at a detail due to the detail itself or due to a fabrication tolerance with the nominal stress as a reference value. The maximum stress is often referred to as the hot spot stress that is used together with S-N data for fatigue life calculation. This hot spot stress is thus derived as the stress concentration factor times the nominal stress. Stress concentration factors for butt welds in plates have e.g. been presented in the DNV rules (1977, 1987) and by Maddox (1985, 1997). For plates and tubular sections SCFs have been presented by Connoly and Zettlemoyer (1993) and for tubular sections taking into account also slope of transition by Lotsberg (1998). The effect of stiffeners on the stress concentration factor for the butt welds in plated structures has been investigated by Lotsberg and Rove (2000). 3.5.2

Eccentric butt welds in plates

The stress at a butt weld between two plates as shown in Figure 13 is considered. It is assumed that the plates are welded together with an eccentricity e (and without angular mismatch). The plates are subjected to a nominal axial loading. Due to the eccentricity there will be secondary bending Sb at the weld. e Sb =3tS,om

and the stress concentration frequently referred to at an unstiffened plate weld joint is obtained from the definition given in the introduction as S C F - S"~ + S~ = 1 + 3 e Snom t

3.5.3

Fabrication tolerances in plated structures

Geometric stress concentration factors for butt welds in stiffened plates were investigated by finite element analysis using 20-node isoparametric elements. The geometry included cope holes, and is typical for stiffened plates in floating production vessels and ships. The geometry is also relevant for semi-submersibles.

Fatigue Strength A s s e s s m e n t

303

6 ~~162162

A

l

e

Notch

Figure 12. Eccentricity of butt welds. The FE analysis results were compared with the following equation for eccentricity with shift in neutral axis due to difference in plates with thickness: 0.5

SCF=I+3(tl_t2

)

t2 t~5 + t21.5

This equation was proposed by Maddox (1985). It is also used by the International Institute of Welding (1996) and in British Standard 7910 (1999). The FE analysis results were also compared with the following equation for eccentricity with shift in neutral axis due to eccentricity of plates in addition to shift in neutral axis due to joining of plates with different thickness: 0.5

S C F = 1 + 6e

tz

t~ 5 + t 21.5

0.5

+ 3 (t~ - t 2 )

t2__~____ t~5 + t 21.5

It was observed from the analyses that the cope holes imply an increased stress at that region. It is likely that a reduced area in way of the cope hole will imply an increase in SCF. This increase can be expressed as SCF = 1 +

Asection Asection - mcope hole

where: Asectio n = Sectional area of the plate and longitudinals without cope hole Acopehole = Area of cope hole in a section normal to the force.

Further details are discussed by Lotsberg and Rove (2000).

Special Task Committee VI.2

304

Tolerances are important for calculation of stress concentration factors for butt welds and cruciform joints. It is not quite obvious what tolerances are accounted for in the S-N data, as eccentricities were not measured for most of the test data that are used as a basis for derivation of the design S-N curves that are used today. In IIW (1996) it is stated that an eccentricity of 0.10t is included in the test data for butt welds and 0.15t for cruciform joints. In testing of cruciform joints the transverse plate is free in terms of boundary conditions. In a real structure it is restrained with respect to rotation in a similar manner as the two other plates. Therefore this corresponds to a tolerance of 0.3t that is accounted for in the S-N data for cruciform joints. This is in the range that normally is being accepted as a fabrication tolerance for cruciform joints. For butt welds an effective eccentricity in the equations given above can be calculated as e

= e m a x fabrication code - - O .

It

The maximum value of eccentricity in the IACS construction standard is the minimum of 0.15t and 3 mm for butt welds and t/3 for cruciform joints. Thus when the IACS standard is fulfilled with respect to these tolerances mainly butt welds at connections with different plate thickness such as shown in Figure 14 need to be investigated with respect to additional stress resulting from eccentricity. [

tlIi

[

i

t2

Figure 14. Worst case combination of eccentricity and transition in thickness.

3.6

Effect of Corrosive Environment

The effect of marine environment on fatigue strength has been researched and debated extensively over the last three decades. A major problem has been the apparent inconsistency of the effects. The following section is based on a literature survey that was made as basis for development of a Norsok standard for fatigue design of offshore structures presented by Lotsberg and Larsen (2001c). This work also formed the basis for a revision of the Norwegian standard, NS3472 (2001) and DNV-RP-C203 (2001). Since 1987 it has been generally accepted that the fatigue life of joints with cathodic protection in seawater environment is not less than that for joints in air for N>107. This is documented in the HSE (1992) background document. GrCvlen (1987) can also be cited: "The overall effect of cathodic protection upon fatigue life will depend on geometrical and loading conditions. Under long life conditions cathodic protection may give longer life than air exposure. Under conditions where shorter fatigue life is to be expected (f. ex. large initial defects), cathodic protection may give even shorter fatigue life than free corrosion." Thus, it may be concluded that there is some reduction in fatigue strength for low values of N, see HSE background document from 1992 and Berge et al. (1987). A similar conclusion was made by Mohaupt et al. (1987): "A comparison of data obtained in air with that obtained in artificial seawater, with or without cathodic protection, indicates that the fatigue life is reduced by a factor of about 2.5 to 3.0, and the beneficial effects, if any of cathodic protection, are too small to be recognised in fatigue design guidelines or codes". It should, however, be noted that this conclusion is based on test data with N < 3" 105 cycles.

Fatigue Strength Assessment

305

Based on this evidence the S-N curves in the Norwegian design standards were modified in the same manner as adopted by HSE (1995), i.e. the curves are shifted to the left in seawater environment with cathodic protection compared to those for air conditions. This means that the fatigue life, in terms of number of cycles N to failure, is reduced by a factor of approximately 2.5 for N < 106 cycles, while the fatigue life is kept equal to that in air for N > 107 cycles. For 106
.......

IltII]

!1

i

f l lllllI

I [ 1II1

Air ....

Seawater with cathodic protection Seawater free corrosion

.

.

.

.

.

.

.

i

.

o,,

=E

8, "

.

.

.

.

.

i

"~,-~. ",,

v

100

'"

~

'"

,,,

\~

'

~

...

!

.

.

.

.

.

.

.

.

.

.

~ 9

li

.

,,,

-,.'~

,,,

~'~ ~,

:

,,, ' ~ ,~,,

.

'"

~

.

.

.

~

i

l

9

I

10 1.00E+04

1.001::+05

1.00E+06

%

1.001:+07

1.00E+08

i

9

1.00E+09

Number of cycles

Figure 15 Basic S-N curve D as function of environment.

3.7

Residual Stress and Effect of Mean Stress

The procedure for fatigue analysis using S-N curves originating from The Welding Institute, Gumey (1976), is generally based on the assumption that in the welds there will be large and tensile residual stresses. Hence, design is based on the cyclic stress ranges only, linked to S-N curves representing pulsating tensile loading. The mean stress is neglected in the analysis. This has been an efficient design approach for bridges and offshore structures for the last 25 years. It is likely that some of the residual stress as introduced by the welding is removed by repetitive load cycles. However, a significant reduction in residual stress should be demonstrated in order to rely on mean stress effect in design. The largest contribution to fatigue damage for ships and offshore structures is related to the small stress ranges in the S-N curve. As pointed out by Maddox (1997) at these small cycles the residual stress has to be very low in order that part of the cycle comes into compression.

