integrated installation for offshore wind turbines …
TRANSCRIPT
INTEGRATED INSTALLATION FOR OFFSHORE WIND TURBINES
ETSU W/61/00617/00/REP
URN 03/1649
ContractorCorns UK Ltd
Prepared byJ Way, H Bowerman
The work described in this report was carried out under contract as part of the DTI New and Renewable Energy Programme, which is managed by Future Energy Solutions. The views and judgements expressed in this report are those of the contractor and do not necessarily reflect those of the DTI or Future Energy Solutions.______________________
First Published 2003 © Crown Copyright 2003
dti
EXECUTIVE SUMMARY
ObjectivesThe overall aim of the project was to demonstrate the viability of integrating the offshore installation of foundation, turbine and tower into one operation. This has been achieved through completion of the following objectives:
• Telescopic tower study. Reversible process incorporating lift and lock-off mechanisms.
• Transportation study. Technical and economic feasibility of transporting and installing a wind turbine unit via a standard barge with minimal conversion
• Self burial system study. To demonstrate the feasibility of self burial of a large slab foundation via controlled jetting beneath the slab.
BackgroundWith current technology, there is a need for a major marine spread to install the foundation, tower and turbine. There are clearly benefits if offshore installation can be integrated into one operation.
The proposed concept was as follows:
• A Bi-Steel gravity foundation (iron ore filled insitu to provide on-bottom stability).
• Tower and turbine are attached onto the foundation at shore. The tower is telescopic in order to increase stability by halving the height of the tower and turbine during transportation, thus lowering the centre of gravity of the nacelle.
• The whole foundation/tower/turbine is loaded out from the quayside, transported to site and set down on the seabed by a standard barge with suitable conversion.
• The foundation unit will bury itself in a controlled manner on sand sites.
The proposed installation technique will minimise offshore spread and maximise the proportion of work carried out onshore with consequent benefits in terms of cost, quality and safety.
Provisional cost studies suggested that the scheme was highly competitive with currently proposed offshore wind farm solutions, with potential savings in excess of £100k for each foundation unit installed. Some of the reasons for this saving are follows:
Barge modifications are anticipated to be straightforward and readily achievable within the build period of the foundation units themselves. In principle a barge could be secured on the spot market, thus minimizing costs.
• Duration of equipment hire is low.• Most assembly work is carried out adjacent to a quayside. Very little work is carried out
offshore. This is a major safety and quality gain as well as a cost benefit.
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• Because so little work is carried out offshore, the risk due to weather is considerablyreduced. In summer conditions it should be possible to install at a rate of one per day using a single barge (subject to distance from the supply base).
Work Carried OutThe project was split into three separate studies - telescopic tower, transportation and self burial.
The telescopic tower study investigated:
1. Methods of raising and lowering the tower. Two methods were studied - two rope hoist- type systems (as proposed in the original proposal) and a hydraulic lift system.
2. Methods of locking-off the tower in the raised position. Again, two methods of locking off were studied - a shrink-fit system (as posited in the original proposal) and a mechanical clamping system. Both systems were compared with the conventional grouted method.
3. Additions and modifications to the onshore fabrication and erection processes. The additional steelwork elements of the telescopic system were evaluated and costed, and a methodology for quayside assembly was developed.
4. Modifications to the offshore fabrication processes. A methodology for the offshore processes was developed.
The transportation study investigated:
1. The suitability of 18 licensed sites around the UK for the integrated concept outlined above.
2. The basic criteria for the development of a converted barge suitable for pick-up, load-out and set-down of the integrated OWEC.
3. Five installation options in sufficient detail to select one option for detailed conceptual design and costing. These comprised 2 A-Frames, two horn and one strand-jack option.
4. The detailed conceptual design of the cantilever / strandjack option.
The self burial study investigated:
1. The basic design parameters for underwater jetting to facilitate the design of the large scale test. This element included small scale testing due to the lack of published information on this subject.
2. A series of medium scale tests (4.8m diameter foundation in a 8.45m diameter sand bed) in three phases between October 2002 and April 2003. The jetting system for first phase comprised an array of jets distributed around the underside of the tank. The second phase introduced eductors and a radial jetting system. The final phase combined the radial and agitating systems.
Main ResultsThe telescopic tower study indicated that:
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• The pulley / hoist mechanism was more expensive than had originally been envisaged. The expected cost reductions for standardisation appear not to be achievable in practice.
• An alternative lifting mechanism, based on hydraulics, was shown to be a more economic alternative.
• The heat shrink lock off mechanism was shown to be technically feasible although expensive both in terms of weight of steel and energy input required.
• Indicative costs for the overall system were approximately £145k more than conventional tower fabrication/erection processes.
The transportation study indicated that:
• Two UK sites are ideally suited to the integrated system, a further six sites are well suited.• The transportation system must be dual-lift, ie combine a structural support and an
independent means of raising and lowering the turbine, for example winch/strand jacks.• Whilst in the interest of predictability it is desirable to lock the turbine and barge together,
very large forces are induced in rigidly fixed cantilevers. Due to the magnitude of these forces the proposed barge modifications were less economic than had been hoped.
The self burial study indicated that:
• Self burial of a loaded flat slab using water jets is technically feasible.• The optimum system combines radial jets and agitating jets (which transport the slurry to
the perimeter of the foundation) with eductors distributed around the perimeter (to remove the slurry from the immediate vicinity of the foundation).
Recommendations• Further development of the fully integrated wind turbine is not recommended.• Systems which build on aspects of the technical development may warrant further study, eg
barge transportation of foundation only, combined with the self burial system.• Although the system developed in this study utilised perimeter eductors, replacing these
with fewer, larger eductors in the centre of the foundation may be more economic at full scale. The feasibility of this alternative arrangement should be investigated at the medium scale.
• Although the proposed system of edge eductors is scalable it requires demonstration at full scale, for example a gravity foundation base slab for a met mast.
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CONTENTSExecutive Summary............................................................................................................................. iContents............................................................................................................................................ iv1 Introduction.................................................................................................................................. 1
1.1 Aim and Objectives.............................................................................................................. 11.2 The Concept.........................................................................................................................21.3 Base Case Design.................................................................................................................4
2 Telescopic Wind Towers .............................................................................................................52.1 Introduction..........................................................................................................................5
2.1.1 Objective of this Study.................................................................................................52.1.2 Study Work Scope .......................................................................................................5
2.2 Design Philosophy...............................................................................................................62.3 Design of Rope Hoist Lift / Lowering Device.....................................................................7
2.3.1 Introduction..................................................................................................................72.3.2 Concept 1 - Multi-Pulley Hoist...................................................................................72.3.3 Concept 2 - 4 Pulley Hoist..........................................................................................82.3.4 Concept 3 - Multi-Pulley Reduced F.o.S...................................................................102.3.5 Optimised Concept..................................................................................................... 102.3.6 Discussion of Rope Hoisting Systems....................................................................... 11
2.4 Hydraulic Lift System........................................................................................................ 132.4.1 Estimated Pressure and Flow Requirements.............................................................. 132.4.2 Cylinder Sealing Requirements................................................................................. 152.4.3 Cylinder Guidance..................................................................................................... 162.4.4 Summary of Hydraulic Lift System........................................................................... 18
2.5 Design of Lock Off System............................................................................................... 192.5.1 Shrink Fit Lock-Off System....................................................................................... 192.5.2 Heating Method and Budget Prices...........................................................................202.5.3 Alternative Clamping Methods..................................................................................212.5.4 Discussion of Lock-Off Systems...............................................................................22
2.6 Access and Services...........................................................................................................242.7 Tower / Turbine Installation Methodology........................................................................25
2.7.1 Conventional Installation Methodology..................................................................... 252.7.2 Telescopic Tower Installation....................................................................................25
2.8 Economic Analysis............................................................................................................282.8.1 Estimated Supply / Installation Cost for Conventional OWEC.................................282.8.2 Additional Structural Costs for Telescopic Tower....................................................292.8.3 Raising Mechanism Costs..........................................................................................292.8.4 Heat Shrink Lock-Off................................................................................................302.8.5 Summary of Cost Options..........................................................................................30
2.9 Scale Model Test Rig.........................................................................................................322.9.1 Design........................................................................................................................322.9.2 Test Procedure............................................................................................................322.9.3 Cancellation of Testing Programme ..........................................................................36
3 Transportation............................................................................................................................373.1 Introduction........................................................................................................................373.2 Site Review........................................................................................................................38
3.2.1 Locations....................................................................................................................383.2.2 Seabed Conditions .....................................................................................................393.2.3 Water Depths..............................................................................................................39
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3.2.4 Access........................................................................................................................403.2.5 Loadout Ports.............................................................................................................413.2.6 Dry dock Availability..................................................................................................433.2.7 Overall Ranking.........................................................................................................44
3.3 Installation Criteria Study..................................................................................................453.3.1 General.......................................................................................................................453.3.2 Base (Cantilever) Condition.......................................................................................453.3.3 Horns..........................................................................................................................483.3.4 A Frame......................................................................................................................493.3.5 Conclusion of Preliminary Study...............................................................................50
3.4 Development of Installation Options.................................................................................513.4.1 General.......................................................................................................................513.4.2 Technical Assessment - Horn Option........................................................................553.4.3 Technical Assessment - A Frame (Unit Hung-Off)..................................................563.4.4 Technical Assessment - A Frame (Unit on Deck).....................................................573.4.5 Environmental Limitations of Dual Lift Systems......................................................583.4.6 Single Point Suspension Option.................................................................................603.4.7 Cantilever / Strand Jack Concept...............................................................................623.4.8 Costing.......................................................................................................................633.4.9 Scoring.......................................................................................................................653.4.10 Conclusions of Option Development.........................................................................65
3.5 Conceptual Design of Preferred Solution..........................................................................673.5.1 Basic Concept............................................................................................................673.5.2 Design Parameters......................................................................................................673.5.3 Hydrodynamic Analysis.............................................................................................673.5.4 Structural Design........................................................................................................723.5.5 Marine Operations......................................................................................................793.5.6 Costings......................................................................................................................82
3.6 Retrospective Overview.....................................................................................................854 Self Burial System ..................................................................................................................... 87
4.1 Introduction........................................................................................................................874.2 Phase 1 - Definition of Test Parameters............................................................................88
4.2.1 Basis Soil Condition ..................................................................................................884.2.2 Detection of Surface and Sub-Surface Boulders by Geophysical Methods...............884.2.3 Jetting System Concept..............................................................................................894.2.4 Small Scale Test 1......................................................................................................934.2.5 Small Scale Test 2......................................................................................................97
4.3 Phase 2: Agitating Jet Tests............................................................................................. 1004.3.1 Introduction.............................................................................................................. 1004.3.2 Jetting Design........................................................................................................... 1004.3.3 Test Piece Design..................................................................................................... 1014.3.4 Test Series 1............................................................................................................. 1054.3.5 Test 2a...................................................................................................................... 1074.3.6 Test 2b...................................................................................................................... 1074.3.7 Conclusions from Phase 2........................................................................................ 109
4.4 Phase 3 - Eductor and Radial Jet Tests............................................................................ 1104.4.1 Introducti on.............................................................................................................. 1104.4.2 Eductor Pumps......................................................................................................... 1104.4.3 Submersible Pumps.................................................................................................1124.4.4 Test 3a...................................................................................................................... 113
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4.4.5 Test 3b...................................................................................................................... 1154.4.6 Conclusions from Phase 3........................................................................................ 119
4.5 Phase 4: Combined Jetting System..................................................................................... 1214.5.1 Introducti on.............................................................................................................. 1214.5.2 Plumbing Design...................................................................................................... 1214.5.3 Pumps....................................................................................................................... 1234.5.4 Manifold................................................................................................................... 1234.5.5 Test 4........................................................................................................................ 1234.5.6 Test 5........................................................................................................................ 1244.5.7 Test 6........................................................................................................................ 1264.5.8 Test 7........................................................................................................................ 1274.5.9 Test 8........................................................................................................................ 1284.5.10 Conclusions from Phase 4........................................................................................ 130
5 Ecomonic Assessment of Integrated Installation.................................................................... 1316 Discussion and Recommendations........................................................................................... 132References........................................................................................................................................ 133Acknowledgements.......................................................................................................................... 134Appendix A: Integrated Installation Concept Drawing................................................................... 135Appendix B: Telescopic Tower Test Rig Drawings....................................................................... 137Appendix C: Horn / Strandjack General Arrangement................................................................... 141
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1 INTRODUCTION
1.1 Aim and ObjectivesThe overall aim of the project was to demonstrate the viability of integrating the offshore installation of foundation, turbine and tower into one operation. This has been achieved through completion of the following objectives:
• Telescopic tower study. To confirm that the winch system and lock-off mechanism work when subjected to an eccentric load as would be applied by a real turbine.
• Transportation study. To investigate the feasibility of transporting and installing a wind turbine unit via forks mounted over the end of a submersible barge and to determine the extent (if any) of barge strengthening that may be required such that costs may be determined. Alternative methods will also be investigated.
• Self burial system study. To demonstrate that the basic concept of controlled jetting under a large slab is feasible and that adequate control of level can be achieved.
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1.2 The ConceptCorus has been developing offshore wind turbine foundation concepts that reduce costs by speeding up installation and reducing the marine spread required. With current technology, regardless of how efficient foundation installation becomes, there is still a need for a major marine spread to install the tower and turbine. There are clearly benefits if offshore installation of the foundation, tower and turbine can be integrated into one operation. The proposed concept met this objective, however some elements needed further research before commercial viability could be established.
The foundation is a gravity structure comprising of a circular cone/pile extending out of the centre of a large base slab. The base slab is made from dense concrete and the cone is filled with a dense blend of iron ore in order to give the required mass for on-bottom stability.
The tower and turbine are attached onto the foundation at shore. A telescopic tower has been proposed in order to increase stability by halving the height of the tower and turbine during transportation. The lower half of the tower would be parallel sided to allow the upper half of the tower to run within it. Methods of raising and lowering the upper section would be developed, for example using a simple system of pulleys and cables, together with a number of alternative methods of providing the joint between the upper and lower halves.
The whole foundation/tower/turbine must be transported upright. It is proposed to pick the unit up about 10m above the base of the foundation using trunnions attached to the foundation unit. These would locate in lifting forks extending over the end of a submersible barge. A detailed study would determine the motions of the various elements to confirm the feasibility of such a method of transportation.
The foundation unit is designed to sit on top of rock and hard soil sites, but to bury itself in a controlled manner on sand sites. Proving the method of burial would be the final area of investigation. Preliminary work suggested this is possible using a system of low cost water jets coupled to pumps. The number and direction of the jets would be selected to ensure that jetting is controllable. By dividing the base area into zones and jetting at different rates (as dictated by level), it should be possible to control the penetration of the base. The pumps to control this operation would be containerised and located on the installation barge. While burial is being undertaken the tower can be winched up and secured in position. The process of lowering and jetting was expected to take less than 12 hours.
The proposed installation technique would minimise offshore spread and maximise the proportion of work carried out onshore with consequent benefits in terms of cost, quality and safety.
Provisional cost studies suggested that the scheme is highly competitive with currently proposed offshore wind farm solutions, with potential savings in excess of £100k for each foundation unit installed. Some of the anticipated reasons for this saving were follows:
Barge modifications were anticipated to be straightforward and readily achievable within the build period of the foundation units themselves. In principle a barge could be secured on the spot market, thus minimizing costs.
• Duration of equipment hire is low.• Most assembly work is carried out adjacent to a quayside. Very little work is carried out
offshore. This is a major safety and quality gain as well as a cost benefit.
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• Because so little work is carried out offshore, the risk due to weather is considerablyreduced. In summer conditions it should be possible to install at a rate of one per day using a single barge.
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1.3 Base Case DesignIn order to ensure consistency the three elements of the project, telescopic tower, transportation and self burial were developed from a common base-case design scenario.
The OWEC Integrated Installation Concept Drawing, BIS152/DRG/90000/A is given in Appendix A.
Concept Specification was as follows:
Nacelle Power Rating Hub Height Water Depth
Wave Height Wave Period Ice loading Boat Impact
2.75MW 75m above MSL Min 15m Max 25m12m (Extreme 50yr wave)11sNone150te displacement vessel at velocity of 2m/s
Table 1.1: Aeordynamic LoadingCondition Moment (MNm) Shear (kN)Non-operating (50 yr return) 48.4 925Operating (1 yr return) 45.4 846Fatigue (5x106 cycles) 12.9 294
The foundation design was based on the Bi-Steel Gravity Base concept comprising a steel plate cone on a 24m wide octagonal Bi-Steel base slab. The cone (plus tower and nacelle) is transported to site ballasted either partially or completely with water, where it is then fully ballasted to lower it to the seabed.
Once on the seabed the self burial system will bury the unit 2.5m into the seabed to reduce or eliminate the requirement for additional scour protection. Once buried the unit is filled with iron ore via ports in the side of the cone.
The weight of the unit during transportation was 1650 tonnes (cone fully submerged and water filled), with the centre of gravity 10.3m above the underside of the base slab.
The insitu nett on-bottom weight of the unit was 4500 tonnes, with the centre of gravity 9m above the underside of the base slab.
Soil conditions were assumed to be medium to dense sand with scattered cobbles at the seabed.
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2 TELESCOPIC WIND TOWERS
2.1 Introduction
2.1.1 Objective of this Study
The objectives of this activity were as follows:
• To confirm that the winch system and lock-off mechanism work when subjected to an eccentric load as would be applied by a real turbine.
2.1.2 Study Work ScopeIn order to meet the objectives set out above, a detailed assessment of the requirements was undertaken covering:
• A mechanical winching system for tower raising.• Alternative methods of tower raising.• A "shrink-fit" type system for tower lock off.• Alternative proposals for tower lock off.• Methods of erection covering both onshore and offshore processes.• Indicative costs.• Design of a scale telescopic tower test rig to demonstrate the raising and lock off concepts.
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2.2 Design PhilosophyThe operating parameters for the tower were as follows:
• The lifting device for the tower would be a mechanical winching device.• A lock off system would be required using a heat shrink method.• Weights of the main components as given in BIS 152ZDRG/90000/A . A nominal 10m
overlap between the upper and lower tower has been assumed in order to provide stability at the end of the tower raising phase.
• Assumed Upper tower weight is 94 Te.• Assumed Nacelle and blades weight is 125 Te.• Wave loading was neglected as it was deemed not applicable as the joint would be above
wave crest.• Aerodynamic loading as per Table 1.1.• The Rotor was assumed to be 5000mm from tower centreline.• The C of G of nacelle was assumed to be 160mm from the tower centreline.• Tower bottom section would be approx. 45000mm long with the capability for the inner
tower to lift a further 35000 m.
A typical arrangement of the tower concept is shown below:
80000
80000
102540
45000
Figure 2.1: Arrangement of Telescopic Tower Unit
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2.3 Design of Rope Hoist Lift / Lowering Device
2.3.1 Introduction
The design criteria for the winching system was assessed using the requirements of BS 2573 [1].
Three concepts were considered:
• Concept 1 consisted of a series of small pulleys located inside the lower tower section lifted by a single rope. In order to minimise the rope force required a multiple pulley arrangement to give large mechanical advantages was considered requiring a single hoisting drive unit.
• Concept 2 consisted of a simplified lifting system using 4 hoisting units arranged circumferentially around the tower.
• Concept 3 was as Concept 1 but reducing the rope safety factor from the British Standard recommendations of 4.5 to 2.5 on the basis that the tower lifting and lowering operation would be very infrequent and that prior to any lifting operation, due consideration would be given to checking/changing the rope.
Following the initial theoretical analysis of the above concepts, the study was expanded to show theeffect of varying the pulley diameter and number of pulleys to find an optimum cost solution. Thiswas carried out in conjunction with a hoisting specialist company, Jeamar Hoists of America.
2.3.2 Concept 1 - Multi-Pulley HoistDetailed calculations have been prepared for the sizing of a multiple pulley hoist arrangement.Using 12 off lifting pulleys to give 24 rope rises, the results are summarised as follows:
• Rope load required = 90kN.• Minimum braking load from BS 2573 = 442kN.• Rope diameter required = 28mm.• Pulley diameter = 560mm.• Hoist drum diameter = 500mm.• Based on a telescopic height of 35m, rope length = 840m to be hoisted.• Based on typical rope drum length of 1750mm, 7 rope layers would be required to coil up
840m of rope.• Estimated hoisting power to complete in 20 minutes = 86kW.
The basic layout of the pulley arrangement is shown in Figure 2.2.
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Top Row of Pulleys fixed radially to the inside of the outer tower
35 M Telescopic Height
Bottom Row of Pulleys fixed radially to the outside of the inner tower
Figure 2.2: Arrangement of Multi-Pulley Reeving System (Concept 1)
Advantages:
• Proven technology.• Simple concept.• Can raise and lower tower.• All equipment can be mounted within the tower giving good aesthetics.• Effects of corrosion should be minimised because of above.
Disadvantages:
• High installed power capacity required to operate, can be overcome by portable generation, diesel driven hoisting units.
• High capital outlay.• Maintenance access to hoist units could be tight.• 840m of rope to be reeved.• Many moving parts.• High manufacturing accuracy and therefore cost required for pulley arrangement.• Due to component size, limitation on manufacturing capabilities.• Complex installation procedure.
2.3.3 Concept 2 - 4 Pulley HoistCalculation have derived the sizing of a pulley arrangement using 4 off lifting pulleys to give 4 rope rises, the results are summarised as follows:
• Rope load required = 534kN.• Minimum braking load from BS 2573 = 2650kN.• Rope diameter required = 64mm (21/2 inch).
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• Pulley diameter = 1280mm.• Hoist drum diameter = 1150mm.• Based on a telescopic height of 35m, rope length = 35m to be hoisted.• Typical rope drum length of 700 mm, 1 rope layers would be required to coil up 35m of
rope.• Estimated hoisting power to complete in 20 minutes = 21kW/drive 4 off required, =84kW.
The basic layout of the pulley arrangement is shown in Figure 2.3.
3200 DIA
Pulley and winch system by 4 off large dia ropes.. 65000
5000 DIA REF
45000
Figure 2.3: Arrangement of Four Winch System (Concept 2)
Advantages:
• Proven technology.• Simple concept.• Easy to maintain.• Could raise and lower tower.• Easier to manufacture than Concept 1.• Easier to install than Concept 1.
Disadvantages:
• High installed power capacity required to operate, could be overcome by portable generation, diesel driven hoisting units.
• Very high capital outlay.• Suitable corrosion protection of the ropes and pulley components would be required.• The aesthetics of this system would be poor as hoist units are so large they would need
mounting outside the tower, most likely in an enclosure on the side of the tower.
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2.3.4 Concept 3 - Multi-Pulley Reduced F.o.S.Calculations have evaluated the sizing of a multiple pulley hoist arrangement using a reduced factor of safety for the rope based on an infrequent operation.
Using 12 off lifting pulleys to give 24 rope rises, the results can be summarised as follows:
• Rope load required = 90kN.• Minimum braking load based on 2.5 safety factor = 246kN.• Rope diameter required = 21mm.• Pulley diameter = 420mm.• Hoist drum diameter = 380mm.• Based on a telescopic height of 35m, rope length = 840m to be hoisted.• Based on typical rope drum length of 1750mm, 7 rope layers would be required to coil up
840m of rope.• Estimated hoisting power to complete in 20 minutes = 86kW.
The basic layout of the pulley arrangement is shown for Concept 1.
2.3.5 Optimised Concept
Following the initial assessment of hoisting requirements it was decided to develop Concept 3 further as this was deemed the most cost effective option for hoist lifting. Contact was made with Jeamar Corporation of America who manufacture both standard and purpose built hoisting equipment.
It was felt that by using their vertical pulley units which comprise an integral pulley and shaft assembly complete with bearings mounted into a housing an optimum solution could be provided.
In order to determine the optimum number of pulleys for the lifting arrangement, a simple spreadsheet was produced to determine the optimum pulley number, size and cost.
The results are tabulated in Table 2.1, based on a nominal tower diameter of 5000 mm, lifting a load of 236 Te through a height of 35 M.
The analysis indicated that a minimum pulley diameter of 305mm (pulley typeHB23000) was required to give the required load carrying capacity. This arrangement involved placing 30 pulleys on each tower section giving a mechanical advantage of 60/1. An efficiency ratio was calculated using manufacturers recommendations based on the number of pulleys and a single efficiency of 98/99 % per pulley to give the rope pull required from the hoist unit.
An assessment was made to estimate the total rope length to be recoiled using each pulley arrangement. This ranges from 700m of 32mm diameter rope to 2100m of 19mm diameter rope of rope dependant on the pulley arrangement.
The optimum pulley arrangement in terms of cost has been calculated. It was seen that the cost did not vary significantly as it involved the cost of either many small pulleys or fewer larger pulleys. However, the least cost method, in terms of pulleys, would involve using 10 off 598mm diameter pulleys around both the inner and outer tower shells which would give a mechanical advantage of 20/1, this gave a purchased pulley cost of $90k (£60k). It may be possible to reduce this cost by
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circa 20% with bulk purchase and negotiation although this would likely be offset by the mounting costs to the tower structure. The cost of the hoisting unit would also need to be added to the pulley cost to gain a total appreciation of the cost of this option.
Table 2.1: Summary of Pulley AnalysisPulley Reference HB23000 HB31000 HB41000 HB52000 HB68000Diameter (mm) 305 357 406 547 598Rope Diameter (mm) 19 22 25 29 32Block width (mm) 508 590 660 844 922Max No. of pulleys ? 30.92 26.62 23.8 18.61 17.04No. of pulleys 30 26 23 18 17Pulley Capacity (kN) 102 138 182 231 302Lift Capacity (kN) 3,060 3,588 4,186 4,158 5,134No. Needed 28.96 21.4 16.23 12.79 9.78Actual No. 30 22 17 13 10Line parts 60 44 34 26 20Efficiency Ratio 46 35 27 21 17Rope Pull Required(1 Hoist) (kN) 51 68 87 111 142Rope Length to lift (m) 2,100 1,540 1,190 910 700Pulley Cost ($) 1,755 2,517 3,345 3,756 4,534O/A Pulley cost/tower ($) 105,300 110,748 113,730 97,656 90,680
Hoist Drum RequiredMinimum Diameter (mm) 342 396 450 522 576Layer 1 rope length (m) 1.07 1.24 1.41 1.64 1.81Layer 2 rope length (m) 1.19 1.38 1.57 1.82 2.01Layer 3 rope length (m) 1.31 1.52 1.73 2 2.21Layer 4 rope length (m) 1.43 1.66 1.88 2.19 2.41Total length per row (m) 5.01 5.81 6.6 7.65 8.44No of rows required 418.83 265.26 180.38 118.91 82.89Drum Width required (m) 7.96 5.84 4.51 3.45 2.65
Alternative Diameter (mm) 1,000 1,000 1,000 1,000 1,000Layer 1 rope length (m) 3.14 3.14 3.14 3.14 3.14Layer 2 rope length (m) 3.26 3.28 3.3 3.32 3.34Layer 3 rope length (m) 3.38 3.42 3.46 3.51 3.54Layer 4 rope length (m) 3.5 3.56 3.61 3.69 3.74Total length per row, (m) 13.28 13.4 13.51 13.66 13.77No of rows required 158.1 114.96 88.09 66.62 50.83Drum Width required, (m) 3 2.53 2.2 1.93 1.63
Minimum Hoist Torque (kNm) 25.63 34.16 43.38 55.61 70.9Maximum Hoist Torque (kNm) 28.56 38.67 49.88 65.28 84.51
2.3.6 Discussion of Rope Hoisting Systems
In the course of carrying out the above assessment, discussions were held with Jeamar in the USA and Clark-Chapman in the UK regarding the cost and practicality of using hoisting units for the telescopic erection of wind turbine towers.
