STRUCTURAL SYSTEMS: OFFSHORE A.1 : Offshore Structures: General Introduction A.2 : Loads (I) : Introduction and Environmental Loads A.3 : Loads (II) - Other Loads A.4 : - Analysis I A.5 : - Analysis II A.6 : Foundations A.7 : Tubular Joints in Offshore Structures A.8 : Fabrication A.9 : Installation A.10 : Superstructures I .11 : - Superstructures II .12 : Connections in Offshore Deck Structures
A.1: Offshore Structures: General Introduction OBJECTIVE/SCOPE To identify the basic vocabulary, to introduce the major concepts for offshore platform structures, and to explain where the basic structural requirements for design are generated. PREREQUISITES None. SUMMARY The lecture starts with a presentation of the importance of offshore hydro-carbon exploitation, the basic steps in the development process (from seismic exploration to platform removal) and the introduction of the major structural concepts (jacket-based, GBS-based, TLP, floating). The major codes are identified. For the fixed platform concepts (jacket and GBS), the different execution phases are briefly explained: design, fabrication and installation. Special attention is given to some principles of topside design. A basic introduction to cost aspects is presented. Finally terms are introduced through a glossary. 1. INTRODUCTION Offshore platforms are constructed to produce the hydrocarbons oil and gas. The contribution of offshore oil production in the year 1988 to the world energy consumption was 9% and is estimated to be 24% in 2000. The investment (CAPEX) required at present to produce one barrel of oil per day ($/B/D) and the production costs (OPEX) per barrel are depicted in the table below. Condition
CAPEX $/B/D
OPEX $/B
Average
4000 - 8000
5
Middle East
500 - 3000
1
Non-Opec
3000 - 12000
8
North Sea
10000 - 25000
5 - 10
Deepwater
15000 - 35000
10 - 15
Conventional
Offshore
World oil production in 1988 was 63 million barrel/day. These figures clearly indicate the challenge for the offshore designer: a growing contribution is required from offshore exploitation, a very capital intensive activity.
Figure 1 shows the distribution of the oil and gas fields in the North Sea, a major contribution to the world offshore hydrocarbons. It also indicates the onshore fields in England, the Netherlands and Germany.
2. OFFSHORE PLATFORMS 2.1 Introduction of Basic Types The overwhelming majority of platforms are piled-jacket with deck structures, all built in steel (see Slides 1 and 2).
Slide 1 : Jacket based platform - Southern sector North Sea
Slide 2 : Jacket based platform - Northern sector North Sea A second major type is the gravity concrete structure (see Figure 2), which is employed in the North Sea in the Norwegian and British sectors.
A third type is the floating production unit. 2.2 Environment The offshore environment can be characterized by:
water depth at location soil, at seabottom and in-depth wind speed, air temperature waves, tide and storm surge, current ice (fixed, floes, icebergs) earthquakes (if necessary)
The topside structure also must be kept clear of the wave crest. The clearance (airgap) usually is taken at approximately 1,50 m, but should be increased if reservoir depletion will create significant subsidence.
2.3 Construction The environment as well as financial aspects require that a high degree of prefabrication must be performed onshore. It is necessary to design to limit offshore work to a minimum. The overall cost of a man-hour offshore is approximately five times that of an onshore man-hour. The cost of construction equipment required to handle loads, and the cost for logistics are also a magnitude higher offshore. These factors combined with the size and weight of the items, require that a designer must carefully consider all construction activities between shop fabrication and offshore installation. 2.4 Codes Structural design has to comply with specific offshore structural codes. The worldwide leading structural code is the APIRP2A [1]. The recently issued Lloyds rules [2] and the DnV rules [3] are also important. Specific government requirements have to be complied with, e.g. in the rules of Department of Energy (DoE), Norwegian Petroleum Direktorate (NPD). For the detail design of the topside structure the AISC-code [4] is frequently used, and the AWS-code [5] is used for welding. In the UK the Piper alpha diaster has led to a completely new approach to regulation offshore. The responsibility for regulatory control has been moved to the Health and Safety Executive (HSE) and the operator has to produce a formal safety assessment (TSA) himself instead of complying with detailed regulations. 2.5 Certification and Warranty Survey Government authorities require that recognized bodies appraise the aspects of structural integrity and issue a certificate to that purpose. The major certification bodies are:
Det norske Veritas (DnV) Lloyds Register of Shipping (LRS) American Bureau of Shipping (ABS) Bureau Veritas (BV) Germanischer Lloyd (GL)
Their requirements are available to the designer [2, 3, 6, 7, 8]. Insurance companies covering transport and installation require the structures to be reviewed by warranty surveyors before acceptance. The warranty surveyors apply standards, if available, on a confidential basis. 3. OFFSHORE DEVELOPMENT OF AN OIL/GAS FIELD 3.1 Introduction The different requirements of an offshore platform and the typical phases of an offshore development are summarized in [9]. After several initial phases which include seismic field surveying, one or more exploration wells are drilled. Jack-up drilling rigs are used for this purpose for water depths up to 100 - 120 m; for deeper water floating rigs are used. The results are studied and the economics and risks of different development plans are evaluated. Factors involved in the evaluation may include number of wells required, fixed or floated production facilities, number of such facilities, and pipeline or tanker offloading. As soon as exploitation is decided and approved, there are four main technical activities, prior to production:
engineering and design
fabrication and installation of the production facility drilling of production wells, taking 2 - 3 months/well providing the off loading system (pipelines, tankers, etc.).
The drilling and construction interaction is described below for two typical fixed platform concepts. 3.2 Jacket Based Platform for Shallow Water First the jacket is installed. The wells are then drilled by a jack-up drilling unit standing close by with a cantilever rig extending over the jacket. Slide 3 shows a jack-up drilling unit with a cantilever rig. (In this instance it is engaged in exploratory drilling and is therefore working in isolation.)
Slide 3 : Cantilevered drilling rig: Self-elevating (jack-up) exploration drilling platform. Design and construction of the topside are progressed parallel to the drilling, allowing production to start soon after deck installation. For further wells, the jack-up drilling unit will be called once again and will reach over the well area of the production deck. As an alternative to this concept the wells are often accommodated in a separate wellhead platform, linked by a bridge to the production platform (see Slide 1). 3.3 Jacket and Gravity Based Platform for Deep Water The wells are drilled from a drilling rig on the permanent platform (see Slide 2). Drilling starts after the platform is built and completely installed. Consequently production starts between one and two years after platform installation. In recent years pre-drilled wells have been used to allow an earlier start of the production. In this case the platform has to be installed exactly above the pre-drilled wells. 4. JACKETS AND PILE FOUNDATION 4.1 Introduction Jackets, the tower-like braced tubular structures, generally perform two functions:
They provide the substructure for the production facility (topside), keeping it stable above the waves. They support laterally and protect the 26-30 inch well conductors and the pipeline riser.
