Damage Resulting From Shuttle Tanker

HSE

Health & Safety Executive

Damage resulting from shuttle tanker: FPSO encounters

Prepared by PAFA Consulting Engineers for the Health and Safety Executive

OFFSHORE TECHNOLOGY REPORT

2002/006

HSE

Health & Safety Executive

Damage resulting from shuttle tanker: FPSO encounters

PAFA Consulting Engineers Hofer House 185 Uxbridge Road Hampton Middlesex TW12 1BN United Kingdom

HSE BOOKS

© Crown copyright 2002 Applications for reproduction should be made in writing to: Copyright Unit, Her Majesty’s Stationery Office, St Clements House, 2-16 Colegate, Norwich NR3 1BQ First published 2002 ISBN 0 7176 2331 9 All rights reserved. No part of this publication may be reproduced, stored in a retrieval system, or transmitted in any form or by any means (electronic, mechanical, photocopying, recording or otherwise) without the prior written permission of the copyright owner.

This report is made available by the Health and Safety Executive as part of a series of reports of work which has been supported by funds provided by the Executive. Neither the Executive, nor the contractors concerned assume any liability for the reports nor do they necessarily reflect the views or policy of the Executive.

ii

CONTENTS EXECUTIVE SUMMARY

v

1.

INTRODUCTION

1

2.

OBJECTIVES AND SCOPE OF WORK

2

3.

COLLISION DAMAGE ESTIMATION METHODOLOGY 3.1 Background to Damage Assessment Method 3.2 Frieze’s Damage Estimation Formulae 3.3 Application of Methodology

3 3 3 4

4.

DETAILS OF THE SHUTTLE TANKER AND FPSO 4.1 Collision Scenarios 4.2 FPSO Stern Shape 4.3 Shuttle Tanker Bow Shape 4.4 Relative Ballast Condition

5 5 6 7 9

5.

RESULTS OF RIGID TANKER BOW IMPACT SCENARIOS

10

6.

SENSITIVITY STUDIES 6.1 Accuracy of the Frieze Energy Equation 6.2 Bow and Stern Shapes 6.3 Energy Calculation and Distribution 6.4 Direction of Vessel Collision 6.5 Impact on Vessel Equipment

15 15 15 17 20 24

7.

SUMMARY AND CONCLUSIONS

25

8.

REFERENCES

26

APPENDICES A Example of Damage Material Estimation B Example of Energy Calculation

iii

Printed and published by the Health and Safety Executive C30 1/98

iv

EXECUTIVE SUMMARY The HSE commissioned P A F A Consulting Engineers to estimate the damage that could occur during a shuttle tanker – FPSO collision in an initial impact energy range of 0 – 100 MJ. A simplistic assessment method was required, allowing a large range of scenarios to be covered. · The preferred approach was the Frieze formula – a variation on the widely used Minorsky energy absorption method; · Dynamic strain rate effects are accounted for in the model; · This formula covers bending, stretching, compression, scraping, buckling, crushing, folding, fracture and tearing damage modes; · The main effort to determine the absorbed energy is in estimating the volume of damaged steel. The eight scenarios specified were: · two types of FPSO stern shape (inclined and flat); · two types of shuttle tanker bow shape (sharp and rounded); · two types of relative ballast condition (full/empty tanker – empty/full FPSO). The following assumptions were made: All the energy was absorbed by the FPSO stern; · All the motion was in the horizontal direction; · The impact was perpendicular onto the stern, i.e. no lateral or angular offsets; · Only structural steel plating and stiffeners were considered to absorb energy; · The mass of the tanker was held constant while the velocity was varied. ·

For these scenarios, expressed in terms of initial impact kinetic energy (KE), the range of stern penetrations was found to be 1.17 m to 2.11 m for 20 MJ, 2.13 m to 4.27 m for 60 MJ and 3.0 m to 5.56 m for 100 MJ. In addition, it should be noted that the tanker bow tip overhangs the FPSO stern deck by 1.0 m - 9.5 m depending upon the loading condition. Sensitivity studies indicate the following: · The formula covers all damage modes with a 97.7% confidence interval of around –17% to +25%; · The one aspect of the bow/stern shape not well covered was large raked or bulbous bow sections; · Due to the relatively small velocities involved, strain rate is not a variable in the model; · One scenario considered energy to be absorbed by the tanker bow as well as the FPSO stern. This case yields 25%-30% reduction in stern penetration compared to the rigid bow scenarios. The tanker bow absorbs around 40% of the initial impact energy in this case; · Relative vertical velocities due to wave action may be substantial, e.g. 2.4 - 3.6 m/s in 6.0 m significant waves. While the ship’s master may have some control over the horizontal motion of his vessel, he will have very little control over pitch and heave. For one scenario of a shuttle tanker to FPSO collision, maximum vertical deformation energies are estimated to be 270 MJ in 6.0 m waves; · Damage/destruction of a rear-sited helideck was likely but not expected to be catastrophic whereas, if process equipment is sited at the stern, this is more likely to lead to an escalating incident and loss of life.

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Printed and published by the Health and Safety Executive C30 1/98

vi

1. INTRODUCTION In 2000, a study performed for the Health and Safety Executive (HSE) highlighted concerns surrounding shuttle tanker – FPSO collision risks (Reference 1). Based on five incidents of contact between attendant tankers and FPSOs, a maximum impact energy of 37 MJ was reported. This maximum historic incident was later shown to be nearer 8 MJ, however, the report did highlight the potential for large energy impact events. It should be noted that a more recent collision involved an impact energy estimated at around 30 MJ. Such events would have potentially serious consequences for hull penetration and also for deck mounted facilities. P A F A Consulting Engineers (PAFA), was asked by the HSE to study a range of collision scenarios between the bow of a shuttle tanker and the stern of an FPSO. These would consider different types of bow shape, stern shape and ballast condition. PAFA had previously derived a modified Minorsky formula that could be used to estimate energy absorption in collisions based on the volume of damaged steel. This method was to be employed to estimate the penetration of the FPSO stern based on a rigid bow and also for simultaneous damage to both bow and stern. The study initially concentrated on encounters involving horizontal motions only. This was based on the perception that the main cause of such incidents was a failure (either temporary or permanent) of the positioning or motive power of the tanker combined with a yaw/sway motion of the FPSO, all of which involve horizontal movement only. However, subsequent consideration of the average sea conditions in the North Sea suggested it would be relatively likely for both vessels to be experiencing heave/pitch motions at the time of such an incident, and that the consequences of such vertical components should also be addressed.

1

2. OBJECTIVES AND SCOPE OF WORK The primary objective of this project was to assess the scale of damage that occurs during a direct impact between the bow of a shuttle tanker and the stern of an FPSO. The estimated damage was required for a range of collision scenarios involving the dissipation of up to 100 MJ of kinetic energy. The method applied to determine the energy dissipation, calculated as a function of the volume of damaged material, was based on the Minorsky equations (Reference 2), modified and applied by Frieze (References 3 and 4) for a series of bow impacts with bridge piers. The methodology is described, in detail, in Section 3. It was decided that eight scenarios should be investigated, comprising: · · ·

two types of FPSO stern shape two types of shuttle tanker bow shape two types of relative ballast condition.

The eight scenarios are referenced as follows: Inclined Stern

Flat Stern

Rounded Bow

Sharp Bow

Rounded Bow

Sharp Bow

Full tanker – Empty FPSO

1

2

5

6

Empty tanker – Full FPSO

3

4

7

8

A conservative assessment of the tanker penetration into the FPSO stern was initially required, whereby all the energy was assumed to be dissipated into the FPSO. Thus, there was no energy absorption and deformation of the tanker bow or loss of energy through friction or waves. In addition, the relative motions were assumed to be horizontal throughout the duration of the impact. The results of these damage scenarios are described in Section 5. A second series of analyses were performed to estimate the effect of varying the type of impact by including a vertical motion component, accounting for damage to the tanker bow and by varying the relative tanker mass and velocity. The effect of these assumptions is addressed within a sensitivity study reported in Section 6. Estimates of the likely proportion of energy absorbed by the FPSO as a proportion of the total collision energy are made based on conservation of both energy and momentum in the short duration impacts predicted. Conclusions from the project are summarised in Section 7.

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3. COLLISION DAMAGE ESTIMATION METHODOLOGY 3.1 BACKGROUND TO DAMAGE ASSESSMENT METHOD Vessel impacts to other vessels or bridge piers and other land-based structures usually involve the bow of the striking vessel. The mechanisms involved are complex and include both structural and hydrodynamic components. The structural mechanisms suffered by the steel components comprising the vessel structure include bending, stretching, compression, scraping (friction), buckling, crushing, folding, fracture and tearing. The hydrodynamic mechanisms include rigid body motions with attendant changes in added mass. All six vessel motions can be involved, although analyses are usually restricted to surge motions in order to maximise the forces involved and, thus, the damage consequences for the struck object. There are a range of solutions available to estimate the level of damage. The most sophisticated analyse such impacts using non-linear finite element analysis (FEA) coupled with a system that accounts for the hydrodynamic response of the vessels involved. While the accuracy of such models is rapidly improving, these assessments remain very time consuming. As the sophistication of the modelling reduces so does the corresponding computational time and effort, although, substantial effort is still required to interpret the vessel’s drawings. For this study, where typical vessels, representative of their type, are being studied, it is impractical to perform sophisticated finite element analyses. Minorsky (Reference 2) published an empirical energy absorption approach based on the depth of penetration and damage width. This was subsequently amended by Vaughan (Reference 5) who replaced these parameters with the actual volume of damaged material, and by Woisin (Reference 6) who replaced the constant term in Minorsky’s formula with one related to the volume of torn material perpendicular to the line of collision. 3.2 FRIEZE’S DAMAGE ESTIMATION FORMULAE Frieze (Reference 3), with the assistance of Woisin, gathered previous research and test results to enable an improved formula to be developed that accounted for structural mechanisms that were expected to be independent of scale such as buckling, from those that are scale-dependent such as tearing. The methodology was originally derived by Frieze to deal primarily with bow damage from impacts with bridge piers. At the time, this was expected to involve many of the structural mechanisms, in particular, bending, stretching, compression, scraping (friction), buckling, crushing, and folding. Tearing was expected to be of some importance, while fracture, other than that resulting from extensive localised tensile strains, was not considered of particular relevance. During impacts, straining occurs at a variety of rates, all of which influence the stress at which yield occurs. Ultimate tensile strength and the strain at rupture are also affected. The impact of strain rate was considered but only in respect of yield stress. Following calibration against test data, the following formula was proposed: Energy (kJ) = 60.5 (sd/sy) Vs + 6.0 t2 l where 60.5 Vs

represents the work done statically in bending, stretching, compression, scraping (friction), buckling, crushing, and folding. 60.5 is the average stress, in N/mm² Vs is the volume of damaged (not torn) material, in m².mm