In terms of Concept 2, the four winch system, Clark-Chapman have given a verbal estimate of £120k per hoist unit, which allowing for mounting structure and installation would require an investment of approximately £500k per wind turbine. This option has not been considered any further because of this high capital outlay.
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Utilising the findings of Concepts 1 and 3, as well as the discussions with Jeamar, an estimate of a multi pulley hoist arrangement has been developed. Jeamar would recommend the use of HB31000 pulleys working in conjunction with a 60 KW hoisting unit. The cost of pulleys and hoisting unit would be £100k excluding installation, electric's or support structure. Whilst not the cheapest option considered in terms of pulley cost, Jeamar considered this the optimum in terms of an overall reliable system, taking into account number of pulleys, together with the size of hoisting unit required.
Following this analysis, it is apparent that the use of a rope hoisted system, whilst simple in concept using tried and tested technology, comes at a significant cost in terms of the mechanical equipment and arrangements to be made to power this facility for a system that will seldom be used.
As such it was considered appropriate to consider alternative proposals.
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2.4 Hydraulic Lift SystemIn reviewing possible alternative lift/lower mechanisms, it became apparent that the telescopic tower could be considered as a simple vertically mounted hydraulic cylinder. The telescopic tower consisted of a lower tower structure which would represent the cylinder body, the upper tower section would form the piston and rod.
As the tower would be in an offshore environment, there was a ready hydraulic medium available to raise the tower - namely sea water.
The tower structure basically comprised a cylinder of some 4.5m internal diameter. Other key dimensions are shown in Figure 2.4:
Upper Tower section is sealed andacts as a pistonLower tower tube is capped at base to act as
a hydraulic cylinder when filled with sea water
80000
80000
102540
T" 7700010482
Sea Level
4500027500
Figure 2.4: Arrangement of Hydraulic Telescopic Tower Unit
In order for the hydraulic lift method to work the system would have to incorporate the following features:
• Pumping system to allow water under pressure to the underside of the upper tower section.• Sealing system to contain the pressurised water within the lower tower, this sealing system
would have to work dynamically as the tower was raised.• A guidance system would have to be provided to allow the upper and lower towers to
remain reasonably concentric during lifting. This guidance system would need to be able to withstand any out of balance forces from the nacelle unit coupled with any wind moments generated during the erection process.
2.4.1 Estimated Pressure and Flow RequirementsBased on this hydraulic cylinder concept, the pressure required to lift the nacelle and upper tower structure was estimated.
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Weight of Nacelle and upper tower Wn ;=2150103 ■ N
Diameter of tube D :=4500mm
Height of lift of upper tube base above sea level Hl := 7500mm
Density of water p ;= 1010—m3
Accel due to gravity g = 9.807—s2
Hydraulic pressure required including 20% for losses
p := 1.2- Wn -t-Hl ■ p■ gi
n.D2\ 4 1
p = 2.481atm
From the above calculation, it was shown that a pressure of some 2.5 bar would be sufficient to raise the telescopic structure.
Substantial plates would be required to be fitted to the base of the upper and lower towers to effectively seal them from water ingress. Preliminary calculations indicate that plates of 80mm thickness would be required to resist the water pressure which would add approximately 12 Te to each tower section. Alternatively a stiffened steel assembly or a steel/concrete composite could be used.
To raise the structure in approximately 20 minutes a flow rate of some 5000 GPM would be required. This coupled with the pressure gave us an estimate of the pumping requirements. For the purpose of this exercise it was assumed that a ship mounted diesel driven pump would be required which would erect each tower in turn on the offshore wind farm. It was felt that this was the minimum cost option for a wind farm and also gave the least environmental and most aesthetic impact for the tower structure.
The self contained diesel driven pump set would comprise the following :-
• Pump.• Diesel engine.• Coupling.• Batteries and charging system.• Control Panel.• Diesel storage tank.• Skid base to contain all items.
The best pump for the application would be a heavy duty end suction slurry pump. Although more expensive and not as efficient as a horizontal centrifugal type it would be best suited to sea water and possible entrainment of sand and other debris.
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As the operating pressure at the structure was required to be approximately 2.5 bar the pump would need to generate approximately 3.5 bar to allow for the suction head and for line losses etc. in the temporary pipework connections.
A volume requirement of 5000 GPM is approaching the maximum allowable limits for sensible handling of temporary pipework / hoses. The pump suction pipework would be 300mm nb and the delivery pipework 250mm nb. Clearly smaller pumps and pipework could be used if the raising time was increased.
For this volume flow rate the diesel engine would be approx. 250 kW with a fuel consumption of approx. 60 litre/hr. Therefore, to cover operation for 8 hours, 480 litres of fuel would be required, which would be easily contained by a tank one metre x one metre x one metre.
The overall cost of this unit would be approx. £60k. The overall dimensions of the pump and skid structure would be approximately 1.5 metres wide by 4 metres length.
This unit would be used to raise all towers within the wind farm and also for subsequent maintenance involving tower raising/lowering. As such based on a wind farm of 30 towers the unit cost is low at £2k per tower.
2.4.2 Cylinder Sealing RequirementsThe main problem to overcome with regard to sealing was the manufacturing tolerances of the inner and outer tubes. Using conventional fabrication techniques, the tolerance on diameter for a 4200mm diameter tube of 15mm wall thickness was +/- 14 mm.
To provide sealing solutions for the upper tower structure during the raising operation by hydraulic pressure, contact was made with specialist seal suppliers Walker Seals, Presray and Oil States, budget proposals were received from Walker Seals and Oil States.
Oil States have much experience in the sealing of offshore structures and have recommended the use of a modified inflatable grout packer which would be inflated by water pressure instead of a grout fill. The inflatable seal would be backed up using a flat rubber back up seal. Oil States consider that this option could work as its capacity was substantially in excess of the requirement but this would need to be balanced against the friction produced by the packer applying pressure to the cylinder walls.
The main advantage of using an inflatable seal was perceived to be that this facility was very capable of accommodating the tolerance on diameter. A secondary advantage was that if two seals were used a distance apart, they could form the guidance mechanism for the tower during raising. This function would work by the application of sufficient inflation pressure to overcome the reaction forces at the inflatable seals generated by the out of balance erection moment due to wind loading and the offset of the nacelle centre of gravity from the tower centerline.
A typical inflatable packer is shown in Figure 2.5 below.
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Iirl.r, hum (iriii i1. Flanker wttn Lmuk irp i nt se,i
Figure 2.5: Oil States Inflatable Packer
Budget costs for an inflatable water seal were obtained from Oil states at £6500 per tower, although no detailed proposals have been forwarded.
Walker Seals on the other hand suggested the use of a "U" ring seal which would be energised on its underside by the sea water under pressure to form the seal, this would be manufactured using their standard extrusion processes. Walkers have provided detailed sizing and prices for the scale test rig but not the full size unit. Like the inflatable seal, this unit could accommodate the manufacturing tolerances of the tower construction, however Walkers also advocated the use of an inflatable back up seal which would consist of a horizontal "U" Ring seal, located between support plates. Bonded to the outer face of this seal would be a wear strip to minimise the friction on the seal during raising operations. This arrangement is shown in Figure 2.6.
Wear Strip Material TEA
Nitrile Rubber "U" Ring Material PB70
Outer Shell showing tolerance on Diameter at +/-14 mmFigure 2.6: Seal and Inflatable Support Arrangement
Nitrile Rubber Horizontal "U" Ring Material: PB 70
Inflation pressure point
Inner Shell
2.4.3 Cylinder Guidance
As with the problems of sealing, any guidance solution would have to overcome the manufacturing tolerances of the inner and outer tubes.
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The design of any guidance system would therefore have to cope with this variation in diameter. As stated earlier, the guidance system would need to be able to withstand any out of balance forces from the nacelle unit coupled with any wind moments generated during the erection process. An assessment was made of the anticipated loads on the tower during the erection process.
Assessment of Erection Loads on the structure
Nacelle and blade mass Mn := 125000kg
Upper tower mass Mm =94000kg
Nacelle CofG from tower Ln := 160mm
Average tower diameter Dt :=3.5- m
Exposed tower height Ht :=80 - m
Lever arm from tower top to joint La := 57.5- m
Estimate of Wind loads using CP3-1972
Basic Wind Speed (NE England) v :=48-—
sTopography factor (offshore) S1 :=1
Size Factor (Class C open country 80m high)
S2:=1.11
Statistical factor for construction S^=0.77
Design Wind speed Vm=v-SLS2-S3 Vs = 41.02°ms
Dynamic Pressure ^=0.613f^i^j-Vs2m3
q = 1.032103 <*-N- m2
Force on the tower Ft :=0.6 -q- Dt- Ht Ft = 1.733-105 N
Moment reaction at joint FtMj :=—- La Mj = 4.98310° »N- m
If We limit wind speed during erection to 20 M/s (45 mph)
Let Vs:=20™
q:=0.613|M -Vs2 q = 245.2^
\m3/ m2
Ft :=0.6-q-Dt-Ht Ft = 4.119104N
Mj :=F-La
2Mj = 1.18410° »N- m
Total Erection Moment Me :=Mj + Mn-g-Ln Me = 1.38-106«N- m
Reaction forces at guidance points assuming 8m appart Rg 1=^1 Rg = 1.726105 N8- m
s
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It can be seen from the above calculation that if raising of the towers was confined to wind speeds of 20m/sec or less, the guidance system would have to withstand a maximum erection moment of 1.38MNm which would generate reactions at the guidance points assuming they were spaced at 8m separation of approximately 17Te.
It is considered that this use of an inflatable packer would provide a suitable guidance mechanism.
Walkers have developed their ideas further than Oil States and would provide two horizontal "U" ring seals together with a low friction wear strip bonded to the seal as shown in Figure 7 above. At this stage they have designed and priced the unit for the scale model but not the full size unit.
2.4.4 Summary of Hydraulic Lift System
Following the assessment of this alternative lifting method it was felt that the hydraulic lifting process offered advantages over the rope and pulley lifting method in terms of:
• Relatively simple construction.• One pumping set could service a whole wind farm with low resultant unit cost.• Mechanically simple in concept with few moving parts.• Sealing solutions proposed would also solve the guidance requirements.• Potential to be significantly cheaper than a rope hoist solution.
Hydraulic lifting costs could be summarised as follows:
• Cost of hydraulic pumping unit £2k per tower (based on spreading cost over 30 turbines), see section 2.4.1
• Cost of seal and guidance system £13k per tower (2 off inflatable seal units per tower) see section 2.4.2.
• Cost of additional plating to seal the towers estimated at 24 Te at a cost of £800 per Te. £20k
This gave an overall cost for the hydraulic system of approximately £35k
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2.5 Design of Lock Off System
2.5.1 Shrink Fit Lock-Off System
The proposed lock-off mechanism is based on the "heat shrink" concept.
During the tower erection, a clamping ring at the top of the lower tower would be heated and thus caused to thermally expand. This would provide clearance for a similar clamping ring at the bottom of the upper tower.
On completion of the raising operation, the heat source would be removed from the ring on the lower tower to allow it to contract to such an extent that it would clamp the upper section in the raised position.
The advantage of this system over other locking mechanisms is that that the process could be potentially be reversed to allow the tower to be lowered for maintenance or decommissioning, without the implications of access to bolted/welded joints etc and without the need for large jack up platforms.
2.5.1.1 Single Clamp RingInitial thoughts based on using two clamping rings nominally 8m apart (to carry the moment) were not developed further as this would have involved the machining of additional large shrink rings and associated heating requirements for assembly. The study therefore developed a mechanism for resisting the design loads using a single clamping ring.
This ring would have to resist the self weight of the nacelle unit and upper tower section and the bending moments due to the aerodynamic loading. An empirical method developed by SKF was used for various wall thickness ratios and the 50 yr. return aerodynamic loading. The results of these calculations are summarised in Table 2.2
Table 2.2: Summary of SFK Shrink Fit CalculationsItem Geometry 1 Geometry 2 Geometry 3Outer Ring o/d (mm) 5,625 5,000 5,300Mating surface dia (mm) 4,500 4,500 4,500Inner ring i/d (mm) 3,600 4,050 3,825ce=d/D 0.8 0.9 0.85ci=di/d 0.8 0.9 0.85Width of ring (mm) 1,125 1,050 1,080Interference (mm) 3.25 7.64 4.68Bending Moment (MNm) 48.4 48.4 48.4Axial Load (Te) 280 280 280StressesOuter Ring, mating surface (N/mm2) Shrink 84 185 117Outer Ring, mating surface (N/mm2) Bending 6.4 16.8 9.6Resultant Range 90.4 - 77.6 201.8 - 168.2 126.6 - 107.4Inner Ring, mating surface (N/mm2) Shrink 68 167 100Inner Ring, mating surface (N/mm2) Bending 7.4 16.1 10Resultant Range 75.4 - 60.6 183.1 - 150.9 110 - 90Inner Ring, Inner surface (N/mm2) Shrink 91 194 125Inner Ring, Inner surface (N/mm2) Bending 7.7 17.2 10.8Resultant Range 98.7 - 83.3 211.2 - 176.8 135.8 - 124.2
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2.5.1.2 Finite Element AnalysisThe geometries in Table 2.2 were then analysed for verification purposes using finite element analysis which produced the results in Table 2.3.
Table 2.3: Summary of FE ResultsStresses Geometry 1 Geometry 2 Geometry 3Outer Ring, mating surface (N/mm2) Shrink 80 182 113Resultant Range Shrink + Bend (N/mm2) 83 - 77 190 - 177 115 - 102Inner Ring, mating surface (N/mm2) Shrink 71 172 104Resultant Range (N/mm2) 77 - 63 186 - 161 111 - 95Inner Ring, Inner surface (N/mm2) Shrink 87 188 121Resultant Range (N/mm2) 83 - 67 206 - 176 130 - 116
It can be seen that the predicted stresses with the finite element analysis give a good correlation when compared to the empirical method developed by SKF, in most cases the finite element stresses are slightly lower than those predicted empirically.
Based on the above, the optimum ring size was calculated. The 1 yr. return loading was used for the design load and the stress was limited to nominally 80% yield stress at this loading value. This approach allowed sufficient reserve capacity to resist the 50 yr. return wind load (an increase of only 6.6% in terms of moment loading).
The fatigue moment loading is only 28% of the one year return loading so using the above criteria, target fatigue stresses on the structure would be the order of 22% yield stress which was deemed acceptable.
Based on a geometry of 4500mm interface diameter with ring wall thicknesses of 150mm and lengths of 1200 mm, together with a ring interference of 4 mm, stresses of 120 N/mm2 were obtained for the shrink fit and peak stresses of 270 N/mm2 when the 1 yr. return loading was applied. In terms of the forged rings which would be used to manufacture the clamping rings, these levels of stress were well within the capabilities of the material.
With the application of the 1 yr. return wind loading, some loss of contact between the inner and outer rings was indicated over a 60 degree arc for a depth of 1/3 of the ring, this was seen as acceptable as it coincided with a corresponding increase in pressure on opposing faces.
2.5.2 Heating Method and Budget PricesCalculations were carried out to assess the heating required to expand the outer clamping ring to provide clearance to lift the tower into the raised position, these calculations indicate a mean bulk temperature of some 200 degrees centigrade is required to expand the ring approximately 11mm.
To gain an appreciation of the cost and time associated with carrying out this heating process contact was made with Beatyheat who specialise in on-site heat treatment technology. They advised that to heat the outer clamping ring and the shell adjacent to the ring to achieve 180 degrees centigrade on the inside surface would require a total power input of 567kW provided by 210 heating elements of 2.7kW. The heating areas would be insulated with a 25mm thick ceramic fibre blanket.
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Beatyheat have estimated a budget price of circa £16000 to carry out this work on one tower. This cost excluded the actual power supply to the tower (which would be substantial and maintained for over 100 hours), and any access provisions to carry out the work.
2.5.3 Alternative Clamping MethodsAlternative clamping options were considered namely:
• Fill the annulus above the seal with an approved grout.• Use a series of mechanical slips located around the tube circumference, when these engage
with the inner tower section the teeth would bite hence the tower self weight would form the radial gripping force.
These two options were discussed with Oil States and the findings were as follows:
2.5.3.1 Grout SealClamping would be achieved by grouting the annulus above the bottom seal - this would provide all the restraint required and the suitability of grouted connections for this application is well proven.
This method is not reversible and would not allow the tower to be lowered for maintenance purposes.
This method would not allow the installation process to be reversed at any time. At the end of the service life, decommissioning would require water / abrasive cutting of the tubulars.
Budget prices quoted by Oilstates indicated costs as follows:
• Grout Seal - £6,500 per tower.• Grouting Services - £130,000 for 8 towers (£16,250 per tower).• Decommissioning cost by cutting - £130,000 for 8 towers.
2.5.3.2 Mechanical Slip JointThis option would involve the use of a mechanical slip and dog type locking mechanism. Once the slips were engaged, the gripper teeth would bite into the inner tube, and the weight of the upper structure would energise the radial gripping force. This method, unlike the grout seal option, is reversible and would allow dismantling rather than destructive decommissioning.
Oil States currently manufacture standard products up to 2.5m diameter but modular gripping units could be manufactured to suit this application. These units are hydraulically operated as standard, however it may not be necessary to provide hydraulic operation due to the fact that the systems of work requiring tower lowering would require infrequent activation. A worm drive system operated by hand, or with a pneumatic wrench could be adequate.
A typical arrangement is shown in Figure 2.7. This method is currently used to secure oil rigs to their piles, although it should be noted that the loading on oil rigs is predominantly axial whereas the loading on a wind turbine tower is predominantly flexural.
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Grippingposition
AiiiujiHiti nifllnd»n Inwiii mu Jirlpfilnn
In flnfri(0* pile «tcirdiiciiiH
Rsleaudposition
A/thmlng qiUmtercWM iirp|Ki-fl
u|ifien1n, 't'Ii irJi.|t IHhf nr oornluflor
Figure 2.7: Mechanical Slip Joint Mechanism (Typical)
2.5.4 Discussion of Lock-Off Systems
It is felt that the shrink fit lock off system could be developed into a workable solution. Close attention to the design of the mating surfaces would need to be considered.
The size of these clamping rings would be such that manufacturing capability would be limited to a few UK companies and would probably involve the use of forgings as a feed stock.
An advantage of using a forged clamp ring is that materials can be selected with sufficient mechanical properties to allow as small a ring wall thickness as possible, whilst still maintaining overall integrity.
The cost of applying the heat to make and release the joint could be prohibitive and the possibility of sea water attack and fretting of the joint interface must be considered. This would be especially pertinent when the joint is released for turbine maintenance.
The detailed arrangement for attaching the clamping rings to the main tower sections needs further development.
Alternative clamping methods are available to lock the tower in the raised position and there are other commercially available options which have been developed by Oil States specifically for the offshore environment. All these options rely on an interface pressure being developed between the upper and lower tower structure, as such for each option local stiffening of the shell by clamping rings would be required.
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The best technical solution may be the use of the mechanical slip joint as described above. This type of device is well proven in terms of resisting axial loads of over 1000 Te and could be developed to resist the required bending loads in this application. The slip joint is a reversible connection and whilst normally being hydraulically operated, it could be manually operated for this application. This solution would be stand alone not requiring the application of heat over long periods to release the joint. At this stage the costs associated with this alternative have not been received from Oil States but they comment that it would be "very expensive".
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2.6 Access and ServicesIrrespective of the mechanism for raising the upper tower, all access and service provisions must be made within the upper tower. For the hydraulic system it is clearly not possible to accommodate electrical services or doors within the pressurised chamber. The case of the winch mechanism, any service provision would be likely to interfere with the cables an/or pulleys.
The main access door would be positioned just above the insitu joint between the upper and lower sections. This would be approximately 20m above mean sea level which is somewhat higher that the access location for conventional towers. At this level it is likely to be necessary to provided an intermediate platform stage in the access ladder - this would be prefabricated onto the outside of the lower tower.
Electrical cables would be run in J-tubes on the outside face of the lower tower, as used in many conventional OWECS. Cables would be drawn up the J-tubes and into the upper tower using a winch located close to the access door and terminated in a break-bar assembly close to this position. The main transformer will be located in a dedicated chamber below the access door, in the area of overlap between the upper and lower towers. Access and services between the transformer and the nacelle would be as for conventional towers.
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2.7 Tower / Turbine Installation Methodology
2.7.1 Conventional Installation Methodology
Experience from previous offshore installations was used at Horns Rev in Denmark for the installation of an 80 turbine wind farm during the summer of 2002.
Work was planned such that maximum pre assembly work was carried out on shore such that the offshore work was kept to a minimum. This resulted in an eight-step turbine installation process which consisted of:
1. Foundation2. Transition piece3. Lower tower4. Upper tower5. Nacelle complete with 2 blades6. Third blade7. Electrical Hook up8. Commission
At Horns Rev the foundation consisted of a 200Te monopile section made up of 4m diameter cans sunk 20-25m into the seabed. Once installed, erosion protection was applied to the seabed around the pile by way of typically 900Te of 3-15cm diameter granite stones.
Once the foundation was installed the transition piece, which contained the tower mounting flange and some ancillary access platforms and equipment, was grouted to the monopile to tight alignment tolerances to counter any inaccuracies during the piling process.
At this point, the tower and turbine were erected. This was carried out using 2 jack up ships that were capable of erecting 1 turbine each over a 24 hour period.
On completion of the turbine installation, electrical hook up was carried out and the units commissioned. During this activity a total of 30 Commissioning personnel were active for the 80 Turbines.
2.7.2 Telescopic Tower Installation
The erection of telescopic wind turbines would again be based on maximising the erection on shore, to this end the following work would be completed prior to floating offshore.
• Manufacture of the separate tower sections including clamping rings and inflatable sealing units.
• Installation of tower furniture (electric's, ladders, platforms etc).• Erection of the inner tower and outer tower sections.• Installation of the pulley system for hoisting or the pipework for the hydraulic lift.• Installation of the Nacelle and blade units.
2.7.2.1 Onshore Erection StageThe onshore erection stage would be carried out at a suitable quayside location.
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Manufacture of the separate tower sections would be broadly as that for conventional turbines. It is not envisaged that the installation of the clamping rings would pose any additional problems as these would be fabricated in a similar manner to current practices.
Installation of the tower furniture would be carried out as on conventional towers. It is envisaged that the upper tower would contain all the necessary furniture of a conventional tower including electrical control panels and step up transformers. Cabling would terminate at a junction box within the upper tower.
Erection of the inner and outer towers could be carried out by lifting the inner section into the outer section using a mast crane. This would involve the use of a crane of sufficient size to lift the upper tower section which would weigh approximately 115Te and be approximately 65m long into the lower tower section which when upright would be approximately 45m high, giving a total lift height of approximately 120m. To allow the clamping rings to pass, the outer ring would have to be heated sufficiently to allow clearance for the inner ring to pass the outer ring.
External furniture would be added to the lower tower structure in the form of access platforms for boat access. Provision would also be required possibly by spiral staircase and work platforms to give access to the clamp ring at the top of the lower tower and to provide access once erected via a suitable access hatch to the upper tower.
Installation of the seal units for hydraulic lift would need particular care to prevent seal damage during the assembly process.
In terms of the hoist lift option, installation of this equipment would pose several problems to be overcome. For example, space between the outer and inner shell would be typically 300 - 500mm maximum, giving very little space to install or maintain pulleys or to thread the steel rope, however it may be possible to carry out this operation using a shuttle type system if the towers are horizontal. This option has not been developed further or costed. Careful consideration to the mounting of the hoist unit and order of construction would have to be developed.
The hydraulic lift has the advantage over hoist lift in terms of ease of assembly.
Installation of the nacelle and blades would be as conventional towers using a crane of sufficient capacity and height (circa 125Te and 80m lift).
Discussions have been held with a specialist erection company , Sarens, who have advised that these onshore erection operations could be carried out using, in the main, a Demag CC2600 crawler crane together with a 200 Te telescopic crane for ancillary lifts and load tailing operations in conjunction with the crawler crane.
Budget hire costs for these cranes are as follows:
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Demag CC2600 200 Te Telescopic
Mobilisation (one off) cost £50000 Hire Charge all in/ day £1500Hire charge per week £9000Provision of Driver per week £1000 Fuel per week £1000
In terms of the hydraulic lift option, assuming continuity of lifts, equipment deliveries etc., Sarens have anticipated some 2 days erection per Turbine. If we allowed an additional 2 days for furniture mounting, electrical hook up and logistical delays, for 30 turbines, on shore hire would be 120 days or 18 weeks. This gave a craneage cost of £430k for 30 turbines or £14k per turbine.
It was anticipated that onshore labour to carry out this work for the hydraulic lift option would be 4 mechanical operatives and 3 electrical operatives per turbine for 4 days. This gave a cost of £8.5k per turbine.
The rope hoist lift option could potentially involve more resources during the onshore erection phase in terms of mounting the pulleys, hoist unit and rope reeving, it is assumed that this would take a further 3 days per turbine, the cost of craneage and labour would increase to some £23k per turbine craneage and £15k per turbine labour.
2.7.2.2 Offshore Erection StageOffshore erection processes would be much simpler than at present as this would involve:
• The submerging of the complete tower unit to the sea bed• Erection of the tower using either hoist or hydraulic means (including the heating of the
outer clamping ring without the need for large jacking structures and cranes• Hook up the electric supply and commissioning as conventional units
It is assumed that the cost of erection at sea would be similar for the hoist or hydraulic lift options and would involve connection of a mobile generator to power the hoist unit or connection of mobile pumps for the hydraulic lift.
Both options would require the heating processes for the shrink fit.
Once the tower was raised and locked into position, electrical hook up would be carried out to the junction box located in the upper tower structure through a convenient access hatch. Following this hook up, commissioning could be carried out as conventional turbines.
In summary it is anticipated that the offshore erection processes assuming adequate weather conditions would involve approximately 2 operatives for 2 days per turbine (excluding the heating of the clamp ring) giving a typical labour cost of £1500 per turbine.
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2.8 Economic Analysis
2.8.1 Estimated Supply / Installation Cost for Conventional OWEC
In carrying out this assessment of the technical requirements for Telescopic Wind Towers, provisional costs of some of the aspects pertaining to these proposals have been developed, the following section is the summary of a literature search into current costs for the supply and installation of offshore wind farms.