The installation methods for the jacket and the piles have a profound impact on the design. 4.2 Pile Foundation The jacket foundation is provided by open-ended tubular steel piles, with diameters up to 2m. The piles are driven approximately 40 - 80 m, and in some cases 120 m deep into the seabed. There are basically three types of pile/jacket arrangement (see Figure 3):
Pile-through-leg concept, where the pile is installed in the corner legs of the jacket. Skirt piles through pile sleeves at the jacket-base, where the pile is installed in guides attached to the jacket leg. Skirt piles can be grouped in clusters around each of the jacket legs. Vertical skirt piles are directly installed in the pile sleeve at the jacket base; all other guides are deleted. This arrangement results in reduced structural weight and easier pile driving. In contrast inclined piles enlarge the foundation at the bottom, thus providing a stiffer structure. 4.3 Pile Bearing Resistance Axial load resistance is required for bearing as well as for tension. The pile accumulates both skin friction as well as end bearing resistance. Lateral load resistance of the pile is required for restraint of the horizontal forces. These forces lead to significant bending of the pile near to the seabed. Number, arrangement, diameter and penetration of the piles depend on the environmental loads and the soil conditions at the location. 4.4 Corrosion Protection The most usual form of corrosion protection of the bare underwater part of the jacket as well as the upper part of the piles in soil is by cathodic protection using sacrificial anodes. A sacrificial anode (approximate 3 kN each) consists of a
zinc/aluminium bar cast about a steel tube and welded on to the structures. Typically approximately 5% of the jacket weight is applied as anodes. The steelwork in the splash zone is usually protected by a sacrificial wall thickness of 12 mm to the members. 5. TOPSIDES 5.1 Introduction The major functions on the deck of an offshore platform are:
well control support for well work-over equipment separation of gas, oil and non-transportable components in the raw product, e.g. water, parafines/waxes and sand support for pumps/compressors required to transport the product ashore power generation accommodation for operating and maintenance staff.
There are basically two structural types of topside, the integrated and modularized topside which are positioned either on a jacket or on a concrete gravity substructure. 5.2 Jacket-based Topsides 5.2.1 Concepts There are four structural concepts in practice. They result from the lifting capacity of crane vessels and the load-out capacity at the yards:
the single integrated deck (up to approx 100 MN) the split deck in two four-leg units the integrated deck with living quarter module the modularized topside consisting of module support frame (MSF) carrying a series of modules.
Slide 4 shows an integrated deck (though excluding the living quarters and helideck) being moved from its assembly building.
Slide 4 : Integrated topside during load out
5.2.2 Structural Design for Integrated Topsides For the smaller decks, up to approximately 100 MN weight, the support structure consists of trusses or portal frames with deletion of diagonals. The moderate vertical load and shear per column allows the topside to be supported by vertical columns (deck legs) only, down to the top of the piles (situated at approximately +4 m to +6 m L.A.T. (Low Astronomic Tide). 5.2.3 Structural Design for Modularized Jacket-based Topsides A major modularized topside weighs 200 to 400 MN. In this case the MSF is a heavy tubular structure (Figure 4), with lateral bracing down to the top of jacket.
5.3 Structural Design for Modularized Gravity-based Topsides The topsides to be supported by a gravity-based substructure (see Figure 2) are in a weight range of 200 MN up to 500 MN.
The backbone of the structure is a system of heavy box-girders with a height of approximately 10 m and a width of approximately 12 - 15 m (see Figure 5).
The substructure of the deck is rigidly connected to the concrete column and acts as a beam supporting the deck modules. This connection introduces wave-induced fatigue in the deck structure. A recent development, foreseen for the Norwegian Troll platform, is to provide a flexible connection between the deck and concrete column, thus eliminating fatigue in the deck [10]. 6. EQUIPMENT AND LIVING QUARTER MODULES Equipment modules (20-75 MN) have the form of rectangular boxes with one or two intermediate floors. The floors are steel plate (6, 8 or 10 mm thick) for roof and lower floor, and grating for intermediate floors. In living quarter modules (5-25 MN) all sleeping rooms require windows and several doors must be provided in the outer walls. This requirement can interfere seriously with truss arrangements. Floors are flat or stiffened plate. Three types of structural concepts, all avoiding interior columns, can be distinguished:
conventional trusses in the walls. stiffened plate walls (so called stressed skin or deck house type). heavy base frame (with wind bracings in the walls).
7. CONSTRUCTION 7.1 Introduction The design of offshore structures has to consider various requirements of construction relating to: 1.
fabrication.
2. 3. 4. 5. 6. 7. 8.
weight. load-out. sea transport. offshore installation. module installation. hook-up. commissioning.
A documented construction strategy should be available during all phases of the design and the actual design development should be monitored against the construction strategy. Construction is illustrated below by four examples. 7.2 Construction of Jackets and Topsides 7.2.1 Lift Installed Jackets The jacket is built in the vertical (smaller jackets) or horizontal position (bigger jackets) on a quay of a fabrication site. The jacket is loaded-out and seafastened aboard a barge. At the offshore location the barge is moored alongside an offshore crane vessel. The jacket is lifted off the barge, upended from the horizontal, and carefully set down onto the seabed. After setting down the jacket, the piles are installed into the sleeves and, driven into the seabed. Fixing the piles to the jacket completes the installation. 7.2.2 Launch Installed Jackets The jacket is built in horizontal position. For load-out to the transport barge, the jacket is put on skids sliding on a straight track of steel beams, and pulled onto the barge (Slide 5).
Slide 5 : Jacket being loaded onto barge by skidding At the offshore location the jacket is slid off the barge. It immerses deeply into the water and assumes a floating position afterwards (see Figure 6).
Two parallel heavy vertical trusses in the jacket structure are required, capable of taking the support reactions during launching. To reduce forces and moments in the jacket, rocker arms are attached to the stern of the barge.
The next phase is to upright the jacket by means of controlled flooding of the buoyancy tanks and then set down onto the seabed. Self-upending jackets obtain a vertical position after the launch on their own. Piling and pile/jacket fixing completes the installation. 7.2.3 Topsides for a Gravity-Based Structure (GBS) The topside is assembled above the sea on a temporary support near a yard. It is then taken by a barge of such dimensions as to fit between the columns of the temporary support and between the columns of the GBS. The GBS is brought in a deep floating condition in a sheltered site, e.g. a Norwegian fjord. The barge is positioned between the columns and the GBS is then deballasted to mate with and to take over the deck from the barge. The floating GBS with deck is then towed to the offshore site and set down onto the seabed. 7.2.4 Jacket Topsides For topsides up to approximately 120 MN, the topside may be installed in one lift. Slide 6 shows a 60 MN topside being installed by floating cranes.
Slide 6 : Installation of 60MN K12-BP topside by floating crane For the modularized topside, first the MSF will be installed, immediately followed by the modules. 7.3 Offshore Lifting Lifting of heavy loads from barges (Slide 6) is one of the very important and spectacular construction activities requiring a focus on the problem when concepts are developed. Weather windows, i.e. periods of suitable weather conditions, are required for these operations. 7.3.1 Crane Vessel Lifting of heavy loads offshore requires use of specialized crane vessels. Figure 7 provides information on a typical big, dual crane vessel. Table 1 (page 16) lists some of the major offshore crane vessels.