3

sd/sy

6.0 t2 l

represents strain rate effects for e < 0.02 = 1.000 for e´ < 0.001 s-1 = 1.393 + 0.131 log(e´) for 0.001 £ e´ < 1.000 s-1 = 1.393 + 0.393 log(e´) for e´ ³ 1.000 s-1 -1 e´ = 5 v t/a², in s v = velocity, in m/s t = thickness of plating, in m a = transverse frame spacing, in m represents tearing effects 6.0 is the average stress, in kN/mm² (= 3.0 for single-sided tearing) t is the thickness of plating, in mm l is the length of the tear in, m

Further details of the derivation of this formula can be found in Frieze and Smedley (Reference 7). 3.3 APPLICATION OF METHODOLOGY The methodology was originally derived on the basis of work done with, as appropriate, strainrate effects averaged over the duration of the encounter. However, in practice, more refined incremental assessments are sought to estimate the penetration, time, forces and velocities involved for a predefined Kinetic Energy (KE). 1.

Calculate the added mass Ma of the vessel. A value of 1.05 times the displacement (in kg) is frequently used for horizontal vessel motions.

2.

Calculate the total kinetic energy Eo = ½ Ma Vo2 , where Vo is the initial impact velocity.

3.

Assume an indentation d1 (in m). In this study, increments of 0.05m were taken – initial calculations suggested that a 100 MJ collision would not penetrate the stern by more than 6.0m and thus this was the largest penetration modelled.

4.

Calculate the volume of crushed material Vs. Ignore material perpendicular to the axis of impact since membrane action will have a secondary effect and can effectively be excluded.

5.

Using the velocity Vo, determine e’ and then sd/sy.

6.

Calculate the tearing term t²1. In the initial eight scenarios it was considered that no structural elements would exhibit tearing.

7.

Using W1 = 60.5(sd/sy)Vs determine the work done (in kJ).

8.

The average force during this indentation is F1 = W1/d1 (in kN).

9.

The kinetic energy at this indentation d1 as E1 = Eo-W1 (in kJ).

10.

The velocity at this indentation V1 = Ö[2E1/Ma] (in m/s).

11.

The time taken is t1 = Ma(Vo-V1)/F1 (in secs).

12.

Repeat steps 2 to 11 for a second increment in indentation d2. The steps are performed incrementally but the total kinetic energy must be used to determine the velocity at the end of a step.

4

4. DETAILS OF THE SHUTTLE TANKER AND FPSO 4.1 COLLISION SCENARIOS It was decided that eight scenarios should be investigated, comprising: · two types of FPSO stern shape · two types of shuttle tanker bow shape · two types of relative ballast condition. For each scenario the relationship between the kinetic energy of the collision and the bow penetration of the FPSO stern was sought for energies in the range 0 – 100 MJ. In these initial scenarios, it was assumed that: · The FPSO stern absorbed all the energy, i.e. there was no damage to the tanker bow and no loss of energy through friction or in the waves. · All the motion was in the horizontal direction. · The impact was perpendicular between the bow and stern, and on the centreline of the FPSO, i.e. no lateral or angular offsets. · Only structural steel plating and stiffeners were considered, i.e. no energy absorption by equipment, piping, etc, was considered. · The mass of the tanker was held at 110,000 tonnes full and 80,000 tonnes empty, while the velocity was varied. · The transverse frame spacing parameter a = 11.0m for the inclined stern FPSO and = 9.5m for the flat stern FPSO. The consequence of these assumptions and a sensitivity study on the model are discussed in Section 6 of this report. The eight scenarios are referenced as follows: 1. Full tanker with Rounded Bow shape colliding with Empty FPSO with Inclined stern 2. Full tanker with Sharp Bow shape colliding with Empty FPSO with Inclined stern 3. Empty tanker with Rounded Bow shape colliding with Full FPSO with Inclined stern 4. Empty tanker with Sharp Bow shape colliding with Full FPSO with Inclined stern 5. Full tanker with Rounded Bow shape colliding with Empty FPSO with Flat stern 6. Full tanker with Sharp Bow shape colliding with Empty FPSO with Flat stern 7. Empty tanker with Rounded Bow shape colliding with Full FPSO with Flat stern 8. Empty tanker with Sharp Bow shape colliding with Full FPSO with Flat stern In Summary, these eight scenarios are as follows:

Full tanker – Empty FPSO Empty tanker – Full FPSO

Inclined Stern Rounded Sharp Bow Bow 1 2 3 4

Flat Stern Rounded Sharp Bow Bow 5 6 7 8

The specifications of these variables are given in the following sections.

5

4.2 FPSO STERN SHAPE The stern shape and stiffening details of the two selected FPSOs were based on detailed drawings supplied by the vessel’s operators. These two vessels were considered to represent significantly different designs and thus gave two different scenarios that have been classified as: 1) Inclined Stern and 2) Flat Stern. For reasons of confidentiality, the vessels are not named in this report and details of the stiffening cannot be described. Figure 4.1 illustrates the two stern shapes. For the inclined stern (Figure 4.1a), it can be seen that the stern is flat across the full width = 45m. There is a small vertical plate (depth = 0.925m), with hull plating inclined at 35° above and below this plate. The plating above the upper deck was not included in the analysis since plate and stiffening details were not available (shown as a dashed line). The flat stern is illustrated in Figure 4.1b. In this case there is a flat region 14.8m wide by 9.8m deep descending from the upper deck. The deck itself widens by 26° inward from the stern throughout its full depth. Below 9.8m, the hull is inclined at 15° to the horizontal. For the scenario where the FPSO is full and the stern is flat, the rudder is 3.0m horizontally from the bulbous bow section at initial impact. While this is not accounted for in the energy calculations, it is possible that the FPSO’s control mechanisms may be damaged at high-energy collisions. In both models, horizontal and vertical plating and horizontal and vertical beams are considered to absorb energy. Calculations for transverse plating and framing, including the bow plate and horizontal stiffening showed these to contribute little and, in accordance with the methodology described in Section 3.3, these elements have not been included in the energy calculations.

45m 26°

5m

14.8m

0.925m

9.8m

18.8m

15°

a) Inclined Stern

b) Flat Stern

Figure 4.1 Isometric Inclined and Flat FPSO Stern Sections

6

4.3 SHUTTLE TANKER BOW SHAPE The two forms of shuttle tanker bow considered were initially based on photographs of shuttle tankers in-service. In the initial analyses, the bow was considered rigid and, consequently, no stiffening was required. In a subsequent analysis, the effect of bow and stern simultaneously crushing was considered, see Section 6.3.2. The bows were defined to be identical down the vertical centre-line and also in the overall tanker width, but different in cross-section. These two bows were considered to represent significantly different designs and thus gave two different scenarios that have been classified as: 1) Rounded Bow and 2) Sharp Bow. In the vertical, the bow was sub-divided into three sections, a raked bow inclined at 35° over a depth of 9m, a vertical bow section over a depth of 10m and a relatively small bulbous bow section over a depth of 9m. The vertical ℄ section of both bow types is illustrated in Figure 4.2. Figure 4.3 illustrates the plan section of the two bow types whilst Figure 4.4 presents corresponding isometric views of the bow sections. 6.3m

9m

10m

9m

2.4m

Figure 4.2 Bow Vertical Section (both bow types)

40m

a) Rounded Bow

b) Sharp Bow

Figure 4.3 Rounded and Sharp Bow Horizontal Sections

7

z y

x

Figure 4.4 Isometric Rounded and Sharp Bow Sections

The bow sections were based on a simple formula, For the rounded bow:

(

x = 20 + k z 400 - y 2

For the sharp bow:

(

x = k z 400 - y 2

)

)

where kz is a function of the height above the bottom of the tanker hull as follows: Height z (m) 28 27 26 25 24 23 22 21 20 19 : 10 9 8 7 6 5 4 3 2 1 0

Rounded Bow - kzr 1.720 1.640 1.560 1.480 1.400 1.320 1.240 1.160 1.080 1.000 1.000 1.000 1.000 1.051 1.102 1.153 1.204 1.255 1.204 1.153 1.102 1.051

8

Sharp Bow - kzs 5.350 5.200 5.050 4.900 4.750 4.600 4.450 4.300 4.150 4.000 4.000 4.000 4.000 4.100 4.200 4.300 4.400 4.500 4.400 4.300 4.200 4.100