Based on an exchange rate of 1.5 Euro/£, analysis of recent and proposed offshore installations gives a typical installed cost of £1.14million/MW. Detailed component costs making up this figure are not widely available, however following the literature search most agencies such as BWEA, EWEA and Greenpeace, refer to Cost Optimisation of Wind Turbines for Large-scale Offshore Wind Farms [2] which quotes the following breakdown of component costs, which when combined with the overall cost estimated above gave an indication of the component cost breakdown for a 2MW machine.
Table 2.4: Cost Breakdown of Wind Turbine ComponentsItem % of Total Cost/MW £K/TurbineTower 12.4 0.14 282Blades 12.9 0.15 293Gearbox 8.8 0.1 200Nacelle 7.6 0.09 173Hub 1.9 0.02 43Shaft 2.9 0.03 66Generator 5.3 0.06 121Yaw control 2.9 0.03 66Controller 2.9 0.03 66Braking 1.2 0.01 27Foundation 17.6 0.2 400Assembly 3 0.03 68Transport 3 0.03 68Grid Connect 17.6 0.2 400Total 100 2,275
The above information was used to arrive at an installation assembly cost of £68000. It should be noted that these costs are based on the first tranche of offshore wind turbines which tended to be situated on more favourable sites. It is speculated that the actual cost may now be higher.
A Greenpeace study [3] gave an additional breakdown of labour costs required for various stages of
the wind turbines development in terms of labour years per installed MW, these figures have been used to estimate labour costs for conventional installation of a 2 MW machine based on a labour year being 45 weeks at £1250/operative/ week. (5 day week, 50 hours @ £25/hr).
Table 2.5: Manpower Requirements for Conventional Turbine Erection & CommissioningItem Labour Effort
(Man-Weeks)Cost/MW
Foundation Structure 27 318Electrical & Connecting cables 4.5 5.6Wind Turbines 27 318Project Management &Commissioning
9.9 13.4
Total 68.4 86.5
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From the above we considered the turbine erection only and assumed 35% of the project management costs was associated with this activity gives a typical installed labour cost of £k38.
This indicates a balance in erection costs to cover craneage and tools of approximately £30k per turbine or £900k for a 30 turbine farm.
Alternative sources of hire costs of jack up rigs are typically the order of £25k to £50k per day which gives a craneage cost of £750k - £1500k based on erecting one turbine per day. It can be seen from the above that using the figures derived above are at the lower end of the range.
This cost is obviously extremely dependant on weather conditions and the sea state at the time of the erection.
2.8.2 Additional Structural Costs for Telescopic Tower
A telescopic wind tower would have a greater supply cost than a monopiled tower due to the following:
• Overlap in length between the lower and upper tower section, estimated at 10m which would give an increased structural weight of some 30Te. Based on a typical manufacturing cost of a tower at £800/Te of steelwork (based on recent quotations for similar section monopiles), the extra cost for this item would be £24k.
• Lower tower would be a larger diameter than upper tower to allow for the telescopic action, likely to be nominally 4.5m dia as opposed to typical 4.2m, increase in weight approximately 15Te. As above based on a typical manufacturing cost of £800/Te of fabricated steelwork, the extra cost for this item would be £12k.
• Using the shrink fit approach, machine clamping rings would be thicker locally to resist the clamp and aerodynamic forces on both the outer and inner tubes giving an increased weight of 42Te. To manufacture these rings which would be nominally 4.8m outside diameter with 4.2m inside diameter and 1.2m deep, it was envisaged that forgings would be required which would need to be machined at the shrink fit interface. Based on a quoted manufacturing cost of £1300/Te together with a pattern cost of £6000 per ring (split between 30 turbines) and a machining cost of £2000 per ring, this gave an estimated cost of £59k.
• It is anticipated that to provide access to the clamping ring and to the upper tower structure additional access stairways and platforms would be required which would cost an estimated £12k.
The extra tower cost therefore based on an increased weight of 87Te per turbine would be £107k.
2.8.3 Raising Mechanism Costs
2.8.3.1 Winch OptionAlthough a hoist driven lift mechanism appears to be the less practical solution, Section 2.3.5 of this report includes an estimate for a hoisting system to erect the turbine at an estimated £100k which does not include for power supplies. Power could be taken from the unit described above or if preferred, a purpose built diesel set of 135 kVA could be supplied for approximately £15k ie approximately £500 per tower based on buying one unit for 30 towers.
Labour costs associated with this option are described in Section 2.6.2.1 of this report.
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2.8.3.2 Hydraulic OptionHydraulic lifting costs could be summarised as follows:
• Cost of hydraulic pumping unit £2k per tower (based on spreading cost over 30 turbines), see section 2.4.1 of the study.
• Cost of seal and guidance system £13k per tower (2 off inflatable seal units per tower) see section 2.4.2 of the study.
• Cost of additional plating to seal the towers 24Te at £800 per Te. £20k
This gave an overall hydraulic cost of approximately £35k
In addition to this would be the onshore and offshore erection costs as detailed in section 2.6.2 of this report.
2.8.4 Heat Shrink Lock-Off
Item 2.5.2 of the study includes an estimate for applying the heat pads to the clamping rings to erect the turbine at a cost of £16k per turbine excluding any access requirements and electric supplies.
Access would be provided by way of the access stairways and platforms described above.
Enquiries were made into the cost of a purpose designed diesel generator which could be ship mounted to service a wind farm, the cost of one of these units to give 600kVA capacity was approximately £55k, with necessary cabling. Fuel costs for each heating cycle would be the order of £7k (based on 121 litres/hr diesel fuel consumption for 100 hours and £0.6/litre).
As stated in Section 2.6.2.1 of this report it is likely that an additional heating cycle would be required to allow the inner and outer tower to be assembled onshore prior to shipping offshore, the estimated cost for this cycle would be circa £20k including 60000 kWh of electricity for the heating cycle.
2.8.5 Summary of Cost Options
Table 2.6: Cost Option 1 - Hoist Lift and Shrink FitItem Cost £KExtra Steel Structure 107Cost of Hoist Drive 100Guidance system* 13Heating pads for shrink fit (Hire) 16diesel fuel cost 7Cost of diesel generator ** 2Heating Cycle during onshore build 20Onshore build costs 38Offshore build costs 1.5Total 334.5* Based on using inflatable guides** If we assume one generator is bought for a wind farm consisting of 30 Turbines
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Table 2.7: Cost Option 2 - Hydraulic Lift and Shrink FitItem Cost £KExtra Steel Structure 107Cost of hydraulic option 35Heating pads for shrink fit (Hire) 16diesel fuel cost 7Cost of diesel generator * 2Heating costs during onshore build 20Onshore build costs 22.5Offshore build costs 1.5Total 211* If we assume one generator is bought for a wind farm consisting of 30 Turbines
From the above cost summaries, it can be seen that Option 2 is the most favourable. This option, with an anticipated erection cost plus the cost of additional steelwork and structure of £211k, is £145k more expensive than the erection cost of an equivalent conventional tower estimated in Section 2.7.1 of the report.
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2.9 Scale Model Test Rig
2.9.1 Design
This activity remit included for the design of a 1/4.5 scale model test rig to enable the hoisting concept and the shrink fit clamping mechanisms to be tested. Following on from the findings in Section 2.7, it was decided that the scaled rig would test the hydraulic lift and shrink fit concept as this was felt to offer the most practical and cost effective full size solution.
2.9.1.1 Scaled LoadingThe first phase of this work was to estimate the forces to be applied to the scaled model to give the physical similarity of the full sized tower. The resuting loads are summarised below:
Loads During Erection Axial 199kN or 24kNShear 2kNMoment 15kN
50 Year Extreme loadings Axial 199kN or 24kNShear 46kNMoment 533kN m at the joint
670kN m at the tower base
The reason for the two axial loads specified, the 199kN load is required to simulate the 2.3 bar water pressure in the tower base to raise the tower. The 24kN load is the scaled weight of the nacelle and upper tower structure.
2.9.1.2 Calculations and Fabrication DrawingsCalculations have been undertaken to facilitate the preparation of detailed fabrication drawings.
In order to limit the height of the overall tower it was decided to set the stroke of the hydraulic lift to 1500 mm, this was deemed a suitable stroke to test the seal and guidance arrangement.
Drawing EX 103910-12 was produced for the scale model showing component parts and erection procedure. This drawing is shown in Appendix B.
Calculations indicated horizontal deflections under the applied maximum load of the order of 2mm at the shrink joint and 4mm at the top of the test tower, these would have been measured relative to a "datum bar" using suitable dial gauges, located during testing to produce a load deflection curve for the structure. From this, drawings SD/Y24/03/01 and SD/Y24/03/02 were developed for the support frame. These drawings are shown in Appendix B.
In order to erect this structure a crane hook height of over 8m would be required, it was agreed that this could be accommodated in the CNES Teesside Structural Workshop buildings.
2.9.2 Test Procedure2.9.2.1 No Load TestThe test tower would be assembled in the test rig with the top cross beam on the test support structure removed. The inflatable seals would be energised with water at a nominal 10 bar and low
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pressure mains water would be admitted to the tower base to raise the tower unloaded. The tower would be raised approximately 1200mm (clamping rings are not engaged). This tower raising/lowering operation would be completed a number of times to check for smooth operation.
A sketch of the test rig layout for this test is shown in Figure 2.8 below:
Tower raise lower
Low Pressure water input
Figure 2.8: No Load Test Rig Configuration
2.9.2.2 Pressure TestThe test weights would be assembled on top of the tower as detailed in drawing EX 102910-12. These test weights of up to 20 Te would be needed to generate the necessary 2.3 bar pressure in the lower tower to simulate actual full scale pressure conditions.
The location of the test weights has been designed with an offset from the tower centreline to simulate the erection moment conditions.
Testing would involve repeated raising and lowering operations to test the seal and guidance system and also a leakage test in the raised position. Again as the previous test, the stroke would be limited to nominally 1200mm
A sketch of the test rig layout for this test is shown in Figure 2.9 below
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Test Weight
Tower raise lower
Low Pressure water input
Figure 2.9: Pressure Test Configuration
2.9.2.3 Raise and ClampThe next test would involve applying the heating pads to the outer clamp ring with the tower in the lowered position, once the heating cycle had been carried out, the tower would be fully raised until the clamping rings are engaged and held in this position by the water pressure until the clamp ring had cooled and the tower was locked in the raised position.
Once cooled, the clamp ring area would be soaked in a strong saline solution for subsequent testing to try to mimic sea water attack. A sketch of the test rig layout for this test is shown in Figure 2.10 below:
Test Weight
Tower raise lower
Outer Clamp Ring Heated
Low Pressure water input
Figure 2.10: Raise and Clamp Configuration
2.9.2.4 Load TestingOnce the tower has been clamped in the raised position, the 20 Te test weight would be removed, the top jacking beam and the support structure top cross beam would be erected.
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Three Hydraulic cylinders would be fitted to the structure and connected to the hydraulic pumps. The datum bar and dial gauges would be fitted and the tower loaded progressively up to the 50 year return scaled moment and shear conditions. At intervals, the loads and deflections would be recorded at a number of points on the tower such that the load deformation curves could be produced and compared with theoretical. Particular attention would be taken with any movements at the clamping ring interfaces.
The loads would be progressively removed, again deflections would be taken to ensure no permanent deformation or movement had occurred.
At this point the tower would be repeatedly loaded up to the 50 year return equivalent moment to demonstrate any creep tendencies of the structure and in particular the joint interface.
The sketch of the test rig layout for this test is shown in Figure 2.11 below:
Jacks
Maximum Loads to apply Axial 24 KN Shear 46 KNMoment at joint 533 KNM
Approx 570( mmTower scale!/4.5
Figure 2.11: Load Test Configuration
2.9.2.5 Unclamp and Low erOn completion of the load tests, the jacks and measuring equipment would be removed and the clamp ring reheated to allow the tower to be lowered. Attention would be paid to ensure the tower lowers smoothly.
At this point more saline solution would be applied to the joint and the unit would be left overnight or preferably over a week end.
At this point tests as described in Sections 2.8.2.3 and 2.8.2.4 would be repeated.
2.9.2.6 Dismantle and Condition AssessmentOnce testing had been completed, the tower and support structure would be dismantled and a thorough inspection of the clamp faces, seals and guides carried out.
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2.9.3 Cancellation of Testing Programme
A cost estimate was generated based on the test plan and drawings described above. Fabrication will be carried out in-house, however a number of elements eg heating of the clamping rings will be external supply (in which case quotations have been sought).
The final cost of the programme has been found to be approximately £7k in excess of the budgetary allowance, despite efforts to minimise costs. Since this cost is based on fixed quotations there is no scope for reducing costs other than to reduce the scale of the test.
It was concluded that a further reduction in scale would not be desirable since this would exacerbate the scale related problems and further reduce the applicability of the results at full scale. Furthermore, the desk study described in the preceding sections has successfully demonstrated that the hydraulic lift / heat shrink combination is technically straightforward, utilising known materials operating within conventional pressures and stress ranges. It is therefore questionable how much new information would be gained from the test programme, particularly at the smaller scale.
It was therefore recommended to DTI that the telescopic tower study be concluded without testing, and that the funding thus made available be redirected to other areas of the project. This was accepted and led to the completion of Phase 4 of the Self Burial study (Section 4.5).
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3 TRANSPORTATION
3.1 Introduction
The objective of this study is as follows:
• To investigate the feasibility of transporting and installing a wind turbine unit via forks mounted over the end of a submersible barge and to determine the extent (if any) of barge strengthening that may be required such that costs may be determined. Alternative methods will also be investigated.
This has been achieved through the following activities -
(a) Site ReviewThe 18 locations licensed for offshore wind farm installations have been investigated and assessed against criteria such as depth, access and limitations imposed by barge and unit size such as draft, width, length. The sites have been subjectively ranked for suitability for the integrated installation concept.
(b) Criteria DevelopmentThis activity looks at the broad options for quayside load-out, transportation and set-down and develops the basic criteria which must be fulfilled by the detailed option study.
(c) Installation OptionsThis activity develops three options in sufficient technical detail to produce preliminary cost estimates for each option, and an overall score based on a subjective risk assessment.
(d) Conceptual DesignThis activity develops a preferred option in detail in order to provide a reliable upper and lower bound cost estimate for the complete load-out, transportation and set-down process. General arrangement drawings have been created which may form the basis of a future barge conversion.
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3.2 Site Review
3.2.1 Locations
The 18 locations licenced for offshore wind farm installation are on 13 sites around the UK, particularly the eastern side of the Irish Sea, the Bristol Channel and the east coast of England from the Thames estuary to Teesside. The sites are shown in the diagram below [4].
TW7I II*1 I IHWJT iVJT i.
, A yVX/WV ^ r
1 fS otway Ffath
jy YXBarrow 3 * sSlwIlFtet a
6oulhpof1 4# J
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S Damn;at her in nda^#
Legend
• W ii.rtnn 41 ■ Tin f ™ u pw j
S FW "Vlln'bHS |5 I E^rH-I^WI
IDtuIb *h |D • dMDH&J
A
i 13 T»JnHt
Inner Downing
r\J it>J f y*
Scroby Ssnde
0untied Sand?
Kcfirtth Flail
j-_ ■■
Fig 3.1: Potential offshore wind farm locations around the UK
The sites and ports are assessed against physical criteria of depth and access together with limitations imposed by barge size such as draft, width length and air draft of the unit. The sites and loadout ports are ranked by subjective review into three categories:
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Table 3.1: Site CategorisationCategory Remark
“A” Prime candidate, likely to meet criteria subject to more detailed review“B” Secondary candidate, may require review of criteria and adaptation of limitations and to be subject
to detailed review.“C” Tertiary candidate, will probably not meet the present criteria; the project objectives, criteria and
possible limitations will need reassessment if this candidate is to be included.
3.2.2 Seabed ConditionsEach site will require detailed investigation, however for this conceptual level of review, on the basis of BA charts and BGS data, those sites with a predominently sand seabed will be acceptable; gravel is less acceptable and will require detailed consideration and rock seabed with no, or insufficient overlay of sand will not be acceptable. On this basis, the sites are (Table 3.2):
Table 3.2: Site Categorisation - Seabed ConditionsNo. Site Acceptability1 a b Solway A2 Barrow A3 a b c Shell Flat A4 Southport A5 Burbo A6 a b N Hoyle, Rhyl A7 Scarweather B8 Kentish Flats B9 Gunfleet A10 Scrooby Sand B11 Cromer C12 a b Lynn, Dowsing A13 Teeside B
3.2.3 Water Depths
The installation water depth criteria are a minimum water depth of 10m and a target water depth range of 15m to 25m.
The sites have been reviewed on the BA charts of the areas and estimates made of the locations and area coverage of the sites based on sites of 1nm by 3nm in extent for 30 units. The least water depth and greatest water depth at chart datum (CD) have been noted.
The rise of tide above chart datum has been assessed, using the BA program Totaltide for two typical seven day periods, one in early May and one in late August; for the representative year 2002. The tides were assessed for ports nearby the sites and then checked against average tide data shown on the charts. The least and greatest high water rise during the periods were noted.
Based on a required 10m available water depth, the site acceptability is (Table 3.3):
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Table 3.3: Site Categorisation - Water Depth (10m)No. Site Acceptability @
10m WDRemark
1 a b Solway B Only in deeper areas2 Barrow A Sufficient water at CD3 a b c Shell Flat C Tidal rise is required4 Southport B Only in deeper areas5 Burbo C Tidal rise is required6 a b N Hoyle, Rhyl B Only in deeper areas7 Scarweather B Only in deeper areas8 Kentish Flats C Tidal rise is required9 Gunfleet C Tidal rise is required10 Scrooby Sand C Tidal rise is required11 Cromer A Sufficient water at CD12 a b Lynn, Dowsing B Only in deeper areas13 Teeside A Sufficient water at CD
Based on a required 15m and greater water depth, the site acceptability is:
Table 3.4: Site Categorisation - Water Depth (15m)No. Site Acceptability @
10m WDRemark
1 a b Solway B Only in deeper areas plus tidal rise2 Barrow A Sufficient water at CD3 a b c Shell Flat C Only in deeper areas plus tidal rise4 Southport B Only in deeper areas plus tidal rise5 Burbo C Only in deeper areas plus tidal rise6 a b N Hoyle, Rhyl B Only in deeper areas plus tidal rise7 Scarweather B Only in deeper areas8 Kentish Flats C Not Possible9 Gunfleet C Not Possible10 Scrooby Sand C Not Possible11 Cromer A Sufficient water at CD12 a b Lynn, Dowsing B Only in deeper areas plus tidal rise13 Teeside A Only in deeper areas plus tidal rise
3.2.4 AccessAccess in this context refers to the ability of vessels of suitable draft to navigate to and from the candidate sites. While there are various navigational obstructions, no site is inaccessible. Certain of the sites would have tidal access only, and require the installation set-up and actual installation operation to be completed without any time loss or equipment failure, within a tidal cycle of 12 hours. At this level of review the installation should be achievable on a rising tide and completed by or at high water and before the tide starts to drop away.
Access to some of the sites is over existing subsea oil and gas pipelines. This is a risk, and will require careful management to avoid danger to the pipelines from hook-carried turbine units. In any event, underwriters will require specific assurances that the carriage is safe and there is no dropped object risk to the pipelines. The degree of acceptability is (Table 3.5):
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Table 3.5: Site Categorisation - AccessNo. Site Pipeline
Crossings1 a b Solway B2 Barrow B3 a b c Shell Flat B4 Southport B5 Burbo A6 a b N Hoyle, Rhyl A7 Scarweather A8 Kentish Flats A9 Gunfleet A10 Scrooby Sand B11 Cromer B12 a b Lynn, Dowsing B13 Teeside A
3.2.5 Loadout Ports
The criteria for acceptable loadout ports relative to this study re:
• Allowable drafts• Allowable widths and lengths• Wide access time windows• Short transit distances to the installation site.
The nominal transport draft is about 8m to 8.5m for carriage at the barge end. If the unit is transported on deck for a lifted installation then the transit draft is only barge draft and estimated at 2m.
The minimum barge size to be considered is 25m beam, as this will allow transport of the structure without projection of the unit base outside the barge dimensions. A likely barge size is 27.5m beam as this is a conventional 300ft by 90ft barge, of which there are many.
The access time windows are typicallly dependent on the stand of the high water and the operating custom of the port and lock, if any. In broad terms exit from the port will be constrained to either side of high water slack if the unit is carried at the barge end and exit windows will be short, and for some ports may be constrained by the spring tides. If the limit is only barge draft then port exit windows will be longer.
Most ports advise maximum drafts and exit windows based on spring tides with the greatest rise. Data on port limitations has been drawn from public domain data and 18 possible loadout ports have been contacted directly for details; of those only 10 actually provided any formal response.The detail of exit windows for any one port will have to be the subject of closer study when the ports and transportation options are better defined.
Ideally the port closest to the installation site should be chosen as the loadout port, however, because of draft, width, length and tidal restrictions, this may not be achievable. Thus the transit time from the port to the installation site may be many hours. For the purposes of this study a tow speed of 5 knots has been used.
If the 8.5m draft constraint is maintained then there are suitable loadout ports in each geographical region (ie Irish Sea, Bristol Channel and East Coast). It should be noted that the deeper draft berths
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are often retained only for commercial shipping traffic and may not be released for long term use as a construction berth. Alternatively a method will have to be devised of reducing the transit draft or taking all cargoes on deck for minimum draft transits to open up all sites.
If the width constraint is 27.5m barge beam then there are suitable ports in each of the regions. If the width constraint is 25m then there are a few more candidate ports to chose from. However, the ports closest to the installation sites, such as Workington and Fleetwood may not be suitable.
The length of the combined barge and unit is also important. A typical 300ft by 90 ft barge is 91.5m long, this plus the overhang of the unit plus the 12m semi-diameter results in a transport at least 105m to 110m long. The required length is even greater as at least two tugs will enter a lock with the barge, giving a length requirement of 130m as a minimum if harbour tugs are used that can lie transversely in the lock and a transit tug can be connected once the outer gate is opened, or 150m if other, larger, tug configurations are used.
There are ports in each installation area that can meet the length requirement, but again, they are not necessarily ports closest to the installation sites. Several lock ports will permit transits to “canal” through, with open lock gates, but this depends on the rise of tide and previous negotiation with the port. This removes the length restriction, but imposes tidal window restrictions.
The summary of loadout port acceptability for deep draft transport based on exit drafts is given in Table 3.6
Table 3.6: Site Categorisation - PortsRegion Port Deep draft Constraint
Workington B Draft, widthBarrow A Draft, width
Irish Sea Fleetwood C Draft, width, lengthLiverpool A OKBirkenhead B Draft, widthBelfast A DraftBarry C Width, draft
Bristol Channel Swansea B Draft, widthCardiff B Draft, widthDover B DraftChatham C Draft, widthHarwich A DraftLowestoft C Draft
East Coast Gt Yarmouth C DraftImmingham B Draft, widthTees B DraftHartlepool B Draft, widthSunderland C Draft, width
If the draft is to be only some 2m or so, then many more ports become acceptable, for instance the apparent anomalies of Great Yarmouth and Lowestoft, which routinely work large barges of project cargo.
A port constraint that is not considered here and will be a matter for detailed review of candidate ports is the quay freeboard, the distance from the water edge to the quay edge.
The summary of the port review must be that either
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• Liverpool and Belfast are the candidate ports for deep draft transits and the potential frequency of entry and exit at deep draft should be further researched. Barrow is also a candidate port. The area of operations should be the Irish Sea, or
• The draft of the transport should be reduced, either by taking the unit on deck or by lifting it higher during transport, opening up more potential loadout ports.
3.2.6 Drydock Availability
An option is to construct and load the units at a ship drydock or offshore construction yard. Of the ports discussed above, those with such a facility include (Table 3.7):
Table 3.7: Site Categorisation - DrydocksRegion Port Deep draft Drydock
Workington B N/ABarrow A ClosedFleetwood C N/Akmsn Sea Liverpool A DrydockBirkenhead B DrydockBelfast A DrydockBarry C Drydock
Bristol Channel Swansea B ClosedCardiff B ClosedDover B N/AChatham C ClosedHarwich A N/ALowestoft C YardGt Yarmouth C YardEast Coast Immingham B ClosedTees B ClosedTees B Caisson basinHartlepool B YardSunderland C Drydock
The drydock market is volatile and drydocks variously go out of business, become generally unavailable or are turned to other uses. The list above gives drydocks that are known to be in existence with their present status.
Barrow drydock is closed, as are Swansea, Cardiff, Chatham and Tees. Immingham has been converted to a wet dock. Tees has a caisson basin, last used for the construction of an offshore concrete structure in about 1987, but the caisson gates have fallen into disrepair and will need refurbishment.
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3.2.7 Overall Ranking
Based on the above parameters the overall ranking of the sites is (Table 3.8):
Table 3.8: Site Categorisation - Overall RankingNo. Site Seabed Water
depth @10m
WaterDepth @15
Access Loadout port within12 hrs
Summary
1 a b Solway A B C B B B2 Barrow A A A B B A3 a b c Shell Flat A C C B A C4 Southport A B C B A B5 Burbo A C C A A C6 a b N Hoyle, Rhyl A B C A A B7 Scarweather B B B A A B8 Kentish Flats B C C A A C9 Gunfleet A C C A A C10 Scrooby Sand B C C B A C11 Cromer C A A B A B12 a b Lynn, Dowsing A B B B A B13 Teeside B A B A A A
Given that of the loadout ports Liverpool and Belfast offer least restrictions for the deep draft loadout and wet tow, and Barrow itself may be a suitable port, then the Barrow site is a prime candidate. All drafts advised by ports at this feasibility level must be reviewed with caution as the commercial berths that can accommodate deeper drafts may not be available for construction activities and the deep draft capability may only be occassionally available subject to the tidal cycle.
However, if the transport draft can be reduced then the Teesside site is also a prime candidate; again, the commercial berths that offer 10m may not be available for construction activities. The Tees has an air draft restriction of 61m up river, and of the construction berths available below this the Able UK yard and caisson dry dock has a depth of only some 3.4m at low water. Alternatively Hartlepool can take barges at construction berths in the outer harbour and has an available depth of 6m to 8m.
If the transport draft can be reduced significantly or the load taken on deck then other sites become attractive. A result of taking the cargo on deck is that the sites which have less than 10m depth of water also become attractive.
It is to be emphasised that the site and port limitations are discussed here at a feasibility level and will need to be checked in site-specific detail. This is because all the geometric limitations are tight and ports typically promulgate maximum capabilities rather than routine operational limits.
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3.3 Installation Criteria Study
3.3.1 General
In assessing the loads and motions imposed on the foundation unit, tower, and barge a discussion on the possible transportation/installation/loadout modes is required.
The simplest support concept is that proposed within the scope of work, and essentially comprises a cantilever arrangement hung over the stern of a barge supporting a pair of saddles into which a pair of trunnions mounted on the turbine column sits. Thus the turbine base is submerged and is free to rotate in Pitch. It is this scheme together with a “unit on barge deck” condition for which motions and loads are derived and it is these loads that will form the basis for any future design work.