7.3.2 Sling-arrangement, Slings and Shackles For lifting, steel wire ropes in a four-sling arrangement are used which directly rest in the four-point hook of the crane vessel, (see Figure 8). The heaviest sling available now has a diameter of approximately 350 mm, a breaking load of approximately 48 MN, and a safe working load (SWL) of 16 MN. Shackles are available up to 10 MN SWL to connect the padeyes installed at the module's columns. Due to the space required, connecting more than one shackle to the same column is not very attractive. So when the sling load exceeds 10 MN, padears become an option.
Table 1 Major Offshore Crane Vessels Operator
Name
Mode
Type
Lifting capacity (Tonnes)
Heerema
Thor
Monohull
Fix
2720
Rev
1820
Fix
2720
Rev
2450
Fix
4536 + 3628 = 8164
Rev
3630 + 2720 = 6350
Fix
3630 + 2720 = 6350
Rev
3000 + 2000 = 5000
Fix
4000
Odin
Hermod
Balder
McDermott DB50
Monohull
Semisub
Semisub
Monohull
DB100
DB101
Semisub
Semisub
Rev
3800
Fix
1820
Rev
1450
Fix
3360
Rev
2450
DB102
Semisub
Rev
6000 + 6000 = 12000
Micoperi
M7000
Semisub
Rev
7000 + 7000 = 14000
ETPM
DLB1601 Monohull
Rev.
1600
Notes: 1. 2.
Rated lifting capacity in metric tonnes. When the crane vessels are provided with two cranes, these cranes are situated at the vessels stern or bow at approximately 60 m distance c.t.c.
1.
3. Rev = Load capability with fully revolving crane.
Fix = Load capability with crane fixed. 7.4 Sea Transport and Sea Fastening Transportation is performed aboard a flat-top barge or, if possible, on the deck of the crane vessel. The module requires fixing to the barge (see Figure 9) to withstand barge motions in rough seas. The sea fastening concept is determined by the positions of the framing in the module as well as of the "hard points" in the barge.
7.5 Load-out 7.5.1 Introduction For load-out three basic methods are applied:
skidding platform trailers shearlegs.
7.5.2 Skidding Skidding is a method feasible for items of any weight. The system consists of a series of steel beams, acting as track, on which a group of skids with each approximately 6 MN load capacity is arranged. Each skid is provided with a hydraulic jack to control the reaction. 7.5.3 Platform Trailers Specialized trailer units (see Figure 10) can be combined to act as one unit for loads up to 60 - 75 MN. The wheels are individually suspended and integrated jacks allow adjustment up to 300 mm.
The load capacity over the projected ground area varies from approximately 55 to 85 kN/sq.m. The units can drive in all directions and negotiate curves. 7.5.4 Shearlegs Load-out by shearlegs is attractive for small jackets built on the quay. Smaller decks (up to 10 - 12 MN) can be loaded out on the decklegs pre-positioned on the barge, thus allowing deck and deckleg to be installed in one lift offshore. 7.6 Platform Removal In recent years platform removal has become common. The mode of removal depends strongly on the regulations of the local authorities. Provision for removal should be considered in the design phase. 8. STRUCTURAL ANALYSIS 8.1 Introduction The majority of structural analyses are based on the linear theory of elasticity for total system behaviour. Dynamic analysis is performed for the system behaviour under wave-attack if the natural period exceeds 3 seconds. Many elements can exhibit local dynamic behaviour, e.g. compressor foundations, flare-stacks, crane-pedestals, slender jacket members, conductors. 8.2 In-place Phase Three types of analysis are performed:
Survival state, under wave/current/wind attack with 50 or 100 years recurrence period. Operational state, under wave/current/wind attack with 1 or 5 years recurrence period, under full operation. Fatigue assessment. Accidental.
All these analyses are performed on the complete and intact structure. Assessments at damaged structures, e.g. with one member deleted, and assessments of collision situations are occasionally performed. 8.3 Construction Phase The major phases of construction when structural integrity may be endangered are:
Load-out Sea transport Upending of jackets Lifting.
9. COST ASPECTS 9.1 Introduction The economic feasibility of an offshore project depends on many aspects: capital expenditure (CAPEX), tax, royalties, operational expenditure (OPEX). In a typical offshore field development, one third of the CAPEX is spent on the platform, one third on the drilling of wells and one third on the pipelines.
Cost estimates are usually prepared in a deterministic approach. Recently cost-estimating using a probabilistic approach has been developed and adopted in major offshore projects. The CAPEX of an installed offshore platform topside amounts to approximately 20 ECU/kg. 9.2 Capital Expenditure (CAPEX) The major elements in the CAPEX for an offshore platform are:
project management and design material and equipment procurement fabrication transport and installation hook-up and commissioning.
9.3 Operational Expenditure (OPEX) In the North Sea approximately 20 percent of OPEX are required for offshore inspection, maintenance and repair (IMR). The amount to be spent on IMR over the project life can add up to approximately half the original investment. IMR is the area in which the structural engineer makes a contribution by effort in design, selection of material, improved corrosion protection, accessibility, basic provisions for scaffolding, avoiding jacket attachments dangerous to divers, etc. 10. DEEP WATER DEVELOPMENTS Deep water introduces a wide range of extra difficulties for the operator, the designer and constructor of offshore platforms. Fixed platforms have recently been installed in water of 410 m. depth, i.e. "Bullwinkle" developed by Shell Oil for a Gulf of Mexico location. The jacket weighed nearly 500 MN. The maximum depth of water at platform sites in the North Sea is approximately 220 m at present. The development of the Troll field situated in approximately 305 m deep water is planned for 1993. In the Gulf of Mexico and offshore California several fixed platforms in water depths of 250 - 350 m are in operation (Cerveza, Cognac). Exxon has a guyed tower platform (Lena) in operation in 300 m deep water. An option for deeper locations is to use subsea wells with flowlines to a nearby (approximately maximum 10 km) fixed platform at a smaller water depth. Alternatively subsea wells may be used with flexible risers to a floating production unit. Subsea wells are now feasible for 300 - 900 m deep water. The deepest wells have been developed off Brasil in moderate weather conditions. The tension leg platform (TLP) seems to be the most promising deepwater production unit (Figure 11). It consists of a semisubmersible pontoon, tied to the seabed by vertical prestressed tethers. The first TLP was Hutton in the North Sea and recently TLP-Jolliet was installed at a 530 m deep location in the Gulf of Mexico. Norwegian Snorre and Heidrun fields have been developed with TLPs as well.
11. CONCLUDING SUMMARY
The lecture starts with the presentation of the importance of offshore hydro-carbon exploitation, the basic steps in the development process (from seismic exploration to platform removal) and the introduction of the major structural concepts (jacket-based, GBS-based, TLP, floating). The major codes are identified.