4.4 RELATIVE BALLAST CONDITION The two ballast conditions considered were based on the extreme conditions of: 1) a full tanker colliding with an empty FPSO and 2) an empty tanker colliding with a full FPSO. These two conditions are illustrated in Figure 4.5 whereby, in the first scenario, the bow of the tanker is 1.5m above the stern of the FPSO while, in the second scenario, the tanker bow is 10m above the FPSO stern. It should be noted that in the full tanker condition the raked section of the bow collides with the FPSO stern. The contact area will increase gradually in this case, leading to a relatively large penetration. For the empty tanker scenarios it can be seen for the flat FPSO stern, in particular, there will be a large contact area shortly after impact, which can be expected to lead to smaller deformations but larger forces. It can be seen in Figure 4.5 that the bow overhangs the main deck of the FPSO by 1.05m and 6.3m for the full tanker and empty tanker scenarios respectively. The effect of this overlap on the FPSO topside is discussed in Section 6.5. 6.3m

1.05m 10m 1.5m

MSL

a) Full Tanker : Empty FPSO

b) Empty Tanker : Full FPSO

Figure 4.5 Relative Ballast Conditions for the Tanker and FPSO

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5. RESULTS OF RIGID TANKER BOW IMPACT SCENARIOS In Sections 3 and 4 of this report, the impact energy absorption calculation methodology and impact scenarios were defined. Initial calculations were performed to meet particular impact energy targets. In these calculations the methodology in Section 3.3 was followed. Latterly, the HSE requested that FPSO stern penetration graphs be derived for all eight scenarios, for the range of impact energies between 0 MJ and 100 MJ. Consequently, it was decided to determine the penetration at 2 MJ intervals for each scenario. For each energy calculation, 10 energy absorption iteration steps were specified. This implied that a total of 8 x 50 x 10 = 4000 damage calculations would be required. To meet this objective some automation of the calculation process was required. Therefore, for each scenario, two Excel worksheets were produced. In the first worksheet, the bow of the tanker was moved horizontally in steps of 0.05m into the FPSO stern, up to a maximum 6.0m penetration. At each interval, the volume of plate in the horizontal (xy-) and the vertical longitudinal (xz-) planes and beams in the horizontal longitudinal (x-) and vertical (z-) directions were estimated using simple 3-dimensional formulations. In addition, the volume of plate in the vertical transverse (yz-) plane and transverse (y-) beams were also estimated, although these were not included in the total volume of damaged (not torn) material Vs. An example of this worksheet for a 0.25m step is given in Appendix A In the second worksheet, the energy calculation is performed for each 2 MJ step. For each of these energy calculations, 10 FPSO penetration iterations were performed, each one linked to the Vs calculated in the first worksheet. The penetrations were manually selected, becoming more refined as all the impact energy was absorbed. The tenth iteration exceeded this value of energy so that the maximum penetration could be interpolated between the ninth and tenth steps. For example, in Scenario No. 1 at 50 MJ, by adjusting the 9th and 10th iterations, it was determined that failure would occur at a penetration of around 3.0m. Consequently, penetration steps were specified as: 0.35m, 0.70m, 1.05m, 1.40m, 1.75m, 2.10m, 2.45m, 2.80m, 3.00m and 3.05m. Damaged steel volumes relating to these penetrations were extracted. The approach described in Section 3.3 was performed for these penetrations and steel volumes, enabling incremental velocities, forces, energies and time steps to be determined. By interpolation, the FPSO stern would absorb the 50 MJ of energy after 3.04m penetration, the process taking around 5 seconds. An example of this energy calculation is given in Appendix B. In reality, 10 iterative steps were more than required since the iterative process appeared relatively insensitive to the penetration values selected in the first six or seven steps. Figure 5.1 presents the results of the energy absorption calculations for all eight scenarios specified. A summary of the minimum and maximum penetrations at 10 MJ intervals is given in Table 5.1. It can be seen in Figure 5.1 that, for most scenarios, there is an initial rapid penetration within the first 1.0m of the stern, following which the increase in penetration with increase in KE is fairly linear. The main exceptions are: 1. The two scenarios concerning the flat stern of an empty FPSO (Scenarios 5 & 6). In these cases, there is relatively stiff resistance at penetrations between 0.5m and 1.5m. This is due to three vertical beams near the top of the stern where the raked bow is indenting. Once these have been fully encountered there is less material to resist additional penetration and these scenarios then have the largest penetration per unit impact energy.

10

2. The two scenarios concerning the flat stern of a full FPSO (Scenarios 7 & 8). In these cases, the stern initially provides little resistance at its base to the top of the extruding bulbous bow. However, following a penetration of 0.5m of this bulbous extrusion, the 9.8m high vertical stern is in contact with the vertical mid-bow section over a relatively large contact area. Consequently, there is a relatively large amount of stiffening resisting penetration and these cases then have the smallest penetration per unit impact energy. The results are otherwise predictable with the sharper bow penetrating further than the rounded bow, although generally by only around 0.5m, despite the extreme bow shapes considered. Furthermore, when the tanker was full, and consequently impacted the stern with the raked section, penetrations were also greater than the case of the empty tanker high out of the water. (Note. the mass of the tanker is not a factor since similar collision energies are being compared). From Table 5.1, it can be seen that the minimum penetration is around half the maximum penetration for the scenarios considered. At a substantial collision energy of 50 MJ, where all this energy is assumed to be absorbed by the stern with no loss to friction or waves, the maximum penetration is estimated to be less than 4m in the horizontal direction. In Figure 5.2, the damage calculations have been plotted to a new base to relate to the amount of overlap that will occur between the tanker and FPSO, i.e., for the full tanker the bow protrudes by 1.05m relative to the point of impact while for the empty tanker the bow protrudes by 6.3m. Thus for the case of an empty tanker colliding with a inclined stern full FPSO at 50 MJ impact energy, it is estimated that the bow would overlap the stern by around 9.5m. Clearly this has the potential to cause considerable damage to topside equipment. The sensitivity of these reported damage calculations will be discussed further in Section 6. Table 5.1 Range of Axial Penetration from the Eight Scenarios

KE (MJ) 10 20 30 40 50 60 70 80 90 100

Min. Pen. Max. Pen. (m) (m) 0.73 1.48 1.17 2.11 1.41 2.67 1.65 3.20 1.90 3.78 2.13 4.27 2.36 4.67 2.58 5.02 2.79 5.30 3.00 5.56

In Figure 5.3, one Scenario is illustrated in detail. Scenario 1 (100 MJ) axial penetration is plotted against: · the energy absorbed in work done, · the collision force, · the tanker velocity throughout the collision · and the time taken to reach a particular indentation Therefore, it can be seen that a full collision history can be produced for any collision event.

11

All Scenarios - Axial Penetration Relative to Stern 6.0

1 Full Tanker - Rounded Bow + Empty FPSO - Inclined Stern 2 Full Tanker - Sharp Bow + Empty FPSO - Inclined Stern 3 Empty Tanker - Rounded Bow + Full FPSO - Inclined Stern

5.0

4 Empty Tanker - Sharp Bow + Full FPSO - Inclined Stern 5 Full Tanker - Rounded Bow + Empty FPSO - Flat Stern 6 Full Tanker - Sharp Bow + Empty FPSO - Flat Stern

Axial Peneration (m)

4.0

7 Empty Tanker - Rounded Bow + Full FPSO - Flat Stern 8 Empty Tanker - Sharp Bow + Full FPSO - Flat Stern

3.0

2.0

1.0

0.0 0

10

20

30

40

50

60

70

80

Kinetic Energy (MJ)

Figure 5.1 Axial Penetration of the Tanker Bow into the FPSO Stern for the 8 Scenarios

12

90

100

All Scenarios - Axial Overlap of Bow Relative to Stern 14.0

12.0

1 Full Tanker - Rounded Bow + Empty FPSO - Inclined Stern

2 Full Tanker - Sharp Bow + Empty FPSO - Inclined Stern

3 Empty Tanker - Rounded Bow + Full FPSO - Inclined Stern

4 Empty Tanker - Sharp Bow + Full FPSO - Inclined Stern

5 Full Tanker - Rounded Bow + Empty FPSO - Flat Stern

6 Full Tanker - Sharp Bow + Empty FPSO - Flat Stern

7 Empty Tanker - Rounded Bow + Full FPSO - Flat Stern

8 Empty Tanker - Sharp Bow + Full FPSO - Flat Stern

Axial Overlap (m)

10.0

8.0

6.0

4.0

2.0

0.0 0

10

20

30

40

50

60

70

80

Kinetic Energy (MJ)

Figure 5.2 Axial Overlap of the Tanker Bow over the FPSO Stern for the 8 Scenarios

13

90

100

100

35

90 30

80 25

Impact Force (MN)

Work Done (MJ)

70 60 50 40 30

20

15

10

20 5

10 0

0

0

1

2

3

4

5

0

1

Stern Penetration (m )

2

3

4

5

4

5

Stern Penetration (m )

1.4

6.0

1.2

5.0

1.0

0.8

Time (secs)

Veocity (m/s)

4.0

0.6

3.0

2.0

0.4

1.0

0.2

0.0

0.0

0

1

2

3

4

5

0

Stern Penetration (m )

1

2

3

Stern Penetration (m )

Figure 5.3 Axial penetration against Energy, Force, Velocity and Time Scenario 1 – 100 MJ