Previous sections of this report discuss the geometric constraints imposed by the various ports of departure, primarily the barge needs to be wider than the 24m unit base. The majority of North Sea barges are of the standard 300’ (91.5 m) x 90’ (27.43 m) x 20’ (6.1m) size; these will fit the required width criteria. There may be barges of intermediate width but numbers will be very limited. Secondly our preliminary calculations using this standard barge size indicate that smaller barges may not have sufficient ballast or strength capacity. Larger barges would have more capacity in both of these areas but would limit the number of departure ports. Thus all calculations are based on using a standard barge of the above dimensions.
Calculations are based on the weights given in Table3.9 (based on Drg BIS152/DRG/90000/A, Appendix A):
Table 3.9: Weights and CofG’sCase Weight (tonnes) VCG (m)Hung off (water fill & buoyancy) 1647 10.30Deck Transport (no water fill & no buoyancy) 2363 7.44Hung off (no water fill & buoyancy) 568 22.19Deck Transport (water fill & no buoyancy) 3442 6.37
3.3.2 Base (Cantilever) Condition
This section examines the limitations, loads and motions obtained with the turbine hung off the end of the end of a standard 300’ x 90’ North Sea barge.
All calculations are based on the support system proposed in the scope of work whereby the turbine is hung off the stern of the barge in a saddle arrangement. It is assumed that the saddles will be housed in a cantilever arrangement suspended over the stern of the barge. Trunnions placed either side of the turbine mast will sit in these saddles. First pass analysis allowed the unit to rotate in the pitch direction, however the analysis showed that there was a high risk of the unit hitting the stern of the barge in even small seastates. Secondly the drag forces on the submerged portion of the unit caused significant rotation of the unit with again the risk of contact for bow hung units. Thus the full analysis was undertaken with the unit held rigid.
As per the originally proposed scheme the cone section is assumed to be filled with water. The cantilever support points are assumed to be 2 m above barge deck level; the longitudinal position is dependent on allowable clearances and is discussed below. The vertical position of the turbine support points was taken to be 13.7m above the base - this provided a maximum 8m submergence and was related to the barge cantilever support positions and barge draught.
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3.3.2.1 Clearance CriteriaThe turbine will pitch and heave relative to the seabed and relative to the barge (if it is free to rotate). Only a detailed motions analysis would provide an accurate assessment of the motions (and hence clearances) involved, and these would be dependent on the support concept.
In harbour areas it is normal to allow a lm clearance between the bottom of the barge and the seabed - some reduction may be allowed if the seabed in way of the route has been surveyed and the seabed is soft. However at installation sites there will be more movement so the clearance has to be increased. It is therefore proposed to increase the minimum clearance to 2 m. This would be applicable to seabed and barge/turbine clearances; see Figure 3.2 below.
tFigure 3.2: Cantilever Option Clearances
The consequences of these limits are follows:
• The immersion of the turbine cannot be greater than 8m if loadout and installation is required in 10m water depth.
• The 8m draught limit brings the cone section in line with the bottom of the barge, so that in order to obtain a 2m horizontal clearance the centreline of the turbine needs to positioned at least 8.4m from the stern of the barge; see figure below. The stern trim moment caused by the overhung weight needs to be counter balanced by water ballast forward. Adding this counterballast means that the minimum level keel barge draught that can be obtained is about 3.5 m.
• The overhung turbine weight and the counter ballast forward will result in a considerable hogging moment in the barge. The longitudinal strength of the barge would need to be verified; it may be that additional ballast water may be needed amidships to reduce moments - leading to an increase in minimum barge draught.
Seabed clearances can be increased (& hence overhang reduced) if the barge were to be trimmed by the head, for a stern mounted turbine. A bow trim may be acceptable for a short duration when passing shallow water areas - but is not advisable, and would require that a ballast engineer board the barge to reset the ballast for deeper water tow, again unacceptable in anything but the most benign conditions.
It is usual to tow a barge with a 18 stem trim to provide better controllability, so hanging the turbine off the bow of the barge would seem a natural solution to bottom clearance, and the natural rake of
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bow would provide additional horizontal clearance between the barge and turbine. But hanging the turbine off the bow has its own problems:
• Interference with towing gear• Increase loads in turbine and supports due to slamming• Loss of lee protection from waves and current, which is likely to result in increased towage
resistance• Increased drag would tend to push bow down
3.3.2.2 Loadout and InstallationThe proposed support concept would need to rely on trimming the barge to pick up and install the turbine unit.
With the turbine unit hung off the stern and with all ballast water (apart from residual) removed the barge has a stern trim of about 6.9 degrees with the water lapping the stern of the barge; see figure below. The bow is about 4.8m out of the water, the condition is barely stable, and the protrusion of the bow would probably be unacceptable in terms of barge longitudinal strength. As a static condition the angle of trim would be unacceptable for barge crew and ballast operators. The immersion of the turbine is about 11 95m so this form of installation could only install in water depths less than 12m.
MAXIMUM TRIM WITHOUT DECK SUBMERGENCE ALL BALLAST WATER REMOVED
Figure 3.3: Barge Trim During Loadout
Similarly the maximum water clearance (freeboard) that could be obtained at the suspension points would be about 5 m, achieved by ballasting the bow. Thus there would be no chance of lifting the turbine (lift points at 13.7 m) off a quayside, and would limit pick up (loadout) of the turbine off the seabed to water depths in excess of 13.7 - 5.0 = 8.5 m.
There are submersible barges and vessels on the market that would enable the stern to submerge further, but only the largest of these is likely to approach the 25m installation water depth criteria.
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Typically, the Giant class of submersible barge (dimension 140m long x 36m beam x 8.5m depth) can submerge to a depth 8m above its deck. Thus if the tow draughts of the turbine unit and barge were set at 8.5m and 3.5m respectively, then the maximum installation water depth would be limited to 8.5+(8.5-3.5)+8.0 = 21.5m assuming that ballast and stability capacity were found to be OK. Using this size of barge will impose additional constraints on the loadout ports.
3.3.2.3 Loads and MotionAn analysis was carried out for seastates ranging from Beaufort 1 to Beaufort 7. The barge draught was determined from the limiting clearance criteria discussed above, which was approximately 3.5m with level keel. The weight of the turbine was taken to be 3442 tonnes (i.e. cone filled with water). The analysis software calculated the buoyancy of the turbine.
Loads were generated at the suspension point using a single support restrained in all translational directions, in conjunction with “virtual” vertical supports placed very far apart to determine moments due to roll and pitch.
For a 2 point support arrangement spaced 5m apart (4m dia. column + 2 x 0.5m trunnions) the vertical loads due to Roll in BF 7 beam seas would be + 5830 tonnes and -3400 tonnes
In BF 5 beams seas the maximum and minimum loads obtained are 3495 tonnes and -1350 tonnes respectively
Thus both the downforce and uplift loads are considerable, and it is likely that additional high and low level restraints will be required to reduce loads due to rotational moments.
3.3.2.4 Tow Drag Forces on Turbine UnitAssuming the stern of the barge does not afford any lee protection to the submerged part of the unit and taking the tow speed to be 5 knots results in a drag load of the turbine of about 60 tonnes. This load is conservative in that it uses a higher drag factor to account for blockage effects in shallow water - loads will reduce with increased water depth.
3.3.3 Horns
The horn option (see Figure 3.4 below) provides a similar concept to the Base Case and addresses a number of the operational issues. The main advantage is that the unit can be picked up from the quayside and lowered without the need for barge trim. For a quay loadout the height of the horns would need to be around 15m and would require a an overhang of at least 14m to plumb the turbine unit’s lift points. The unit will be transported out of the water, thus eliminating the drag force from the turbine foundation. The main disadvantage is that the hang off distance is increased resulting in increased bending moments, and also it is more difficult to apply constraints (if required) to prevent roll and pitch of the unit.
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Position for quay pick up & installation
>\ /
Figure 3.4: Horn Option
Position for transportation and installation
3.3.4 A Frame
The main advantage of using an “A” frame system is that there is potential to lift the turbine unit from the quayside, deposit it on the barge deck, and install it in any of the proposed water depths. The barge would not require excessive trim.
However there are a number of disadvantages to this system. Primarily, the turbine structure cannot “pass through” the “A” frame unless the frame were made higher than the turbine itself, which would require a minimum lift height of 80.5 m. Thus one might as well use a proprietary sheerleg. Even with this height of “A” frame the turbine blades would have to be aligned with the barge and rotated during load out to avoid the A-Frame. Thus the “A” frame has to be moved forward so that the turbine unit does not have to swing through the frame; see Figure 3.5 below.
Figure 3.5: A Frame Option
Assuming that the turbine base can be placed to within lm of the quay edge, making an allowance of lm clearance between quay edge and barge stem, and assuming that the base edge can be placed on the stern of the barge, would require an outreach of (24/2 + 1 +1 +24/2 ) = 26m plus an
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allowance to ensure that the “A” frame balance is stable. If the lift padeyes/trunnions were positioned at 13.7m above the base then the minimum required boom length would have to be about 34m.
With a 24m wide base there is only limited space either side of the base section, (27.43-24)/2 = 1.7 m, to place the “A” frame support steelwork, unless the hinge points were moved further forward that would necessitate a further increase in boom length. Positioning support steelwork at the barge sides would invariably require the removal and repositioning of ballast tank vents. “A” frame loads into the deck will be considerable, and may be difficult to distribute without internal modifications to the barge.
Using the smallest boom length would result in a lead angle to the winch of only about 12 degrees resulting in a required winch capacity of about 4.5 times the lift capacity.
Thus indications are that the required size of a “bolt on” “A” frame system is getting close to the size of system already provided by proprietary sheerlegs.
3.3.5 Conclusion of Preliminary Study
• A standard 300’ x 90’ x 20’ North Sea barge would be the preferred candidate barge in terms of geometric considerations, availability, and ballast capacity. Thus, all of the following reviews are based on this size of barge.
• Any lift system has to be a dual component system (i.e. Boom plus wire, beam plus ram, etc) in order to meet the range of installation water depths required. For the purpose of this study, the secondary component will always be a wire.
• The lift system has to have sufficient outreach to plumb the lift points of the turbine unit. In the “criteria” report it was established that the minimum outreach required was as least 14m for a lift system mounted directly on the stern of the barge.
• The vertical centres of gravity (VCG) of the unit would be 10.3m for installation with water fill and 22.2m for installation without water fill - both include the effect of buoyancy.
• The “hung-off’ case assumed that the turbine unit was supported by a simple two beam cantilever arrangement off the end of the stern of the barge. Initial analysis was carried out with the turbine free to pitch but it was found that the turbine base was likely to impact the stern of the barge even in relatively small seastates, so the analysis was run assuming the turbine to be rigidly connected. The loads obtained were considerable (& indicated uplift), mainly driven by the large roll moment acting on a short support span. It is probable that alternative support arrangements need to be considered in order to reduce the effect of the roll moment.
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3.4 Development of Installation Options
3.4.1 General
Following on from the preliminary study it was decided that three options would be investigated in further detail. These option are:
• Horns with turbine unit suspended beneath the horns.• “A” frame with turbine unit transported suspended from the frame.• “A” frame with turbine unit transported on the barge deck.
These three option were chosen for their potential to lift (loadout) a unit from a quayside.
3.4.1.1 Loadout from QuaysideIn order to loadout from the quayside the lift system has to have sufficient outreach to plumb the lift points of the turbine unit. In the “criteria” report it was established that the minimum outreach required was as least 14m for a lift system mounted directly on the stern of the barge.
The minimum height requirement is dependent on a number of variables, namely:
a) The height of the unit’s lift points.b) The dimensions of the lifting gear (e.g. hook blocks, slings, etc).c) The freeboard of the barge.d) The distance between the water level and the top of the quay.
Variable a) is largely dependent on the stability of the unit. In the “criteria” report is was determined that the vertical centres of gravity (VCG) of the unit would be 10.3m for installation with water fill and 22.2m for installation without water fill - both include the effect of buoyancy. Thus for stability the minimum height of the lift point(s) needs to be a least 10.3 m. There are a possible number of variations in submerged weight, water fill weight, etc and so for conservatism for this phase of the study the height of the lift points will be taken as 15.0m (approximately half way between the two installation loadcases).
Variable b) In order to allow space for blocks, lifting gear, and a 1m quayside to underside base clearance, the height of the suspension point will be taken as 6m above the lift points.
Variable c) is dependent on the amount of ballast water required to counter ballast the weight of the overhung turbine unit, and if required, the amount of additional ballast required to counter any excessive bending moment in the barge.
Variable d) will vary from port to port and will be tide dependent. The possibilities are too numerous to consider at this stage of the study, so for this phase of the study it will be assumed that the deck of the barge can always be made level with the quayside during a loadout operation.
Thus the minimum design dimensions for a quayside loadout are:
• • Overhang : 14.0 m• • Height (above barge deck) : 21.0m
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3.4.1.2 Design LoadsIn order to simplify the comparison of the three options it was decided to base the basic structural design (and hence the determination of weights and prices) for each option on one single loadcase and then, for each option, discuss the viability of achieving this loadcase.
Since a quay loadout is preferable to a seabed loadout, the design loadcase has been taken as the unballasted “lift in air” weight with a 5% increase to account for dynamic effects during a loadout. Thus the design weight was as given in Table 3.4.
Table 3.10: Design LoadUn-ballasted “lift in air” weight of turbine unit
Hooks, blocks & rigging Dynamic factor Design vertical load at boom tip
2363 tonnes 87 tonnes 1.05 2575 tonnes
For this comparative exercise transverse and longitudinal loading at the boom tip was assumed to be negligible (i.e. zero).
All three lift options were initially assumed to have dual fall systems i.e. two hooks with two sets of blocks, falls, and rigging. The lift points on the turbine unit were assumed to be 2 trunnions each located either side of the column with a mean offset of 2.5m off the centreline of the turbine unit.
For the determination of dynamic loads in the lift system the stiffness of the fall/sling combination (per hook) was based on using an 18-part sheave system using 60mm diameter wire with an elastic modulus of 70000 N/mm2. Slings were assumed to have similar stiffness properties.
Where there is a load imbalance due to dynamic roll moments acting on the dual fall lift system, the largest fall load is assumed to pass directly down one side of the lift structure, i.e. down one column of the A frame and down one stiffened plate structure for the horn arrangement. This is not considered unreasonable as the sheave blocks (see diagram below) are each located to one side of the boom tip. As the loads will oscillate from one fall to the other as the turbine unit rolls, the load used for comparison against the design load of 2575 tonnes will be taken as 2 times the greatest fall load, as each side of the lift structure needs sufficient steel to accommodate this dynamic load.
For all three options the arrangement at the boom tip will be roughly the same, and is depicted in Figure 3.6.
3.4.1.3 Design ParametersAll preliminary structural design of the lift system(s) was based on Chapter 3 of Lloyds Lifting Appliances Code [5].
Any preliminary barge modification design work was done in accordance with Lloyds ship classification rules [6], in particular Chapter 4 of Part 3.
For operation in sheltered water or short voyages higher permissible bending moments and shear forces are permitted based on reduced vertical wave bending moments and shear forces. Wave moments and shear forces can be reduced by the following reduction factors, but the reduction should never be greater than 50% in any event:
• Sheltered water : 50%• Short voyages : : 20%
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PLAN VIEW
Grey - Plated structure, continuing down for Horns
Orange - A-frame columns Figure 3.6: Structural arrangment for dual lift system
Short voyages are defined as voyages of limited duration in reasonable weather, which is categorised as seastates associated with a Beaufort wind scale of six or less. Sheltered water applies to sea areas where the fetch is limited to six nautical miles or less. Thus the following reduction factors will be applied:
• During Loadout Operations 50%• During transport or installation 20%
All lift system steel was assumed to be Grade 50, with a yield stress (oy) of 345 N/mm2.
All steel material for the barge (both existing and new) was assumed to be Grade 43, with a yield stress (oy) of 235 N/mm2.
Lloyds “Code for Lifting Appliances in a Marine Environment, January 1987” gives the following allowable stress criteria:Tension, Ob = 0.67 oyCompression, oc = 0.67 oyShear, t = 0.39 oyBearing ob = 0.67 oy
Von Mises Combined Stress = (OXX2 + Oy / +/" OxOy + 3^) <= 0.74 Oy
3.4.1.4 Structural CalculationsThe majority of the preliminary structural design was undertaken by hand, supported where necessary by Excel spreadsheets. Bending moments and shear forces acting on the barge were derived using the software MOSES (see below). The software balances the static weights acting on
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the barge (barge self weight, ballast water, lift system, and turbine unit) against the hydrostatic forces due the barge buoyancy and determines the resulting moments and shear forces.
The hull girder properties of the barge (section modulus, inertia, and cross-sectional shear force distribution) were determined by taking scantlings from a midship section of a typical 300’ x 90’ barge and entering them into a spreadsheet. Sectional properties were adjusted at the raked ends of the barge.
The calculations assume that where necessary the above deck structure supporting the lift system will transfer shear evenly into the longitudinal bulkheads and sideshell.
3.4.1.5 Motions AnalysisIn order to determine the hydrodynamic loads and motions acting on the turbine unit, the barge, and the lift system, a motions analysis was undertaken using the computer program MOSES (Marine and Offshore Structural Engineering Simulator), version 6.027, developed by Ultramarine Inc. of Houston.
The analysis requires the solution of a multi-body dynamic system using flexible connectors for the falls and slings. The submerged parts of the barge and turbine unit are modelled with 3-D diffraction mesh elements for the computation of the hydrostatic and hydrodynamic loadings. Drag plates are added to model the damping effect of the submerged base of the turbine unit.
The motion’s analysis is performed in the frequency domain. The barge’s and the turbine unit’s response to regular waves are determined from the hydrodynamic analysis and reported in the form of Response Amplitude Operators (RAO’s). An RAO is the response (motion or force) to a regular wave of 1m amplitude (i.e. A wave height of 2 m) - it is a measure of the dynamic amplification factor of a particular response function under wave action. RAO’s are examined for a range of periods typically 3 to 25 seconds in 1 second increments. Motion RAO’s, together with phasing are reported for 6 degrees of freedom (Surge, Sway, Yaw, Roll, Pitch, and Yaw).
RAO’s can give a quick insight into the behaviour of a floating body, and if the responses are directly proportional to wave height (i.e there is a linear relationship) then RAO’s can simply be multiplied by the wave height ratio to determine the response for a particular regular wave environment. In reality, the wave environment is likely to be confused and spread; if this is the case then a regular wave approach (RAO) will be conservative. Similarly roll damping is a function of wave steepness, and therefore one would expect roll response/wave height ratio to reduce with increasing wave height. Thus, an irregular wave analysis was undertaken to give more realistic results.
An irregular wave analysis is carried out by applying a wave climate in the form of an energy spectrum, defined by direction, significant wave height (Hs), and associated period (Tp or Tz). Normally for the fetch limited UK environs a Jonswap spectrum is most appropriate. By combining the RAO spectrum and the wave energy spectrum the statistical response can be determined. It is normal to base the maximum responses on a 3 hour exposure period.
Thus the irregular wave analysis derives maximum responses for the same matrix of loadcases that are used for the regular wave analysis. In addition to the maximum motion response, the maximum velocities, and acceleration are reported, together with maximum forces in the connectors (slings).
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The irregular sea analysis was run for a series of increasing seastates based on the Beaufort scale. Motions at the turbine unit centre of gravity, and forces in the slings were reported for each Beaufort number rising up to Beaufort 6. The analysis was carried out for head, quartering, and beam seas.
3.4.2 Technical Assessment - Horn Option3.4.2.1 Structural ConceptThe horn concept is depicted in Figure 3.4.
The horn is rigidly connected to the barge, and due to the large loads involved requires that the base of the structure mate with the structural bulkheads of the barge. It is therefore envisaged that the structure will be essentially a stiffened plate box section. It could be possible (to reduce weight) to design the middle section of the horn as a truss, but both ends will need to be plated to ensure shear transfer of sheave block loads (upper ends) and compression/tension forces into the barge bulkheads at the bottom end.
It is envisaged that the fabrication costs of integrating a truss section would be similar to fabricating the whole structure of stiffened plate - so the only advantage would be a slight reduction in weight.
The plan layout of the stern of a typical 300’ x 90’ barge, with bulkheads indicated as dotted lines, is as follows:
Figure 3.7: Plan Layout
The base of the horn will be approximately 13500 x 13500mm square to mate with the barge bulkheads. The horn will taper in both planes so that it is deep and wide enough to accommodate the sheave blocks at the upper end. Construction would be similar to that of a barge/ship structure and would suit a fabricator versed in shipbuilding techniques.
Preliminary plate thicknesses were determined from a global examination of combined bending and compression stresses. The structure as a whole was examined for buckling.Stiffeners and deeper transverse webs will be spaced to ensure that local buckling of plate does not occur.
The estimated steel weight of the horn structure is approximately 145 tonnes.
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3.4.2.2 Barge StrengthAs the horns are rigidly connected to the barge the moment induced by the overhung weight of the turbine unit will be transferred locally into the barge structure (as opposed to a sheerleg arrangement where some of the moment can be transferred via a tieback structure to the other end of the barge).
A longitudinal strength analysis has been carried out for the “horn” option, without any additional barge strengthening. The bending moments have been found to exceed the permissible by a considerable margin and the excedance extends to almost half the length of the barge, thus some considerable reinforcement of the barge is required to increase the section modulus of the barge.
The shear forces exceed the permissible only for a distance of about 7m at the stern. As the horns will comprise a plated structure that will extend forward to a transverse bulkhead at least 13.5m forward of the stern, it is expected that the horn structure itself will contribute to the barge’s shear area.
The estimated weight of barge reinforcement steel is 135 tonnes.
3.4.3 Technical Assessment - A Frame (Unit Hung-Off)3.4.3.1 Structural ConceptFigure 3.8 below depicts the expected arrangement with the turbine unit suspended from the boom tip of an “A” frame system mounted on the end of the barge:
Figure 3.8: A Frame (Unit Hung-Off)
An “A” frame is essentially a floating boom that is pinned at its base, and supported by tieback wires at its upper end. The frame can only rotate in elevation; its attitude being controlled by the length of the tiebacks. The tieback wires and the falls are independently controlled; usually by winches.
The construction of the A frame is assumed to be as per proprietary floating sheerlegs with two tubular or box section columns tapering from base to tip to form an “A”. A mid height cross bracing forms the of the “A”. At the top the columns will lead into a plated structure that supports the sheave blocks, and at the bottom of each leg the “hollow” section leads into a “plate” and “pin” hinge arrangement.
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The hinge arrangement will sit on top of an “I” beam grillage arrangement that will span the sideshell and the off-centre longitudinal bulkheads. The ends of the “I” beam will be supported by large wing plates that will direct loads into the sideshell and bulkhead. The schematics are given in Figure 3.9.
Figure 3.9: Hinge Schematic
The tieback structure will consist of a braced “goalpost” arrangement, fabricated of “hollow” section. The base of sections will sit on top of transverse bulkheads that will receive load via shearplates slotted into the “hollow” sections.
It has been assumed that columns and braces of the A frame and tieback structure do not see any moment i.e. that, in the case of the A frame, the moment resulting from the need to provide a horizontal extension to enable the blocks to hang directly over the turbine unit lift points is balanced by a counter moment resulting from correct positioning of the tieback connections.
The calculations indicate that some underdeck strengthening of the barge may be required in way of the wingplates placed over the barge longitudinal bulkheads. This will comprise additional stiffening to prevent localised buckling of the bulkhead plating.
The estimated steel weight of the A frame structure including support grillage and backstay structure is approximately 230 tonnes, of which 5 tonnes includes underdeck strengthening.
3.4.3.2 Barge StrengthWith a sheerleg arrangement some of the moment can be transferred via backstay wires to a support structure at the other end of the barge; this has the effect of reducing the hogging moment, which a weight overhanging the barge end(s) causes.
Bending moments fall within permitted boundries. As for the horns the shear forces exceed the permissible for a distance of about 7m at the stern. It is expected that the required barge shear material can be easily incorporated into “A frame” support structure.
The steel weight estimate given above includes some allowance for additional barge shear material.
3.4.4 Technical Assessment - A Frame (Unit on Deck)3.4.4.1 Structure ConceptFigure 3.10 depicts the expected arrangement with the turbine unit suspended from the boom tip of an “A” frame system mounted on the end of the barge:
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Figure 3.10: A Frame (Unit on Deck)
In this case the turbine unit is loaded out directly from the quayside on to the deck of barge, where it will be seafastened. This could provide a major advantage in that transportation could be undertaken in more arduous seastates, and could enable the barge to wait offshore for a reasonable installation window. Previous analysis for ctr “g” which covered criteria indicated that if the turbine base was structurally capable, the unit would just need to land on timber packing for vertical forces; roll and pitch stops would provide horizontal restraint.
The disadvantage is that the A frame needs to be set back to accommodate the unit on deck and thus the boom length needs to be longer to enable sufficient outreach for loadout. To accommodate both positions with a lift height of 21m would require boom angles of about 29degrees (loadout) and 60 degrees (placement on deck). The former angle is too low and results in excessive tie-back loads, so on the basis that quantity usually reduces cost it was decided to ensure that the tie-back wires took the same design load as the falls i.e. approximately 2500 tonnes. This ensures common sizes for sheave blocks, winches, and wire. This equated to a loadout angle of 60 degrees that results in a boom length of about 75m and a vertical lift height of about 64 m, compared with 21m for the other two options.
The structural concept is identical to that of the stern mounted A frame. The weight of required steel increases to accommodate the longer boom and the extra loads into the support grillage. A similar amount of underdeck strengthening will be required, but there is no requirement for extra shear material for the barge (see below).
The estimated steel weight of the A frame structure including support grillage and backstay structure is approximately 360 tonnes, of which 5 tonnes includes underdeck strengthening.
3.4.5 Environmental Limitations of Dual Lift Systems
When comparing environmental limits there are two modes to consider, transportation and installation. When considering transportation, weather from any heading is applicable; for installation it is standard practice to align the installation vessel into the prevailing weather so only the head seas case is applicable .
The preliminary study indicated that the transportation limits for the unit on deck could be in excess of Beaufort 6 using only a small amount of steel for seafastening only.
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Motions analyses were run for 3 different cases in order to determine loads in the lift slings. The results are presented in the graphs below; the benchmark is the 2575 tonne vertical load used to size the lift systems.
Case a) Unit submerged - 8.4m overhangCovers transportation and installation for the “seabed loadout ” option horn and stern mounted A frame. The 8.4m overhang was determined from the preliminary study where a minimum clearance of 2m between unit and barge was set.
10000
Hs (m)
Figure 3.11: Submerged Lift - 8.4m Overhang
—♦—Beam —■—Quarter
A Head Benchmark
Case b) Unit submerged - 14.0m overhangCovers transportation and installation for the horn and stern mounted A frame “Loadout from Quay” options, and the submerged part of the installation for the “unit on deck” option (Fig 3.12).
Figure 3.12: Submerged Lift - 8.4m Overhang
♦ Beam —■—Quarter —A— Head
Bench mark
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Case c) Unit in airCovers the installation case for the unit on deck option prior to submergence (Figure 3.13).