For the fixed platform concepts (jacket and GBS), the different execution phases are briefly explained: design, fabrication and installation. Special attention is given to the principles of topside design. A basic introduction to cost aspects is presented. Finally terms are introduced within a glossary.
12. GLOSSARY OF TERMS AIR GAP Clearance between the top of maximum wave and underside of the topside. CAISSONS See SUMPS CONDUCTORS The tubular protecting and guiding the drill string from the topside down to 40 to 100m under the sea bottom. After drilling it protects the well casing. G.B.S. Gravity based structure, sitting flatly on the sea bottom, stable through its weight. HOOK-UP Connecting components or systems, after installation offshore. JACKET Tubular sub-structure under a topside, standing in the water and pile founded. LOAD-OUT The operation of bringing the object (module, jacket, deck) from the quay onto the transportation barge. PADEARS (TRUNNIONS) Thick-walled tubular stubs, directly receiving slings and transversely welded to the main structure. PADEYES Thick-walled plate with hole, receiving the pin of the shackle, welded to the main structure. PIPELINE RISER The piping section which rises from the sea bed to topside level. SEA-FASTENING The structure to keep the object rigidly connected to the barge during transport. SHACKLES Connecting element (bow + pin) between slings and padeyes. SLINGS Cables with spliced eyed at both ends, for offshore lifting, the upper end resting in the crane hook. SPREADER Tubular frame, used in lifting operation. SUBSEA TEMPLATE Structure at seabottom, to guide conductors prior to jacket installation. SUMPS Vertical pipes from topside down to 5-10 m below water level for intake or discharge. TOPSIDE Topside, the compact offshore process plant, with all auxiliaries, positioned above the waves. UP ENDING Bringing the jacket in vertical position, prior to set down on the sea bottom. WEATHER WINDOW A period of calm weather, defined on basis of operational limits for the offshore marine operation. WELLHEAD AREA Area in topside where the wellheads are positioned including the valves mounted on its top. 13. REFERENCES [1] API-RP2A: Recommended practice for planning, designing and constructing fixed offshore platforms.
American Petroleum Institute 18th ed. 1989. The structural offshore code, governs the majority of platforms. [2] LRS Code for offshore platforms. Lloyds Register of Shipping. London (UK) 1988. Regulations of a major certifying authority. [3] DnV: Rules for the classification of fixed offshore installations. Det Norske Veritas 1989. Important set of rules. [4] AISC: Specification for the design, fabrication and erection of structural steel for buildings. American Institute of Steel Construction 1989. Widely used structural code for topsides. [5] AWS D1.1-90: Structural Welding Code - Steel. American Welding Society 1990. The structural offshore welding code. [6] DnV/Marine Operations: Standard for insurance warranty surveys in marine operations. Det norske Veritas June 1985. Regulations of a major certifying authority. [7] ABS: Rules for building and classing offshore installations, Part 1 Structures. American Bureau of Shipping 1983. Regulations of a major certifying authority. [8] BV: Rules and regulations for the construction and classification of offshore platforms. Bureau Veritas, Paris 1975. Regulations of a major certifying authority. [9] ANON: A primer of offshore operations. Petex Publ. Austin U.S.A 2nd ed. 1985. Fundamental information about offshore oil and gas operations.
[10] AGJ Berkelder et al: Flexible deck joints. ASME/OMAE-conference The Hague 1989 Vol.II pp. 753-760. Presents interesting new concept in GBS design. 14. ADDITIONAL READING 1.
BS 6235: Code of practice for fixed offshore structures. British Standards Institution 1982. Important code, mainly for the British offshore sector.
2.
DoE Offshore installations: Guidance on design and construction, U.K. Department of Energy 1990. Governmental regulations for British offshore sector only.
3.
UEG: Design of tubular joints (3 volumes). UEG Offshore Research Publ. U.R.33 1985. Important theoretical and practical book.
4.
J. Wardenier: Hollow section joints. Delft University Press 1981. Theoretical publication on tubular design including practical design formulae.
5.
ARSEM: Design guides for offshore structures welded tubular joints. Edition Technip, Paris (France), 1987. Important theoretical and practical book.
6.
D. Johnston: Field development options. Oil & Gas Journal, May 5 1986, pp 132 - 142. Good presentation on development options.
7.
G. I. Claum et al: Offshore Structures: Vol 1: Conceptual Design and Hydri-mechanics; Vol 2 - Strength and Safety for Structural design. Springer Verlag, London 1992. Fundamental publication on structural behaviour.
8.
W.J. Graff: Introduction to offshore structures. Gulf Publishing Company, Houston 1981.
Good general introduction to offshore structures. 9.
B.C. Gerwick: Construction of offshore structures. John Wiley & Sons, New York 1986. Up to date presentation of offshore design and construction.
10. T.A. Doody et al: Important considerations for successful fabrication of offshore structures. OTC paper 5348, Houston 1986, pp 531-539. Valuable paper on fabrication aspects. 11. D.I. Karsan et al: An economic study on parameters influencing the cost of fixed platforms. OTC paper 5301, Houston 1986, pp 79-93. Good presentation on offshore CAPEX assessment
A.2: Loads (I) : Introduction and Environmental Loads OBJECTIVE/SCOPE To introduce the types of loads for which a fixed steel offshore structure must be designed. To present briefly the loads generated by environmental factors. PREREQUISITES A basic knowledge of structural analysis for static and dynamic loadings. SUMMARY The categories of load for which a pile supported steel offshore platform must be designed are introduced and then the different types of environmental loads are presented. The loads include: wind, wave, current, earthquake, ice and snow, temperature, sea bed movement, marine growth and tide generated loads. Loads due to wind, waves and earthquake are discussed in more detail together with their idealizations for the various types of analyses. Frequent references are made to the codes of practice recommended by the American Petroleum Institute, Det Norske Veritas, the British Standards Institution and the British Department of Energy, as well as to the relevant regulations of the Norwegian Petroleum Directorate. 1. INTRODUCTION The loads for which an offshore structure must be designed can be classified into the following categories: 1. 2. 3. 4. 5.
Permanent (dead) loads. Operating (live) loads. Environmental loads including earthquakes. Construction - installation loads. Accidental loads.
Whilst the design of buildings onshore is usually influenced mainly by the permanent and operating loads, the design of offshore structures is dominated by environmental loads, especially waves, and the loads arising in the various stages of construction and installation. This lecture deals with environmental loads, whilst the other loadings are treated in Lecture 15A.3. In civil engineering, earthquakes are normally regarded as accidental loads (see Eurocode 8 [1]), but in offshore engineering they are treated as environmental loads. This practice is followed in the two lectures dealing with loads, Lectures 15A.2 and 15A.3. 2. ENVIRONMENTAL LOADS Environmental loads are those caused by environmental phenomena such as wind, waves, current, tides, earthquakes, temperature, ice, sea bed movement, and marine growth. Their characteristic parameters, defining design load values, are determined in special studies on the basis of available data. According to US and Norwegian regulations (or codes of practice), the mean recurrence interval for the corresponding design event must be 100 years, while according to the British rules it should be 50 years or greater. Details of design criteria, simplifying assumptions, required data, etc., can be found in the regulations and codes of practice listed in [1] - [8]. 2.1 Wind Loads Wind loads act on the portion of a platform above the water level, as well as on any equipment, housing, derrick, etc. located on the deck. An important parameter pertaining to wind data is the time interval over which wind speeds are averaged. For averaging intervals less than one minute, wind speeds are classified as gusts. For averaging intervals of one minute or longer they are classified as sustained wind speeds.