14

6. SENSITIVITY STUDIES 6.1 ACCURACY OF THE FRIEZE ENERGY EQUATION The Frieze energy absorption equation described in Section 3.2: Energy (kJ) = 60.5 (sd/sy) Vs + 6.0 t2 l had its coefficients adjusted to give a best fit to 41 test data. These test data covered longitudinal members, framing and bulbous bow sections. Following screening from an original data set of 52 data points (Coefficient of Variation (COV) = standard deviation / mean = 33%), the data was reported to have a COV = 19%. The tearing term was based on the average of coefficients proposed by Vaughan (Reference 8), Reckling (Reference 9) and Woisin (Reference 10), the COV of these also being around 20% Since most data excluded the dynamic strain-rate term and the tearing term, the mean energy expression becomes 60.5Vs. Assuming data is Lognormally distributed about the mean, then 2 standard deviations to this mean expression gives the 97.7% confidence interval KE = [43.5Vs, 83.5Vs]. The mean energy-penetration curve and 97.7% confidence interval is illustrated in Figure 6.1. For 50 MJ impact energy the 97.7% confidence interval to a mean penetration of 3.0m is [2.5m, 3.75m] The energy model has considered all potential damage modes, although some modes, for example crushing, are better represented than others in the test database are. It is felt, however, that any bias in damage type is representative of the damage likely to occur in bow collisions. Scale effects were considered to be significant, although only tearing was ultimately found to be particularly scale-dependent. These scale effects have been represented in the model. Overall, while the database is a compilation of failure modes and smaller in number than would be desirable, it is felt than all gross uncertainties will have been represented in the model. 6.2 BOW AND STERN SHAPES The bow and stern shapes, along with the relative ballast conditions of the two vessels have been described in Section 4. These were required to be typical of offshore tankers and FPSOs, yet at the more extreme ends of the design spectrum. The FPSO sterns being based on drawings are accurate and cover two different designs of vessel that have significantly different shaped sterns. The tanker plan view (Rounded or Sharp bow) also appears representative. The vertical bow shape was not varied. Other shuttle tankers may have longer raked sections and larger bulbous bows. The effect of a longer raked section would be insignificant for the vessel in full or near-full ballast condition since the raked section is in contact with the FPSO stern throughout the duration of any impact (up to 100 MJ) in the current model. However, in lighter condition, it will become more likely that a raked rather than flat section will contact the FPSO stern. This would imply that in the less full condition the large raked bow tanker would behave more like the condition specified for a full tanker, i.e. higher penetrations, but correspondingly lower forces.

15

Scenario 1 - Mean Energy Equation and 97.7% Confidence Interval 6.0 Lower confidence interval Mean Upper confidence interval

5.0

Axial Peneration (m)

4.0

3.0

2.0

1.0

0.0 0

10

20

30

40

50

60

70

Kinetic Energy (MJ)

Figure 6.1 Mean Energy Equation and 97.7% Confidence Interval

16

80

90

100

Similarly, a larger bulbous bow would have more penetration when the tanker is less full. This acts in a similar manner to the raked bow, leading to higher penetrations (if it contacts prior to the raked bow section) and smaller forces. If the tanker is comprised of two sections rather than three (i.e. no vertical bow section between the raked and bulbous bow) then both the raked bow and bulbous bow could come into contact with the FPSO stern during the collision. In such cases, the energy absorption of these two elements would combine leading to an overall reduction in bow penetration. Appurtenances on the bow, such as winching gear, have not been considered. In general these would be within the upper 1m – 2m of the bow, and while they could increase the damage to the stern, this would be to topside equipment above the main or upper decks on the FPSO. 6.3 ENERGY CALCULATION AND DISTRIBUTION 6.3.1 Vessel Mass and Velocity In the energy calculations reported in Section 5, the vessel mass was maintained constant at 110,000 tonnes full and 80,000 tonnes empty. A 5% added mass was applied to this value and based on KE = ½mv², the velocity for any specified energy level was calculated. The dynamic strain rate factor is a function of velocity such that an enhancement in energy absorption can occur at high strain rates where the enhancement factor, sd/sy, applies for values of 5 v t/a² > 0.001 s-1, in which v is the velocity, in m/s; t is the thickness of plating, in m; and a is the transverse frame spacing, in m. In our Scenarios, t = 0.013m and the main transverse frames are widely spaced a = 9.5m or a = 11m. Therefore, the dynamic strain rate factor is only effective in Scenarios 7 & 8 where the KE > 80 MJ, the vessel mass is 80,000 x 1.05, and the transverse framing is 9.5m. In these cases the dynamic factor is less than 1%. Taking an extreme case of a 40,000 tonne vessel with transverse framing (15mm thick) spaced at 5m intervals, the 100 MJ impact would yield an enhancement factor of sd/sy = 1.10, or the non-tearing term would become 66.5Vs. Consequently, given the other uncertainties in the model, it appears reasonable to use the graphs produced in Section 5 of this report, with no adjustment for dynamic strain rate effects, for all shuttle tanker masses and velocities. 6.3.2 Energy Absorbed by the Bow To estimate the proportions of damage to the tanker bow and FPSO stern in the Scenarios performed in this project, simultaneous bow and stern damage were considered for one Scenario. Drawings of a typical shuttle tanker were obtained, and plate and stiffening details were arranged within the original bow shape defined in Section 4.3. Since the drawings were more similar to the sharp bow this was preferred and Scenario 2 Full tanker – Empty FPSO inclined stern was selected. Due to confidentiality restrictions, the stiffening details of the shuttle tanker are not presented in this report. The calculation method used was a simplification of work performed at MIT by Kim (Reference 11) and Gooding (Reference 12). In this approach, force-deformation curves are independently produced for the bow and the stern, assuming each impacts a rigid stern and bow, respectively. An example of this is given in Figure 6.2 based on the Scenario 2, the 60 MJ collision event.

17

30

Damage Force (MN)

25 20 15 10 5 0 -5.0

-4.0

-3.0

-2.0

Shuttle Tanker

-1.0

0.0

1.0

2.0

Deformation (m)

3.0

4.0

5.0

FPSO

Figure 6.2 Force-Deformation Relationship in 60 MJ Impacts with Rigid bodies

The bow energy was calculated in an identical manner to that of the stern, previously described. However, in this instance since the hull of the tanker is 1.5m above the FPSO, the hull plate is considered to tear and consequently tearing is included in the calculations for damaged area. The combined damage calculation then requires the minimum force to be determined that gives the work done in damage to the bow + work done in damage to the stern = target KE. By linear interpolation between the measurement points on the graph it was determined that at a maximum force = 20 MN, the tanker bow will have absorbed 25.45 MJ with a deformation of 2.31m, while the FPSO stern will have absorbed 34.55 MJ with a deformation of 2.87m. i.e., the FPSO stern will absorb 58% of the energy. In Table 6.1, the results are summarised for Scenario 2 at 20, 40, 60, 80 and 100 MJ energies. It can be seen that the FPSO consistently absorbs around 60% of the energy and that the stern penetration is reduced by around 25%-30% by considering simultaneous bow crushing. Table 6.1 Distribution of Absorbed Energy Between the Bow and Stern

Total Energy (MJ) 20 40 60 80 100

Shuttle Tanker Bow Def. Energy (m) (MJ) 1.41 8.66 1.86 16.46 2.31 25.45 2.63 33.31 2.84 38.29

FPSO Stern Def. Energy (m) (MJ) 1.56 11.34 2.31 23.54 2.87 34.55 3.40 46.69 4.03 61.71

Force (MN) 13.0 16.8 20.0 22.4 23.9

Original Stern Def. % reduction (m) in def. 2.11 26 % 3.12 26 % 3.96 28 % 4.77 29 % 5.53 27 %

Therefore, it has been shown that calculations can be performed, at twice the effort, giving the damage to both the tanker bow and FPSO stern during a collision event. For the one Scenario considered the tanker absorbs around 40% of the energy for around this 25% - 30% reduction in the maximum stern penetration compared to the results presented in Section 5.

18

6.3.3 Distribution of Energy The calculations performed in this study, in accordance with the specification by the HSE, are based on the assumption that all the impact energy is absorbed by the mechanics of the collision (e.g. crushing, tearing, etc). This section gives consideration to this level of conservatism in this assumption. 6.3.3.1 Proportion of energy absorbed For a 90° collision with no angular momentum then the amount of absorbed energy can be calculated as follows. Before the collision: KE1 = ½ m1 v12 KE2 = ½ m2 v22 = 0

Momentum1 = m1 v1 Momentum1 = m1 v1 = 0

striking ship motionless struck ship

After the collision: KE1+2 = ½ (m1+m2) v1+22

Momentum1 = (m1+m2) v1+2

both ships free to move

From conservation of momentum: m1 v1 = (m1+m2) v1+2 or

v1+2 = m1 v1 /(m1+m2)

The total energy before impact = The total energy after impact = =

½ m1 v12 ½ (m1+m2) . [m1 v1 /(m1+m2)]² ½ (m1 v1)²/(m1+m2)

The total energy absorbed

½ m1 v12 - ½ (m1 v1)²/(m1+m2) ½ m1 v12 [1- m1/(m1+m2)]

= =

Therefore, where the mass of the two vessels is nearly identical, around half the kinetic energy of the collision would be expected to be absorbed, assuming that there was no resistance to motion. For an FPSO this is not the case since the FPSO is moored, although the mooring will allow some motion with very little, if any resistance. Therefore, it can be estimated that in some 50% of the initial KE would be absorbed during the collision. Proportionally, more energy being absorbed when the tanker is empty and the FPSO full. 6.3.3.2 Distribution of impact energy In Section 4.1, it was noted that, conservatively, the FPSO stern would absorb all the kinetic energy. Estimates of the proportion of energy dissipated in collisions will vary depending upon the scenario considered. However, as an example, Suzuki et al (Reference 13), quote two collision scenarios between a large 65,000 tonne tanker and a 7,000 tonne ship in full and ballast condition. The distribution of absorbed energy based on finite element analyses were:

Striking ship in ballast condition Striking ship in fully loaded condition

Striking Ship 61%

Struck Ship 25%

Friction

Waves

13%

1%

7%

61%

30%

2%

It can be seen that the difference in the distribution of energy absorption between striking and struck ships in this case is considerable. Friction accounts for around a quarter of the energy, while the wave absorb very little energy.