6000
5000
4000
3000
2000
Figure 3.13: Lift in Air
—♦—Beam —■—Quarter —A— Head
^^^Benchmark
It can be seen that beam and quartering seas are very limiting in all three cases (under 0.5m Hs) with the submerged cases producing larger loads over the full range of Hs. To be able to transport the turbine suspended (either in water or in air) in Beaufort 6 weather conditions would be prohibitive in terms of extra weight and cost.
The head sea case gives an acceptable installation seastate of about 1.25m Hs for all three cases; thus there appears to be no advantage gained from utilising the submerged buoyancy of the unit. Although having a shorter overhang (8.4m vs. 14 m) resulting from a seabed loadout should improve the bending moment capability of the barge (and therefore less reinforcement for the horn case), it does not improve the installation seastate.
3.4.6 Single Point Suspension Option
The forgoing study indicates that the two-point suspension system results in high sling loads in quartering and beam seas, which severely restricts the seastate for load-out and transportation. For this reason it was decided to carry out a comparative motion analysis for a single point suspension system.
Case a), submerged - 8.4m overhang, was re-run with a single point lift system. For this comparative exercise, the lift point was taken to be on the centreline of the 4m diameter tube.
In reality the lift point would have to be at least 3m off the centreline (to avoid contact of the lifting gear with the unit) which means that the centre of gravity of the unit would have to be shifted to ensure that the unit stays almost vertical during installation. It is unlikely that much “out of verticality” would be tolerated during the set-down phase.
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Lift off at loadout quay would also have to be nearly vertical to ensure adequate clearance and to avoid setting up a pivot loads on one edge of the base. As the unit started to submerge the buoyancy uplift will change the attitude of the unit so selective ballasting may be required during submergence.
Figure 3.14 shows the results for 8.4m submerged case with a single point lift
This indicates that although the head sea cases is still limited to just over lm Hs there is a marked improvement in beam and quartering sea capability.
On a note of caution - In the 3m Hs beam sea case the transverse load acting on the boom tip was found to be 1275 tonnes. This would be acting at a height of 21m which would result in an additional vertical load acting down one plated side (Horn) or strut ( A frame), for a 13.5m span, of 1983 tonnes. Based on the premise propounded in Section 3.4.1.2 the actual vertical load in the boom to be used for comparison should be around 6540 tonnes which is not that dissimilar to that achieved for the two point system.
Thus one could conclude that the although direct vertical loads are reduced using a single point system, the moment induced into the horns by transverse loads will require a lift structure design of similar structural capacity (and hence steelweight) to that of a dual point system. A single lift point on the turbine unit is likely to require increased stiffening of the column. In addition, the requirement to have an offset lift point (to accommodate larger hook dimensions) is likely to induce a considerable moment into the turbine unit. A progressive ballasting system is also likely to be required to keep the unit vertical as the unit adopts different attitudes in and out of the water. Single point suspended systems tend to “tramp” more than dual point systems thus requiring more motion controls in the form of bumpers and tugger wires to avoid impact with the barge and lift systems.:
SUBMERGED LIFT (8.4 M OVERHANG)SINGLE POINT LIFT
4000 n
2 2500
< 2000
O 1500
♦ S.P Beam —S—S.P Quarter
S.P Head Benchmark
Hs (m)
Figure 3.14: Submerged Lift (Single Point Lift) - 8.4m Overhang
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3.4.7 Cantilever / Strand Jack Concept
An alternative option to improve the weather window whilst fulfilling the loadout and set-down criteria is to utilise a cantilever and strand-jack concept
3.4.7.1 Structural ConceptThe basic proposed concept is shown in Figure 3.15.
Figure 3.15: Cantilever / Strand Jack Concept
The support for the unit would be provided by 4 x 600 tonne strand jacks. The main body of the jack would be mounted on top of a cantilever beam overhanging the stern of the barge. The strands would pass through the beam to an end link, which would be pin connected to the base of the turbine unit.
The procedure would be as follows:
• Loadout: The barge would be manoeuvred such that the beams overhung the quayside. The end link for each jack would be lowered, such that each link could be connected into the unit base housing. The strand jacks would then lift the unit until the top of the unit base mates with the underside of the cantilever beams.
• Transportation: It is envisaged that roll and pitch shear stops located on the unit and cantilever beams would engage during the final strokes of the mating operation, and provide horizontal restraint. Once mated the strand jacks would be locked off, providing vertical restraint during transportation. It is expected that transportation in Beaufort 6 sea conditions would be possible.
• Installation: Installation would be the reverse of loadout. The strand jacks would lower the turbine unit onto the seabed. Due to the slow movement of the jacks, ~ 1.5 m/hour, the wires will receive considerable dynamic load due to barge and turbine movement. From discussions with strand jacks suppliers heave compensation devices can be fitted to each jack to take out dynamic movement. Once on the seabed the wires will be slackened until the endlinks drop out of the pins. Recovery of the endlinks would be by pre-connected tugger lines. Manufacturers feel confident that a strand jack system could install this turbine unit in reasonable weather conditions (vessel heading into weather).
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The same level of barge reinforcement would be required as for the conventional horn option. It is expected that the steel weight of the lift structure would be about 25% greater than the conventional horn due to the need for two cantilevers.
3.4.7.2 Barge ConversionThe global conversion of a barge to a horns and single point lift/strandjack system is similar to the horns and two point lifting wire/winch arrangement.
With regard to the strandjack system, the specific differences are the deletion of the lifting hooks, sheaves wires winches and power systems and the cost saving this implies. With the single point fall system, it is expected that lift system costs will be similar to a dual point system. The strandjack system requires the horns to be fitted with saddles or securing points for the strandjacks and wire recovery reels. The strand jacks are assumed to be a rental item and included in the operational costing.
3.4.7.3 OperationThe lifting and lowering operation using strand jacks will take a similar time to the wire and winch system, the difference is that it will include a four point lift from the base rather than a two point lift from a point above the centre of gravity.
Because the strandjacks are a rental item they are included in the operational costing with elements for mounting and commissioning, seasonal hire, dismounting and demobilisation.
3.4.8 Costing
3.4.8.1 Cost ModelsTwo cost models have been developed, firstly for the barge conversion requirements and secondly for the operational costs. These have been developed for the four candidate systems assuming the locked port and longer transit time. The cost models have been based on a common format with items and activities deleted from the model to reflect differences in the requirements or operations.
All costs are shown as pounds sterling and based on prices in 2002 either drawn from in-house data or budget discussions with manufacturers and suppliers. No attempt has been made to drive
suppliers budget cost estimates downwards, and this means that better deals may be struck in a live project. High cost contributor items, such as tug dayrates, have been taken at market rates and lower cost deals could probably be made taking into account the long period of hire; additionally cheaper tugs, such as Eastern block flag could be hired at lower dayrate though with an increased
supervision cost.
It is important to appreciate that the costs have been applied consistently to the different options, so the comparative differences in system cost will be demonstrated.
The models have been developed with minimum, likely and maximum costs in @Risk to allow Monte Carlo sampling of the costs in a 1000 iteration simulation and provide a distribution of results. The present models allow sampling of each of the inputs rather than the subtotals of conversion or operational blocks. Our experience is that sampling of all items results in a closer distribution of results than when working with larger blocks of subtotal costs, and this should be borne in mind when reviewing results.
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Table 3.11: Cost ComparisonSystem Horns
£ millionA frame aft
£ millionA frame mid deck
£ millionCantilever/StrandJack
£ millionConversion costMin 3.3 4.9 5.5 1.9Max 4.2 5.9 7.1 2.7Mean 3.7 5.4 6.2 2.2
Operational costMin 6.5 6.5 5.5 6.8Max 8.3 8.3 7.0 8.9Mean 7.4 7.4 6.2 7.7
Installation costsMean conversion cost 3.7 5.4 6.2 2.7Mean operational cost 7.4 7.4 6.2 7.7Mean installation cost 11.1 12.8 12.4 10.4
Total 30 unitsCost per unit 0.37 0.43 0.41 0.35
Number of days 215 215 186 215Nominal dayrate 52,000 60,000 67,000 48,000
The cantilever / strand jack system is the cheapest conversion cost because of its simplicity, the robust design and construction of the cantilevers compared to the A frame and the elimination of expensive winch equipment (the jacks are considered to be an operational item). The A frame aft is
cheaper than placing the A frame mid deck because of the physically smaller A frame and associated support and topping systems.
The stand-jacks, horns and A frame aft have similar operational costs and the A frame mid deck operational cost is less because of the improved uptime and hence shorter campaign period.
The installation costs are cheapest for the strand jack system because of the lower conversion cost, while the A frame aft system gives the highest total installation cost because of the relatively higher conversion cost than the horns plus the penalty of weather downtime. The A frame at mid deck has the highest conversion cost but the weather uptime advantages lead to the least operational cost and the mid range overall cost.
For comparison a simple costing has been made for a large sheerleg vessel with an installation
spread, resulting in a dayrate of about £77,000 per day. For an installation campaign such a vessel would not have the long set-up and clean-off period and a nominal 150 days hire time has been assumed. This results in an installation cost of £11.5 million, or a cost of £0.39 million per unit. When compared with the cost of installation by the barge systems, then hiring in a ready made sheerlegs unit is about the same cost per unit installed. This suggests there may be little cost advantage in using the converted barge for a single season, though there would be a cost advantage over two or more seasons.
This line of investigation of sheerleg comparison has not been addressed in detail, either for operational requirements or dayrate, and it is important to consider whether an existing sheerleg vessel can lift the proposed fully outfitted turbine unit in terms of lift geometry and crane hook location.
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3.4.9 Scoring
3.4.9.1 Scoring ModelThe three systems have been scored on a weighted model that incorporates technical capability, operational requirements and flexibility and commercial implications.
The model includes the several technical requirements of the barge structure, conversion requirements and geometry; the lifting systems are characterised by the nature of the support structure, the proven ancillary systems and the winch and power requirements. The operational requirements include the loadout operation, seafastening, routine and contingency transport loads, system uptime assessment and the simplicity and flexibility of the system. The commercial aspects address the relative conversion and operational costs.
Professional judgement has been used with the analyses presented in the sections above to score each aspect on a scale of 1-3 and to weight each aspect on a scale of 1-3 (Table 3.12)
Table 3.12: Scoring BasisScoring
Technical1 Capable with reservation2 Capable with tight margins or major investment3 Capable though needing design verification
Commercial 1 Market cost2 Conversion or operational benefits
Weighting1 Routine2 Major importance3 Critical significance
3.4.9.2 Scoring ResultsThe summary of the scoring reduced to a percentage of the highest possible score is given in Table 3.13.
Table 3.13: Scoring ResultsWeighted score as a percentage of total possible
Horns (Dual Fall) Horns (Single Fall) A frame aft A frame mid deck Cantilever / Strand Jack
57% 52% 49% 75% 64%
• The A frame aft is the least attractive option because of the relatively high conversion cost, downtime penalty and operational cost.
• The horns has benefits of lower conversion cost and simplicity but has downtime penalties and hence additional cost.
• The cantilever / strand jack option combines the low conversion cost and simplicity of the horn option with improved operational characteristics.
• The A frame mid-deck with the unit transported on deck scores highest overall.
3.4.10 Conclusions of Option Development
• The “A frame at mid deck” scheme offers the most attractive option for unit installation based on the scoring model, however it is felt that this is comparable with a standard commercial sheerleg working with a dumb barge.
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• The horns option, having a simpler and cheaper conversion requirement in the basic layout, should be further investigated to assess whether by changing the lifting mechanism to a strand-jacking system and/or changing the unit lift geometry the installation operations and costs could be reduced and the weather workability could be increased.
• The horns and strand jack system will be taken through to the conceptual design stage.
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3.5 Conceptual Design of Preferred Solution
3.5.1 Basic Concept
This Section details the conceptual design of a horn/strand jack system as shown diagrammatically in Figure 3.15.
The concept intent is to use a strand jack system supported by cantilever beams overhung the stern of a barge, to enable the following:
• Pick up of a turbine unit from a quayside (loadout) by connecting the cable anchors to the base of the unit, and raising the unit to mate with the underside of the cantilever beams.
• Transport of a unit secured to the underside of cantilever beams (by strand jacks) from loadout port to installation site. Additional structural stops will be required to seafasten against transverse loads.
• Install a unit in water depths ranging from 10m to 25m by lowering on the strand jacks until the base is fully placed on the seabed.
• Utlise temporary fittings and systems that are fitted above barge deck and removable at the end of an installation season.
3.5.2 Design Parameters
The findings from the first two phases of the study determined that the conceptual design had to meet the following design parameters to be a viable alternative to existing methods of installation:
Barge Type and Size:Installation water depth range: Minimum Overhang:Transportation seastate:Installation seastate:Quay Loadout weight:Max' allowable level barge draught: Minimum transit time (port to field) Maximum transit time (port to field)
Standard 300' x 90' x 20' barge 10 to 25m 14m from stern Beaufort 6 (minimum) Beaufort 3 (minimum)2363 tonnes (no water fill) 4.85m (loadline)4 hours @ 5 knots (example) 12 hours @ 5 knots (example)
Design Standards and steelwork design parameters are as described in Section 3.4.1.3.
For the design concept the following was considered:
• The global and local strength of the barge• The ballast capacity of the barge• Sizes of the supporting steelwork sufficient to estimate a price
3.5.3 Hydrodynamic Analysis 3.5.3.1 MethodologyThe previous phases identified that the concept needs to be able to transport and install the units in Beaufort 6 and Beaufort 3 weather conditions respectively to be a viable alternative to existing
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installation vessels. A hydrodynamic analysis needs to be undertaken to determine these weather limits.
In order to determine the hydrodynamic loads and motions acting on the turbine unit, the barge, and the lift system, a motions analysis was undertaken using the computer program MOSES (Marine and Offshore Structural Engineering Simulator), version 6.027, developed by Ultramarine Inc. of Houston.
Where the response is expected to be non-linear, i.e. in the case when the turbine base is touching down, then time domain analysis needs to be run in order to account for the time varying stiffness of the touchdown scenario.
3.5.3.2 Transportation AnalysisAn irregular sea analysis was run for a series of increasing seastates based on the Beaufort scale. Motions at the turbine unit centre of gravity, and forces in the connectors were reported for each Beaufort number rising up to Beaufort 6. The analysis was carried out for head, quartering, and beam seas.
For the 1m seastate the sea was assumed to be almost unidirectional (i.e. very little spreading) reflecting the likelihood of swell content at this small wave height. All other seastates with significant wave heights greater than 1.0m were assumed to be fully developed and confused i.e. well spread. It should be noted that spread seas produce roll motions in head seas.
Motions and accelerations are calculated for the turbine unit centre of gravity. The maximum values, for Beaufort 6 conditions, are presented in the tables below.
Table 3.14: Transportation MotionsMOTIONS Surge(m) Sway (m) Heave(m) Roll (degs) Pitch (degs) Yaw (degs)Beam 0.00 2.02 4.50 14.76 0.00 0.00Head-Qtr 1.35 2.40 5.12 11.69 0.85 0.07Head 1.57 1.88 5.65 7.44 0.98 0.04
Table 3.15: Transportation AccelerationsACC'l'Ns Surge (m/s2) Sway (m/s2) Heave (m/s2) Roll
(degs/ s2)Pitch(degs/ s2)
Yaw (degs/ s2)
Beam 0.10 1.02 2.30 7.66 0.10 0.02Head-Qtr 0.30 0.90 2.40 6.05 0.21 0.04Head 0.35 0.56 2.53 3.81 0.24 0.02
The maximum forces, for Beaufort 6 conditions, into either cantilever beam are presented in Table 3.16.
Table 3.16: Transportation ForcesForces Static
Vertical(tonnes)
DynamicVertical(tonnes)
MaximumVertical(tonnes)
MinimumVertical(tonnes)
Transverse(tonnes)
Longitudinal(tonnes)
Beam 1182 708 1890 473 340 14Head-Qtr 1182 607 1789 574 262 17Head 1182 518 1670 663 145 20
The Safe Working Load (SWL) of the proposed strand jack units is 600 tonnes, so if 3 jacks are used per beam the highest utilisation is (1890 x 100)/(3 x 600) = 105%. Design codes normally allow an overstress (typically 1/3rd on allowable) where the maximum response is an extreme
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event, as it is in this case, especially as the barge will only sail on a forecast of Beaufort 5 decreasing.
The forces indicate that the minimum vertical force is substantially positive i.e. there is no apparent risk of uplift and slackening of the tieback cables (that could result in snatch loading).
The above forces will be used to size the supporting steelwork.
3.5.3.3 Slamming AnalysisAs the unit is hanging a considerable distance (26m to its extreme end) over the stern of the barge, there will be a strong possibility of slamming occurring on the underside of the turbine unit during transportation. The design height of the cantilever beams above the static waterline needs to be a compromise between an acceptable level of slamming occurrence and the additional steel required to support the raised beams.
As part of the transportation analysis heave RAOs were determined for a point at the extreme end of the underside of the turbine base, offset to one side by 4m (an apex of the hexagon base). These RAOs (and their phasing) were then combined, in a spreadsheet, with the predicted elevation of unit amplitude regular waves (effectively wave RAOs), calculated for a range of periods, at the same global plan position beneath the turbine base. The difference between base heave RAOs and the wave RAOs at the point in question gives the clearance RAOs. By integrating these RAOs with a wave energy spectrum, clearance statistics can be determined.
It is expected that stern seas would be more likely to cause slamming so additional analyses were
run for stern and stern quartering seas.
The static clearance height (between underside of point on base and still waterline) was adjusted until an acceptable level of slamming was achieved. Thus for static clearance height of 4m the following slamming statistics were obtained for Beaufort 6 sea conditions ( 3m Hs & 8 s Tp):
Table 3.17: Slamming StatisticsStern Stern-Qtr Beam Head-Qtr Head
MaxSubmergence(m)
1.35 3.3 -0.98 >040 1.25
No of Slams/3 hrs
36 218 0 2 31
Significant Slam Velocity (m/s)
2.23 3.09 0.00 1.86 2.25
The number of slam occurrences in stern quartering seas reduces to 60 slams/3 hours in Beaufort 5 conditions, and almost to zero for all other weather headings. Bearing in mind that a condition of
the barge sailing will be a weather forecast of Beaufort 5 decreasing with an indication that the installation weather criteria (Beaufort 3) is predicted within the 3 day forecast, then the predicted level of slamming is considered acceptable.
3.5.3.4 Installation Analysis (Prior to Touchdown)As for the transportation analysis an irregular sea analysis was run for a series of increasing seastates based on the Beaufort scale, starting with Beaufort 1 and ending in Beaufort 3. For this intermediate installation case the turbine unit was assumed to be immersed to a draught of 7 m.
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Based on the transportation analysis it was assumed that the turbine unit would be suspended from 6 strand jacks, 3 per cantilever beam.
It was assumed that during installation the barge would always be moored head to the weather so the analysis was carried out for head sea, and seas quartering 22.5° off the bow (to account for some variation in wave direction), assuming uni-directional seas.
The 600 t strand jacks support 38 x 18mm diameter "dyform" steel cables with an elastic modulus of 195,000 N/mm2. For modelling purposes the 38 cables were grouped into one cable with an equivalent single diameter to give the same stiffness properties.
Motions and accelerations are calculated for the turbine unit centre of gravity. The maximum values, for Beaufort 3 conditions, are presented in Tables 3.18 and 3.19 below.
Table 3.18: Installation MotionsMOTIONS Surge(m) Sway (m) Heave(m) Roll (degs) Pitch (degs) Yaw (degs)Head±22.5° 1.69 0.24 3.76 0.71 3.53 1.82Head 1.84 0.10 3.81 0.29 3.57 0.73
Table 3.19: Installation AccelerationsACC'L'Ns Surge (m/s2) Sway (m/s2) Heave (m/s2) Roll
(degs/ s2)Pitch(degs/ s2)
Yaw (degs/ s2)
Head±22.5° 1.38 0.08 0.97 0.37 0.94 1.64Head 1.50 0.03 0.98 0.15 0.95 0.66
The maximum forces, for Beaufort 3 conditions, into any strand jack cable are given in Table 3.20
Table 3.20: Installation ForcesForces Static (tonnes) Dynamic (tonnes) Maximum (tonnes) Minimum (tonnes)Head-22.5° 310 269 579 41Head 310 270 580 40
Thus the maximum utilisation on any strand jack is 97%; and the minimum load is still positive indicating that snatch loads are not likely to occur.
Thus the design of any anchor point on the turbine base should be based on the SWL of the jacks which is 600 tonnes.
3.5.3.5 Installation Analysis (at Touchdown)In order to assess the non-linear effects of bottom contact during touchdown, frequency and time domain analysis was run. The primary concern of this analysis is to assess "snatch" loads arising from loss of tension in the cables as the base grounds, rather than base/seabed impact loads.
Conservatively the seabed was assumed to be rigid, and was modelled as a "gap" element through which the turbine unit cannot pass. The turbine base/seabed contact loads were not calculated as these are expected to be small due to the large added mass and damping effect resulting from the "squeeze" of fluid between the base underside and the seabed. Similarly it is expected that the seabed surface will be relatively soft.
The analysis was carried out for head sea, and seas quartering 22.5° off the bow, assuming unidirectional seas.
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The maximum lowering speed of the strand jacks is 20 m/hour, and it is expected that it will be during the final 2m of lowering that bottom impact will occur, so the time taken to lower to the seabed could be (60 x 2/20) x 2 = 12 minutes if the ballast system could keep up. However, some of the 300’ x 90’ barges only have a ballast capacity of about 600 tonnes/hour.
MOSES indicates that the nett submerged weight of the unit just prior to setdown (in 22m of water) is 1680 tonnes. Thus the trimming moment due to the unit (about midships) is about 100,000 tm. The distance between the centres of tanks 1 (bow) and tanks 5 (stem) is about 76.4 m. Thus about 1300 tonnes of ballast needs to be transferred to keep the barge level i.e. About 2 hours of ballasting time . It would probably be advisable to commence installation with a stern trim (using water in no. 4 tanks) to reduce the amount of ballasting time required.It is expected that by a combination of correct ballasting and continual lowering of the cables that the final 2m sequence could be achieved in 1 hour. Thus the time domain analysis was run for 1 hour real time.
The unit was analysed for a number of positions just prior to set down and also for a case where about 250 tonnes (static) was taken by the seabed. In both cases the turbine unit appeared to act as a fixed vertical mooring which caused the barge to rotate about the unit rather than midships. Thus heave of the turbine unit appeared to be minimal, and there were no apparent non-linearities in the load response, an indicator of the presence of snatch loads. None of the reported cable loads was greater than the design load of 600 tonnes.
It is expected that there would be some snatch loading when sufficient weight was taken off the end of the barge to cause the point of pitch rotation to move forward towards midships. To ascertain this would require another separate analysis with the barge effectively moored to the turbine unit - this is beyond the scope of work of this study.
The cantilever system has been designed so that the forward end of the beam and the weight of the tiebacks provide a counterweight about the stern hinge point so that as soon as the tension in the lowering jacks reduces it is immediately compensated for by the “counter” tension/weight. The system cannot overrun (i.e. the towers fall backwards) because the forward end of the beams will stop further movement when it makes contact with its support grillage.
Lowering Strandjacks to articulate to accommodate articulation of beams
Back Tension in Stays providing counterweight
Transportation Tieplates cut for installation
Figure 3.16: Tied-back Cantilever
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3.5.3.6 ConclusionIt was originally envisaged that the horn/strandjack system would take the simple form of two cantilever beams suspended over the stern of the barge and tied into the barge deck structure (see Figure 3.16). Although the installation options study showed that the horn/falls system was severely weather limited it was hoped that by suspending from the base of the unit across a width of 22m rather than across the diameter of the main tube (~ 5 m) that the vertical loads resulting from dynamic overturning moments would be so reduced as to make this simple cantilever concept viable in the Beaufort 6 transit conditions.
However the hydrodynamic analysis shows that the barge is subject to considerable roll motion (~ 158) in beam seas BF 6 sea conditions, more than one would normally expect for a conventional barge transport. The large rotational inertia of the turbine unit appears to have a considerable influence on natural roll period and the support loads induced by the rotational acceleration.
The transportation analysis shows that the maximum load into either cantilever is around 1900 tonnes. The installation options study showed that the shear forces and bending moments induced in the hull girder by the 2575t “nominal design” load for the horn/falls system could be accommodated by mating the horn structure with the longitudinal bulkheads/sideshell and transverse bulkheads/stern shell, and adding additional deep “T” beams to improve the bending moment capability. The proposed conceptual structural design was able to distribute load across the full width of the barge girder by external reinforcement.
The 1900 tonnes generated by the BF 6 transport is for one cantilever only and therefore can only be effectively distributed across one side of the barge. Thus typically if the main supports for each cantilever were at the stern and at a transverse bulkhead 13.5m forward the compression load into the half the stern plating would be about 3800 tonnes and half the underdeck welds at bulkhead 13.5 would need to take 1900 tonnes in tension. The stern shell/bulkheads would also be subject to torsional stresses induced by the dynamic loading. Rudimentary stress checks indicated that the existing hull girder was grossly overstressed and would require a large amount of internal and external reinforcement. The requirement to provide extensive internal reinforcement goes against the concept of a “bolt-on” arrangement, and would probably require the barge to be dry-docked. Preliminary structural checks indicated that the size of the cantilever beams would also have to be very substantial to accommodate these large loads and to provide torsional rigidity.
The “bolt-on” restriction and the requirement to transport in Beaufort 6 conditions meant a re-think on the structural concept, as described in the following Section.
It is strongly recommended that if the concept is taken further that model tests be carried out to confirm the motions and loads obtained.
3.5.4 Structural Design3.5.4.1 IntroductionThe arrangement used to support the turbine unit during transportation and installation is shown in drawing No. GM-44339-01 (Appendix C). The cantilever and compression tower arrangement can be regarded as a hybrid of horns and "A" frame concepts.
The normal expectation of horns is that of a substantial steel structure overhanging the bow or stern of a vessel. The base of the horns would be expected to marry straight into the hull of the vessel.
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However this produces considerable direct bending moment into the hull. Previous work during phases 1 and 2 showed that the proposed structural concept for the horns was limited to transportation in seastates less than Beaufort 3 in spite of including some 135 tonnes of barge reenforcement to take the induced bending moment, and shear forces. Thus in order to transport in Beaufort 6 sea states the re-enforcement is expected to be prohibitive, and it is expected that the barge would have to be dry-docked to basically have a new stern end put on. This extensive hull modification is not considered in keeping with the study premise of designing a “bolt-on” system.
"A" frames are guyed back to a tieback structure located at the other end of the lift vessel thereby reducing the moment into the hull girder. The "A" frame system requires the use of sheaves, blocks, wires, and winches to control the lift and attitude of the frame (as discussed in previous reports).
The alternative of using strand jack cables as tiebacks to support two cantilever beams which in turn support the turbine unit via strand jacks, was hoped would achieve a cheaper and simpler arrangement. It was expected that the proposed support arrangement would provide the structural efficiency similar to that of a suspension bridge (albeit a cantilever).