The wind velocity profile may be taken from API-RP2A [2]: Vh/VH = (h/H)1/n (1) where: Vh is the wind velocity at height h, VH is the wind velocity at reference height H, typically 10m above mean water level, 1/n is 1/13 to 1/7, depending on the sea state, the distance from land and the averaging time interval. It is approximately equal to 1/13 for gusts and 1/8 for sustained winds in the open ocean. From the design wind velocity V(m/s), the static wind force Fw(N) acting perpendicular to an exposed area A(m2) can be computed as follows: Fw = (1/2) r V2 Cs A (2) where: r is the wind density (r » 1.225 Kg/m3) Cs is the shape coefficient (Cs = 1,5 for beams and sides of buildings, Cs = 0,5 for cylindrical sections and Cs = 1,0 for total projected area of platform). Shielding and solidity effects can be accounted for, in the judgement of the designer, using appropriate coefficients. For combination with wave loads, the DNV [4] and DOE-OG [7] rules recommend the most unfavourable of the following two loadings: a. 1-minute sustained wind speeds combined with extreme waves. b. 3-second gusts. API-RP2A [2] distinguishes between global and local wind load effects. For the first case it gives guideline values of mean 1hour average wind speeds to be combined with extreme waves and current. For the second case it gives values of extreme wind speeds to be used without regard to waves. Wind loads are generally taken as static. When, however, the ratio of height to the least horizontal dimension of the wind exposed object (or structure) is greater than 5, then this object (or structure) could be wind sensitive. API-RP2A requires the dynamic effects of the wind to be taken into account in this case and the flow induced cyclic wind loads due to vortex shedding must be investigated. 2.2 Wave Loads The wave loading of an offshore structure is usually the most important of all environmental loadings for which the structure must be designed. The forces on the structure are caused by the motion of the water due to the waves which are generated by the action of the wind on the surface of the sea. Determination of these forces requires the solution of two separate, though interrelated problems. The first is the sea state computed using an idealisation of the wave surface profile and the wave kinematics given by an appropriate wave theory. The second is the computation of the wave forces on individual members and on the total structure, from the fluid motion. Two different analysis concepts are used:
The design wave concept, where a regular wave of given height and period is defined and the forces due to this wave are calculated using a high-order wave theory. Usually the 100-year wave, i.e. the maximum wave with a return period of 100 years, is chosen. No dynamic behaviour of the structure is considered. This static analysis is appropriate when the dominant wave periods are well above the period of the structure. This is the case of extreme storm waves acting on shallow water structures. Statistical analysis on the basis of a wave scatter diagram for the location of the structure. Appropriate wave spectra are defined to perform the analysis in the frequency domain and to generate random waves, if dynamic analyses for extreme wave loadings are required for deepwater structures. With statistical methods, the most probable maximum force during the lifetime of the structure is calculated using linear wave theory. The statistical approach has to be chosen to analyze the fatigue strength and the dynamic behaviour of the structure.
2.2.1 Wave theories Wave theories describe the kinematics of waves of water on the basis of potential theory. In particular, they serve to calculate the particle velocities and accelerations and the dynamic pressure as functions of the surface elevation of the waves. The waves are assumed to be long-crested, i.e. they can be described by a two-dimensional flow field, and are characterized by the parameters: wave height (H), period (T) and water depth (d) as shown in Figure 1.
Different wave theories of varying complexity, developed on the basis of simplifying assumptions, are appropriate for different ranges of the wave parameters. Among the most common theories are: the linear Airy theory, the Stokes fifth-order theory, the solitary wave theory, the cnoidal theory, Dean's stream function theory and the numerical theory by Chappelear. For the selection of the most appropriate theory, the graph shown in Figure 2 may be consulted. As an example, Table 1 presents results of the linear wave theory for finite depth and deep water conditions. Corresponding particle paths are illustrated in Figures 3 and 4. Note the strong influence of the water depth on the wave kinematics. Results from high-order wave theories can be found in the literature, e.g. see [9].
2.2.2 Wave Statistics In reality waves do not occur as regular waves, but as irregular sea states. The irregular appearance results from the linear superposition of an infinite number of regular waves with varying frequency (Figure 5). The best means to describe a random sea state is using the wave energy density spectrum S(f), usually called the wave spectrum for simplicity. It is formulated as a function of the wave frequency f using the parameters: significant wave height Hs (i.e. the mean of the highest third of all waves present in a wave train) and mean wave period (zero-upcrossing period) To. As an additional parameter the spectral width can be taken into account.
Wave directionality can be introduced by means of a directional spreading function D(f,s), where s is the angle of the wave approach direction (Figure 6). A directional wave spectrum S (f,s) can then be defined as: S (f,s ) = S(f).D (f,s ) (3)
The response of the structure, i.e. forces, motions, is calculated by multiplication of the wave energy spectrum with the square of a linear transfer function. From the resulting response spectrum the significant and the maximum expected response in a given time interval can be easily deduced. For long-term statistics, a wave scatter diagram for the location of the structure is needed. It can be obtained from measurements over a long period or be deduced from weather observations in the region (the so-called hindcast method). The scatter diagram contains the joint probability of occurrence of pairs of significant wave height and mean wave period. For every pair of parameters the wave spectrum is calculated by a standard formula, e.g. Pierson-Moskowitz (Figure 6), yielding finally the desired response spectrum. For fatigue analysis the total number and amplitude of load cycles during the life-time of the structure can be derived in this way. For structures with substantial dynamic response to the wave excitation, the maximum forces and motions have to be calculated by statistical methods or a time-domain analysis. 2.2.3 Wave forces on structural members Structures exposed to waves experience substantial forces much higher than wind loadings. The forces result from the dynamic pressure and the water particle motions. Two different cases can be distinguished:
Large volume bodies, termed hydrodynamic compact structures, influence the wave field by diffraction and reflection. The forces on these bodies have to be determined by costly numerical calculations based on diffraction theory. Slender, hydrodynamically transparent structures have no significant influence on the wave field. The forces can be calculated in a straight-forward manner with Morison's equation. As a rule, Morison's equation may be applied when D/L £ 0.2, where D is the member diameter and L is the wave length.