306

Special Task Committee VI.2

Most fatigue tests are performed under pulsating tension loading. Thus the S-N curves are applicable for this loading condition. Most ship structures are subjected to some permanent loads, which in some cases results in significant mean stress. When the mean stress is high and in compression, the calculated fatigue damage might be over estimated. The Japanese Ship Classification Society (NK) performed a survey of 30 second generation VLCCs (very large crude oil carriers) built in 1987-90 where a number of side longitudinals had cracked during the first years of service. It was pointed out by Yoneya et al. (1993) that most cracks were found in cargo tanks at locations 2-5 m below the loaded water line, at details with high stress concentrations. It was stated that there was practically no damage in wing ballast tanks. As explained in the report, in a full load condition, the end connections of the side longitudinals in the empty wing ballast tanks would only be subjected to compressive stresses under the action of sea loads acting on ship's sides. The end connections of the side longitudinals in the loaded cargo tanks would be subjected to constant tension under internal pressure, and the side pressure from sea would induce stress fluctuation on the tension side. As a result, the end connections of the side longitudinals in cargo tanks became more susceptible to fatigue damage. Thus in many classification companies there has been a strong drive to include a mean stress effect in the fatigue design procedure. At present the mean stress effect has been reflected in the fatigue assessment procedure by Bureau Veritas, the China Classification Society, DNV, Germanischer Loyd, Korean Register of Shipping and RINA. A typical procedure is shown in Figure 16. The mean stress effect has not been included by ABS and Lloyd's Register. The effect of mean stress was also investigated in the full scale fatigue tests of side longitudinals reported by Lotsberg et al. (2001b). Two specimens with double sided brackets were tested respectively in ballast and loaded conditions. The test results from the loaded condition, giving compression at the hot spot, were well above the S-N curve for tensile load cycling. The results are presented in terms of reduction factors on compressive stress in Figure 16. It is shown that the effect of mean stress in these tests is larger than that given in the DNV CN (2001) for fatigue assessment. Most design standards for land and offshore structures allow for mean stress effects for base material and some for Post Weld Heat Treated (PWHT) structures. The Japanese Society of Steel Construction is one of few standards that include a beneficial effect of compressive mean stress in fatigue assessment in addition to the aforementioned classification companies. A different practice with respect to mean stress effects has developed in the ship industry as compared with the offshore industry. One reason for this is difference in material strength used in the two industries. A typical material yield strength for offshore structures is in the range 3 5 0 - 500 MPa. The material yield strength used in ships is in the range 235-320 MPa. A lower yield strength in ship structures implies that local yielding at hot spots will more likely occur than if a higher material yield strength is used and that residual tensile stresses are removed during early service life. Thus, it is considered acceptable to include some beneficial effects from compressive load cycles when a fatigue assessment of a ship is performed. Also the consequence of a fatigue crack is traditionally viewed differently in a sailing ship compared with an offshore platform as a repair can easily be performed after each 5-year survey in a dock.

Fatigue Strength Assessment

307

................................................................................................................... -1--.-2...................................................................................

c m L

1

Compressive stress at hot spot

I/1 L

Curve for design used by one Classification Company

{tl C 0 0 't"' 0 t"

._o 0 "0

0.6

/

g

0.4

Full scale fatigue test result

0.2

IX:

i.

-0.5

.

.

.

.

i

0

0.5

1

Mean stress per stress range

Figure 16 Reduction factor on stress range from test compared with a design curve, DNV CN (2001).

3.8

Plate Thickness Effects

Design S-N curves are in general derived on the basis of small scale laboratory tests with plate thickness typically in the range 1 0 - 30 mm. In some ship structural details, the plate thickness may be substantially larger. Experiments and fracture mechanics analyses have shown that the plate thickness effect in fatigue of welded joints is a geometric effect, essentially affected by the following parameters (Berge, 1984): i) The magnitude of the stress concentration at the weld toe, mainly determined by the local weld geometry. ii) The gradient of stress in the plane of crack growth, mainly determined by the plate thickness. iii)The number of cycles in crack growth through the region of a steep stress gradient, relative to the total number of cycles to failure, mainly determined by the size of the initial crack and the crack ellipticity. The thickness effect may be described by the following formula

where AS0 is fatigue strength for thickness to. The exponent n will vary according to the parameters i)iii) above. For transversely loaded fillet welds n = 0.25 appears to fit data quite well. For welds with a smaller stress concentration at the weld toe (butt welds, longitudinal welds), the exponent is somewhat less. From 1984 the thickness effect has been invoked in fatigue design codes for welded steel structures. A matter of dispute is the assessment of a reference thickness to, which should represent a "mean" of the thicknesses represented in the S-N data base underlying the design curves. In most codes a thickness in the range 16 - 25 mm is assumed.

308

3.9

Special Task Committee VI.2 Improvement of Fatigue Life by Fabrication

Weld toe improvement methods have been widely investigated and have in most cases been found to give substantial increase in fatigue strength. However, there are large variations in the actual improvements achieved, and the results obtained by various methods are not always ranked in a consistent manner. One explanation for the observed variations is the lack of standardisation of the optimum method of application, but variation in the material, type of loading and type of specimens may also have influenced the results. The effectiveness of the treatment also depends heavily on the skill of the operator. Therefore it is important to perform the improvement according to a well defined procedure. Reference is made to IIW Recommendations on Post Weld Improvement of Steel and Aluminum Structures, Haagensen and Maddox (2002). The weld toe is a primary source of fatigue cracking because of the severity of the stress concentration it produces. Apart from the relatively sharp transition from the plate surface to the weld, dependent on the weld profile, the stress concentration effect is enhanced by the presence of minute crack-like flaws, extending to depths (below undercut) of a few tenths of a millimetre. Fatigue cracks readily initiate at these flaws. The weld toe improvement methods described in the IIW Recommendations rely on two main principles: i) Reduction of the severity of the weld toe stress concentration - Two methods are given, grinding and remelting by TIG dressing. The primary aim is to remove or reduce the size of the weld toe flaws and thus extend the crack initiation part of the fatigue life. A secondary aim is to reduce the local stress concentration due to the weld profile by achieving a smooth blend at the transition between the plate and the weld face. ii) Introduction of beneficial compressive residual stress - the other main approach to improving the fatigue lives of welded joints that are most likely to fail from the weld toe is to introduce compressive residual stresses in the weld toe region. Thus a tensile stress must first overcome the residual stress before it becomes damaging. It should be noted that the occasional application of high stresses, in tension or compression, can be detrimental in terms of relaxing the compressive residual stress. It should be noted that improvement of the toe will not improve the fatigue life if fatigue cracking from the root is the most likely failure mode. The considerations made in the following are for conditions where the root is not considered to be a critical initiation point. The effect from different improvement methods can not be added. Where local grinding of the weld toes below any visible undercuts is performed the fatigue life may be increased by a factor corresponding to improving the hot spot S-N curve from FAT90 to FAT120, Figure 17. Grinding a weld toe tangentially to the plate surface, as at A, will produce only little improvement in fatigue strength. To be efficient, grinding should extend below the plate surface, as at B, in order to remove toe defects. Grinding is normally carried out by a rotary burr. The treatment should produce a smooth concave profile at the weld toe with the depth of the depression penetrating into the plate surface to at least 0.5 mm below the bottom of any visible undercut (see Figure 17). The grinding depth should not exceed 3 mm or 7 % of the plate thickness, whichever is smaller. In general grinding has been used as an efficient method for reliable fatigue life improvement after fabrication. Grinding also improves the reliability of inspection after fabrication and during service life. However, experience indicates that it may be a good design practice to exclude this factor at the design stage. The designer is advised to improve the details locally by other means, or to reduce the stress range through design and keep the possibility of fatigue life improvement as a reserve to allow for possible increase in fatigue loading during the design and fabrication process.

Fatigue Strength Assessment

309

It should also be noted that if grinding is required to achieve a specified fatigue life, the hot spot stress is likely to be rather high. Due to grinding a larger fraction of the fatigue life is spent during the initiation of fatigue cracks, and the crack grows faster after initiation. This implies use of shorter inspection intervals during service life in order to detect the cracks before they become dangerous for the integrity of the structure.

1

|

A ~ ~ , B

Depth of grinding should be 0.5mm below bottom

~ofanyvisibleundercut.

/~//////A~//////~__~t ~ r Figure 17 Grinding of welds.

4.

FRACTURE MECHANICS

4.1

Introduction

To ensure that ship structural members survive the expected service life, typically 20 years and 108 stress cycles, designers apply the classification society rules to ship design. However, fatigue crack damage still occurs, often starting very early in a ship's life, with some cracks growing quickly to considerable size. Moreover, for two ships sailing the same route, fatigue crack damage can occur in one and not in the other, expressing the implicit uncertainty of fatigue design. Fatigue design rules are based on empirical S-N data obtained by constant amplitude loading, and the Miner rule. Fracture mechanics provides a more general method for fatigue damage calculation applying fatigue crack growth analysis, instead of the Miner rule. This enables evaluation of the effect of both load sequence and geometric irregularities of welds on fatigue life variability.

4.2

Fatigue Design Based on Fracture Mechanics Approach

When applying fracture mechanics approach instead of the Miner rule to fatigue design of ship hull, the following need to be determined: 1) Working stress and its sequence induced by wave load acting on structural members. 2) Crack growth law and calculation of stress intensity factor of cracked structural members. 3) Initial condition of surface crack appearance along weld toe.

4.3

Working Stress and Sequence Effects

Long-term distribution of working stress is the sum of the working stress induced by each wave during ocean-going service. It can be expressed as a Weibull distribution, and is usually approximated to be exponential.