Original concepts were to pin the cantilever beams at the stern of the barge and have all the vertical load taken by the tie backs. However there were some concerns that that having two pinned beams capable of moving independently of each other might cause torsion problems even though the turbine base itself would provide some rigidity in the transport mode. The other concern was that in such a dynamic environment as the sea (as opposed to a suspension bridge) snatch loading might occur in the cables, which would be highly undesirable. This would still be the preferred option but the torsional analysis is beyond the scope of this study, and therefore a more rigid connection system has been considered.
The revised option was to carry the cantilever beams further forward and provide two supports, one at the stern and one at the next bulkhead position forward. The majority of the load would still be taken by the cables but by not having the beams free to rotate the problem of torsion and snatch loading should be eliminated. This is discussed later.
One of the features of strand jacks is that they can only lower at 20 m/hr , which means that there is a strong likelihood of “snatch” loadings occurring as the turbine unit settles on the seabed. With this in mind it is the intention, on arrival at the installation site, to release the forward end of the beams and lower the tiebacks so that the beams are now hinged. Any slackening of tension in the lowering cables will be “countered” by the weight of the forward tiebacks and the forward end of the beam. See Figure 3.16.
3.5.4.2 Barge StrengthBending moments and shear forces acting on the barge were derived using the software MOSES (see above). The software balances the static weights acting on the barge (barge self weight, ballast water, lift system, and turbine unit) against the hydrostatic forces due the barge buoyancy and determines the resulting moments and shear forces.
The hull girder properties of the barge (section modulus, inertia, and cross-sectional shear force distribution) were determined by taking scantlings from a midship section of a typical 300’ x 90’ barge and entering them into a spreadsheet. Sectional properties were adjusted at the raked ends of the barge.
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The calculations assume that where necessary the above deck structure supporting the lift system will transfer shear evenly into the longitudinal bulkheads and sideshell.
Figures 3.17 to 3.20 below show the results of the longitudinal strength analysis. The first set of graphs are the results for the still water equilibrium condition when the cantilever beams are carrying the 2363 tonne static weight of the turbine unit. The second set of graphs shows the results for the "quasi" dynamic case, which includes the dynamic load (applied as an additional static load) for the transportation case in head seas, Beaufort 6 conditions.
The upper and lower level pairs of curves indicate the permissible moments for “hog” and “sag” respectively. “Sheltered” and “Short” refer to sheltered water applications (e g. Loadout) and short voyages (transportation and installation) as defined in section 2.2. “SW” refers to still water i.e. Static loadcases.
PERMISSIBLE SW BENDING MOMENTS STRAND JACKS - STATIC
LU5O5(3
QZLUCO
600000
400000
200000
mo on an on
-400000
-600000
POSITION ON BARGE FROM AFT (m)
------ SHELTERED------ SHORT------ SHELTERED------ SHORT^—STATIC
Figure 3.17: Permissible SW Bending Moments - Static
PERMISSIBLE SHEAR FORCES STRAND JACKS - STATIC
„ 40000 g 30000w 20000£o
100000
< -20000 g: -30000
-40000
DC -100000.-66---- 20t6& 66—60.00
-SHELTERED
-SHORT
-SHELTERED -SHORT
• STATIC
POSITION ON BARGE FROM AFT (m)
Figure 3.18: Permissible SW Shear Forces - Static
"Quasi" dynamic results are given in Tables 3.19 and 3.20
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HImomozQZIIIm
PERMISSIBLE SW BENDING MOMENTS STRAND JACKS - DYNAMIC
•SHELTERED
• SHORT
■SHELTERED
• SHORT
■DYNAMIC
Figure 3.19: Permissible SW Bending Moments - Dynamic
PERMISSIBLE SHEAR FORCES STRANDJACKS - DYNAMIC
40000SHELTERED
20000 SHORTSHELTEREDSHORT20.00 60.00 80.00< -20000
-------DYNAMIC-40000
POSITION ON BARGE FROM AFT (m)
Figure 3.20: Permissible SW Bending Shear Forces - Dynamic
It can be seen that both the static and "quasi" dynamic analysis indicate that the bending moments and shear forces fall well within permissible values, and therefore no additional global strengthening of the barge is required with this concept.
3.5.4.3 Cantilever BeamsThe cantilever beams will comprise two box section girders, each positioned approximately 1 lm off the centreline of the barge. At the aft end, each beam will house 3 lowering strand jacks; these will be directly supported by 6 tieback cables that connect into the beams via double eyes mounted either side of each lowering jack.
Each beam runs forward to a hinge at the stern of the vessel, and thence onto a support 14m forward of the barge stern.
In the static case, the 2363 tonne weight of the turbine unit is taken directly off the lowering strand jacks into the tieback cables. The only structural consideration for the beams at this stage is the large axial load induced by the lead of the tieback cables.
However in the dynamic case, particularly in beam seas there is a load imbalance between the two beams of +/- 708 tonnes. The tieback cables are relatively soft (compared with a barge deck) and this results in large relative deflections of the two beams (one beam deflecting downwards and the
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other deflecting upwards), which are likely to induce racking of the turbine unit. Thus, the supports need to be stiffened to minimise this effect.
In order to size the additional steelwork required for torsional stiffening a maximum 50mm difference in deflection (between the two beams) across 22000mm was set as criteria.
The proposed way of doing this is to utilise the stiffness of beams. During the loadout operation, the unit will be lifted to mate with the underside of the beams. The tieback jacks will then shorten the tieback cables so that the forward end of each beam just touches its support grillage. Tensions in the tieback jacks will be equalised and recorded. A tie connection (see figure 3.1) will then be made between the forward end of the beam and its grillage support - this can take the form of welded strap plates, or possibly bolts. Thus the end of the beam is now fixed and any deflection due to any additional load fluctuations will now be influenced by the beam stiffness.
It is bending stiffness that drives the size of the beams (rather than bending moment or shear capability). This requires a large inertia value, and is best served by a truss arrangement connected into the original “box” section which houses the lowering jacks. Linking the two trusses transversely will provide additional torsional rigidity.
Figure 3.21: Braced Cantilever Structure
Unfortunately due to the large lever arm the truss arrangement has to be substantial in order to meet the previously stipulated deflection tolerance.
The total predicted weight of the cantilever truss arrangement (including barge support grillage) is about 572 tonnes. This could be considerably reduced if the deflection tolerance could be increased. It may be a cheaper option to go for more or stiffer tieback cables.
Calculations indicate that the maximum and minimum loads into the barge deck due to the additional loads taken by the beams are 945 and -1400 tonnes at the stern hinge point, and 708 and -467 tonnes at the forward end of the beam. These loads are within acceptable global and local limits for the barge (assuming an appropriate grillage to distribute loads).
Once at the installation site (and the weather is suitable) the ties between the forward end of the beam and its grillage support should be removed so that the beams can now hinge to maintain tension in the lowering cables (to avoid slackening).
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As the beams will be lowered slightly during the final stages of installation then the offlead angle from strandjacks could become greater than the permissible 3°. This means that the lowering stand jacks will have to sit in a gimbal arrangement so that they remain upright in spite of the articulation of the cantilever beams. The lowering cables lengths will differ but the loads in the jacks will be monitored and the cable lengths adjusted accordingly.
3.5.4.4 Tieback CablesThe tiebacks will support the majority of turbine unit vertical load. Each beam will require 6 ties, comprising 6 groups of 38 x 18mm diameter "dyform" steel cable. Each tie will connect into the beams as shown on drawing no. GM-44339-01 (Appendix C) and will lead via a compression tower (one each side of the barge) to a tieback structure at the bow of the barge.
The system will be designed so that the tieback cables are operating at 70% of their SWL at maximum loads (derived from hydrodynamic analysis) into the cantilever beams i.e. 6 x 600 x 0.70 = 2520 tonnes. The maximum vertical design load carried will be 3 x SWL of 600 t jacks i.e. 1800 tonnes.
The goalposts of the forward tieback structure will be located on bulkhead 82.296 m, and the cables will connect into it at a height of 10m above the deck. The height of the compression tower to provide a 458 lead off the beams needs to be 30m above deck. Thus the length of each tie needs to be approximately 110 m; 12 ties are required in total giving a total required length of 1320 m.
It is the intention to pre-tension the cables to ensure the correct division of load between the ties and the cantilever support structure, so 12 x 600 tonne jacks will be required for this purpose. These will connect into the forward goalpost arrangement, and as they will also be used to lower the beams (during installation), they will be required to be mounted in hinged racks (so that the jacks can rotate as the tower is lowered/raised).
3.5.4.5 Support TowersIn order to support the tieback cables two support towers are required. These are positioned on the aftermost transverse bulkhead about 14m from the stern to ensure that the shear force capability of the barge is not compromised.
The towers need to be about 30m high in order to get a 45° lead onto the cantilever beams. The eyes connecting the tieback cables need to be positioned so that there is no residual bending moment induced in the towers, i.e. they are designed to see compression loads only.
The structures will be built as lattices, hinged at the bottom, and with tension plates in way of the eyes connecting the tieback cables. The towers will sit on a grillage arrangement that will distribute loads into the barge.
Best estimates of weight of the towers were made based on in house data - in particular jack-up legs. The estimated weight of the compression towers is 80 tonnes. The estimated weight of the support structure is 165 tonnes.
3.5.4.6 Tieback StructuresThe tieback structures will consist of two (one per side) “goalpost” arrangements, fabricated of “hollow” section. The forward end of the structure will sit on top of transverse bulkhead 82.296 and it will be braced back by 6 tubular braces to frames/bulkheads aft. The tubular sections will
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receive load via shearplates slotted into the “hollow” sections that will mate with the barge structure below.
It has been assumed that columns and braces of the A frame and tieback structure do not see any moment. The structural concept is shown in sketches included within the hand calculations.
The estimated steel weight of the tieback structures is approximately 68 tonnes in total.
3.5.4.7 SeafasteningOnce the turbine unit base is located underneath the cantilever beams and the jacks locked off, shear plates will be welded between the box beams and the top plating of the turbine base. Box beam stubs perpendicular to the main beams will allow shear plates to take out transverse loads - approximately 340 tonnes per side.
Transverse forces will be taken to the barge deck by two tubular braces welded to the box beams and led to a “floating” connection on the barge deck. The “floating” connection is designed to provide restraint only in the transverse direction, and comprises a box stub into which the braces connect. This box stub is housed within a box that is restrained to deck. The contact plates will be greased to allow the braces to hinge and “breathe” with the movement of the cantilever beams.
Longitudinal forces will be taken to the barge deck as axial load in the cantilever beams.
The braces are permanent items and do not need to be cut to release the turbine unit - it is only the shear plates that need to be cut.
The estimated weight of the seafastening structure (permanent & temporary) is 25 tonnes.
3.5.4.8 Turbine Unit Lift PointsThe proposed lift point arrangement on the turbine unit is as shown on GM drawing nos. GM- 44339-001 sheets 3 and 4 (Appendix C).
The arrangement requires 6 recesses to be made in the turbine unit base (3 per side) and at each point the integration (into the base structure) of two reverse hooks of 75mm thickness.
The design safe working load of each point is 600 tonnes.
The arrangement requires a link plate to be fabricated to connect the reverse base hooks of the unit base to the standard two eye clevis end connection of the strand jack cable.
At loadout the cables will be lowered so that the linkplates hang over the hooks. Further lowering and manhandling will be required to house the linkplates within the hooks. The cables will then be tensioned, checked and then the unit will be lifted. The unit will be lifted until the base mates with the underside of the cantilever beams. At this point the cable end connection will be housed within a cut-out in the box section.
During installation the unit will be lowered until it sits on the seabed. The cables will be lowered until they are totally slack enabling the linkplate to drop out of the hooks under its own selfweight. Once clear of the hook (will probably require ROV inspection) pre-connected tugger lines will pull the link and the end connection clear. The jacks would then haul in the cables.
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The estimated weight of fabricated steel (linkplate and hooks) is 6 tonnes. This excludes internal re-enforcement that will be needed as part of the turbine unit’s structural designer.
3.5.4.9 Fabricated Steel WeightCalculations indicate that the total weight of fabricated steel required is just over 1000 tonnes. No internal reinforcement of the barge is required.
Note that these studies refer to the loaded transportation. The loads and motions of the compression tower will also be of interest in the un-loaded return transportation case. These have not been analysed here as, by inspection, the compression towers can be cross-stayed and guyed off. Naturally this would have to be included as a design case if the concept were to proceed.
3.5.5 Marine Operations3.5.5.1 ProceduresThe marine operations sequence considers:
• Loadout• Preparation for transport• Clear port and pilotage• Sea transport• Mooring and setup at installation site• Installation of the unit• Unmooring, quit site and return to load next unit.
The operations assume the options of a closed lock system and a tidal river port as load-out locations with variously short medium length tow distances.
The operational criteria include:
• Sailaway on a forecast of Bf 5 wind conditions decreasing over the period of the tow and installation operation together with.
• Forecast of minimum 24 hour installation at site of Bf 3 or less within the three day forecast window.
3.5.5.2 Crew and Barge SystemsThe barge complement is taken as a maximum of 20 persons working a single, possibly extended, shift. The barge riding crew for sea passage is restricted to two persons, the others being brought out to the installation site by fast launch, unless the sea passage is short.
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Compression tower strand jacks
Strand jacks Compression tower
0 m 10 m 50 m□
35
Key Key1 Mooring winch 13 Hydraulics2 Strand jacks 14 Spare3 Air dive station 15 Meetings and4 Decompression 16 Superintendent5 Rigging store 17 Client6 Workshop 18 Field engineer7 Winch control 19 Barge engineering8 Survey 20 Spare9 Welding store 21 Spare10 Power gen 22 Mess room11 Power gen 23 Messroom12 Spare 24 Galley
Key25 Power gen26 Spare27 Hydraulics28 Spare29 Spare30 Fuel31 Portaloos32 Grey water33 Garbage34 Z boat and davit35 10 man rafts and boarding ladders36 20 man rafts and davits
General arrangement of cantilever and strand jack system
Figure 3.22: GA of Cantilever and Strand Jack System
The barge conversion and outfitting includes various systems as detailed in Table 3.21
Table 3.21: Outfitting SystemsMooring winches Four winches, one at each comer of the barge. Local control for simplicity. No
redundancy of line numbers but operations will be of short duration and conducted in benign conditions.
Power supply Self contained power systems with supply cross-overs for redundancy.Air dive Air dive station costed here, alternatively an ROV operation to minimise diver
exposure.Operational hardware Rigging, welding and workshops for self contained operations.Operations management Containerised offices for installation superintendent, client and field engineers.Communications Stand alone VHF systems with hand held radio systems for work
communications.Quarters Minimum manning on sea passage and functional site quarters for galley and
messing.Fuel and water Tote tanks on deck with crane change-out.Toilets and drainage Self contained chemical toilets with shore change-out.Z boat Contingency work boat.Life saving and firefighting Probably to Class IX SOLAS using liferafts with davit and throw-over
launching.
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3.5.5.3 Lifting SystemThe lifting system comprises a pivoted cantilever at the stern, guyed to a pivoted compression tower and tied back to the barge deck at the forward end. The guys can be dead lengths, though this will require care in manufacture, or adjustable via strand jacks. The tie backs will be adjustable strand jack assenblies.
The lifting falls will be adjustable strand jacks, three on each cantilever beam.
The system requires one diesel power unit per jack with an automated recoiling system to tend the wires. The strand jacks will be controlled by a supervisor or foreman from a central location with a good view of the operational area and supported by instrumentation and a control panel.
The exposure to sea water is detrimental to wire conditions. It will be prudent to provide a fresh water washdown hose at the cantilevers for washing the wires as they are retrieved. Even so, the wash water will not clean the central core where corrosion can be expected. It will be necessary to exchange the lifting cables every month as planned maintenance to avoid failure of the wires under load. The wire change-out can be undertaken in rotation at the loadout port.
3.5.5.4 Barge OutfittingBarge outfitting is assumed to take place at the intended loadout port and with no need to drydock the barge. The structural members of cantilever beams and compression tower togerther with the foundation structures will be prefabricated. Control cabins, workshop, office and quarters containers will be pre-outfitted and interconnection service requirements pre-designed. The various strand-jack, guy and tie-back components will be designed and ordered under a rental contract.
The schedule for erection of the components will be a matter for detailed design, however this study assumes:• Major steelwork installation• Rigging of guys, tiebacks and compression tower• Installation of lifting strandjacks in parallel with• Main deck positioning and fixing of containerised units• Interconnection, hookup and commissioning.
The labour force, cranage and general supervision will be provided by the conversion yard. The strand-jack contractor will provide a two man team for supervision and testing of the strand jacks.
The conversion and outfitting schedule should allow one month.
3.5.5.5 Installation ScheduleThe candidate ports are assumed to be Liverpool as a locked port on a tidal river with a 12 hour transit to an installation site and Tees as an open tidal river port with a shorter one hour transit to site.
Operational times have been assessed as minimum, likely and maximum operational times, and include:
• Pilotage inwards• Locking inwards and mooring• Loadout of the unit• Securing for sea
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• Waiting for the next locking outward or pilots• Unberthing• Locking outwards or pilotage• Transit to site• Mooring at site• Positioning and installing the unit• Unmooring• Transit return to the pilotage boarding area.
The particular features of the strand jack lifting system compared to the systems previously assessed are the speed of operation, at:
• Lifting, 24 m/hr• Lowering , 20 m/hr.
These times, about an hour for the bodily lift and lower operation were incorporated in the earlier analysis of operational times and the analysis has not been repeated.
3.5.6 Costings3.5.6.1 Cost ModelsThe two cost models (Appendix C) of barge conversion requirements and operational costs are based on models used in previous phases and maintain a common format. Costs are shown in pounds sterling and based on prices in 2002 either drawn from in-house data or budget discussions with manufacturers or suppliers. No attempt has been made to drive suppliers’ budget cost estimates downwards and this means that better deals may be struck on a live project. Higher cost items such as tug dayrates have been taken at market rates and lower cost deals could probably be made taking into account the long period of hire; additionally cheaper tugs, such as Eastern block flag could be hired at a lower dayrate, though at increased supervision cost.
The models have been developed with minimum, likely and maximum costs in @Risk to allow Monte Carlo sampling of the costs in a 1000 iteration simulation and provide a distribution of results. The present models allow sampling of the inputs rather than the subtotals of conversion of operational blocks. Our experience is that sampling of all items results in a closer distribution of results than when working with larger blocks of subtotal costs, and this should be borne in mind when reviewing results.
3.5.6.2 Cost ComponentsThis installation option has particular changes to the cost components in two areas of steelwork and strandjack system when compared to previous studies. The steelwork requirements are given in Table 3.22:
Table 3.22: Cost ComponentsItem WeightCantilever beams and support grillage (2 total) 572 tCompression tower support grillage 165 tCompression towers (2 total) 83 tTieback structure 174 tUnit seafastening 25 tAnchor points and link plates 6 t
Total steelwork 1025 t
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The strand jacks are a rental item to be provided for the season. They are typically rented for an agreed base period at a lump sum, with an additional rate for extra time. Supervision and operators are provided on a dayrate.Aspects included in the lump sum price include:
• Design of the jacking system, operational manuals, QA and HSE statements• Certification of the jacks and strand wires• Mobilisation and demobilisation transport• Agreed hire period to cover assembly and operation.
Aspects not included in the rental are:
• Design and supply of support steelwork• Design and supply of anchorage steelwork• Labour, cranage and plant for assembly and disassembly
Contractors provided system costs based on a lump sum for a period of hire plus rates for subsequent optional hire periods. These have been developed into a dayrate for various subsystems.
The use of rental strand jacks includes power packs, hydraulics and control systems. The cost estimate includes reductions in barge conversion and outfitting to reflect rented systems.
3.5.6.3 Cost Summary The results show:
Table 3.23: Cost SummaryAspect Min Mean Max
£ Million £ Million £ Million
Barge conversion cost 5.4 6.5 8.1Rental and operational cost 7.4 8.1 9.1
Installation cost for 30 units 12.8 14.6 17.2
Installation cost for 1 unit 0.49
Number of daysNominal dayrate
186£78,000
This third phase of the study set out to investigate a chosen concept, to identify the technical and practical requirements and from that to develop a cost of installation per unit. Specific intentions were to utilise strandjacks on a horns or cantilever arrangement and to increase the weather workability. These specific requirements have resulted in the concept changing from a simple and robust concept to a more complicated mechanical and operational design with significant structural support steelwork and more of the expensive rental equipment than had been estimated in the installation options study.
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The cantilever and strandjack option results in a cost per unit installation of £0.49 million compared with a range of £ 0.37 through £ 0.43 million for the simple, though more weather limited, horns and A frame options.
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3.6 Retrospective OverviewThis section provides a retrospective view of the activities carried out under the ‘installation options’ study in the light of the findings of the detailed conceptual design.
Three different installations options were studied.
• “A” frame with turbine unit transported suspended from the frame• “A” frame with turbine unit transported on the barge deck
For each option a basic conceptual design was “sketched out” and considered for the following:
• Global and local strength of the barge• Ballast capacity of the barge• Basic sizes of the associated steelwork• Seastate limitations for transportation and installation• Required marine spread and marine operability• Purchase and operational costs
Each option was scored against each of the above criteria to determine a final overall score. The scoring system determined that the mid-deck A frame system with the unit on deck scored most highly. One of the main advantages was that with the turbine unit carried on deck, the unit could be transported in much more severe weather conditions (> BF 6) than the other two options which required transportation conditions less than BF 3.
The option of an “A frame at mid deck” can be compared to a commercial sheerleg vessel working with a dumb barge and was accepted as workable based on the analyses to date. It was also concluded that to be a viable alternative to the sheerleg/barge combination the candidate system had to be able to transport in Beaufort 6 conditions, and install in Beaufort 3 conditions. Similarly, it was agreed that the ability to use the vessel’s lift system to loadout was a major “plus” point.
The possibility of using a potentially low cost lift system, such as strand jacks, in conjunction with cantilevered horns was warranted consideration. It was considered that there was little point in taking the mid- deck “A” frame option forward for more detailed review as sheerleg technology was already well proven. With this in mind the strandjack/horn concept was chosen as the candidate for the detailed study in the expectation that it might provide a simple and cost effective alternative to sheerlegs.
The expectation from the findings of the option study was that by lifting and securing the turbine unit by the base as opposed to the column, the dynamic vertical loads induced by the barge roll motions would be taken out by a much larger couple lever (22m as opposed to 5m lever for the column lift) and therefore reduce the induced vertical loads and improve the ability to transport in larger seastates. Unfortunately, the hydrodynamic analysis for BF 6 conditions indicated that the dynamically induced vertical loads into the cantilever beams were still very high (in the order of 1900 tonnes per beam); this ruled out the “simple” cantilever concept because of the extensive external and internal reinforcement required.
Constrained by the requirement for a system that could be “bolted-on” to any barge of “fortune”, the design had to resort to more complex methods of directing loads into the barge without the need for extensive internal reinforcement - hence the scheme shown in drawing No.GM-44339-01. The
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main principle behind the revised concept being that if the majority of the load could be carried further forward by tiebacks (cables) then shear forces and bending moments into the barge could be accommodated with minimal internal reinforcement. Whilst the concept ensured that all of the static load would be taken by the cables, dealing with the fluctuating dynamic load required a support platform with more torsional rigidity, hence the need to stiffen the cantilever structure, which in turn meant the beams attracted almost all of the dynamic load with the resulting increase in steel weight (and cost).
If the other two, discounted, "hung transportation" options had been considered for Beaufort 6 transit conditions, similar increases in steel weight and complexity would have been expected with a similar affect on costs.
If serious structural alterations to the barge itself (inevitably requiring dry-docking) were acceptable, it is considered that it would be best to put a shaped “slot” in the stern of the barge into which a seabed/drydock supported turbine unit could be lifted, secured and then transported.Putting a “slot” in the existing stern, rather than adding an extension should enable the induced hull bending moment to be accommodated without reinforcement. It is suggested that any “rake” of stern be removed to provide a square end to give additional buoyancy and hence reduce shear forces at the stern. It is assumed that the buoyancy of the submerged unit would compensate for the loss of buoyancy due to the “slot”. Using a “slot” should allow the induced bending moment to be transferred directly into the deck and keel of the barge rather than through shear via a deck connection reducing the amount of additional steel required. The connecting of the strandjack link- plates (to the base) underwater would have to be carefully considered. This scheme would rule out the use of a barge of "fortune" hired for a particular campaign of installations.
Another alternative is the possibility of suspending the unit form a single lift point attached to the main column of the unit; the idea being that by allowing freedom in pitch and roll, dynamic forces could be minimised. In theory it could be promising but manufacturing a single 2400t universal joint (or similar) would need serious consideration.
Costs could be reduced if a lifting/lessening of some of the restrictions/criteria such as dry-docking, weather, loadout from seabed, etc were acceptable.
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4 SELF BURIAL SYSTEM
4.1 IntroductionThe objective of this study is as follows:
• To demonstrate that the basic concept of controlled jetting under a large slab is feasible and that adequate control of level can be achieved
This has been achieved through the following activities -
(a) Definition of test parametersThe purpose of this activity was to define the practical limitations of the self burial concept and to ensure that the experimental work builds on current understanding and was representative of full scale. Because of the lack of existing knowledge some small scale experimental work has also been carried out.
(b) Large scale testingThe only way to ensure that the results are representative is to test at large scale. Large scale testing has been carried out in three phases between October 2002 and April 2003. The phased testing has permitted technological step-improvements to be implemented in order to fulfil the objective of the study.
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4.2 Phase 1 - Definition of Test Parameters
4.2.1 Basis Soil Condition
Seabed soil conditions have been defined for base case consideration in this study, by reference to two British Geological Survey [7]’ [8] publications.
By examination of the near-surface sediments around the UK coastline, (and in the absence of site specific or geotechnical data) it would appear appropriate to consider a d50 range of 0.1mm to 0.6mm (fine to medium sand), with a d90 estimated coefficient of 1.4*d50. The occurrence of these sediments can be readily ascertained by conventional survey techniques (drop sampling, boreholes etc).
Sands of this size should also be readily obtainable close to the Pembroke Dock test site, which will facilitate the experimental programme.
4.2.2 Detection of Surface and Sub-Surface Boulders by Geophysical Methods
The self burial concept excavates below the base slab by transporting fluidised seabed material away from the base area. Realistically it will only be possible to fluidise sands and possible smaller gravels using water jets of reasonable power. The British Geological Survey shows that sediments size around UK near and medium shore waters are predominantly in the range 0.1mm - 0.6mm (fine to medium sand), and this will therefore constitute the base range for the self-burial jetting system.
Larger materials, e.g. cobbles or boulders, are likely to obstruct or prevent burial of the base slab. Isolated cobbles (particle size circa 100mm) may not prevent burial because the water jets are likely to remove the surrounding sands, leading to burial of the cobble. This mechanism is not likely with larger boulders. It is therefore important to be able to identify the presence, and ideally the size and depth, of these larger materials to establish the suitability of the site for the self burial system.
Independent advice has been sought from experienced geotechnical engineers working for Fugro Survey, Thales, and Gardline. The most reliable method for this type of investigation is seismic reflection/refraction. This technique provides a profile of surface or sub-surface features by measuring the time for seismic waves to return to the surface after reflection or refraction by subsurface formations. The method is very effective in detecting surface formations, but resolution decreases with increasing depth. Views vary on the maximum detectable size for different systems, however an overall ‘best-estimate’ is summarised in Table 4.1.