19

6.4 DIRECTION OF VESSEL COLLISION In the Scenarios considered in this report, the impact velocities between FPSO and shuttle have been assumed to occur in a horizontal plane, along the fore and aft direction of the two vessels. This model has ignored the vertical component due to relative pitch and heave motions between the bow of the shuttle and the stern of the FPSO. In this Section, the objective is to determine the ranges of possible impact velocities associated with likely operating sea states. A simple estimate of absorbed impact energy is also provided for the purpose of comparison with the horizontal impact energies used elsewhere in this report. 6.4.1

Analysis Method

The heave and pitch RAOs (Response Amplitude Operators) for each vessel are used to calculate the RAOs for the vertical motion of the bow of the shuttle tanker and a possible point of impact towards the stern of the FPSO. Two impact points were considered, one on the stern and another at some fraction of the vessel’s length forward of the stern. A fraction of 10% was chosen to investigate whether the results might prove sensitive to the assumed separation between the vessels at impact. The RAOs of both vessels are converted to a common phase reference point so that the difference between the two motions corresponds to the relative motion. Using the standard spectral calculation, the spectrum of relative motion is obtained by multiplying the wave height spectrum by the square of the modulus of the RAO, and the spectrum of relative velocities by multiplying the motion spectrum by the square of the wave frequency (in radians per second). The ranges of the relative motions and velocities are obtained by integrating the response spectra with respect to frequency. Significant values are obtained on the basis of a narrow band assumption and the associated Rayleigh distributions as 4 times the root mean square values. Finally, the significant value of the impact velocity is determined as half the significant value of the velocity range. The sea states are taken (according to DnV RP 2, 1985) as Pierson Moscowitz spectra with a specified significant wave height and a range of possible zero crossing periods. Three calculations have been implemented for the lower, midpoint and upper limits of the specified range of zero crossing periods. Three values of significant wave height have been used, 3.0 m, 4.5 m and 6.0 m. The lowest value corresponds to the maximum sea state at which a shuttle might approach an FPSO to begin connection procedures. The value of 4.5 m corresponds to a normal upper limit for loading operations. The highest value could arise in a rising storm if separation of shuttle and FPSO is delayed for any reason. The values quoted as the results of this calculation are significant values, corresponding to the mean of the highest third of the cycles. Hence, in any 9 wave cycles, over a period of between 40 seconds and two minutes, there would normally be at least one and perhaps two wave cycles in which the relative velocity exceeds the quoted value. The maximum relative velocity over an extended period would be nearly double the significant values. 6.4.2

Sources of RAO Data

Obtaining appropriate and fully documented RAO data proved extremely problematic. Although the HSE had thought they would be able to provide suitable data from submitted safety cases, this was not possible. The HSE was able to provide this study with one recent stern collision damage report, see Reference 1, although this did not report specific RAOs. Various sources of data were available for FPSO RAOs based on diffraction calculations but there was no corresponding data on associated shuttle tankers.

20

Navion reviewed various model test reports for their shuttle tankers but these data were presented in a form suitable for assessing whether defined thresholds had been exceeded and not suitable for the types of calculation specified below. One operator provided tabulated values for an FPSO and its associated shuttle tanker, referred to as FPSO1. These values appeared to have been derived from model tests but it was evident from the tabulations that different axes systems, motion reference points and phase conventions had been used for the two vessels and these conventions were not fully specified. By inspection of the low frequency trends of heave and pitch RAOs, it was possible to deduce a set of underlying conventions that agreed (at least approximately) with physical expectations and these were adopted as true for one series of calculations. Given the difficulty in interpreting phase data for heave and pitch RAOs, an alternative source was sought as confirmation that the calculated values were realistic. MARIN have published a fully documented set of RAOs for one FPSO, within the confidentially restrictions of a JIP, referred to as FPSO2. Given that FPSOs are frequently converted from tankers, it was decided to implement a second calculation in which the RAOs of FPSO2 be used as those of both the FPSO and those of the shuttle. Whatever the approximation involved in this assumption, there is no uncertainty in the conventions for definitions of the RAOs. For the purposes of making the calculations in a well defined system of co-ordinates and phase conventions, all data was transformed to equivalent values corresponding to the axes and phase conventions used in the CMPT publication on the design and analysis of floating production systems. 6.4.3

Results

Table 6.2 presents the results of the calculations for the impact velocities between an FPSO and a shuttle, based on data from an operator, that appear to have been obtained from model tests. Table 6.3 is based on assuming both shuttle and FPSO have RAOs similar to those evaluated for FPSO2 using a diffraction calculation. Table 6.2 Impact between FPSO and Shuttle: Results for various values of significant wave height, zero crossing period and impact point location (based on FPSO1 RAOs)

Range of spectral crossing periods (s)

Significant value of impact velocity between the bow of a shuttle and the stern of an FPSO (in m/s) Impact 10% fwd of stern of FPSO Impact on stern of FPSO

Significant wave height = 6.0 m tzlower = 6.87 tzmid = 10.14 tzupper = 13.42

1.37 3.64 3.54

1.34 3.46 3.55

Significant wave height = 4.5 m tzlower = 5.85 tzmid = 8.74 tzupper = 11.62

0.53 2.23 2.84

0.70 2.04 2.78

Significant wave height = 3.0 m tzlower = 4.66 tzmid = 7.08 tzupper = 9.49

0.17 0.77 1.70

0.23 0.73 1.59

21

Table 6.3 Impact between FPSO and Shuttle: Results for various values of significant wave height, zero crossing period and impact point location (Based on FPSO2 RAOs)

Range of spectral crossing periods (s)

Significant value of impact velocity between the bow of a shuttle and the stern of an FPSO (in m/s) Impact 10% fwd of stern of FPSO Impact on stern of FPSO

Significant wave height = 6.0 m tzlower = 6.87 tzmid = 10.14 tzupper = 13.42

1.58 2.37 1.85

1.57 2.59 2.10

Significant wave height = 4.5 m tzlower = 5.85 tzmid = 8.74 tzupper = 11.62

0.72 1.74 1.63

0.71 1.84 1.82

Significant wave height = 3.0 m tzlower = 4.66 tzmid = 7.08 tzupper = 9.49

0.20 0.85 1.20

0.19 0.85 1.29

6.4.4

Discussion and Estimate of Deformation Energy

6.4.4.1 Impact velocities Initial calculations provided to the HSE, based on a simple box model gave relative velocities in the range 0.6 - 2.1 m/s for significant wave heights between 4.5 m and 6.0 m. In Tables 6.2 and 6.3, velocities range between 0.5 - 3.6 m/s. These higher velocities in part reflect the inclusion of dynamic amplification of pitch or heave and, further, the initial model had clearly defined cancellation frequencies that yield zero response at precise values of frequency. Full diffraction calculations and model test data include various levels of dynamic amplification depending on the relationship wave frequency and natural periods in heave and pitch. Model tests tend not to show zero but rather reduced values of RAOs at cancellation frequencies. Both effects will increase relative motion and relative velocity in the present calculation. The impact velocities calculated for the middle and upper values of wave zero up cross period for the FPSO and shuttle are particularly large. Inspection of the input RAOs shows that this is entirely due to the values of the pitch RAO for the shuttle which show a significant effect of dynamic amplification for the assumed loading condition (fully loaded). It is entirely possible for either an FPSO or shuttle to exhibit such amplification in the frequency ranges of interest to this study. If a ship’s master is totally familiar with the characteristics of his vessel, he may anticipate this type of response and make arrangements to leave location before it occurs. However, it is apparent from the data received that the relative velocity is a complex function of the response of both vessels and so it is difficult to be certain that it will always be possible to avoid this type of behaviour. It is clear that the relative velocity depends as strongly on wave period as it does on wave height and generally the velocity is lowest for the shortest wave periods. The trend with wave height may be examined by comparing the velocities for the central value of Tz at each wave height, as follows.

22

Significant wave height (m) 3.0 4.5 6.0

Significant relative velocity (m/s) 0.7 – 0.8 1.7 – 2.2 2.4 – 3.6

6.4.4.2 Deformation energy Estimates of the energy involved in the collision and the maximum energy available for deformation damage to the hulls can be evaluated on the basis of fairly simple considerations of conservation of energy, momentum and angular momentum. It is easiest to implement this calculation for FPSO2 because the mass of both vessels is defined and there is an inherent symmetry in the calculation. Impact energy will arise from both the relative heave and the relative pitch of both vessels. Examining the relative motion and velocity spectra shows that the main contribution arises for frequencies between 0.45 and 0.55 radians per second. In this frequency range, the contributions to bow heave are 0.4m per m wave height due to vessel heave and 1.5m per m wave height due to vessel pitch. Further these motions are not fully in-phase. Consequently, the pitch motion dominates the relative velocity and the associated energy will be evaluated on this basis. The total energy available will be approximately at a maximum when the downward velocity of the bow is equal to the upward velocity of the stern (or visa versa). The energy associated with the pitch motion for this situation is E = [0.5 (I + IA) (dθ/dt)2] * 2 L/2 dθ/dt = vbow = vrelative/2 Where I + IA dθ/dt L vbow and vrelative

is the moment plus added moment of inertia of the vessel is the angular velocity is the length of the vessel is the velocity of the bow of the shuttle(equal & opposite to the FPSO stern) is the relative velocity

Based on the approximations that I + IA = (m + mA )L2/12, m = 115000 t and mA = 1.4 m, this formula gives total energy in the kinematics of the motion as follows. Hs (m) 6.0 4.5 3.0

Tz (=Tzmid) (s) 10.1 8.7 7.1

vrelative (m/s) 2.60 1.84 0.85

Total Energy (MJ) 360 180 42

However, after the impact occurs, a certain fraction of the energy will remain in the form of angular velocity of each vessel (rotating about the point of impact). In the simplest terms, angular momentum will be conserved while energy will be shared between the kinematics of the motion and deformation of the structure. It may easily be shown that in the idealised situation 25% of the total energy remains in the form of kinetic energy while a maximum of 75% may be absorbed in deformation of both structures. Hence the maximum available deformation energy in these cases is summarised as follows.