310

Special Task Committee VI.2

The conventional procedure of fatigue design has employed the time independent random process to describe the sequence of wave loads (Figure 18a). However, it has been shown that the real sequence of wave load is not a time independent random process, but time dependent (Tomita et al., 1992). The storm model has been proposed as a new time dependent random process (Figure 18b). WoJe height

Wave height

1-" o:lmsm

I "-]~

I ozlms•'-F"

(a)

stcrm

I I "'-I" cctmsmVl

(b)

Figure 18. Time independent random process (a) and storm model (b).

4.4

Crack Growth Law and Stress Intensity Factor

The Paris-Elber law, which is based on linear elastic fracture mechanics theory, is in general employed to estimate the fatigue crack growth rate. There are some approximate equations for the stress intensity factor of surface-cracked ship structural members, and usually the modified Raju-Newman equation is used (Ship Research Panel 219, 1996; Tomita et al, 2002).

4.5

Initial Conditions of Surface Crack Appearance along Weld Toe

For fatigue crack growth analysis, the initial conditions of surface crack appearance along weld toe must be defined. In general, welding irregularities exist on the weld toe, and as a result multiple surface cracks initiate there, not a single surface crack. These cracks are different in terms of size, shape, place and number, and then grow to become long, shallow, elliptical surface cracks, interacting with each other and coalescing (Ship Research Panel 245, 2001, Kada et al., 2002).

4.6

Application of Fatigue Crack Growth Analysis to Fatigue Strength of Ship Hull

The following analysis is taken from Tomita et al. (2002) and Kada et al. (2002) and shows how fracture mechanics may be applied to predict fatigue crack growth in a ship structure. The analysis is based on the VLCC Bilge Knuckle Joint test, which is reported in the comparative analysis, Section 6.1. The test results were analysed based on crack growth analysis making use of a modified Paris-Elber's law, Equations below. Crack opening point Kop, Equation 3, was obtained from variable amplitude loading fatigue tests, Tomita et al., (2002). d a / d N = C [ ( A K ee ) ~

AK e:: =

K,,p =

K max -

- (AK eff,th/ ]

K op

0.75Km,,x -7.0

-31.5< Kmax<38.5

- 30.6

Km,x< -31.5

21.9

Kin,x__>38.5

The material constants were: C = 1.45.10 11, m = 2.75, 1996).

Unit: MPam'/2

AKeff,th

-

2.45 (Ship Research Panel 219,

Fatigue Strength Assessment

311

The Raju-Newman equations were used as an approximate equation for stress intensity factor of surface cracks at weld toe. To account for multiple crack initiation and interaction/coalescense, an interaction factor was introduced, Wang et al. (2001). In order to analyse fatigue crack growth behaviour, the initial surface crack along the weld toe must be assessed, such as the size, shape, place and numbers of surface cracks. In the analysis, the initial situation of surface crack occurrence at weld toe was determined so as to the same situation as the experimental result at 3.8.105 loading cycles, Figure 19. Then, fatigue crack growth behaviour is analysed taking the interaction effect into consideration when the multiple surface cracks become elliptic shape crack, joining with each other. The resulting calculated surface crack profile is shown in Figure 19. Figure 20 shows a comparison of surface fatigue crack contours in the through thickness direction obtained from experiment and calculation. Experimental results are good agreement with calculated crack profile and crack growth rate. Next, a single surface crack assumption with the initial crack size of 0.4mm in length and 0.2mm in depth was applied. The result is also shown in Figure 20 by a gray line. It is clear that the result is far from the experimental result. Experiment PIJ~IE BO~'OM

(:-FLOOi~ ....................

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+ ~

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UNII:

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Calculation XlO +

L0ading cycles (,104) '

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93

121

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12 i~ 20 ;~ ++.! ]2 3+ 40

Transverse Position (ram)

Figure 19. Crack growth behaviour.

In general, we have little data on the initial condition of surface crack appearance and no method devised for arriving at such determinations. It is strongly desired that the method arrived at the determinations of initial surface crack appearance is established for improving qualitatively the fatigue strength evaluation. 10 t

,g 4,

4 - Y

1

' / ' ..............

~

~it~Surface C raek oEx p e r i n e.n. t a l D a t ? )

q

_

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(0.4m m *0.2m m ) I

I

2

a

I

4

I

I

6

I

I

8

I

I

10

I

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12

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r

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14

Loading Cyc]es

I

16

I

f

18

I

I

20

I

] I

22

I

I

24

I

26 ('* 105)

Figure 20. Comparison of surface crack growth curves in the through thickness direction.

312 4. 7

Special Task Committee VI. 2 Fracture Mechanics Analysis - Conclusions

Fracture mechanics provides a tool for more detailed fatigue analysis compared to the S-N curve approach. Good agreement with relevant experimental results has been demonstrated. However, for more general applications in design, several factors should be examined more precisely, such as crack growth law, calculation of stress intensity factor for surface cracked structural members, mean stress effects, threshold stress intensity factor under random loading conditions, and shape, size and distribution of initial cracks.

5.

CUMULATIVE DAMAGE

Many conditions and circumstances that exist during the service life of a structure influence the rate and extent of fatigue crack initiation and propagation. Material, fabrication techniques and quality, post-fabrication treatments, environment, and size all affect and contribute to the way in which fatigue damage accumulates see Lieurade (2001), Infante (2000), Barro (1999). Marine structures pose additional challenges as many potential crack initiation sites are manually welded and difficult to inspect. Sites are often concealed by distributive systems, insulation, underwater, or located within large voids/tanks with limited accessibility. The problem of fatigue can be broken down into at least three basic components that must be defined in order to assess fatigue damage accumulation: the structural details (fatigue behavior), the service stresses, and the fatigue damage accumulation theory that relates the two. Although a detailed stress analysis can be performed on the structure, and the loading can usually be quantified in some manner, the fatigue phenomenon is still empirically based and requires experimental data and a hypothesis of how to relate the different pieces of information in order to assess fatigue damage. Although, the consequences of experiencing a fatigue failure in service needs to be addressed, the structural redundancy and toughness of contemporary steels allows most marine structures to be crack tolerant and repaired by convenience rather than necessity. Assumptions and factors of safety are also employed to ignore issues that complicate the assessment process (e.g. mean stress, endurance limit, nominal stress) so long as the fatigue life estimates results are not overly conservative or non-conservative. Justification for using such simplifications is generally supported by testing. Analysis of existing or retired ships to benchmark and/or adjust the fatigue assessment procedure is often a reasonable step, see Sieve (2000), Sielski (2001), Maherault, (1999). 5.1

Constant Amplitude S-N Curve

Because structural service loadings vary immensely both in complexity and form, specimens are typically tested under sinusoidal loadings to construct a basic relationship between fatigue strength and endurance. The form of the S-N curve is commonly taken as a power function that plots as a straight line in log-log space. Linear regression analysis is used to represent the log(stress) and log(life) fatigue data by a straight line. Assuming the logarithms of the failure lives follow a normal probability distribution with constant variance at any stress level, the statistical scatter of the data about the regression line defines the standard deviation. S-N curves associated with a specific probability of failure can then be represented by a line offset a certain number of standard deviations from, but parallel to, the regression line (50% probability of failure). The resulting S-N-P (applied Stress, Number of cycles and Probability of failure) curve then serves as the basic input to the fatigue damage accumulation model to predict fatigue life under a (non-sinusoidal) service loading at a specific probability of failure. Often, other loadings that reflect service conditions are used in a characterization to assess the ability of the fatigue damage accumulation model to predict the fatigue life of service loading using the constant amplitude S-N curve. Although seemingly a simple task, issues associated with the construction of an S-N or S-N-P curve such as effects of high (near yield) or low (endurance

Fatigue Strength Assessment

313

limit) stresses, addressing runouts, suspensions, and outliers, selection of probability of failure distribution, even definitions of stress and failure, can lead to difficulties interpreting and comparing independent efforts. Most every fatigue damage accumulation theory makes use of constant amplitude S-N curves. The S-N curve is usually based, quite subjectively, on experience, time and resources, end use, and availability of other existing fatigue data to supplement the effort. Even with small specimens, it is not feasible to generate an accurate S-N curve by testing a large number of specimens at many different stress levels. Compromises must be made. Recently however, a methodology has been developed that optimizes the sample size and distribution of specimens between two constant amplitude stress levels to minimize the mean squared error in predicting the service loading (Lutes et al, 2002). General formulas for optimal S-N testing are found that apply to any variable amplitude service load. Examples are provided for service loadings that have Weibull or log-normally distributed stress ranges and allows one to consider effects of sample size and proportion of specimens tested at each of two stress levels. The methodology applies to situations where the range of service load stresses lies between the two S-N curve stress levels (interpolation) as well as the situation where the range of service load stresses lies partially below the two S-N curve stress levels (extrapolation, or accelerated testing). This work may be viewed as a major extension of the work of Nelson (1990) who developed optimal S-N curve procedures for the special case in which the service load is constant amplitude. Although research continues to advance the understanding of fatigue damage accumulation, most design codes have adopted documented S-N curves that have been associated with a number of different detail configurations commonly encountered in ships and offshore structures. The S-N curves used in the design codes are well established and accepted by the community. They are based on fatigue data gathered from many sources and detail configurations, analysed, and separated into categories or classes to form a family of S-N curves. Most S-N curves are usually expressed with convenient values of slope (-3.0) and standard deviations to allow determination of S-N curves associated with a desired probability of failure.