Table 4.1: Detection range for geophysical survey methodsDepth of Investigation
On seabed 1m 5m 10mType of seismic system Side scan Pinger Boomer Mini airgunMinimum size detectable <0.1m 0.3m 0.8m 3m
The optimum method for shallow water is currently pinger and/or boomer type seismic profiling. If the water depth is greater than 10m then it is unlikely that boulders will be detectable deeper than 10m below the seabed - this should not be a major issue since the soils of interest are likely to be in the depth range <5m. The lateral extent of a sub-surface profile is very limited. For 10m water depth the geophysical survey lines would have to be relatively closely spaced, say 5m, in order to provide complete coverage i.e. 200 line-km of survey for each square kilometre of survey area. For
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this reason it is more cost-effective if the geophysical survey can be targeted at foundation locations predetermined on the basis of a desk study.
Useful outcomes of a comprehensive desk study also include:
• An indication of the likely age and environment of deposition of the sands. Many of the southern North Sea sand formations are extremely unlikely to contain boulders, either because of their age or because of how they were deposited. For example, if the sands were deposited in the last 10000 years then they are very unlikely to have boulders within them.
• Better survey design e g. target particular risky horizons, or aim to acquire a representative sample of data covering each formation and adopt a risk minimisation approach.
Gravel and cobble layers are difficult to detect using seismic techniques unless there is good correlation with vibrocore or shallow boreholes indicating a distinctive layer containing substantial quantities of coarse gravels and cobbles with occasional boulders which can be traced as reflectors. Without sampling these layers their reflection patterns cannot be definitively confirmed as containing material likely to cause installation problems.
Overall, the detection of buried boulders appears to be a notoriously difficult problem. However using a combination of desk study and seismic survey it should be possible to identify the presence of materials likely to cause a problem to the self burial system, or at the very least minimise the risk to acceptable levels.
It is also worth noting that if the self burial system proves to be viable then survey methods may be developed to enable better definition (eg. frame lowered to seabed with multi-seismic sources).
4.2.3 Jetting System Concept
4.2.3.1 Concept 1 - Inclined Jets, Towards PerimeterThe basic concept is to use high pressure water jets on the underside of the foundation to scour away channels to the perimeter thus allowing the foundation to settle. By altering the rate of flow and the zones of operation it should be possible to keep the foundation vertical (Fig.4.1 and 4.2, below).
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Initial research into the viability of this approach has been undertaken by ODL and Global Maritime via a review of existing published material. The principal issue that remain unanswered by the existing research is a definitive solution to the ‘under foundation’ sediment transport equations.
The mechanism for the controlled formation of sufficient channels radially out from the centre of the foundation, together with the real time knowledge of behaviour under the base if channels collapse, requires further investigation.
The process relies on transferring the kinetic energy of the water jets to the sand particles to maintain a radial transport of the sediment. Another issue to resolve is how to maintain sediment flow when the base has already sunk into the sand as in Figure 4.2.
Quantification of the water jet velocity and flow rate to achieve this movement is being investigated. The paper Sand Bed Response to Submerged Water Jet[9] by T. O’Donoghue and B. Trajkovic reports experimental data on vertically impinging water jets which has facilitated the formulation of a number of equations describing the dimensions of the scour pit formed as a function of the water jet diameter and velocity and the particle size. From these formulae tables of water jet number, size, flow rate and erosion characteristics may be produced.
One of the papers referenced in the T. O’Donoghue et al paper has provided relevant and useful information. Erosion of loose beds by submerged circular impinging vertical turbulent jets [10] has been received from the International Association of Hydraulic Engineering and Research (IAHR) in Madrid. This paper, together with the first paper, has provided an insight into the operation of these submerged jets.
The problem at this stage is that both these practical studies used water jets with a maximum velocity of about 5 m/s. The formulas suggest that very high velocities should provide a much better scour regime with a far smaller number of jets and much lower pumping requirements. At this stage there seem to be no published papers where these high values of water velocity are used with cohesionless sediment.
Calculations derived from the O’Donoghue formulae have shown that a very large, but not unfeasible, pump would be required at full scale.
4.2.3.2 Concept 2 - Vertical JetsThe concept of total fluidisation under the base has been considered. Much research has been carried out into the fluidisation of sands in a variety of circumstances (from fluidised beds to volcanic pyroclastic flow). It seems that a potential option may be to utilise the weight of the foundation to expel a fluidised slurry from around the perimeter.
This would mean that the channel forming jets (utilisation of the kinetic energy of the jets to create and maintain scouring channels under the base) would not be required. With the new concept the kinetic energy of the jets would be used fundamentally to fluidise the sand down about 1 meter below the base. If the perimeter was also fluidised then the resistance to upflow should be minimal.
Due the weight of the foundation, the sand slurry beneath it will be at a higher pressure that the surrounding water. As long as the edge is fluidised or the centre is excavated there will be a net movement of fluid to the edge and centre.
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To control the burial of the foundation the fluidising jets are slowly powered up over the entire area of the base. The central slurry excavator is powered up to create a small crater in the middle. At a certain stage the weight of the foundation would overcome the remaining ‘solid’ areas of sand and force the slurry into the centre and around and out of the perimeter.
By maintaining just enough fluidisation jetting control of burial will be maintained. Should there be any tendency to tilt then the jetting which would be controlled in zones around the foundation can be reduced to create area of higher solidity and thus prevent burial on one side of the foundation (Figs.4.5 and 4.6).
Figure 4.5: Initial jetting Figure 4.6: Partially buried base
The advantages of the above system are:
• Pump power, jet number and therefore plumbing seem to be less than for the other two systems.
• By fluidising the entire area and slowly increasing the fluidisation flow it should be possible to control the burial more successfully. To counter any tilt that appears all that is required is to turn down the flow on the ‘down’ side zone or zones to solidify the sand in that area and the base will level.
• When the fluidisation jets are stopped there are no voids beneath the base, the weight will consolidate the sediment.
• Possible additions to the system would involve addition of jets to the topside to help spread the sediment on top as the base buries. This would also help with reversal of the system, i.e. reversal of the installation.
Figure 4.7 shows a diagram of a sand bypassing system in use in USA to augment a stationary slurry pump, taken from Fluidizer System Design for Channel Maintenance and Sand Bypassing1111. By installing a buried jetting pump an area of fluidised sand can be made which feeds into the crater created by the slurry pump:
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---------------- ----------------- V____________________ftwUMEflSUfiflli F jfr PuW
Mtf-y i > -wp-t i Vtii
Figure 4.7: Fluidiser pipe used in conjunction with a fixed slurry pump; hatched area indicated fluidised zone 151
The above diagram is included to illustrate the powerful effect of sand fluidisation in practice; the sediment is transported to the crater by gravity on a shallow incline. The orientation of the holes is horizontal because vertical downward facing holes “would tend to self bury”.
Although it has many interesting parallels with the concepts under consideration, the fundamental difference is that fluidisation of the bed is performed (as in fluidised beds) by injecting fluid from underneath. The formulae developed are not helpful in the vertical downward jet scenario.
4.2.3.3 Extrapolation of Research to DateAs previously stated all the work published to date involves the use of relatively low velocity jets. There is a need to extend that experimental base to establish the relationship between jet velocity/flow rate and the dynamic scour hole diameter and depth.
T. O’Donoghue concludes that “like dynamic depth, scour hole diameter for a given sediment size depends on the jet momentum flux, with little dependence if any on jet height”. Again, it is stressed that this conclusion applies only to cases of “highly erosive jets.”
This is exactly the regime we would expect to be utilising so the expectation that high velocity jets would be the direction we should be investigating seems correct.
There is one more detail that could be important to the operation of these jets, it is that all the data so far has been in terms of free flowing jets, Figure 4.8 below is a diagram from the paper by Aderibigbe and Rajaratnam describing the different types of jet regime.
In the cases we would be interested in, ‘strongly deflected jet regime’ is the nearest description.The point, however is that none of these tests show what would happen with jets applied below a large plate.
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The strongly deflected jet is characterised by a 180 degree change in the velocity vector of the jet with an eventual egress of that jet water into the clear water above - in so doing creating the circular berm of material you see in the diagram.
Fig. 4a Strongly deflected ;ei regime 1 fSDJEt l). Fig. 4b Strongly deflectedregime II (SDJR Hj.
Fig. 4c. Wujldv eellsLLEci jet regime E i."yVD.5"R !). Fig. -til. Weakly dafleered jer regime [I (WDJR TTJ,
Figure 4.8: Sketches of flow regimes (from Aderibigbe and Rajaratnam)
It would seem likely that the plate above would if anything add to the fluidisation zone in terms of diameter and depth of the dynamic scour volume for a given jet velocity and diameter.
4.2.4 Small Scale Test 1
4.2.4.1 Test ArrangementIt was decided to undertake a small scale test of the possible jetting operations to investigate the operation of high velocity water jets into underwater sands.
The size of the experimental tank used by Aderibigbe was an octagonal box, with sides of 0.235m by 0.6m high. It is therefore reasonable to undertake a representative experiment using a tonne bag of coarse builders sand, delivered in a cubic shaped sack about 0.85m x 0.85m x 0.85m.
The test rig is shown illustratively in Figure 4.9.
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Figure 4.9: Schematic Representation of Small Scale Test Rig
As a stout concrete wall was available at the site location it was decided to simplify the design by attaching the hinge of the loading beam to a variable height wall plate.
The sand container is a multi layer assembly of strong canvas sand bags and a plastic waterproof liner. The bottom of the bag incorporates a layer of gravel and a plastic pipe manifold to allow water to be drained and filled from the bottom of the test sample so as to minimise disturbance to the sand.
From initial extrapolations of the equations developed by T. O’Donoghue and B. Trajkovic a set of possible pump and jet sizes was decided. The purpose of the test series was to establish the accuracy of these equations at much higher jet velocities and to assess the practicality of various jetting methods to perform jetting installation of a foundation.
The components of the small scale rig were as follows:
1. Measurement equipmentAn in-line water flowmeter was installed as were pressure gauges on the foundation plate itself.
The load on the plate was measured with a tension link electronic load cell.
A laser level was sighted across to a scale on the foundation to measure the vertical displacement of the foundation as the jetting proceeded.
The measurements taken during the tests were thus time, foundation plate height, water flow rate
and water pressure at the plate. The diameter of the jet nozzle and the applied foundation loads were variables and a number of graphs subsequently plotted to show the results.
2. Pump equipmentThe water was supplied by a high pressure positive displacement water jet pump giving about 7litres/m.
3. Jet sizes and orientation
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The jets were of various sizes and orientation as described later. Because the pump was positive displacement the flow rate was always about the same, but the pressure and thus jet velocity varied.
4. Loading equipment
The application of load onto the foundation plate was by means of a metal tray on a lever arm that could be loaded as necessary. The load on the foundation was monitored via a tension link digital load cell and recorded at the start of each test.
4.2.4.2 ResultsPhotographs showing the results of a series of loaded tests are given in Figure 4.10.
The first test undertaken was with a positive displacement pump giving about 6.75 litres/minute allowing variable line pressures dependent on jet size and sand bed resistance. The initial jet size chosen was 2mm diameter acting vertically giving a nozzle exit velocity of about 36 m/s.
With the full 1 tonne of applied load (equivalent to about 5 tonnes/sq.m) the installation of the disc was disappointing, achieving only about 20mm of insertion over 5 minutes of pumping with the system ‘stalled’. Observations of the disc showed a small amount of fluidisation apparent around the periphery of the disc.
A number of tests with smaller and larger jets (from 1mm to 3mm diameter) were also done with no improvement.
Subsequently some experimentation on unloaded jets (with no disc) were performed to gain some idea of the scour hole size and shape. It became apparent that the cone of fluidisation was deep and narrow - not the effect expected from the extrapolation of the theoretical equations mentioned earlier.
Some thought was given to using a nozzle that would give a cone shaped jet rather than the effectively parallel jet observed from the plain drilled jet. However due to the difficulty of fabricating such a nozzle this was not pursued.
Rotating the shallow cone of fluidisation 90 degrees so that the nozzle is pointing horizontally seemed the next best method of approach.
A nozzle was drilled with 6 equally angularly spaced 1mm diameter holes with the result that in all but the highly loaded case there was evidence that ‘piping’ occurred. Initially the sand moved out from the edge of the disc fairly evenly around the perimeter of the disc, however in nearly all cases a sudden change to piping in one two or three places occurred. This did not prevent continued disc insertion.
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0.55 tonnes/sq.m
1.18 tonnes/sq.m
Figure 4.10: Behaviour of Disk Test Under Various Loadings
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This piping may occur due to initial conditions around the disc to be not quite even or that the manually drilled jets vary slightly in attitude and size.
The graphs shown in Figure 4.11 indicate a slight trend for the disc to bury more readily at lower applied loads. The pressure exerted by the disc thus appears to influence the performance of the fluidising jets.
Displacement v. Time for 6 radial 1mm jets
0.0 1.0 2.0 3.0 4.0 5.0 6.0 7.0 8.0 9.0 10.0Time, minutes
—♦—5.30 tonnes/sq.m —3.26 tonnes/sq.m 2.73 tonnes/sq.m 2.30 tonnes/sq.m—1.63 tonnes/sq.m —1.18 tonnes/sq.m —l—0.55 tonnes/sq.m
Figure 4.11: Small Scale Test 1 - Displacement vs Time
4.2.5 Small Scale Test 24.2.5.1 Test ArrangementWhile the tests described in the preceding section provided useful information regarding the likely mode of sediment removal and the effect of loading, it was considered that more data was required to enable the large scale test to be designed. In particular the question of jet spacing needed to be resolved.
It was therefore necessary to study the relationship between sediment behaviour and distance from a jet nozzle.
A segment that would just fit into our existing test tank was thus constructed. Fundamentally a 60° triangle of side length 800mm was built. A single jet was installed at the apex with an adjustable angle so that jets of various diameters and with exit angles variable from horizontal down to 90°.
Figure 4.12 shows a photo of the unit with 2mm jet working:
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The steel sidewalls are 150mm deep, (the curve at the end to allow fitting into the square test tank at an angle).
Holes were drilled through the test piece and plugged with the plugs (seen white in the photo). The measurements of fluidisation depth were performed with a close fitting wooden dowel pushed downwards until there was an obvious bottom to the fluidisation pit. The insertion depth below the bottom of the test piece plate was then measured and recorded.
The total area of this test piece is double that of the 500mm diameter circular piece and so the applied weight at the lever arm was adjusted accordingly to give a total of 3.45 tonnes/sq.m, (the expected pressure of the full size offshore unit prior to ballasting with ore ballast).
A number of experiments have been conducted to determine the fluidisation depth and extent with various jet sizes, fluid flow rates and angle of jet to horizontal.
4.2.5.2 Results
Figure 4.13 shows the fluidisation depth v. distance from the jet for various fluid flow rates for a 3 mm jet. The exit velocity of the jet is shown in the annotation.
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Fluidisation depth v. Distance from jet 3.0 mm jet diameter
Distance from jet, mm
—♦—2.00 I/m, 4.7 m/s —■—3.00 I/m, 7.1 m/s 4.00 I/m, 9.4 m/s5.00 I/m, 11.8 m/s —*—6.00 I/m, 14.1 m/s —e— 7.00 I/m, 16.5 m/s
Figure 4.13: Fluidisation Depth vs Distance from Jet
Figure 4.14 shows Flow rate v. distance from the jet to achieve at least 60mm depth of fluidisation plus extrapolation to show likely extent of fluidisation at increased flow rates.
Flow rate v. Distance from jet to achieve at least 60mm fluidisation depth
10.00
Flow rate, litres/minute
----- Jet diameter 2.00 mm ------Jet diameter 3.00 mm
Figure 4.14: Flowrate vs Distance from Jet
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4.3 Phase 2: Agitating Jet Tests
4.3.1 Introduction
This section describes the design and manufacture of the large scale test, and the first series of tests, which were carried out on the 17th and 18th of October 2002 at the Port of Pembroke.
4.3.2 Jetting Design
The analysis of the experiments described in Section 4.2 led us to believe that a flow rate from each 3 mm diameter jet of around 12 lires/minute will produce a minimum of a radius of 750mm by about 50 degree fan of fluidisation.
Assembling this onto an octagonal grid gives the following array (Figure 4.15):
712 mm
629 mm
723 mm629 mm
34C*
>98 mm57\nm
4800 mm
1988 mm
Figure 4.15: Octagonal Base Plate Jet Positions
The full picture of the jetting array is shown in Figure 4.16.
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Figure 4.16: Jetting Zones
There is a central zone which would always be fluidised, and four edge zones. The outer zones are to aid control of piping to specific points.
The plumbing of the jets is above the base plate which allows individual V2” male threaded spigots to be inserted through the plate at the appropriate positions. These spigots then have ‘iron’ end caps screwed on. These end caps are drilled with the appropriate jet hole. This gives maximum flexibility in jet arrangement.
4.3.3 Test Piece Design4.3.3.1 Insitii LoadingIn order to perform the burial tests at the correct ground pressure of about 3.45 tonnes/sq.m it is necessary that the total weight of the test piece as described above weighs approx. 66 tonnes.
To achieve this weight it was initially decided to construct a frame capable of taking large test weights. It is not easy to find sufficient equally sized test weights locally and eventually a supplier
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in Cardiff was approached. The costs involved in transporting about 60 tonnes both ways was considerable.
To load the test piece with these weights requires a substantial frame arrangement to allow plumbing runs under the weights. This double bottom approach would also require that the volume between the weights and the bottom plate must be allowed to fill with water as the test piece buries otherwise with 0.5m of burial this would amount to a reduction in loading of around 10 tonnes. However sand must not get into this volume because of the difficulty in removing it after the test. The volume would have to have water draining facility for the same reason. The extent of framing needed to evenly distribute the 5 tonne test weights would also be significant and was found to interfere with possible plumbing runs.
A different approach was used: the construction of a tank-type test piece. This is 4m tall to contain enough water to give the 60 tonnes extra weight required, plus extra volume of water as the test piece buries.
The advantages are several;
1. Allows easy loading/unloading of applied weight using pumps (already on site).2. No sand ingress into the test piece so easy removal of test piece after each test. This allows
a faster turnaround with the possibility of several tests during a working day.3. The centre of gravity more closely approaches the proposed full size foundation - the test
becomes more representative of a real situation.4. Plumbing inside the test piece is available for change if required.5. Loading of the test piece is more controlled - no tendency for large offset loads to cause
initial inclination of the test piece.6. The load is totally evenly distributed so internal frame work is minimised.7. No time/cost constraints on how long the weights are available.
Figure 4.17 shows a graphical representation of the test piece sitting on the sand in the test tank with the overall dimensions of the installation.
Figure 4.17: Test Rig Dimensions
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Jet nozzles were created using screw-on spiggots with pre-drilled 3 and 4mm jet holes (Figure 4.18)
Figure 4.18: View of underside of Test Rig (showing jet nozzles)
An array of internal pipes and channels welded to the inside of the tank ensure the correct distribution of water to the jetting zones (the faint rust lines in Figure 4.18 show the positions of the steel channels on the inside of the test piece).
4.3.3.2 Insertion Depth Measuring SystemThe system that has been devised for the insertion depth measuring will also give real time indication of the tilt of the test piece.
The system works on the simple principle of 2 pairs of water tanks placed orthogonally at four ‘corners’ of the octagonal test piece. These tanks are connected by PVC tubing to four sight glasses at the operations point next to the manifold control (Figure 4.18).
Test piece levelFigure 4.18: Insertion Depth Measuring System
Test piece tilted
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The scales alongside each sight tube give a direct reading of the height of each corner. By noting the height reading as the test program proceeds a real value of the insertion of the test piece is recorded. To make the water in the levels more easily visible a small quantity of Fluorescein dye is introduced to give a luminous green colour.
A number of calculations were performed to establish the safe angle of tilt for the test piece. Even with the tank full the centre of gravity will not move outwards more than about 17% of the radius at 10 degrees of tilt.
4.3.3.3 PumpsTwo pumps were used, the first was a hydraulically powered submersible pump with the pump head lowered into the dock at about mid-water position. This pump was selected to supply the main power pump with up to 150 cu.m/hour water at a head of up to 12m.
The second pump, The Selwood 100SH, is the main power pump, this pump is theoretically capable of supplying up to 200 cu.m/hour at a head of 80m.
4.3.3.4 Flow ControlThe pumps deliver water to a distribution manifold. The design allows for the connection of up to three inlet/outlets and 5 measureable supply lines. There are 5 butterfly valve controlled supply lines each equipped with a variable area flowmeter and supply pressure gauge. These are numbered and colour coded for easy referral to the test piece jetting zones (see Figure 4.19).
Figure 4.19: Flow Control Manifold
This system allows full control of the water flow to each of the 5 zones with the extra facility to stop flow immediately to the 5 zones and dump the pumped water to the dock. There is an extra rear flanged outlet to a flexible overhead pipe that allows the test piece tank to be filled with water.
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4.3.4 Test Series 14.3.4.1 DescriptionThe pumps were started and full flow was gradually applied to all jetting zones. The flow rates and pressures of each zone were recorded at 1 minute intervals initially, later 2 minute intervals were used. The levels of the four ‘comers’ were recorded as displayed on the water level gauge.
As the jets started to cause piping and the sand was observed to be forced out from the test piece base, the level of the test piece was continually monitored. Alterations to the flow to the zones allowed the relative levels to be altered. The maximum out of level was less than 0.5 degrees throughout the test.
It was observed that mounds of sand with an active central piping point were being produced at the centre of each panel of the test piece - as was hoped.
However as the test proceeded the insertion rate declined. It seems that the ‘overburden’ of the mounds at the piping points was reducing the ability of the jets to remove sand from beneath the test piece.
Eventually the test piece insertion stopped. At that stage the pumps were stopped and the tanks drained down.
4.3.4.2 ResultsObservations of the pattern in the sand after the test piece had been removed indicated that not all the jets seemed to be performing fully (See Figure 4.19).
Figure 4.19: Residual Sand Pattern - Test 1
Lifting of the test piece from the sand disturbed the edge, however a circle of small pits nearest the disturbed perimeter could be seen, this circle represents jets that are on the same circuit as the perimeter jets. These perimeter jets had been previously converted to 4mm diameter whereas this
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circle was still set at 3mm diameter. Flow to these was thus greatly reduced and the importance of balanced jet sizes was thus illustrated.
Flow to some of the central 3 mm jets was also seen to be variable so a full jet test was undertaken whilst the test piece was on stands to allow checks and alteration of the jets.
These tests revealed a wide variability in the jets performance. Removal of the jets revealed some had clogged or partially clogged with sand and debris. The 4mm jets were hardly affected, whereas the 3mm jets were all affected to some degree.
The inlet to the dock sump pump did have a coarse filter, a system of finer filters would help, although it seemed that some of the debris inside the jets may have actually been forced in through the jet nozzles as the test piece was positioned on the sand of the test tank
The insertion depth with time for Test 1 is shown in Figure 4.20.
Time, minutes
Figure 4.20: Test 1 - Insertion Depth vs Time
From the graph of insertion depth v. time for Series 1 it can be seen that initially a rate of insertion of about 100mm in the first 6 minutes. At that rate, if it could be maintained, insertion of 1 meter would take about an hour.
Figure 4.21 shows tilt angle vs time, illustrating that control was maintained through the test.
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1.000
0.800
0.600
g 0.400
I| 0.200
0.000
-0.200
-0.400
Time, minutes
Figure 4.21: Test 1 - Tilt angle vs time
4.3.5 Test 2a4.3.5.1 DescriptionSince most of the nozzle blockage problems had occurred with the 3mm jets it was decided to make all the jets 4mm. In addition it was decided to add extra jets to the outside of the test piece to try to create dispersal of the sand mound.
The plumbing system was thoroughly flushed through and all jets were tested to ensure that none were blocked.
The operation was the same as the first test, however soon after the start it became apparent that the insertion rate was very much reduced. The data from the flow rates showed a reduction in flow, and eventually the main supply pump at the dock failed. A split hydraulic hose was discovered and the test was terminated.
Measured flow rates indicated that the pump was performing well below the specified levels. Pump performance is critical to the effectiveness of the self burial system. The advantage of utilising “off-the-shelf’ civil engineering pumps is the availability and cost. The disadvantage of these pumps is that they are mostly used for emptying water from building sites and inevitably pump mostly abrasive mixes of water/sand/gravel. The result is that impellers wear, seals degrade and generally performance is a long way from the published pump curves.
4.3.6 Test 2b4.3.6.1 DescriptionIt was decided to weld up the extra jets that were installed to help clear the overburden of sand :
The final arrangement of jets is as per Series 1 but with all jets 4mm diameter. The test was recommenced and all parameters recorded until insertion rate reduced to virtually zero.
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4.3.6.2 ResultsOn completion of the test the pumps were stopped and the test de-rigged as before. The water was drained down from the test piece and the test tank to reveal the sand mound pattern (Figure 4.22):
Figure 4.22: Residual Sand Mounds - Test 2
The insertion depth with time for Test 2 is shown in Figure 4.23.
Figure 4.23: Test 2 - Insertion Depth vs Time
Test 2 showed a slower initial rate of about half the rate of Test 1. This could be due to the larger jet sizes reducing the flow exit velocity from the jets.
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4.3.7 Conclusions from Phase 2
• The tests show that in principle jetting under the foundation to create fluidisation and piping of the sand to the perimeter of the foundation is possible. The tests also show that the tilt of the foundation can be continually monitored and altered as the insertion process takes place.
• The rate of insertion is too slow after the initial period and larger flow rates are therefore required, especially if 4mm diameter jets are used for practical reasons of avoiding blockage.
• The final jet pattern generally shows good even distribution of scour holes. The problem with the existing pump set-up is that to achieve sufficient flow rates to move material not all zones could be pumping simultaneously. The result was that the foundation inevitably ‘hung’ on the non-pumped zones.
• The overburden of the sand mounds must be a limiting factor in the performance of the jetting. A system of extra jets at the appropriate places could disperse some of the material and keep the ‘pipes’ open to allow more easy removal of the sand from under the foundation.
• The next phase therefore is to obtain a larger capacity pump and to develop a method of removing the sand mounds. The capacity of the pump must take into account the inevitable variation between theoretical pump performance and observed performance
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4.4 Phase 3 - Eductor and Radial Jet Tests
4.4.1 Introduction
In consideration of the results of Phase 2 it was decided that a number of modifications would be made to the test rig in order to facilitate deeper and more rapid self burial:
• A series of external eductor type pumps, one on each face, to remove the overburden of sand as it was being ejected from the side.
• An alternative jetting system to test the channel creating potential of a series of in-line jets.
• Source more efficient pumps to operate the jets and the eductor pumps.
• Plumbing for the above system to allow changeover from the original system to the new system and back again if necessary.
The Phase 3 tests were carried out over the 12th and 13th of December 2002.
4.4.2 Eductor Pumps
The use of eductor pumps to move large quantities of marine sand and aggregates is a well proven technology. Eductors have the advantage that there are no moving parts in the pump, the motive power being a ship or shore based water pump.