23

Hs (m) 6.0 4.5 3.0

Deformation Energy (MJ) 270 135 32

In so far as the effects of vertical impact can be compared with those of horizontal impact these deformation energies should be used for comparison with the results provided elsewhere in this report. 6.5 IMPACT ON TOPSIDE EQUIPMENT A significant part of the HSE project on safety case reviews of FPSO stern collisions (Reference 1) was concerned with vulnerable topside equipment. It was noted that the equipment concerned could only be assessed on a vessel-by-vessel basis, although two very different generic types were noted: · accommodation aft · accommodation forward In many cases where the accommodation is aft, the helideck will be sited aft, and will usually extend beyond the stern. Where an empty tanker impacts the stern, it has been shown that the bow could be around 10m above the stern and may overlap by 9.5m in extreme cases. More realistically, the rear supports of the helideck at the stern would be dented or destroyed. This would render the helideck inoperable, although it would be very unlikely that a helicopter would be stationed on the helideck. The helideck may collapse, partly onto the tanker, but unless personnel were in the area the outcome would not be catastrophic. Where the accommodation is forward, process equipment is usually sited near the FPSO stern. This may include a flare stack or turret which, depending upon design and impact severity would be vulnerable to damage and collapse. This situation is far more serious with flammable material potentially released and an ignition source provided through impact friction in the same area. The effect of a fire or blast would impact on both the FPSO and tanker and has the potential to escalate into a major incident. As stated, these would need to be considered on a case-by-case basis. However, it is hoped that the energy-displacement curves and the energy absorption method described in this report will assist in safety case studies into the risks associated with shuttle tanker – FPSO collisions.

24

7. SUMMARY AND CONCLUSIONS In this project, the penetration of a shuttle tanker bow into the stern of an FPSO has been estimated. The formula for estimating this penetration is considered reasonably accurate although additional test data of all sizes would reinforce the accuracy claimed. The main inaccuracies have been through ignoring deck equipment and smaller strengthening components, by assuming a fully rigid bow and by assuming that the struck vessel does not move as a result of the impact. All of these assumptions will lead to larger estimated penetrations than may be seen in practice. Eight Scenarios have been considered with variations in bow shape, stern shape and relative ballast. The stern penetration was calculated for these scenarios for impact energies absorbed between 0 MJ and 100 MJ. As expected, the sharper bow penetrated by more than the rounded bow – an extra 0.5m penetration being a typical value. The flat stern would generally be more resistant than an inclined stern – the exception being the full tanker and empty FPSO at energies > 35 MJ. The tanker lower in the water also yielded larger penetrations since the raked bow was in contact with the stern rather than the vertical mid-bow section. For two-section bows, comprising a raked section and a bulbous section – with no near-vertical section – travelling high in the water, energies between those reported for a full and an empty tanker could be anticipated. The effect of the bow crushing simultaneously with the stern was considered for one highpenetration Scenario. This suggested that at all energy levels, the bow would take around 40% of the total energy absorbed. This would reduce the penetration into the stern by 25%-30%. It was also noted that not all impact energy would be absorbed by the two colliding vessels. The amount would depend upon the mooring stiffness and the mass of the two vessels concerned. The effect of relative vertical velocities arising at shuttle tanker bow to FPSO stern impacts may arise from the vessels’ pitch and heave motions. Relative vertical velocities at impact may be substantial and the ranges 0.7 – 0.8 m/s in 3.0 m waves, 1.7 – 2.2 m/s in 4.5 m waves and 2.4 – 3.6 m/s in 6.0 m waves. All wave values quoted being significant values, i.e. the average 1/3rd highest values. For one scenario, using median values from the range of impact velocities, maximum deformation energies are estimated as 32 MJ, in 3.0 m waves, 135 MJ in 4.5 m waves and 270 MJ in 6.0 m waves. It should be remembered that these energies cannot be related to the energy : penetration curves produced for horizontal impacts due to the different stiffening arrangements in a vertical collision. Topside equipment is likely to be significantly damaged in high-energy collisions, while the hull could be holed, no high-risk storage areas would be damaged. This is in part due to the large overhang that may occur if the collision was between an empty tanker and a full FPSO. In these cases the bow would overlap by 6.3m before the hulls collided. For accommodation sited at the aft of the FPSO, this is likely to lead to significant damage to the helideck, although it would be unlikely that this event would escalate. For the more common situation of aft sited process equipment, in some vessel designs the flare stack or turret is vulnerable and could lead to a fire/blast escalating event. Such scenarios should be considered in the FPSO safety case and mitigation measures such as bumpers should be designed accordingly.

25

8. REFERENCES 1

BOMEL CONSORTIUM FPSO/Shuttle Tanker Offloading Safety Case Review (Stern Collision Study). HSE Task No B\0035, C854\01\014R, Rev C, August 2000.

2

MINORSKY, V. U. An Analysis of Ship Collision with Reference to Protection of Nuclear Power Plants. Journal of Ship Research, 3, 1-4, 1959.

3

FRIEZE, P.A. Dartford Crossing: An Assessment of the Likely Forces Involved and Extent of Overrun Arising from Ship Collision with a Main Pier. Report to T H Engineering Services Ltd, Croydon, 1988.

4

FRIEZE, P. A. Second Severn Crossing Bridge Pier Impacts. Report to Halcrow-SEEE Joint Venture, 1991.

5

VAUGHAN, H. Bending and Tearing of Plate with Application to Ship-Bottom Damage. The Naval Architect, London, May, 97-99, 1978.

6

WOISIN, G. Design Against Collision. International Symposium on Advances in Marine Technology, Trondheim, Norway, 1979.

7

FRIEZE, P. A. AND SMEDLEY, P. A. Ship Bow Damage During Impacts with Ships and Bridge Piers. ICCGS 2001, Copenhagen, Denmark, 2001.

8

VAUGHAN, H. The Tearing Strength of Mild Steel Plate. Journal of Ship Research, 24, No 2, 96-100, 1980.

9

RECKLING, K. A. Beitrag der Elasto- und Plastomechanik zur Untersuchung von Schiffskollisionen. Jahrbuch der Schiffbautechischen Gesellschaft, 70, 443-464, 1976.

10

WOISIN, G. Comments on Vaughan - The Tearing Strength of Mild Steel Plate. Journal of Ship Research, 26, No 1, 50-52, 1982.

11

KIM, J. Y. Crushing of a Bow: Theory vs. Scale Model Tests. Joint MIT-Industry Program on Tanker Safety, Report #66, 1999.

12

GOODING, P. W. Collision with a Crushable Bow. Marine Technology, Vol. 38, No. 3, pp. 186-192, July 2001.

13

SUZUKI, K., OHTSUBO, H., ENDO, H. AND KITAMURA, O. Development and Application of Simplified Formula for Predicting Collision Damages. ICCGS 2001, Copenhagen, Denmark, 2001.

26

APPENDIX A EXAMPLE OF DAMAGE MATERIAL ESTIMATION

A1

Full Tanker - Empty FPSO - Sharp Bow - Flat Stern Step width = Overlap - x-axis

m

XY Horizontal Deck Plate - Crush y=0, z = 26.5m m y main deck m Area main deck m² Thickness mm

0.25

(adjusted from 0.05m for this example)

0.00

0.25

0.50

0.75

1.00

1.25

1.50

1.75

2.00

45.28 0.01 0.000 13.0

45.28 2.10 0.700 13.0

45.28 2.96 1.976 13.0

45.28 3.63 3.625 13.0

45.28 4.18 5.574 13.0

45.28 4.67 7.778 13.0

45.28 5.11 10.211 13.0

45.28 5.51 12.849 13.0

45.28 5.88 15.676 13.0

y=0, z = 20.8m y s/gear deck Area s/gear deck Thickness

m m m² mm

41.33 0.00 0.000 12.0

41.33 0.00 0.000 12.0

41.33 0.00 0.000 12.0

41.33 0.00 0.000 12.0

41.33 0.00 0.000 12.0

41.33 0.00 0.000 12.0

41.33 0.00 0.000 12.0

41.33 0.00 0.000 12.0

41.33 0.00 0.000 12.0

y=0, z = 18.0m y A/B deck Area A/B deck Thickness

m m m² mm

40.00 0.00 0.000 14.0

40.00 0.00 0.000 14.0

40.00 0.00 0.000 14.0

40.00 0.00 0.000 14.0

40.00 0.00 0.000 14.0

40.00 0.00 0.000 14.0

40.00 0.00 0.000 14.0

40.00 0.00 0.000 14.0

40.00 0.00 0.000 14.0

Volume

m².mm

0.00

9.10

25.69

47.13

72.46

101.12

132.74

167.04

203.79

0.00 0.000 16.5 0.00

0.36 1.499 16.5 24.74

0.71 4.234 16.5 69.87

1.07 7.768 16.5 128.18

1.43 11.943 16.5 197.07

1.79 16.668 16.5 275.02

2.14 21.880 16.5 361.02

2.50 27.533 16.5 454.30

2.86 33.592 16.5 554.27

14.0 0.00

14.0 0.00

14.0 0.00

14.0 0.00

14.0 0.00

14.0 0.00

14.0 0.00

14.0 0.00

14.0 0.00

0.00

0.00

0.00

0.00

0.00

0.00

0.00

0.00

0.00

2.24 15.32 0.00 0.00

5.04 19.84 0.00 0.00

7.84 24.36 0.96 0.00

10.64 28.88 3.71 0.00

13.44 33.40 6.46 0.00

10.80

17.56

24.88

33.16

43.23

53.30

3680

3680

3680

3680

3680

3680

YZ Transverse Stern Plate - Ignore Depth m Area m² Thickness mm Volume m².mm XZ Longitudinal Plate - Crush Thickness mm C/L Overlap m² Others too small to model Volume