5.2

Stress Histories and Cycle Count

Whether stresses are defined by hand calculations or finite element models in terms of hot-spot or nominal far field stress, fatigue damage models require anticipated service loadings (stresses) to be defined in some format as input. Although fatigue characterizations are generally performed using sinusoidal constant amplitude (time history) loadings, actual service stresses are usually much more complex and defined in many different ways. Fatigue service stresses can be generated from analyses by first defining an operational profile and assembling a piecewise collection of cyclic stresses in the form of exceedance curves. The stresses are typically defined by response amplitude operators in the frequency domain and generally assumed to be of short duration and Rayleigh distributed. Measured full- or model-scale wave-induced ship response data are typically collected in a time history format (time domain) and then analyzed to form power spectral density (PSD) curves (frequency domain). Stresses are usually a combination of wave-induced and slamming ship responses (Hansen, 1995). Filtering may be applied to the data to separate low frequency responses from high frequency slamming and whipping responses before post-possessing the data. In the time domain, identifying and isolating service stress cycles can be difficult when conducting fatigue damage assessments. Various analysis methods have been developed, such as Rainflow, peak counting, and mean-level crossing to address this issue; although Rainflow cycle counting is generally accepted as the method of choice for time history cycle definition.

314

Special Task Committee VI.2

Service stresses are also analyzed in histogram form to define a probability density function or to determine whether a theoretical probability density function can adequately represent the distribution. Many design codes consider a two-parameter Weibull probability distribution to represent the long term (lifetime) distribution of service stresses.

5.3

Fatigue Damage Accumulation

Fatigue damage models that can accept the service fatigue loadings directly in different formats can be most useful, provided they predict fatigue life with some degree of accuracy. Miner's Rule (Miner 1945), is one such damage model that has been adapted to service loadings in multiple formats and has become the most widely known and used fatigue damage accumulation model. Miner's rule is based on work performed much earlier by Palmgren (1924) who studied the fatigue life of ball bearings. The Palmgren-Miner hypothesis, or simply Miner's Rule, as it has come to be known, assumes fatigue damage is, for a given stress level, the ratio of applied cycles to the number of cycles to failure. Accumulated fatigue damage is then the summed contribution of cycle ratios associated with other stress levels in the service loading. When the summation of cycle ratios exceeds unity, fatigue failure is assumed to occur. Summation of cycle ratios below unity indicates the structure could remain in service longer than its anticipated service life by an amount of time equal to the original service life divided by the summation of cycle ratios. Consequently, summation ratios less than unity indicate premature failure by an amount of time equal to the original service life divided by the summation of cycle ratios. Miner's Rule is considered to be a linear fatigue damage accumulation model because the cycle ratios are not raised to any power and the stress levels are independent of one another, that is, there are no sequence effects. Miner's Rule generally produces satisfactory results for a wide range of loadings and is the basis for many fatigue design codes for welded structures used throughout the world. An attractive feature of Miner's Rule is that it, and the form the S-N curve, are mathematically consistent, and can be derived from Paris's crack growth formulation (Paris, 1973), as shown by Maddox (1991). Miner's rule can easily be applied to service loadings of many forms. The service stresses can be in the form of a time history and damage calculated at each stress level. The service stresses can be represented by a series of blocks or an exceedance curve. Alternatively, the stresses can be represented in the frequency domain by power spectral density (PSD) curves, see Lutes et al. (1990) and Rizzuto (2000), or even by a probability density function, as was demonstrated using a Rayleigh distribution by Miles (1954) and used in many design codes to determine the fatigue life for ship response to particular seaway loadings of short term exposure. Design fatigue factors (factors on fatigue life) for ship structure can range from 1 for secondary and 3 for primary structure which can be inspected, to 10 for primary structure that cannot be inspected. Summation constants for FPSOs can range from 1 for structure above the waterline to 2 or 3 for submerged structure. Owners may choose to specify additional factors of safety, e.g. NORSOK (1998). Variations on Miner's Rule can be found in the literature that like Miner's Rule, are stress (level) independent and interaction (sequence) independent, but include the ability to amplify or attenuate the accumulated damage depending on the choice of model parameter(s). Such models are therefore referred to as non-linear fatigue damage accumulation models. Other models are stress level dependent and have parameters that are dependent on the applied stress condition. Still others are both stress level dependent and stress interaction dependent. The problems associated with the stress

Fatigue Strength Assessment

315

dependent and/or stress interaction dependent models that are mainly two-fold. First, instead of relying on the basic sinusoidal constant amplitude S-N curve, a much larger experimental effort must be invested to evaluate the stress dependent parameters of the particular model to determine relationships between damage, applied stress, and previously applied stress cycles. Such an effort can quickly become formidable. Second, it is often difficult, if not impossible to apply these models to complex or random service loadings. Ship loadings are not repeatable in the time domain; they are however stationary over reasonably short periods of time and assumed to be ergodic. However, many can be adequately characterized in the frequency domain in terms of power spectral density curves (PSDs), or in the probabilistic domain in terms of probability density functions (PDFs). Although both PSDs and PDFs can be determined from time domain analysis of measured data, the original time histories are not and need not be unique. Therefore fatigue damage accumulation models that depend on unique time history analysis offer little value for evaluating the fatigue life of marine structures. 6.

COMPARATIVE STUDY

6.1

Bilge KnuckleJoint of VLCC

6.1.1 Descriptionof test and experimental results The model was a bilge knuckle section for a double hull VLCC in approximately 1/3 scale, Ship Research Panel 245 (2001). The model was about 6m in length, 5m in width, 3.6m in height and 20tons in steel weight. This is a connection part between inner bottom and bilge hopper plates. To make stress distributions of the model similar to those of the actual ship, a three-floor space in the longitudinal direction was modelled. The model was fixed to a rigid wall at the double hull side, with the ship's bottom being upside and the inner bottom being downside. The load was applied by three syncronised hydraulic jacks on the centreline of the double bottom, see Figure 21. The model was built from mild steel in accordance with NK rules. In order to initiate cracks only at the bilge knuckle section of the centre floor, the weld leg length at the bilge knuckle section at the other floors was increased, and the weld toes were ground smooth.

730L 1000,

~

~

1ill i ~ 1 1 1 1 ~

~[

I~HYDRAULIC f /

/ / /I /I/I/I) ,).) / / 4250

.,.

Figure 21. Bilge knuckle model. The fatigue loading was constructed from a block loading simulating variable amplitude loading conditions during ocean-going service. The typical mean stress alternation for a VLCC is a tensile mean stress in full load condition and a compressive mean stress in ballast condition. The applied load cycles were therefore composed of five blocks with a tensile mean stress and five similar block with a

316

Special Task Committee VI.2

compressive mean stress. The load history is shown in Table 2 and Figure 22. One loading set consisted of 10 blocks, and the number of cycles in each block was 10,000, i.e., the number of cycles of each set was 105.

p

...... ~

.......................................

Figure 22. Block program loading TABLE 2. DETAILS OF THE BLOCKLOADINGPROGRAM Block no.