The steel jackets of offshore structures can usually be regarded as hydrodynamically transparent. The wave forces on the submerged members can therefore be calculated by Morison's equation, which expresses the wave force as the sum of an inertia force proportional to the particle acceleration and a non-linear drag force proportional to the square of the particle velocity:
(4)
where F is the wave force per unit length on a circular cylinder (N) v, |v| are water particle velocity normal to the cylinder, calculated with the selected wave theory at the cylinder axis (m/s) are water particle acceleration normal to the cylinder, calculated with the selected wave theory at the cylinder axis (m/s 2) r is the water density (kg/m3) D is the member diameter, including marine growth (m) CD, CM are drag and inertia coefficients, respectively. In this form the equation is valid for fixed tubular cylinders. For the analysis of the motion response of a structure it has to be modified to account for the motion of the cylinder [10]. The values of CD and CM depend on the wave theory used, surface roughness and the flow parameters. According to API-RP2A, CD » 0,6 to 1,2 and CM »1,3 to 2,0. Additional information can be found in the DNV rules [4]. The total wave force on each member is obtained by numerical integration over the length of the member. The fluid velocities and accelerations at the integration points are found by direct application of the selected wave theory. According to Morison's equation the drag force is non-linear. This non-linear formulation is used in the design wave concept. However, for the determination of a transfer function needed for frequency domain calculations, the drag force has to be linearized in a suitable way [9]. Thus, frequency domain solutions are appropriate for fatigue life calculations, for which the
forces due to the operational level waves are dominated by the linear inertia term. The nonlinear formulation and hence time domain solutions are required for dynamic analyses of deepwater structures under extreme, storm waves, for which the drag portion of the force is the dominant part [10]. In addition to the forces given by Morison's equation, the lift forces FD and the slamming forces FS, typically neglected in global response computations, can be important for local member design. For a member section of unit length, these forces can be estimated as follows: FL = (1/2) r CL Dv2 (5) FS = (1/2) r Cs Dv2 (6) where CL, CS are the lift and slamming coefficients respectively, and the rest of the symbols are as defined in Morison's equation. Lift forces are perpendicular to the member axis and the fluid velocity v and are related to the vortex shedding frequency. Slamming forces acting on the underside of horizontal members near the mean water level are impulsive and nearly vertical. Lift forces can be estimated by taking CL » 1,3 CD. For tubular members Cs » p. 2.3 Current Loads There are tidal, circulation and storm generated currents. Figure 7 shows a wind and tidal current profile typical of the Gulf of Mexico. When insufficient field measurements are available, current velocities may be obtained from various sources, e.g. Appendix A of DNV [4]. In platform design, the effects of current superimposed on waves are taken into account by adding the corresponding fluid velocities vectorially. Since the drag force varies with the square of the velocity, this addition can greatly increase the forces on a platform. For slender members, cyclic loads induced by vortex shedding may also be important and should be examined.
2.4 Earthquake Loads Offshore structures in seismic regions are typically designed for two levels of earthquake intensity: the strength level and the ductility level earthquake. For the strength level earthquake, defined as having a "reasonable likelihood of not being exceeded during the platform's life" (mean recurrence interval ~ 200 - 500 years), the structure is designed to respond elastically. For the ductility level earthquake, defined as close to the "maximum credible earthquake" at the site, the structure is designed for inelastic response and to have adequate reserve strength to avoid collapse. For strength level design, the seismic loading may be specified either by sets of accelerograms (Figure 8) or by means of design response spectra (Figure 9). Use of design spectra has a number of advantages over time history solutions (base acceleration input). For this reason design response spectra are the preferable approach for strength level designs. If the design spectral intensity, characteristic of the seismic hazard at the site, is denoted by amax, then API-RP2A recommends using amax for the two principal horizontal directions and 0,5amax for the vertical direction. The DNV rules, on the other hand,
recommend amax and 0,7 amax for the two horizontal directions (two different combinations) and 0,5 amax for the vertical. The value of amax and often the spectral shapes are determined by site specific seismological studies.
Designs for ductility level earthquakes will normally require inelastic analyses for which the seismic input must be specified by sets of 3-component accelerograms, real or artificial, representative of the extreme ground motions that could shake the platform site. The characteristics of such motions, however, may still be prescribed by means of design spectra, which are usually the result of a site specific seismotectonic study. More detail of the analysis of earthquakes is given in the Lectures 17: Seismic Design. 2.5 Ice and Snow Loads Ice is a primary problem for marine structures in the arctic and sub-arctic zones. Ice formation and expansion can generate large pressures that give rise to horizontal as well as vertical forces. In addition, large blocks of ice driven by current, winds and waves with speeds that can approach 0,5 to 1,0 m/s, may hit the structure and produce impact loads. As a first approximation, statically applied, horizontal ice forces may be estimated as follows: Fi = CifcA (7) where: A is the exposed area of structure, fc is the compressive strength of ice, Ci is the coefficient accounting for shape, rate of load application and other factors, with usual values between 0,3 and 0,7. Generally, detailed studies based on field measurements, laboratory tests and analytical work are required to develop reliable design ice forces for a given geographical location. In addition to these forces, ice formation and snow accumulations increase gravity and wind loads, the latter by increasing areas exposed to the action of wind. More detailed information on snow loads may be found in Eurocode 1 [8]. 2.6 Loads due to Temperature Variations Offshore structures can be subjected to temperature gradients which produce thermal stresses. To take account of such stresses, extreme values of sea and air temperatures which are likely to occur during the life of the structure must be estimated. Relevant data for the North Sea are given in BS6235 [6]. In addition to the environmental sources, human factors can also generate thermal loads, e.g. through accidental release of cryogenic material, which must be taken into account in design as accidental loads. The temperature of the oil and gas produced must also be considered. 2.7 Marine Growth Marine growth is accumulated on submerged members. Its main effect is to increase the wave forces on the members by increasing not only exposed areas and volumes, but also the drag coefficient due to higher surface roughness. In addition, it increases the unit mass of the member, resulting in higher gravity loads and in lower member frequencies. Depending upon geographic location, the thickness of marine growth can reach 0,3m or more. It is accounted for in design through appropriate increases in the diameters and masses of the submerged members. 2.8 Tides Tides affect the wave and current loads indirectly, i.e. through the variation of the level of the sea surface. The tides are classified as: (a) astronomical tides - caused essentially from the gravitational pull of the moon and the sun and (b) storm surges - caused by the combined action of wind and barometric pressure differentials during a storm. The combined effect of the two types of tide is called the storm tide. Tide dependent water levels and the associated definitions, as used in platform design, are shown in Figure 10. The astronomical tide range depends on the geographic location and the phase of the moon. Its maximum, the spring tide, occurs at new moon. The range varies from centimetres to several metres and may be obtained from special maps. Storm surges depend upon the return period considered and their range is on the order of 1,0 to 3,0m. When designing a platform, extreme storm waves are superimposed on the still water level (see Figure 10),
while for design considerations such as levels for boat landing places, barge fenders, upper limits of marine growth, etc., the daily variations of the astronomical tide are used.