m².mm

X Beam(s) - Longitudinal - After 0.8m penetration only - (C/L & @ intervals) - Crush C/L Overlap m².mm 0.00 0.00 0.00 0.00 y = 1.64m Overlap m².mm 0.00 1.76 6.28 10.80 y = 3.28m Overlap m².mm 0.00 0.00 0.00 0.00 y = 7.40m Overlap m².mm 0.00 0.00 0.00 0.00 Volume

m².mm

Y Beam(s) - Transverse - Next Area per beam mm² (z for Tanker) 26.500 no stiff. 25.680 m 24.860 m 24.040 m 23.220 m 22.400 m 21.580 m (z for FPSO) 21.380 no stiff. 20.560 m 19.740 m 18.920 m 18.100 m 17.280 m 16.460 m Volume

0.00

1.76

6.28

main frame >11m in x-direction - Ignore 3000 3680 3680 X o'lap X o'lap X o'lap X o'lap X o'lap X o'lap

0.00 -0.57 -1.15 -1.72 -2.30 -2.87

0.25 -0.32 -0.90 -1.47 -2.05 -2.62

0.50 -0.07 -0.65 -1.22 -1.80 -2.37

0.75 0.18 -0.40 -0.97 -1.55 -2.12

1.00 0.43 -0.15 -0.72 -1.30 -1.87

1.25 0.68 0.10 -0.47 -1.05 -1.62

1.50 0.93 0.35 -0.22 -0.80 -1.37

1.75 1.18 0.60 0.03 -0.55 -1.12

2.00 1.43 0.85 0.28 -0.30 -0.87

45.28 44.73 44.18 43.62 43.05 42.47 41.89

0.00 0.00 0.00 0.00 0.00 0.00

4.22 0.00 0.00 0.00 0.00 0.00

5.96 0.00 0.00 0.00 0.00 0.00

7.29 3.57 0.00 0.00 0.00 0.00

8.41 5.54 0.00 0.00 0.00 0.00

9.39 6.97 2.73 0.00 0.00 0.00

10.27 8.15 5.07 0.00 0.00 0.00

11.08 9.17 6.62 1.44 0.00 0.00

11.83 10.08 7.87 4.54 0.00 0.00

0.00

15.54

21.95

39.97

51.34

70.27

86.44

104.19

126.28

11

11

11

11

11

11

11

11

0.00 0.00

0.25 0.00

0.50 0.00

0.75 0.14

1.00 0.39

1.25 0.64

1.50 0.89

1.75 1.14

2.00 1.39

m².mm

Z Beam(s) - Vertical - C/L & @ intervals - Crush Width per beam mm 11 0.00 3.28

m m

X o'lap X o'lap

C/L Overlap y = 3.28m Overlap

m² m²

0.00 0.00

1.43 0.00

2.85 0.00

4.28 0.78

4.56 2.21

4.56 3.63

4.56 4.56

4.56 4.56

4.56 4.56

Volume

m².mm

0.00

15.68

31.35

64.20

98.69

130.04

150.48

150.48

150.48

A2

APPENDIX B EXAMPLE OF ENERGY CALCULATION

B1

Bow crushing calculation Initial KE Penetration

MJ m

Mass of boat Added mass Frame spacing Thickness

t t m mm

Penetration step 1 Penetration step 2 Penetration step 3 Penetration step 4 Penetration step 5 Penetration step 6 Penetration step 7 Penetration step 8 Penetration step 9 Penetration step 10

m m m m m m m m m m

Penetration step 1 Vel. of boat KE Crushed vol.

m m/s MJ m².mm -1 s

Impact rate factor Work done Average force KE at penetration

MJ MN MJ

Vel. at penetration time increment Cummulative Work Cummulative Time

m/s sec MJ sec

Penetration step 2 Vel. of boat KE Crushed vol.

m m/s MJ m².mm -1 s

Impact rate factor Work done Average force KE at penetration

MJ MN MJ

Vel. at penetration time increment Cummulative Work Cummulative Time

m/s sec MJ sec

Penetration step 3 Vel. of boat KE Crushed vol.

m m/s MJ m².mm -1 s

Impact rate factor Work done Average force KE at penetration

MJ MN MJ

Vel. at penetration time increment Cummulative Work Cummulative Time

m/s sec MJ sec

Penetration step 4 Vel. of boat KE Crushed vol.