1

2

3

4

5

6

7

8

9

10

AP/2 (kN)

68

136

203

136

68

68

136

203

136

68

Pm (kN)

248

248

248

248

248

-248

-248

-248

-248

-248

Measured hot spot stress amplitude during the test is shown in Figure 23. Surface fatigue cracks only initiated at the weld toe of bilge knuckle part of centre floor. After 2.105 cycles the hot spot stress at this location started to decrease, indicating crack initiation. Applying the 5 % drop criterion, fatigue crack initiation had taken place at 4.8.105 cycles. The test ran for 1.23.106 cycles. At this stage the fatigue crack surface was investigated. Contours of the crack at various stages are shown in Figure 24. At 38 or 48.104 cycles of fluctuating load, many surface fatigue cracks had been initiated at the weld toe. Subsequently they were combined into a shallow and broad surface crack, which propagated from around the centre of the transverse floor in the depth direction. At end of test the crack was nearly through thickness. 1

Z \

0.8 0.6 ~

w~.--

~. 0.4

"~i:,.~

-~PLATE BOTTOM ~ ~ ' x ~

"

~9 0.2 ~.. <1 0 \ to -0.2 m : -0.4

UNIT: x 10(

q' b" % ~ ..,, I b +-.-

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12

16 20

24

Transverse Position (mm)

<1 -0.6 -0. 8 -1

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40

"5 60

80

100

120

140

NXl0+ Figure 23. Hot Spot Stress Amplitude

Figure 24. Crack growth behaviour

21

32

16

lO

Fatigue Strength Assessment 6.1.2

317

Fatigue strength assessment according to various codes - discussion

The VLCC bilge knuckle model was analysed using design procedures of ABS, DNV, GL, KR and NK. The material is quite extensive and a summary is given in Table 3. In all the models the size of the FE mesh close to the intersection was t by t. All participants applied a hot spot stress assessment using 1.5t - 0.5t linear extrapolation to the assumed weld toe, the location varying somewhat between the models. DNV carried out a comparison between 8-node and 4-node shell elements. The 8-node model gave a somewhat smaller SCF. DNV is recommending this model for hot spot stress analysis. A mean stress correction similar to the one shown in Figure 16 was applied by GL and DNV. NK has developed a special procedure for mean stress correction, in which shake-down of residual stress is taken into account. With this model only the three largest blocks with tensile stress ranges (Figure 22) became damaging. The procedure is linked to an S-N curve with slope parameter 4.63, and a relatively large FAT value. DNV applied a weld notch factor Kw = 1.5, and a FAT 130 S-N curve. ABS, GL and KR did not apply a correction for weld shape, and uses lower FAT curves. Neither did NK apply a correction for weld shape, but used an S-N curve on a different format, due to the mean stress correction. The criterion applied in the test for detection of fatigue initiation is much too sensitive compared to practical inspection criteria for fatigue cracks, and represents a very conservative measure of fatigue life. It is interesting to note that all estimates based on design criteria were on the conservative side with respect to fatigue initiation life. The NK prediction, which is based on mean life S-N curve is somewhat non-conservative, and within a reasonable scatter-band for fatigue life. Assuming a through thickness crack to represent design fatigue life, the design calculations show a utilisation factor on life in the range 0.14 - 0 . 2 3 . The mean life prediction by NK corresponds to a utilisation factor of 0.67. For most fatigue strength classes the design curve and the mean life curve are approximately a factor of three apart in fatigue life, Gurney (1976). Hence, the utilisation factor for the NK mean life prediction is in the same range as for the design predictions. It is observed that the fatigue life predictions are remarkably uniform, in spite of the significant discrepancies through the individual stages of each calculation. TABLE 3. SUMMARYOF RESULTSFROMCOMPARATIVESTUDY, BILGEKNUCKLEJOINTOF VLCC. Class FEM soc. software

FEM SCFcalc element (MPa/kN)

SCFcalc/ SfFtest

Kw

ABS

4-node shell 8-node shell 4-node shell 4-node shell3) 8-node shell3)

0.62

0.80

1.0

0.72

0.93

0.78

1.01

0.80

1.03

0.82

1.06

ANSYS SESAM

DNV SESAM ANSYS GL ANSYS

S-N curve 1)

for D= 1.0

Initiation

Through thickness

FAT 80

2.1.105

0.44

0.17

1.5

FAT 130

2.2.105

0.46

0.18

1.5

FAT 130

1.7.105

0.35

0.14

1.0

FAT 100

2.4.105

0.50

0.20

318

Special Task Committee VI.2

MARC

KR NK

20-node solid4) NASTRAN 4-node shell NASTRAN 4-node shell

0.81

1.05

0.64

0.81

1.0

FAT 91

2.8.105

0.58

0.23

0.61

0.79

1.0

FAT 1792)

8.2.105

1.71

0.67

~) All curves are design curves with slope parameter m = 3, except NK curve. FAT designation is stress range at 2.106 cycles. 2) An equivalent stress range was calculated from mean and residual stress corrections, and applied with an S-N curve with m = 4.63. The mean stress correction was applied under the assumption that local residual stress will be reduced due to shake-down if maximum stress exceeds yield stress. The curve represents mean fatigue life. 3) Hot spot stress determined at 0.5t line parallel to the centre line. 4) Hot spot stress determined at the centre line.

6.2

Side Longitudinal~Frame connection of an FPSO

6.2.1

Description of test and experimental results

Two test models were built to the same dimensions and specifications. The essential feature of the test models was the transition between longitudinal stiffener and transverse frame, Figure 25. The global geometry of the test models is shown in Figure 26. The models were fabricated according to shipbuilding specifications.

lon F

/

/

/

50

/

/

/

4

HS2

/

/

/

/I z

"'] ~

/J

LAP

4s'

J

-t2

I

:

t:1

.

Figure 25. Dimensions of stiffener/frame transition. The figure is oriented with ship side shown as top horizontal plate. Location of Hot Spot 1 (stiffener to lug weld) and Hot Spot 2 (stiffener to web) are marked as HS 1 and HS2 respectively. Fatigue initiation took place at HS 1.

Fatigue Strength Assessment

319

Dimensions, cf. Figures 25 and 26: 9 Shipside: 6400 x 800 x 16 mm (L x B x T) 9 Frames: 1300 x 800 x12 mm (H x B x T) 9 Longitudinal stiffener: HP 320 x 14 9 Stiffeners on frames: HP 260 x 12 Lug: 12 mm ~'---f

, ....

.......... F = 100 kN

......F -

loo

k.

Figure 26. Global geometry of test model and loading during static tests. In the fatigue test the load range was twice as large. The models were welded to a steel plate which was bolted to the strong floor. The actual load on a ship side is hydrostatic pressure. As an approximation, two equal point loads at mid-span were applied, Figure 26. The effect on hot spot stress due to the approximate load application and other effects, was assessed by finite element analysis. Strain gauges were applied to measure strain at different locations, on top of the stiffener bulb (nominal stress), and at various locations on the stiffener/frame transition (hot spot stress locations). Static stress was measured with an applied load of 200 kN (2 x 100 kN point load, cf. Figure 26). Fatigue testing was carried out with constant amplitude loading, with a load range 2 x 200 kN at R-ratio R = 0.1. In both models fatigue cracking was initiated at HS1 (Figure 25), at the weld toe at the edge of the lug. However, the mode of failure was different in the two models. In Model 1 the crack at HS 1 grew to a surface length of approximately 10 mm. At that stage a second crack was initiated at HS2 (Figure 25), at the weld toe. When the crack at HS 1 had attained a size of "through weld" (10 mm of weld was cracked) it was still the dominant crack. After some 500.000 cycles the HS2 crack started growing at a much faster rate. The crack at HS 1 stopped growing at a length of 20 mm, and detachment took place at the web side of the stiffener. In Model 2 only one crack at HS 1 was initiated, and grew until the lug was detached from the stiffener. As discussed below the through weld crack was considered to be the most relevant criterion for fatigue life. Hence data for HS 1 is used for comparison between test results and analysis. Key results for the two models are shown in Table 4.

320

Special Task Committee VI.2 TABLE 4. RESULTS FROM STATIC AND DYNAMIC TESTS (AP = 400 KN)

Model

ASmeas at

Nini

Nthrough

Ndetach

HS1 (MPa) 1) 200 190 195

2 Mean

122.000 72.000 97.000

420.000 404.000 412.000

1.422.000 2.065.000 1.744.000

1) ASmeaswas measured statically at AP = 200 MPa, and multiplied by a factor of two. Nin i - c r a c k initiation, surface crack at weld toe, depth 1-2 mm

Nthrough - c r a c k through weld, surface length approximately 10 mm Ndetach -- lug completely detached, end of test 6.2.2

Fatigue strength assessment according to various codes - discussion

The side longitudinal model for an FPSO was analysed using design procedures of ABS, BV, DNV, GL, KR and NK. The material is quite extensive and a summary is given in Table 5.

TABLE 5. SUMMARY OF RESULTS FROM COMPARATIVE TEST, SIDE LONGITUDINAL OF F P S O .