Eductor pumps operate on the Venturi principle as shown below:
1. Fluid under pressure enters nozzle and velocity increases as passage narrows.
2. Pressure is fully converted to maximum velocity as water leaves nozzle tip, discharging into venturi tube.
3. Partial vacuum created in upper chamber. Velocity of the stream entering the tube entrains additional water.
4. Recovered liquid/vapor constantly flows into eductor to fill partial vacuum continuously created by high velocity stream.
5. Tapered venturi tube allows velocity of jet stream and well mixture to be gradually reconverted to pressure.
6. At greatly reduced velocity, water continues out of eductor, losing pressure as it approaches discharge piping
Figure 4.24: Principle of the Eductor Pump
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Two possible systems were investigated to undertake the removal of the sand slurry, either one central large eductor pump (which would involve re-directing the slurry forming jets towards the centre of the foundation), or 8 circumferentially mounted eductors. As the existing jet geometry was showing promise and the unknown effects of moving the sand to the centre of the foundation was considered to be high risk it was decided to develop the 8 outside eductor plan.
Because of the geometry and location of the ejected sand pile it was decided to alter the basic shape of the eductor to allow the inlet of motive flow to be to the side allowing the suction input to be at the bottom.
The design is shown in Figure 4.25 below:
Eductor - GA
00.00 mm
1108.80 mm
400.00 mi l
101.60 ml
Figure 4.25: Self-Burial System Eductor
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The motive water flow is introduced from the bottom right in the diagram and the suction flow is through the bell mouth at the bottom. This is set to be level with the bottom of the foundation.
The components used were all taken from existing pipe line fittings so that the complete fabrication of the eductor cloud be achieved economically. If the components were cast and machined the efficiency of the venturi would be increased as the flow through the device would be less turbulent.
The outlet at the top was angled to allow the slurry to be deposited away from the edge of the foundation.
To control the eductors independently of the foundation water jets a separate pumping system is used with a valved distribution system to allow control of pairs of eductors independently. On one side of the foundation is the array of quick connect motive power pipes, marked with the appropriate zone numbers: 4,3,2 and 5. The height of the discharge and the height of the quick connect array was chosen to allow the foundation to bury by about 1 metre without interfering with them. The complete assembly as fitted to the foundation is shown in Figure 4.26
Figure 4.26: Eductor Assembly
4.4.3 Submersible Pumps
As previously discussed, there were some problems with the original pumps that resulted in performance lower than anticipated. It seems likely that these pumps, normally used for civil engineering operation on large building sites for draining building sites had probably experienced some wear to the centrifugal vanes preventing fully efficient performance.
A source of well maintained local pumps was found in Milford Haven. These pumps are hydraulically powered submersibles owned by the Pollution Control Company, part of D. V.
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Howells of Pembroke Dock. Their usual duty is for the transfer of oil cargo from stricken vessels etc, but most of the time they are ‘on standby’ in a warehouse where they are regularly cleaned and maintained. The advantage of this pump is that only one unit is required to supply the jets. A second pump of the same sort was then used to power the eductors.
The pump curves show they are capable of supplying about 150 cu.m/h @ 70m head.
4.4.4 Test 3a
4.4.4.1 DescriptionBecause of the various changes to the underside of the foundation, the addition of eductor pumps and the pump equipment it was decided to start the tests using the original jet system to create some comparison data for the radial channel tests.
The foundation was set up as before and the test piece filled with 60 tonnes of water. The water jetting pumps and the eductor pumps were started up and the burial process recorded as before.
As soon as the eductors were powered up it was apparent from the dark colour of the effluent that they were pumping a good proportion of sand in the slurry away from the centre of each perimeter facet.
Initial burial was quite fast and as before the flow control valves were used to alter the insertion angle as required.
After about fifteen minutes it became apparent that the burial was becoming ‘stuck’ on one side. At that stage attempts were made to pump more water to the jets at that zone as well as to the corresponding eductors. However these attempts were not successful in starting the burial again on this zone. Despite this it was decided to carry on with the operation to see how far the foundation would bury in these conditions. Careful watch was kept on the increasing angle of the foundation and all the usual parameters were recorded.
Once the foundation reached an angle of about 3 degrees it was decided to halt the operation even though the foundation was still burying.
The eductor outlet pipes were removing material from the immediate vicinity of the foundation perimeter, but it was noted that an extension to the outlet would be advantageous as the spoil heaps angle of repose was allowing material to build up against the foundation sides again.
It was decided to drain down the foundation tank as soon as possible and remove it to allow inspection of the pit.
As soon as the foundation was removed it was apparent from the patterns in the sand that there was a whole area of jets that were not operating correctly. Foreign matter had clogged the 4mm jets and prevented the fluidisation of the sand in a large area.
The problem jets were removed and inspected. The offending material was composed mostly of remnants of ‘Dead Mans Finger’, a kind of jelly fish that was later observed in the area. This gelatinous material was able to collect small grains of sand (that would normally travel through the jets) and so blocked the jets.
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It was then decided to remove all jets and flush the entire system with high pressure water to try to remove any more contamination from pumps, pipes and hoses.
4.4.4.2 ResultsFigure 4.27 shows pictures taken at the end of the test.
(c) Close-up of sand mound (d) Residual sand patternFigure 4.27: Condition at end of Test 3a
Figure 4.27(a) clearly shows a left to right tilt across the test piece. The reason for this is illustrated by Figure 4.27(d) where there is a complete absence of jet pits on the far side in contrast to the near side where there is an even and regular residual jetting pattern. The test piece was ‘sitting-up’ on undisturbed sand. The outlets in this area were later found to be clogged.
Figures 4.28 and 4.29 show insertion depth and tilt angle respectively against time.
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300.00
250.00
200.00
■o 150.00
100.00
50.00
Time, minutes
Figure 4.28: Test 3a - Insertion Depth vs Time
0.000
-0.500
-1.000
2 -1.500
c -2.000
-2.500
-3.000
-3.500
Time, minutes
Figure 4.29: Test 3a - Tilt Angle vs Time
The increasing tilt is clearly shown in Figure 4.29, demonstrating the importance of ensuring that outlet nozzles remain unblocked
4.4.5 Test 3b
For the second test of this phase it was decided to investigate the radial water jet system
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4.4.5.1 Radial Channel ConceptThe radial channel system is shown diagrammatically in Figure 4.30.
Figure 4.30: Radial Channel Jetting System
The jets used were again 4mm diameter.
As with the original system the jets were zoned, each zone being a quadrant of the octagon in order to facilitate level control during the burial process.
The jets were arranged along a 2” delivery pipe attached to the underside of the foundation. To reduce damage to the jets as the foundation was placed on the sand they were offset to one side of the pipe. These jets were angled downwards by 20 degrees.
The design concept of this arrangement was to create a channel from the centre to the perimeter, the eductors at each perimeter position then removing the slurry that emerges.
To improve the quantity of water flowing along these channels, a central 4” hole has been introduced, that is directly connected to the water surrounding the foundation. The intention is that the flow of water in the channels will cause some suction at the central hole inducing extra flow from outside.
The hope is that the flow of water along the channel will continually erode the sides of the channels and thus remove the sand from beneath the foundation to a great enough extent that the weight of the foundation will collapse the remaining sections of sand into the channels. This collapsed sand will then be continually removed and the foundation will slowly bury into the pit created.
The plumbing of these new radial pipes was connected up to the same delivery and control system as the original system and all other controls and level measurements were also the same.
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(a) Arrangement of pipes on underside of unit Figure 4.31: Photographs of radial system
(b) Pre-installation test of radial jets
4.4.5.2 DescriptionThe tank was installed and the test started once the water in the sand tank had reached the level feed pipe to the central supply.
Firstly the eductor pumps were started and as in the previous test they immediately started pumping sand slurry.
Almost immediately after the radial jets were started the Zone 2 flowmeter float surged through the flowmeter into the delivery pipe. The effect on the flow was unknown and the flow through the flowmeter had to be extrapolated from the pressure and valve position.
As the foundation began to bury it was noted that water was being drawn into the central supply pipe thus augmenting the flow through the radial channels although to a more limited extent than had been hoped. However the burial rate became very slow and eventually the test was stopped when burial of the foundation had stopped.
4.4.5.3 ResultsThe condition at the end of the test is shown in Figure 4.32
(a) Drained Down Condition (b) Residual Sand PatternFigure 4.32: Condition at end of Test 3b
Deep trenches along the lines of the radial jets were clearly visible.
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A central ‘pillar’ indicated shows that there was very little slurrification of the sand near the centre, and not very much flow through the central duct.
Flat pads of sand were clearly visible on the far side where the radial jet supply lines had left indentations in the sand. The jets in this area were not functioning correctly.
Although the Zone 2 probably had a restricted water flow it appears that the Zone 3 jets were operating inefficiently because there is very little evidence of trench creation. It seems likely that foreign bodies in the supply line were again responsible for the lack of jet power.
Also visible are the flattened areas in each ‘corner’ of the pit where no jet energy could reach. This is the fundamental reason that the foundation could not continue to self bury.
Figures 4.33 and 4.34 show insertion depth and tilt angle respectively against time.
Time, minutes
Figure 4.33: Test 3b - Insertion Depth vs Time
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Ang
le, d
egre
es
0.600
0.400
0.200
0.000
-0.200
-0.400
-0.600
Time, minutes
Figure 4.34: Test 3b - Tilt Angle vs Time
4.4.6 Conclusions from Phase 3
• Despite the problems of maintaining consistent flow through the jets in all the tests, the tests have shown that sub foundation jetting with peripheral removal of the sand will allow self burying.
• During the tests the application of differential flow rates through the various jetting zones allowed control of the angle of the foundation to the horizontal. It would seem most probable that if the foundation had been set down on a moderately sloping site that this could also be levelled through zone control.
• It became evident that removal of the spoil from the perimeter the foundation was essential to allow continual burial. This will become more important the deeper the foundation is required to go.
• Eductor pumps were designed and built to remove the soil accumulation at the perimeter of the foundation and worked well. The eductor pump has the advantage that the motive power is remote and the pump itself can be built economically.
• Foreign bodies whether organic (seaweed, jellyfish etc) or inorganic (large sand grains, rust flakes etc) can cause severe restriction to the operation of the jets.
• Comparisons of the two systems tried with Test 3a (widespread zoned jets) and 3b (radially oriented and zoned jets) shows positive traits in both approaches. Certainly the 3a system could have achieved much greater depth had one zone not become blocked. The 3b system also had jet blockage problems, but the comers furthest from the jets were undisturbed and stable enough to hold the foundation up. The depth of the fluidised sand in the trenches indicates the potential to create channels through which large quantities of sand could be removed.
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• Comparisons between Test 1 and Test 3a show similar initial insertion rates of about20mm/minute rapidly dropping down to around 5mm/minute. Eventually the insertion rate of the first system tailed off as the overburden prevented sand effluent around the perimeter. The insertion rate of the Test 3 a would have continued as the overburden was being removed continually by the eductor pumps.
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4.5 Phase 4: Combined Jetting System
4.5.1 Introduction
As a result of Phase 3 a number of improvements were proposed for the fourth and final testing phase:
• Combine the radial system with the agitating system - The radial system was highly effective in transporting material to the eductors but the extent of its influence was limited (the sand under the corners of the unit was undisturbed). The agitating system was effective over the full base area but relied on pressure variation to transport slurry to the edge of the base. A combined system should prove more effective than either system operating individually.
• Replace steel pipework and channels with ABS plastic pipework - Much of the pipework required replacing in order to construct the combined system. It was decided to use ABS plastic pipework in order to eliminate the risk of blockages from corrosion products. In addition it is likely that plastic pipework would be used at full scale, and useful experience could be gained working with the material, forming connections, joints etc.
Phase 4 testing took place between 31st March and 5 th April 2003.
4.5.2 Plumbing Design
To simplify installation all the jetting pipes were installed on the underside of the foundation and configured radially.
Two basic circuits were installed (Figure 4.35). The dark circular patterns at the perimeter show the positions of the eductor pumps.
The intention was that the radial jets (Fig 4.35(a)) would establish channels transferring the sand to these eductors which would then transport the sand away from the foundation. The jets themselves were situated on the side of the 65mm pipe and being constructed of soft steel 4mm id pipe could be angled outwards and also slightly downward to help create fluidised channels. The pipes were connected in pairs resulting in 4 pairs of pipes to allow zoning of the jets.
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(b) Agitating Jet Pattern(a) Radial Jet PatternFigure 4.35: Theoretical Radial and Agitating Jet Patterns
The agitator system (Fig 4.35(b)), was also based on an array of radial pipes but with straight jets normal to the pipes. As with the Radial array extra jets holes had been provided and blanks inserted where no jets were needed. This allowed rapid changing of the jet geometry as the test series progressed should evidence suggest that there were dead areas or conversely areas with too much fluidisation.
These pipes were also connected in pairs giving 4 zones. The manifold arrangement (described later) allowed the original control valves to control 5 zones in total and so it was decided that at any one time either 4 radial jets zones and 1agitator zone (ie all 8 agitator pipes connected to one control valve) or 4 agitator zones and 1 radial zone would be operational.
(a) Combined Jetting System (All nozzles) (b) Combined System (Central Agitators Off)Figure 4.36: Theoretical Jetting Pattern - Combined System
The theoretical combined radial and agitating pattern is shown in Figure 4.36. There is a highly concentrated fluidised zone in the centre of the foundation. Figure 4.36(b) shows the pattern when the inner jets of the agitating system are replaced by blanking caps. This also has the advantage of maintaining the pressure for the remaining jets.
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4.5.3 Pumps
The pumps used were as described for Phase 3.
Because of previous problems with the pump units picking up debris from the harbour wall a bracket was fabricated to hold the two pump units and it was suspended between opposite walls to hold the units out in open water.
4.5.4 Manifold
A new manifold system was built and installed on the outside of the foundation (Figure 4.37). The purpose of the manifold is to allow variations in the control of the piping systems shown above but using the 5 main control valves situated at the upper level of the test site.
Figure 4.37: Jet Manifold System
The water supply from the 5 valve controlled zones comes in from the top. The yellow levered valves can be oriented to divert, for example, the first four zones to the corresponding four agitator zones with the fifth zone controlling all the radial pipes. (In this photo all valves are actually closed). The situation of the control valves on the outside allowed them to be altered during a test if required.
In addition to this manifold another addition was a 200mm diameter inlet pipe leading from the outside of the foundation below water level to a central hole under the foundation. This was fitted to allow water to be entrained in the radial outward flow of slurry if the flows and pressures allowed. This pipe was fitted with a removable blanking plate to allow the pipe to be closed if necessary.
4.5.5 Test 4
4.5.5.1 Description Test set-up:
• Upper pump manifold Zones 1-4 supplying Radial jets 1-4
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• Upper manifold Zone 5 supplying all Agitator jets• Central 200mm supply pipe open
4.5.5.2 ResultsThe test started well, but it was apparent that the pressure build up under the foundation was sufficient to cause a significant proportion of the introduced water to flow out of the central hole rather than the other way round. After 20 minutes it was decided to stop the test. The residual sand pattern is shown in Figure 4.38.
Figure 4.38: Test 4 - Residual Sand Pattern
It is evident from the photo above that the total fluidisation in the centre is more than required, it was decided that some of the jets in the centre would be blocked to allow more water jet power to the perimeter jets (Figure 4.36(b)).
4.5.6 Test 5
4.5.6.1 Description Test set-up:
• Upper pump manifold Zones 1-4 supplying Radial jets 1-4• Upper manifold Zone 5 supplying all Agitator jets• Central 200mm supply pipe closed
4.5.6.2 ResultsThe residual sand distribution is shown in Figure 4.39. Note how the sand has been graded (Fig 4.39(a)). It is thought that the larger stones are moved closer to the foundation than the lighter sand particles and roll back down to the eductor pump and are recirculated. Eventually this builds up to a concentration of stones. This is a function of the available space outside the foundation to deposit this material.
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(a) Eductor Produced Sand Pile (b) Residual pattern below foundationFigure 4.39: Test 5 - Residual sand distribution
Insertion rate and tilt angle against time are shown in Figure 4.41 and 4.42 respectively.
Time, minutes
Figure 4.40: Test 5 - Insertion Rate vs Time
The test was largely successful and the foundation was buried deeper and faster than on any previous attempt.
As with all the test it was possible to alter the level of the foundation as it was buried by altering the flow to the various zones.
The test was terminated when eductor pumps were unable to remove sand from the perimeter and the burying rate slowed to an extent that no more useful information was being generated.
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2.000
Time, minutes
-------Xaxis------- Y axis
Figure 4.41:Test 5 - Tilt vs Time
4.5.7 Test 6
4.5.7.1 Description Test set-up:
• Upper pump manifold Zones 1-4 supplying Radial jets 1-4 and Agitator jets 1-4• Central 200mm supply pipe closed
On this test it was decided to start with the 4 control zones connected both to the Radial and to the Agitator jets. It was thought that this may give more control to the zones and promote sand movement in areas that may ‘stick’.
4.5.7.2 ResultsThe test proceeded well and the initial insertion rate was very good.As with previous tests the operation of the eductors was a little erratic, sometimes clogging and then freeing again. Figure 4.43 and 4.44 give plots of the insertion depth and tilt angle respectively.
As illustrated by the tilt angle graph, the foundation became stuck on one side. The eductor on this side was found to have picked up pieces of gravel just too large to go through the venturi.
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Time, minutes
Figure 4.42: Test 6 - Insertion Depth vs Time
1.500
1.000
0.500
< 0.000
-0.500
-1.000
Time, minutes
|1 X axis Y axis |
Figure 4.43: Test 6 - Tilt Angle vs Time
4.5.8 Test 7
4.5.8.1 Description
In order to investigate the effect of large cobbles on the burial characteristics it was decided to repeat Test 6 with large rocks placed under the foundation (Figure 4.44). It was hoped that the fluidisation would allow these rocks to settle as burial progressed
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(a) Rocks used under foundation (b) Distribution of rocks prior to testFigure 4.44: Test 7 - Simulation of ‘cobbly’ seabed
4.5.8.2 Results
The residual sand pattern is shown in Figure 4.45. Some of the rocks are evident, some have been buried fairly deeply. This is evidence of the difference between the static picture after the foundation is removed and the dynamic fluidisation effect that is actually taking place during the burial process.
Figure 4.45: Test 7 - Residual sand Pattern
Overall embedment was approximately 450mm although the rate of embedment was marginally slower than for Test 6 - overall it was clear that the cobbles did not significantly inhibit the burial process.
4.5.9 Test 8
4.5.9.1 DescriptionAs the last test in the series it was decided to try the best solutions to achieve the best possible burial of the foundation.
The set up for the test was:
• Upper pump manifold Zones 1-4 supplying Radial jets 1-4 and Agitator jets 1-4.
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• Central 200mm supply pipe closed.• The inductor inlet was modified.
4.5.9.2 ResultsThe residual sand patterns following drain-down are shown in Figure 4.46:
a) General view of buried foundation (b) Detail of eductor mound
(c) Burial Pit (d) Residual sand patternFigure 4.46: Test 7 - Residual Sand Distribution
It appears that the total amount of material deposited outside the foundation by the eductors was creating an overburden effect that was caused by the limited size of the sand tank. Within the limitations of the test tank it would be difficult to bury the foundation further unless the eductor discharged outside the tank. This was not possible on this site as the sand would have washed into the dock.
Figures 4.47 and 4.48 show insertion depth and tilt respectively against time.
The burial achieved was approximately 700mm in 90 minutes (burial was continuing at this point albeit at a slow rate)- this is the best result of all the tests and represents the significant improvement in the self burial technology over the duration of the test programme. Level control across the foundation was also maintained.
Figure 4.47: Insertion Depth vs Time
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Figure 4.48: Tilt vs Time
4.5.10 Conclusions from Phase 4
• This test series has shown that the principle of water jetting under a loaded foundation can create a self burial effect that is controllable in rate and level.
• The evolution of the jets, zones and pumps has allowed great improvements in the performance of the burial rate. It is thought that in large scale foundations a rate of burial greater than this would be difficult to control.
• It is possible with the data from these tests to extrapolate the requirements of pumps and jets for a much larger foundation.
• Many of the problems encountered in this series were to do with eductor performance. Two issues here have been identified:
1. The throat of the eductor must be large enough to take the largest stones that can be sucked into the eductor. A suitably sized mesh guard over the inlet could easily be added.
2. The inlet tube caused blockage by ramming of un-fluidised sand. This could be remedied by designing an eductor without an inlet tube but with the jet itself placed lower. In addition the inclusion of a localised eductor fluidisation jet would ensure that at all times there was a pool of fluidised material at the mouth of the eductor.
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5 ECOMONIC ASSESSMENT OF INTEGRATED INSTALLATION
Table 5.1 presents a process and cost comparison for an integrated installation OWEC (gravity type) and a conventionally installed OWEC (gravity type).
Table 5.1: Process and Cost Comparison (2MW turbine, 30 unit campaign)Integrated Installation Conventional Installation
Process Cost Estimate (£k)
Process Cost Estimate (£k)
Manufacture turbine 1055 Manufacture turbine 1055Fabricate foundation 250 Fabricate foundation 250Self burial pipework 20 Self burial pipework N/ATelescopic tower 424 Tower 282Quayside Assembly 42.5 Quayside Assembly N/ALoadout / transportation / setdown of integrated unit
370 Loadout of foundation 100Loadout of tower and turbine 68
Self burial of foundation 5 Seabed preparation 20Scour protection 20
Iron ore fill 40 Iron ore fill 40Raise Tower 26.5 Offshore installation of tower and
turbine68
Grid Connection 400 Grid Connection 400Total 2633 Total 2303
It can be seen that for a 2MW OWEC, 30 unit campaign, the integrated installation concept is approximately £300k more expensive than ‘conventional’ methods, and is therefore not demonstrably competitive (whilst re-stating the proviso that the installation conditions on which this study is based are more onerous than those for which case-study data exists).
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6 DISCUSSION AND RECOMMENDATIONS
The aim of the project was to demonstrate the viability of the integrated installation concept. This aim has been pursued on two fronts - technical viability and economic viability. Technical viability has been demonstrated for all three components, however the technical solution for both transportation and telescopic tower elements have been shown to be uneconomic. Further
development of the integrated concept as presented in this report cannot be recommended.
However, this research has resulted in findings which should prove extremely valuable in non- integrated or partially integrated schemes, e.g. installation of a telescopic tower on a pre-installed foundation, or transportation and installation of the foundation component only. The self burial technology has wide application in most types of deep water sand/gravel founded structure, beyond simply the renewable energy sector.
Time and budgetary constraints naturally limit the number of alternative options that may be investigated. Potentially economic systems may be suggested by a retrospective review of an uneconomic study.
For example, the telescopic tower study demonstrated the technical feasibility of the heat shrink mechanism, and of a guidance mechanism utilising flexible packers. By combining the packer system and heat shrink mechanism together to resist the applied moments as a couple, it may be possible to significantly reduce the size of the heat-shrink ring.
Similarly for the transportation study, following a logical development path based on ‘simple’ mechanisms and ‘conventional’ maritime practices has led to increasingly load attracting, rigid, and ultimately uneconomic structures. The ultimate conclusion is that the more unconventional concept of a ‘slotted’ barge and/or the suspension of the unit for a universal joint may be significantly more economic - and therefore warrant further investigation.
The potential of the self burial system has been demonstrated beyond question. The combined agitator/radial jet system should be readily scalable to provide coverage of the significantly larger area of the full size foundation. This needs to be demonstrated.
One potential issue with the self burial system as proposed is the control / removal of the ‘spoil’ from around the perimeter of the foundation. Judicious distribution / alignment of eductors should overcome this problem - infact the spoil could be used to beneficial effect as overburden and scour protection.
A possible, albeit radical, alternative would be to replace the numerous perimeter eductors with a large central eductor (such those use in marine dredging operations). This system eliminates the need to remove spoil from the perimeter of the foundation and may improve the definition of the burial pit. However it may be more difficult to ensure that slurrified sand is transported from near to the edges of the foundation to the central eductor. A medium scale test programme would resolve these issues.
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REFERENCES
1. BS 2573 Part 2: Rules for the design of Cranes - Part 2: Specification for classification, stress calculations and design of mechanisms, 1980
2. Cost Optimisation of Wind Turbines for Large-scale Offshore Wind Farms. Riso National Laboratory, Denmark, 1998
3. Offshore Wind Energy - Building a New Industry for Britain. Greenpeace
4. British Wind Energy Website: www.britishwindenergy.com
5. Code for Lifting Appliances in a Marine Environment. Lloyds Register, January 1987
6. Rules and Regulations for the Classification of Ships. Lloyds Register, January 1998
7. Sea-bed sediments around the UK: bathymetric, physical environment, grain size, mineral composition and associated bedforms. Research Report SB/90/01, Offshore Geology Series, British Geological Survey
8. Sea-bed sediments around the UK. North & South shetts (pair) Scale 1:1000000. Offshore Geological Series, British Geological Survey.
9. Sand Bed Response to Submerges Water Jet - T. O’Donoghue and B. Trajkovic. Proc. 11th Int. Offshore and Polar Engineering Conf. Stavanger, Norway, June 17-22, 2001.
10. Erosion of loose beds by submerged circular impinging vertical turbulent jets - O. O. Aderibigbe and N Rajaratnam. J. Hydr. Research, Vol. 34, 1996
11. Fluidizer System Design for Channel Maintenance and Sand Bypassing - US Army Engineer Waterways Experiment Station
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ACKNOWLEDGEMENTS
Corns wish to acknowledge the significant contribution of the following individuals and organisations:
Telescopic Wind Tower Study:Paul Leonard, John Crossling - Corus Northern Engineering Services
Transportation Study:Chris Colling, Simon Bennett - Global Maritime Consultancy Ltd
Self Burial Study:Rob Ellis - Offshore Data LtdLucas Boissevain - Mustang Marine (Wales) Ltd
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APPENDIX A: INTEGRATED INSTALLATION CONCEPT DRAWING
Drawing Name Drawing NumberOWEC Integrated Installation Concept Drawing BIS152 / DRG / 90000 / A
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APPENDIX B: TELESCOPIC TOWER TEST RIG DRAWINGS
Drawing Name Drawing NumberTelescopic Windtower - Detail of Scale Model1 : 4.5
EX103910-12
Test Rig For Telescopic WindtowerGeneral Arrangement & Details of Test Rig for Telescopic Wind Tower
SD / Y24 / 03 / 01
Test Rig For Telescopic WindtowerDetails of Test Rig for Telescopic Wind Tower
SD / Y24 / 03 / 02
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APPENDIX C: HORN / STRANDJACK GENERAL ARRANGEMENT
Drawing Name Drawing NumberHorn / Strandjack ArrangementSheet 1 of 4
GM - 44339 - 001
Horn / Strandjack ArrangementSheet 2 of 4
GM - 44339 - 001
Horn / Strandjack ArrangementSheet 3 of 4
GM - 44339 - 001
Horn / Strandjack ArrangementSheet 4 of 4
GM - 44339 - 001
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