m m/s MJ m².mm -1 s

Impact rate factor

m= ma = x= t=

v= E0 = Vs = e= f= W= F= E1 = v1 = t=

v= E0 = Vs = e= f= W= F= E1 = v1 = t=

v= E0 = Vs = e= f= W= F= E1 = v1 = t=

v= E0 = Vs = e= f=

10.0 0.910

20.0 1.582

30.0 2.442

40.0 3.197

50.0 3.780

60.0 4.270

70.0 4.667

80.0 5.020

90.0 5.303

100.0 5.562

110000 115500 9.5 13

110000 115500 9.5 13

110000 115500 9.5 13

110000 115500 9.5 13

110000 115500 9.5 13

110000 115500 9.5 13

110000 115500 9.5 13

110000 115500 9.5 13

110000 115500 9.5 13

110000 115500 9.5 13

0.20 0.30 0.40 0.50 0.60 0.70 0.80 0.85 0.90 0.95

0.15 0.30 0.45 0.60 0.75 0.90 1.10 1.30 1.55 1.60

0.30 0.60 0.90 1.20 1.50 1.80 2.10 2.25 2.40 2.45

0.40 0.80 1.20 1.60 2.00 2.40 2.80 3.00 3.15 3.20

0.50 1.00 1.50 2.00 2.50 3.00 3.30 3.50 3.75 3.80

0.60 1.20 1.80 2.40 3.00 3.60 4.00 4.15 4.25 4.30

0.60 1.20 1.80 2.40 3.00 3.60 4.20 4.45 4.65 4.70

0.60 1.20 1.80 2.40 3.00 3.60 4.20 4.75 5.00 5.05

0.70 1.40 2.10 2.80 3.50 4.20 4.90 5.20 5.30 5.35

0.70 1.40 2.10 2.80 3.50 4.20 4.90 5.45 5.55 5.60

0.20 0.416 10.0 19.91 0.0003 1.00 1.2 6.0 8.8

0.15 0.588 20.0 13.63 0.0004 1.00 0.8 5.5 19.2

0.30 0.721 30.0 33.43 0.0005 1.00 2.0 6.7 28.0

0.40 0.832 40.0 47.95 0.0006 1.00 2.9 7.3 37.1

0.50 0.930 50.0 63.32 0.0007 1.00 3.8 7.7 46.2

0.60 1.019 60.0 79.46 0.0007 1.00 4.8 8.0 55.2

0.60 1.101 70.0 79.46 0.0008 1.00 4.8 8.0 65.2

0.60 1.177 80.0 79.46 0.0008 1.00 4.8 8.0 75.2

0.70 1.248 90.0 107.20 0.0009 1.00 6.5 9.3 83.5

0.70 1.316 100.0 107.20 0.0009 1.00 6.5 9.3 93.5

0.390 0.50 1.20 0.50

0.576 0.26 0.82 0.26

0.696 0.42 2.02 0.42

0.802 0.49 2.90 0.49

0.894 0.55 3.83 0.55

0.978 0.60 4.81 0.60

1.062 0.55 4.81 0.55

1.141 0.52 4.81 0.52

1.203 0.57 6.49 0.57

1.273 0.54 6.49 0.54

0.30 0.390 8.8 13.52 0.0003 1.00 0.8 8.2 8.0

0.30 0.576 19.2 19.80 0.0004 1.00 1.2 8.0 18.0

0.60 0.696 28.0 46.03 0.0005 1.00 2.8 9.3 25.2

0.80 0.802 37.1 89.27 0.0006 1.00 5.4 13.5 31.7

1.00 0.894 46.2 125.38 0.0006 1.00 7.6 15.2 38.6

1.20 0.978 55.2 162.86 0.0007 1.00 9.9 16.4 45.3

1.20 1.062 65.2 162.86 0.0008 1.00 9.9 16.4 55.3

1.20 1.141 75.2 162.86 0.0008 1.00 9.9 16.4 65.3

1.40 1.203 83.5 190.68 0.0009 1.00 11.5 16.5 72.0

1.40 1.273 93.5 190.68 0.0009 1.00 11.5 16.5 82.0

0.372 0.26 2.02 0.76

0.558 0.26 2.02 0.52

0.660 0.44 4.81 0.87

0.741 0.52 8.30 1.01

0.817 0.58 11.42 1.13

0.886 0.64 14.66 1.24

0.979 0.59 14.66 1.14

1.064 0.54 14.66 1.06

1.116 0.60 18.02 1.17

1.191 0.57 18.02 1.11

0.40 0.372 8.0 14.52 0.0003 1.00 0.9 8.8 7.1

0.45 0.558 18.0 22.10 0.0004 1.00 1.3 8.9 16.6

0.90 0.660 25.2 83.22 0.0005 1.00 5.0 16.8 20.2

1.20 0.741 31.7 105.11 0.0005 1.00 6.4 15.9 25.3

1.50 0.817 38.6 127.67 0.0006 1.00 7.7 15.4 30.9

1.80 0.886 45.3 127.59 0.0006 1.00 7.7 12.9 37.6

1.80 0.979 55.3 127.59 0.0007 1.00 7.7 12.9 47.6

1.80 1.064 65.3 127.59 0.0008 1.00 7.7 12.9 57.6

2.10 1.116 72.0 129.07 0.0008 1.00 7.8 11.2 64.2

2.10 1.191 82.0 129.07 0.0009 1.00 7.8 11.2 74.2

0.351 0.28 2.90 1.04

0.537 0.27 3.36 0.80

0.591 0.48 9.84 1.35

0.662 0.57 14.66 1.58

0.731 0.65 19.14 1.78

0.807 0.71 22.38 1.95

0.908 0.64 22.38 1.78

0.999 0.58 22.38 1.64

1.054 0.65 25.83 1.82

1.133 0.60 25.83 1.71

0.50 0.351 7.1 15.38 0.0003 1.00

0.60 0.537 16.6 23.93 0.0004 1.00

1.20 0.591 20.2 79.65 0.0004 1.00

1.60 0.662 25.3 91.49 0.0005 1.00

2.00 0.731 30.9 91.19 0.0005 1.00

2.40 0.807 37.6 117.26 0.0006 1.00

2.40 0.908 47.6 117.26 0.0007 1.00

2.40 0.999 57.6 117.26 0.0007 1.00

2.80 1.054 64.2 145.11 0.0008 1.00

2.80 1.133 74.2 145.11 0.0008 1.00

B2

Bow crushing calculation (cont) Initial KE

MJ

Penetration step 5 (cont) Work done Average force KE at penetration

MJ MN MJ

Vel. at penetration time increment Cummulative Work Cummulative Time

m/s sec MJ sec

Penetration step 6 Vel. of boat KE Crushed vol.

m m/s MJ m².mm -1 s

Impact rate factor Work done Average force KE at penetration

MJ MN MJ

Vel. at penetration time increment Cummulative Work Cummulative Time

m/s sec MJ sec

Penetration step 7 Vel. of boat KE Crushed vol.

m m/s MJ m².mm -1 s

Impact rate factor Work done Average force KE at penetration

MJ MN MJ

Vel. at penetration time increment Cummulative Work Cummulative Time

m/s sec MJ sec

Penetration step 8 Vel. of boat KE Crushed vol.

m m/s MJ m².mm -1 s

Impact rate factor Work done Average force KE at penetration

MJ MN MJ

Vel. at penetration time increment Cummulative Work Cummulative Time

m/s sec MJ sec

Penetration step 9 Vel. of boat KE Crushed vol.

m m/s MJ m².mm -1 s

Impact rate factor Work done Average force KE at penetration

MJ MN MJ

Vel. at penetration time increment Cummulative Work

m/s sec MJ

10.0

20.0

30.0

40.0

50.0

60.0

70.0

80.0

90.0

100.0

W= F= E1 =

1.0 9.8 5.2

2.6 17.2 12.6

4.5 14.9 10.9

4.5 11.2 15.3

6.1 12.1 19.3

7.8 13.0 22.7

7.8 13.0 32.7

7.8 13.0 42.7

10.4 14.9 45.0

10.4 14.9 55.0

v1 =

0.300 0.32 4.81 1.65

0.467 0.31 7.39 1.39

0.434 0.63 19.14 2.52

0.515 0.73 24.66 2.95

0.578 0.81 30.73 3.30

0.627 0.89 37.29 3.62

0.753 0.75 37.29 3.22

0.860 0.67 37.29 2.93

0.883 0.75 45.02 3.26

0.976 0.69 45.02 3.03

0.70 0.300 5.2 27.74 0.0002 1.00 1.7 16.8 3.5

0.90 0.467 12.6 40.55 0.0003 1.00 2.5 16.4 10.2

1.80 0.434 10.9 53.54 0.0003 1.00 3.2 10.8 7.6

2.40 0.515 15.3 79.61 0.0004 1.00 4.8 12.0 10.5

3.00 0.578 19.3 108.41 0.0004 1.00 6.6 13.1 12.7

3.60 0.627 22.7 156.63 0.0005 1.00 9.5 15.8 13.2

3.60 0.753 32.7 156.63 0.0005 1.00 9.5 15.8 23.2

3.60 0.860 42.7 156.63 0.0006 1.00 9.5 15.8 33.2

4.20 0.883 45.0 221.04 0.0006 1.00 13.4 19.1 31.6

4.20 0.976 55.0 221.04 0.0007 1.00 13.4 19.1 41.6

0.247 0.37 6.49 2.02

0.419 0.34 9.84 1.73

0.363 0.75 22.38 3.27

0.427 0.85 29.47 3.80

0.469 0.96 37.29 4.26

0.479 1.09 46.76 4.71

0.634 0.87 46.76 4.09

0.759 0.74 46.76 3.67

0.740 0.86 58.39 4.12

0.849 0.77 58.39 3.80

0.80 0.247 3.5 30.02 0.0002 1.00 1.8 18.2 1.7

1.10 0.419 10.2 52.58 0.0003 1.00 3.2 15.9 7.0

2.10 0.363 7.6 57.03 0.0003 1.00 3.5 11.5 4.2

2.80 0.427 10.5 84.88 0.0003 1.00 5.1 12.8 5.4

3.30 0.469 12.7 71.94 0.0003 1.00 4.4 14.5 8.4

4.00 0.479 13.2 121.48 0.0003 1.00 7.3 18.4 5.9

4.20 0.634 23.2 192.20 0.0005 1.00 11.6 19.4 11.6

4.20 0.759 33.2 192.20 0.0005 1.00 11.6 19.4 21.6

4.90 0.740 31.6 298.41 0.0005 1.00 18.1 25.8 13.6

4.90 0.849 41.6 298.41 0.0006 1.00 18.1 25.8 23.6

0.171 0.48 8.30 2.49

0.348 0.52 13.02 2.25

0.269 0.95 25.83 4.22

0.306 1.09 34.61 4.89

0.380 0.71 41.64 4.96

0.319 1.00 54.11 5.71

0.448 1.11 58.39 5.19

0.612 0.88 58.39 4.55

0.484 1.14 76.45 5.27

0.639 0.94 76.45 4.74

0.85 0.171 1.7 12.66 0.0001 1.00 0.8 15.3 0.9

1.30 0.348 7.0 54.61 0.0003 1.00 3.3 16.5 3.7

2.25 0.269 4.2 29.73 0.0002 1.00 1.8 12.0 2.4

3.00 0.306 5.4 44.27 0.0002 1.00 2.7 13.4 2.7

3.50 0.380 8.4 55.85 0.0003 1.00 3.4 16.9 5.0

4.15 0.319 5.9 52.18 0.0002 1.00 3.2 21.0 2.7

4.45 0.448 11.6 98.91 0.0003 1.00 6.0 23.9 5.6

4.75 0.612 21.6 229.12 0.0004 1.00 13.9 25.2 7.7

5.20 0.484 13.6 150.80 0.0003 1.00 9.1 30.4 4.4

5.45 0.639 23.6 316.02 0.0005 1.00 19.1 34.8 4.4

0.127 0.33 9.07 2.83

0.252 0.67 16.33 2.91

0.203 0.64 27.63 4.86

0.217 0.77 37.29 5.65

0.294 0.59 45.02 5.56

0.217 0.56 57.27 6.27

0.312 0.66 64.38 5.85

0.366 1.12 72.25 5.67

0.277 0.79 85.57 6.05

0.277 1.20 95.56 5.94

0.90 0.127 0.9 12.80 0.0001 1.00 0.8 15.5 0.2

1.55 0.252 3.7 55.18 0.0002 1.00 3.3 13.4 0.3

2.40 0.203 2.4 30.50 0.0001 1.00 1.8 12.3 0.5

3.15 0.217 2.7 33.96 0.0002 1.00 2.1 13.7 0.7

3.75 0.294 5.0 73.24 0.0002 1.00 4.4 17.7 0.5

4.25 0.217 2.7 37.53 0.0002 1.00 2.3 22.7 0.5

4.65 0.312 5.6 85.53 0.0002 1.00 5.2 25.9 0.4

5.00 0.366 7.7 118.04 0.0003 1.00 7.1 28.6 0.6

5.30 0.277 4.4 71.67 0.0002 1.00 4.3 43.4 0.1

5.55 0.277 4.4 65.54 0.0002 1.00 4.0 39.7 0.5

0.052 0.56 9.84

0.076 1.52 19.67

0.095 1.01 29.47

0.107 0.93 39.34

0.098 1.28 49.45

0.089 0.65 59.54

0.088 1.00 69.55

0.102 1.07 79.40

0.041 0.63 89.91

0.090 0.54 99.53

t=

v= E0 = Vs = e= f= W= F= E1 = v1 = t=

v= E0 = Vs = e= f= W= F= E1 = v1 = t=

v= E0 = Vs = e= f= W= F= E1 = v1 = t=

v= E0 = Vs = e= f= W= F= E1 = v1 = t=

B3

Printed and published by the Health and Safety Executive C30 1/98

Printed and published by the Health and Safety Executive C30 1/98

Printed and published by the Health and Safety Executive C0.50 4/02

ISBN 0-7176-2331-9

OTO 2002/006

£10.00

9 780717 623310