Class FEM soc. software

FEM ASc~c element (MPa)

ABS

4-node shell 8-node thick shell 4-node shell 20-node solid 4-node shell 4-node shell

ANSYS

BV

NSO (New Strudl Offshore) DNV ABAQUS GL KR

MSC Marc Mentat NASTRAN

NK

NASTRAN

AScalc/

Kw

S-N 1Vdesk8" cal curve 1) "

ASmeas

190

0.97

1.0

232.8

1.19

1.84

154

0.79

1.5

349.2

1.79

1.0

187.1

0.96

1.0

1962)

1.01

1.0

FAT 80 FAT 143 FAT 130 FAT 100 FAT 91 FAT 1793)

N yal" I No~p

9

Initia tn

Through weld

Detachmt

1.5.105

1.65

0.37

0.087

7.3.104

0.75

0.18

0.042

4.0.105

4.1

0.97

0.23

4.7.104

0.48

0.11

0.027

2.3.105

2.4

0.56

0.13

2.1.105

2.2

0.51

0.12

All curves are design curves with slope parameter m = 3, except NK curve. FAT designation is stress range at 2.106 cycles. 2) An equivalent stress range ASeq = 240 MPa was calculated from mean and residual stress corrections. 3) The S-N curve has a slope parameter m = 4.63 and represents mean fatigue life. 1)

In all the models the size of the FE mesh close to the intersection was t by t. All participants applied a hot spot stress assessment using 1 . 5 t - 0.5t linear extrapolation to the weld toe. Quite different FE models were applied. Four participants used 4-node shell elements. The stress results were relatively uniform, within the range 0.79 - 1.01 of the measured hot spot stress. The 8-node thick shell model of BV gave a hot spot stress that was significantly larger. The model of GL with 20-node solid elements gave a hot spot stress nearly twice as large as the shell models using 4-node elements.

Fatigue Strength Assessment

321

It should be borne in mind that depending on the FE model, type of element, etc., extrapolation lines and location of hot spot at the intersection could be different from the location of the measured hot spot stress. This appears to be an inherent uncertainty in hot spot stress analysis using FEM. In this case, with constant amplitude loading with R-ratio R = 0.1, only the NK mean stress correction would apply. The equivalent stress range corrected for mean and residual stress by the NK model was ASeq = 240 MPa, for a hot spot stress of 196 MPa. The procedure is linked to a S-N curve with slope parameter 4.63, and a relatively large FAT value. DNV applied a weld notch factor Kw = 1.5, and a FAT 130 S-N curve. BV applied a weld notch factor Kw = 1.84 corresponding to the type and shape of the weld, and a FAT 143 curve. ABS, GL and KR did not apply a correction for weld shape, and used lower FAT curves. Neither did NK apply a correction for weld shape, but used an S-N curve on a different format, due to the mean stress correction. For the stiffener/web connection, fatigue cracking from initiation till full detachment of the lug progressed very slowly, and at a steady rate. This is typical for welded joints in which the stresses are mainly due to displacement controlled constraint effects, and where load shedding takes place due to fatigue cracking. The selection of a criterion for fatigue failure is an open question, as discussed below: - The criterion applied in the test for detection of fatigue initiation is much too sensitive compared to practical inspection criteria for fatigue cracks in ships, and a very conservative measure of fatigue life. -

-

Through weld cracking in this case means that a crack, initiating at the toe of the weld, had penetrated the weld throat of the fillet. At this stage the crack had a surface length of approximately 20 mm, and approximately 10 mm of the weld throat was cracked. This is still very small compared to tolerable cracks in a ship structure. The through weld crack is thus a conservative measure of structural fatigue life. Full detachment of the lug/web (end of test) is clearly a crack that should be detected and repaired. Even though the structural model at this stage was fully load carrying, as a measure of fatigue life this crack size is on the non-conservative side. The comparisons made with the calculated lives using this criterion are included in Table 5 for illustration only.

Based on this discussion, the through weld crack was selected as experimental fatigue life and target for fatigue life predictions. The design calculations show a utilisation factor on life in the range 0.11-0.97, a factor of nine from the lowest to the highest. The analyses based on 4-node shell elements are relatively uniform, with utilisation factors in the range 0.51-0.97. The NK calculation, which is based on a mean life S-N curve, is not showing any less conservatism compared to the analyses based on design S-N curves. The analyses by GL (20-node solid elements) and BV (8-node thick shell) give much lower predictions of fatigue life, due to calculation of a larger hot spot stress. For the BV analysis the assumed weld notch factor Kw also contribute significantly to a low life prediction. It appears from the analyses that for this type of connection the assessment of a fatigue life is more uncertain compared to the case of a bilge knuckle joint. The rather complicated geometry, and the fact that local stresses to a large extent are displacement controlled, may explain this discrepancy.

322

6.3

Special Task Committee VI.2 Conclusions from the Comparative Study

The two test models represent quite different stress systems. The bilge knuckle joint is a primary loadcarrying element, with relatively direct load transfer and essentially a two-dimensional local geometry at the hot spot. The side longitudinal joint may be termed a secondary element, with a complicated three-dimensional local geometry in which the loading is largely displacement controlled, and with stress concentrations that are mainly due to constraint effects and with large secondary bending. This difference could be reflected in some of the results. For the bilge knuckle joint the stress analyses came out with quite uniform results. Relative to the measured hot spot stress the calculated stress was 0.91 + 15%. The GL exercise comparing three different types of element models, and correspondingly different definitions of hot spot stress, gave results in nearly perfect agreement. Considering the inherent variability that is observed in measurement of stress in welded structural models, it may be concluded that the overall agreement between analysis and test is remarkably good. For the side longitudinal, there appears to be a difference between results obtained with 4-node shell analysis and those obtained with higher order elements. The 4-node shell models gave results in good agreement with measured hot spot stress. The 8-node thick shell and the 20-node solid element models overpredicted the hot spot stress quite considerably. Part of the discrepancy could be explained by the fact that the extrapolation path and the location of the hot spot stress are not the same in analysis and test. The design procedure, starting from the hot spot stress, exhibit some significant differences related to the following factors: -

-

Corrections for mean stress and residual stress Weld notch factor Kw Choice of S-N curve

A comparison on each step of the procedures makes no sense, since each procedure appears to be calibrated. The predictions for fatigue life (design and mean life) for the bilge knuckle joint are very uniform and in very good agreement with the test result. The predictions for the side longitudinal are more scattered, obviously due to the differences in hot spot stress assessment as discussed above. Design using thick shell and solid elements came out on the conservative side, with utilisation factors 0.18 and 0.11 respectively, relative to "through weld" cracking. The 4-node shell element models gave utilisation factors in the range 0.37-0.97 on design and 0.51 on mean life prediction (NK procedure). Design predictions with utilisation factor in the range 0.5-1.0 may indicate some lack of conservatism. However, it should be borne in mind that the "through crack" criterion is conservative with respect to fatigue design, also that the joint in question is a secondary detail. Allowing a larger crack as failure criterion will increase the conservatism in the predictions. With this qualification it may be concluded that the design procedures are validated by the test data.

7.

CONCLUSIONS AND RECOMMENDATIONS

As stated in the introduction, a large number of fatigue design rules, recommendations and practices are being applied for ship design and structural design in general. Over the years the rules and recommendations have gradually become more and more prescriptive, detailed and voluminous. A report of 40-50 pages cannot be expected to pay justice to all the issues that should be addressed with regard to fatigue design.

Fatigue Strength Assessment

323

Through the work of the committee, and in particular through the comparative studies, some particular issues have come out as important, as summarised below:

Fatigue design by fracture mechanics analysis Fracture mechanics provides a tool for more detailed fatigue analysis compared to the S-N curve approach. Good agreement with relevant experimental results has been demonstrated. However, for more general applications in design, several factors should be examined more precisely, such as crack growth law, calculation of stress intensity factor for surface cracked structural members, mean stress effects, threshold stress intensity factor under random loading conditions, and shape, size and distribution of initial cracks.

Calculated stress vs. S-N curve stress The essential of fatigue strength analysis is the link between stress analysis and fatigue strength defined by an empirical S-N curve. Fatigue is a local phenomenon, and life prediction requires calculation of very detailed stress components in complex geometries. There seems to be agreement that calculations with an FE mesh with t by t size elements and extrapolation 1.5t-0.5t to the weld toe will give a stress which can be equated to the definition of stress in the empirical S-N curves. For simple joints, like the ones found in catalogues of welds and classified in design rules, this is reasonable. For complex joints with large stress gradients and constraint effects this may not be generally true. There is a lack of fatigue test data to verify the hot spot stress concept for ship details.

FE modelling and hot spot stress The FE modelling and extrapolation procedure for assessment of hot spot stress is a step beyond what engineers normally do in a ship design phase. More simplified procedures would be beneficial to the industry. There is also a concem for the quality of such analyses, which in general require hand calculations through time consuming procedures which may be prone to error. The outcome of the analysis is also dependent on properties of the specific computer code, which most designers accept as a "black box." It has been proposed to base fatigue assessment on an element stress that is calculated at some location close to the weld, together with a lower FAT S-N curve, thus avoiding the extrapolation. If such a procedure were properly calibrated, the fatigue prediction would be the same, at less expense and possibly with less uncertainty.