2.9 Sea Floor Movements Movement of the sea floor can occur as a result of active geologic processes, storm wave pressures, earthquakes, pressure reduction in the producing reservoir, etc. The loads generated by such movements affect, not only the design of the piles, but the jacket as well. Such forces are determined by special geotechnical studies and investigations. 3. CONCLUDING SUMMARY
Environmental loads form a major category of loads which control many aspects of platform design. The main environmental loads are due to wind, waves, current, earthquakes, ice and snow, temperature variations, marine growth, tides and seafloor movements. Widely accepted rules of practice, listed as [1] - [13], provide guideline values for most environmental loads. For major structures, specification of environmental design loads requires specific studies. Some environmental loads can be highly uncertain. The definition of certain environmental loads depends upon the type of analysis used in the design.
4. REFERENCES [1] Eurocode 8: "Structures in Seismic Regions - Design", CEN (in preparation). [2] API-RP2A, "Recommended Practice for Planning, Designing and Constructing Fixed Offshore Platforms", American Petroleum Institute, Washington, D.C., 18th ed., 1989. [3] OCS, "Requirements for Verifying the Structural Integrity of OCS Platforms"., United States Geologic Survey, National Centre, Reston, Virginia, 1980. [4] DNV, "Rules for the Design, Construction and Inspection of Offshore Structures", Det Norske Veritas, Oslo, 1977 (with corrections 1982). [5] NPD, "Regulation for Structural Design of Load-bearing Structures Intended for Exploitation of Petroleum Resources", Norwegian Petroleum Directorate, 1985. [6] BS6235, "Code of Practice for Fixed Offshore Structures", British Standards Institution, London, 1982.
[7] DOE-OG, "Offshore Installation: Guidance on Design and Construction", U.K., Dept. of Energy, London 1985. [8] Eurocode 1: "Basis of Design and Actions on Structures", CEN (in preparation). [9] Clauss, G. T. et al: "Offshore Structures, Vol 1 - Conceptual Design and Hydromechanics", Springer, London 1992. [10] Anagnostopoulos, S.A., "Dynamic Response of Offshore Structures to Extreme Waves including Fluid - Structure Interaction", Engr. Structures, Vol. 4, pp.179-185, 1982. [11] Hsu, H.T., "Applied Offshore Structural Engineering", Gulf Publishing Co., Houston, 1981. [12] Graff, W.J., "Introduction to Offshore Structures", Gulf Publishing Co., Houston, 1981. [13] Gerwick, B.C. Jr., "Construction of Offshore Structures", John Wiley, New York, 1986. Table 1 Results of Linear Airy Theory [11] Phase q = kx - w t
Deep water
Finite water depth
d/L ³ 0,5
d/L < 0,5
za cos q
za cos q
Relative water depth d/L Velocity potential q
Surface elevation z Dynamic pressure
r gza ekz cos q
pdyn = Water particle velocities za w ekz cos q horizontal u = za w ekz sin q vertical w = Water particle accelerations za w2 ekz sin q horizontal u' = -za w2 ekz cos q vertical w' =
Wave celerity c =
co =
c=
Group velocity cgr=
Circular frequencyw =
Wave length L =
Wave number k =
cgr =
cgr =
w=
Lo =
ko =
w=
L=
kd tanh kd =
Water particle displacements horizontal x
-za ekz sin q
vertical z
za ekz cos q
Particle trajectories
Circular orbits Elliptical orbits
Where z a =
A.3: Loads (II) - Other Loads OBJECTIVE/SCOPE To present and briefly describe all loads, except environmental loads, and the load combinations for which a fixed offshore structure must be designed. PREREQUISITES A basic knowledge of structural analysis for static and dynamic loadings. SUMMARY The various categories of loads, except environmental, for which a pile-supported steel offshore platform must be designed are presented. These categories include permanent (dead) loads, operating (live) loads, loads generated during fabrication and installation (due to lifts, loadout, transportation, launching and upending) and accidental loads. In addition, the different load combinations for all types of loads, including environmental, as required (or suggested) by applicable regulations (or codes of practice) are given. The categories of loads described herein are the following: 1. 2. 3. 4.
Permanent (dead) loads Operating (live) loads Fabrication and installation loads Accidental loads
The major categories of environmental loads are not included. They are dealt with in Lecture 15A.2. 1. PERMANENT (DEAD) LOADS Permanent loads include the following: a. Weight of the structure in air, including the weight of grout and ballast, if necessary. b. Weights of equipment, attachments or associated structures which are permanently mounted on the platform. c. Hydrostatic forces on the various members below the waterline. These forces include buoyancy and hydrostatic pressures. Sealed tubular members must be designed for the worst condition when flooded or non-flooded. 2. OPERATING (LIVE) LOADS Operating loads arise from the operations on the platform and include the weight of all non-permanent equipment or material, as well as forces generated during operation of equipment. More specifically, operating loads include the following: a. The weight of all non-permanent equipment (e.g. drilling, production), facilities (e.g. living quarters, furniture, life support systems, heliport, etc), consumable supplies, liquids, etc. b. Forces generated during operations, e.g. drilling, vessel mooring, helicopter landing, crane operations, etc. The necessary data for computation of all operating loads are provided by the operator and the equipment manufacturers. The data need to be critically evaluated by the designer. An example of detailed live load specification is given in Table 1 where the values in the first and second columns are for design of the portions of the structure directly affected by the loads
and the reduced values in the last column are for the structure as a whole. In the absence of such data, the following values are recommended in BS6235 [1]: a. crew quarters and passageways: 3,2 KN/m2 b. working areas: 8,5 KN/m2 c. storage areas: gH KN/m2 where g is the specific weight of stored materials, not to be taken less than 6,87KN/m3, H is the storage height (m). Forces generated during operations are often dynamic or impulsive in nature and must be treated as such. For example, according to the BS6235 rules, two types of helicopter landing should be considered, heavy and emergency landing. The impact load in the first case is to be taken as 1,5 times the maximum take-off weight, while in the second case this factor becomes 2,5. In addition, a horizontal load applied at the points of impact and taken equal to half the maximum take-off weight must be considered. Loads from rotating machinery, drilling equipment, etc. may normally be treated as harmonic forces. For vessel mooring, design forces are computed for the largest ship likely to approach at operational speeds. According to BS6235, the minimum impact to be considered is of a vessel of 2500 tonnes at 0,5 m/s. 3. FABRICATION AND INSTALLATION LOADS These loads are temporary and arise during fabrication and installation of the platform or its components. During fabrication, erection lifts of various structural components generate lifting forces, while in the installation phase forces are generated during platform loadout, transportation to the site, launching and upending, as well as during lifts related to installation. According to the DNV rules [2], the return period for computing design environmental conditions for installation as well as fabrication should normally be three times the duration of the corresponding phase. API-RP2A, on the other hand [3], leaves this design return period up to the owner, while the BS6235 rules [1] recommend a minimum recurrence interval of 10 years for the design environmental loads associated with transportation of the structure to the offshore site. 3.1 Lifting Forces Lifting forces are functions of the weight of the structural component being lifted, the number and location of lifting eyes used for the lift, the angle between each sling and the vertical axis and the conditions under which the lift is performed (Figure 1). All members and connections of a lifted component must be designed for the forces resulting from static equilibrium of the lifted weight and the sling tensions. Moreover, API-RP2A recommends that in order to compensate for any side movements, lifting eyes and the connections to the supporting structural members should be designed for the combined action of the static sling load and a horizontal force equal to 5% this load, applied perpendicular to the padeye at the centre of the pin hole. All these design forces are applied as static loads if the lifts are performed in the fabrication yard. If, however, the lifting derrick or the structure to be lifted is on a floating vessel, then dynamic load factors should be applied to the static lifting forces. In particular, for lifts made offshore API-RP2A recommends two minimum values of dynamic load factors: 2,0 and 1,35. The first is for designing the padeyes as well as all members and their end connections framing the joint where the padeye is attached, while the second is for all other members transmitting lifting forces. For loadout at sheltered locations, the corresponding minimum load factors for the two groups of structural components become, according to API-RP2A, 1,5 and 1,15, respectively.