Element type In the comparative study several types of elements were used and compared. The results indicate that 4-node shell elements give satisfactory results, even for a complicated geometry like the side longitudinal/frame joint. There was no obvious benefit from using higher order elements.

Definition of fatigue life A large majority of the fatigue tests which are underlying the fatigue design curves were carried out with small scale specimens tested in either axial or in bending loading. For these geometries fatigue life is essentially initiation and growth of very short cracks (a few millimetre). Ship structures are in general crack tolerant - cracks of length 100 mm and more are quite commonly found and repaired during regular surveys. As shown by the tests used for the comparative studies, crack growth in a large and redundant ship structure is much more gradual compared to small-scale specimens. The definition of "failure" for a ship detail may be over-conservative.

Lack of commonality The Tables 3 and 5 express a striking lack of commonality between the different procedures. The weld notch factor is applied in various ways, for the same detail different S-N curves are advised, and the mean stress correction is applied with very different effects on the results. However, the variability in the fatigue life predictions is not so large, certainly not for the bilge knuckle joint where the stress

Special Task Committee VI.2

324

analyses came out reasonably in agreement. Apparently many roads lead to the same goal. From a practical point, the industry should try and find a common road, and to make it as simple as possible. ACRONYMS ABS BV DNV FAT FEM FPSO GL ffW KR LR NK PWHT RINA SCF VLCC

American Bureau of Shipping Bureau Veritas Det Norske Veritas Fatigue life (or design life) at N = 2.10 6 cycles Finite element method Floating production and storage ship (for offshore oil production) Germanischer Lloyd International Institute of Welding Korean Register of Shipping Lloyd's Register of Shipping Nippon Kaiji Kyokai Post weld heat treatment Registro Italiano Navale Stress concentration factor Very large crude carrier

REFERENCES

ABS (2002). Guidance Notes on Spectral-Based Fatigue Analysis for Floating Production, Storage and Offloading (FPSO) Systems. American Bureau of Shipping, Houston, USA. ABS (2003). Rules for Building and Classing Steel Vessels. American Bureau of Shipping, Houston, USA. Barro, G., Labesse-Jied, F., Lieurade, H.P., Recho, N., 1999,"Lack of Penetration Effect on the Stress Profile at the Weld Toe of Fillet Welds", IIW-Doc XIII-1790-99, Int. Institute of Welding. Berge, S. (1980). On the effect of plate thickness in fatigue of welds, Engineering Fracture Mechanics, 21:2, 423-435. Berge, S., Eide, O., Astrup, O. C., Palm, S., W~istberg, S., Gunleiksrud, ~. and Lian, B. (1987). Effect of Plate Thickness in Fatigue of Welded Joints in Air and in Sea Water, Steel in Marine Structures, edited by C. Noorhook and J. deBack Elsevier, Science Publishers B.V., Amsterdam, 799-810. Berge S, Johansen A and BjCrheim L G. (2001). Fatigue strength assessment of hull details for an FPSO, Practical Design of Ships and other Floating Structures, Y.-S. Wu, W.-C. Cui and G.-J. Zhou, eds., Elsevier. BS 7910:1999 Guidance on methods for assessing the acceptability of flaws in fusion welded structures. BSI. Connoly, L. M. and Zettlemoyer, N. (1993). Stress Concentration at Girth Welds of Tubulars with Axial Wall Misalignment. Proceedings Int. Conf. on "Tubular Structures", E & F N Spon, London. DNV Classification Note 30.7 (2001). Fatigue Assessment of Ship Structures. Det Norske Veritas, HCvik. DNV-RP-C203 (2001). Recommended Practice. Fatigue Strength Analysis of Offshore Structures. Det Norske Veritas, Hcvik. Fricke, W. (2001). Recommended Hot Spot Analysis Procedure for Structural Details of FPSO's and Ships Based on Round-Robin FE Analyses. ISOPE Stavanger. Proc. 11th ISOPE, Stavanger. Also International Journal of Offshore and Polar Engineering. 12, No. 1, March 2002. GrCvlen, M. (1987): Some Aspects of Corrosion Fatigue or Structural Steels. Steel in Marine Structures, edited by C. Noorhook and J. deBack Elsevier Science Publishers B. V., Amsterdam, 719-727.

Fatigue Strength Assessment

325

Gurney, T. R. (1976). Cumulative damage calculations taking account of low stresses in the spectrum. The Welding Institute Report 2/1976/E. Gurney, T. R. (1976). Fatigue Design Rules for Welded Joints. The Welding Institute Research Bulletin. 17:5. Haibach, E. (1970): Discussion paper, The Welding Institute Conf. on Fatigue of Welded Structures, Brighton, UK. Hansen, P.F. and Thayamballi, A.K., (1995), "Fatigue Damage Considering Whipping Arising from Slamming", Proceedings of the 14th International Conference on Offshore Mechanics and Arctic Engineering, Part 2 of 5, 1995. Hobbacher, A. (1996). Fatigue Design of Welded Joints and Components. Recommendations of IIW Joint Working Group XIII-XV. Abington Publishing, Abington, Cambridge. HSE (1992), Fatigue Background Guidance Document, OTH 92 390. HSE (1995). Guidance on Design, Construction and Certification. IIW Recommendation (1996). "Fatigue design of welded joints and components", Abington Publishing. Infante, V., and Branco, C.M. (2000)."A Comparative Study of the Fatigue Behavior of Repaired Joints by Hammer Peening", IIW-Doc XIII- 1836-2000, Int. Institute of Welding. ISSC (1997). International Ship and Offshore Structures Congress, Report of Comm. 111.2 Fatigue and Fracture, T. Moan and S. Berge (eds.), Elsevier. ISSC (2000). International Ship and Offshore Structures Congress, Report of Comm. 111.2 Fatigue and Fracture, H. Ohtsubo and Y. Sumi (eds.), Elsevier. Kada, K., Matsuoka, K., Tomita, Y. (2002). Fatigue Strength of Large Structural Model of Double Hull VLCC Bilge Knuckle Part under Variable Amplitude Loading Conditions. Extended Abstract, Symposium on Fatigue Testing and Analysis Under Variable Amplitude Loading Conditions. Kang, S. W., Kim, W. S. and Paik, Y. M. (2002) Fatigue Strength of Fillet Welded Steel Structure under Out-of-Plane Bending, International Welding / Joining Conference - Korea. Kim, W. S, Tomita, Y., Hashimoto, K. and Osawa, N. (1997). Effects of Static Load on Fatigue Strength of Ship Structure. Proceedings of the seventh Int. Offshore and Polar Engineering Conference Honolulu. Kim, W. S. (2001). Fatigue Strength of Load-Carrying box Fillet Weldments in Ship Structure. Proceedings of the Eighth International Symposium on Practical Design of Ships and Other floating Structures. Shanghai, China. Lieurade, H.P., and Huther, I. (2001). "Fatigue Strength Improvement Solutions for Welded Structures and Components", nW-Doc X111-1904-01, International Institute of Welding. Lotsberg, I. (1998) Stress Concentration Factors at Circumferential Welds in Tubulars. Marine Structures, 11,207-230. Lotsberg, I and Rove, H. (2000). Stress Concentration Factors for Butt welds in Stiffened Plates. OMAE, New Orleans. Lotsberg, I. (2001a). Overview of the FPSO Fatigue Capacity JIP. OMAE, Rio de Janeiro. Lotsberg, I., Askheim, D. 0., Haavi, T. and Maddox, S. (2001b) Full Scale Fatigue Testing of Side Longitudinals in FPSOs. ISOPE Stavanger. Lotsberg, I. and Larsen, P. K. (2001c). Developments in Fatigue Design Standards for Offshore Structures, ISOPE, Stavanger. Lutes, L.D., and Larson, C.E. (1990). "Improved Spectral Method for Variable Amplitude Fatigue Prediction", Journal of Structural Engineering, ASCE, Vol. 116, No. 4. Lutes, L.D., Sarkani, S., Kihl, D.P., and Beach, J.E. (2002). "Optimal S-N Testing for Variable Amplitude Fatigue Applications", accepted for publication in the Journal of Structural Engineering, ASCE. Maddox, S. J. (1985). Fitness for purpose assessment of misalignment in transverse butt welds subjected to fatigue loading. London: International Institute of Welding, IIW Document XIII-11801985.

326

Special Task Committee VI.2

Maddox, S.J. (1991). "Fatigue Strength of Welded Structures", 2na edition, Abington Publishing, Cambridge, England. Maddox, S. J. (1997). Developments in Fatigue Design Codes and Fitness-for-service Assessment Methods. International Conference on Performance of Dynamically Loaded Welded Structures. Welding Research Council, Inc. Maddox, S. J. (2001). Recommended Design S-N Curves for Fatigue Assessment of FPSOs. ISOPE Stavanger. Maherault, S., Huther, M., Parmentier, G., and Recho, N. (1999). "Semi-probabilistic Fatigue Calibration of the Partial Safety Factors", IIW-Doc XIII-1794-99, Int. Institute of Welding. Miles, J.W. (1954). "On Structural Fatigue Under Random Loading", Journal of Aeronautical Science, Vol. 21 pp. 753-762, November. Miner, M.A. (1945). "Cumulative Damage in Fatigue", Journal of Applied Mechanics, September 1945. Mohaupt, U. H., Bums, D. J., Kalbfleisch, J. G., Vosikovsky, O. and Bell, R. (1987) Fatigue Crack Development, Thickness and Corrosion Effects in Welded Plate to Plate Joints. Steel in Marine Structures", edited by C. Noorhook and J. deBack Elsevier Science Publishers B.V., Amsterdam, 269-280. Nelson, W. (1990). "Accelerated Testing: Statistical Models, Test Plans and Data Analysis, Wiley. Niemi, E. (2001). Structural Stress Approach to Fatigue Analysis of Welded Components. Designer's Guide. XIII-1819-00, XV-1090-01, XIII-WG3-06-99. Niemi, E. (1997). Stress determination for fatigue analysis of welded components, Abington Publishing, Abington, UK. NORSOK N-004 Design of Steel Structures (1998). NTS, Oslo, Norway. NS 3472 (2001) "Design of Steel Structures", NBR Oslo. Palmgren, A. (1924). "Die Lebensdauer von Kugellagern", ("The Life Expectancy of Ball Bearings"), Z. Vereines Deutscher Ingenieure, 68, 1924. Paris, P., and Erdogan, F. (1973). "A Critical Analysis of Crack Propagation Laws", Transactions of the ASME, Journal of Basic Engineering, Vol. 85, Series D, No. 3, December 1973, pp. 528-534. Radenkoviq, D. (1981). Stress analysis in tubular joints, Proc. Int. Conf. Steel in Marine Structures, Institut de Recherches de la Sidrrurgie Franqaise, Plenary Session, pp. 53-96, Paris, France. Rizzuto, E., and Friis Hansen, P. (2000). "Stress Range Distributions for the Computation of Fatigue Cumulated Damage in Marine Structures", Proc. of the Intemational Conference on Ship and Shipping Research (NAV 2000). Ship Research Panel 219. (1996). Study on Practical Use for Fatigue Crack Growth Analysis. Shipping Research Association of Japan. (in Japanese). Ship Research Panel 245. (2001). Study on Ship Structural Life of the Double Hull Tanker. Shipping Research Association of Japan. (in Japanese). Sielski, R.A., Wilkins, J.R. Jr., and Hults, J.A. (2001). "Supplemental Commercial Design Guidance for Fatigue", Ship Structure Committee report #419. Sieve, M.W., Kihl, D.P., and Ayyub, B.M. (2000). "Fatigue Design Guidance for Surface Ships", Naval Surface Warfare Center, Carderock Division, NSWCCD-65-TR-2000/25, November. Tomita, Y., Kawabe, H., and Fukuoka, T. (1992). Statistical Characteristics of Long-Term WaveInduced Load for Fatigue Strength Analysis for Ships. Practical Design of Ships and Mobile Units, 2, 2792-2805. Tomita, Y., Hashimoto, K., Osawa, N., Terai, K., and Wang, Y. (2002). Study on Fatigue Design Load for Ships Based on Crack Growth Analysis. Fatigue Testing and Analysis Under Variable Amplitude Loading Conditions, ASTM STP 1439. Wang, Y., Tomita, Y., Hashimoto, K., and Osawa, N. (2001). A Study on Fatigue Crack Growth under Considering Coalescence of Multiple Surface Cracks. Asian Technical Exchange and Advisory Meeting on Marine Structures, 319-328. Yoneya, T., Kumano, A., Yamamoto, N. and Shigemi, T. (1993). Hull Cracking of Very Large Ship Structures. IOS' 93, Glasgow.

KEYWORD INDEX Accident 1 Accidental limit state (ALS) design 71 Accuracy in production 189 Adhesive bonding 189 Advanced composite materials 109 Anchor 149 Assessment 37

High speed vessels 109 High strength steel 189 Hot spot stress 285 Hull 149 Hull girder collapse 71 Impact 71 Imperfections 189 Improvement techniques In-service 37 Inspection 37 Internal mechanics 71

Benefit 1 Buckling 37 Classification rules 109 Collision/grounding scenarios 71 Comparative calculation 235 Consequence 1 Corrosion 37 Corrosion resistance 189 Corrosive environment 285 Cost 1 Crack 37 Crushing 71 Cumulative damage 285 Design wave

Laser welding LNG 149 Loads 109

Otttoading 149 Offshore 37 Oil outflow 71 Operational envelope

71

Fabrication technologies 109 Fatigue 109, 285, 37 Fatigue design 285 Fatigue load 235 Fatigue strength 285 Finite element analysis 285 Floating production 149 Formal safety assessment 1 Forming 189 FPS 149 FPSOs 149 Fracture 71 Frequency 1 FRP 189 Guidelines Hazard

189

Mean stress 285 Measurements 37 Monitoring 37 Monohull 149 Mooring 149

235

External mechanics

285

109

Pipe in pipe 149 Post-accident mechanics Probability 1 Procedures 37

71

Raking 71 Residual strength of damaged ships Residual stresses 189 Riser 149 Risk analysis 1 Risk assessment 1 Robots 189 S-N curves 285 Semi-submersible 149 Ship collisions 71 Ship grounding 71

37

1 327

71

328 Ship structures 285 Ships 37 Simulation 189 Spar 149 Steel catenary riser 149 Steel tube umbilical 149 Stochastic load description 235 Strength 109 Stress concentration factors 285 Stress transfer function 235 Strip theory 235 Structural design 109 Structural response 109 Structural stress 285 Structure 37 Survey 37

Keyword Index Tearing 71 Tension leg platform Tether 149 Titanium 189 Uncertainties

149

109

Validation 235 Vortex induced vibration,

149

Wave-induced hull girder loads Wave pressure 235 Welded structures 285 Welding consumables 189

235

REPORT MEMBERS INDEX

Amdahl, J.

71

Lotsberg, I. 285 Ludolphy, H. (Deceased)

Bai, Y. 149 Barltrop, N. 71 Berge, S. 285 Boonstra, H. 149 Borzecki, T. 189 Branner, K. 235 Brown, D.T. 149 Bruce, G.J. 37

Maherault, S. 285 Mavrakos, S. 149 Mikkola, T.P.J. 285 Moore, W. 1 Nedergaard, H. 149 NeRo, T.A. 149 Nielsen, L.R 285

Cariou, A. 235 Chen, Y. 1 Chung, T.Y. 149

Olson, D.

Dinovitzer, A. 1 Donner, E.R. 71 Duan, M. 37

Paetzold, H. 285 Paik, J.K. 71 Park, J.Y. 189 Pedersen, RT. 71 Prevosto, M. 1

Egorov, G.V

189

37 Rizzuto, E. 235 R6hr, U. 71

Ferraris, S. 109 Folso, R. 37 Fonseca, N. 109 Fujimoto, Y. 37 Fukasawa, T. 235

Shenoi, R.A. 189 Shin, B.-C. 37 Shin, C.-H. 285 Sun, H.-H. 285 Suzuki, H. 149

Garbatov, Y. 37 Gu, X.K. 235 Gu, Y. 71 Hayman, B. 109 Heinemann, M. 189 Hughes, O. 109

Takeda, Y. 189 Thiberge, E. 109 Tomita, Y. 285 Tonelli, A. 1 Toyama, Y. 109

Ito, H.

Vfirdal, O.T. 37 Vredeveldt, A. 109

71

Kapsenberg, G. Kihl, D. 285

Wang, G. 71 Watanabe, I. 235 Wilckens, H. 189 Wu, N. 189

235

Lallart, E 189 Le Hire, J.-C. 37 Li, R. 149 Litonov, O. 1 Loeken, A. 149

Yang, E 109 Yang, Y.S. 1 Yoshida, K. 1 329

71

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