3.2 Loadout Forces These are forces generated when the jacket is loaded from the fabrication yard onto the barge. If the loadout is carried out by direct lift, then, unless the lifting arrangement is different from that to be used for installation, lifting forces need not be computed, because lifting in the open sea creates a more severe loading condition which requires higher dynamic load factors. If loadout is done by skidding the structure onto the barge, a number of static loading conditions must be considered, with the jacket supported on its side. Such loading conditions arise from the different positions of the jacket during the loadout phases, (as shown in Figure 2), from movement of the barge due to tidal fluctuations, marine traffic or change of draft, and from possible support settlements. Since movement of the jacket is slow, all loading conditions can be taken as static. Typical values of friction coefficients for calculation of skidding forces are the following:
steel on steel without lubrication............................................ 0,25 steel on steel with lubrication............................................... 0,15 steel on teflon.................................................................. 0,10 teflon on teflon................................................................. 0,08
3.3 Transportation Forces These forces are generated when platform components (jacket, deck) are transported offshore on barges or self-floating. They depend upon the weight, geometry and support conditions of the structure (by barge or by buoyancy) and also on the environmental conditions (waves, winds and currents) that are encountered during transportation. The types of motion that a floating structure may experience are shown schematically in Figure 3.
In order to minimize the associated risks and secure safe transport from the fabrication yard to the platform site, it is important to plan the operation carefully by considering, according to API-RP2A [3], the following: 1. 2. 3. 4. 5.
Previous experience along the tow route Exposure time and reliability of predicted "weather windows" Accessibility of safe havens Seasonal weather system Appropriate return period for determining design wind, wave and current conditions, taking into account characteristics of the tow such as size, structure, sensitivity and cost.
Transportation forces are generated by the motion of the tow, i.e. the structure and supporting barge. They are determined from the design winds, waves and currents. If the structure is self-floating, the loads can be calculated directly. According to API-RP2A [3], towing analyses must be based on the results of model basin tests or appropriate analytical methods and must consider wind and wave directions parallel, perpendicular and at 45° to the tow axis. Inertial loads may be computed from a rigid body analysis of the tow by combining roll and pitch with heave motions, when the size of the tow, magnitude of the sea state and experience make such assumptions reasonable. For open sea conditions, the following may be considered as typical design values: Single - amplitude roll: 20° Single - amplitude pitch: 10°
Period of roll or pitch: 10 second Heave acceleration: 0,2 g When transporting a large jacket by barge, stability against capsizing is a primary design consideration because of the high centre of gravity of the jacket. Moreover, the relative stiffness of jacket and barge may need to be taken into account together with the wave slamming forces that could result during a heavy roll motion of the tow (Figure 4) when structural analyses are carried out for designing the tie-down braces and the jacket members affected by the induced loads. Special computer programs are available to compute the transportation loads in the structure-barge system and the resulting stresses for any specified environmental condition.
3.4 Launching and Upending Forces These forces are generated during the launch of a jacket from the barge into the sea and during the subsequent upending into its proper vertical position to rest on the seabed. A schematic view of these operations can be seen in Figure 5.
There are five stages in a launch-upending operation: a. Jacket slides along the skid beams b. Jacket rotates on the rocker arms c. Jacket rotates and slides simultaneously d. Jacket detaches completely and comes to its floating equilibrium position e. Jacket is upended by a combination of controlled flooding and simultaneous lifting by a derrick barge. The loads, static as well as dynamic, induced during each of these stages and the force required to set the jacket into motion can be evaluated by appropriate analyses, which also consider the action of wind, waves and currents expected during the operation. To start the launch, the barge must be ballasted to an appropriate draft and trim angle and subsequently the jacket must be pulled towards the stern by a winch. Sliding of the jacket starts as soon as the downward force (gravity component and winch pull) exceeds the friction force. As the jacket slides, its weight is supported on the two legs that are part of the launch trusses. The support length keeps decreasing and reaches a minimum, equal to the length of the rocker beams, when rotation starts. It is generally at this instant that the most severe launching forces develop as reactions to the weight of the jacket. During stages (d) and (e), variable hydrostatic forces arise which have to be considered at all members affected. Buoyancy calculations are required for every stage of the operation to ensure fully controlled, stable motion. Computer programs are available to perform the stress analyses required for launching and upending and also to portray the whole operation graphically. 4. ACCIDENTAL LOADS According to the DNV rules [2], accidental loads are loads, ill-defined with respect to intensity and frequency, which may occur as a result of accident or exceptional circumstances. Accidental loads are also specified as a separate category in the NPD regulations [4], but not in API-RP2A [3], BS6235 [1] or the DOE-OG rules [5]. Examples of accidental loads are loads due to collision with vessels, fire or explosion, dropped objects, and unintended flooding of bouyancy tanks. Special measures are normally taken to reduce the risk from accidental loads. For example, protection of wellheads or other critical equipment from a dropped object can be provided by specially designed, impact resistant covers. According to the NPD regulations [4], an accidental load can be disregarded if its annual probability of occurrence is less than 10 -4. This number is meant as an order of magnitude estimate and is extremely difficult to compute. Earthquakes are treated as an environmental load in offshore structure design.
5. LOAD COMBINATIONS The load combinations used for designing fixed offshore structures depend upon the design method used, i.e. whether limit state or allowable stress design is employed. The load combinations recommended for use with allowable stress procedures are: a. Dead loads plus operating environmental loads plus maximum live loads, appropriate to normal operations of the platform. b. Dead loads plus operating environmental loads plus minimum live loads, appropriate to normal operations of the platform. c. Dead loads plus extreme (design) environmental loads plus maximum live loads, appropriate for combining with extreme conditions. d. Dead loads plus extreme (design) environmental loads plus minimum live loads, appropriate for combining with extreme conditions. Moreover, environmental loads, with the exception of earthquake loads, should be combined in a manner consistent with their joint probability of occurrence during the loading condition considered. Earthquake loads, if applicable, are to be imposed as a separate environmental load, i.e., not to be combined with waves, wind, etc. Operating environmental conditions are defined as representative of severe but not necessarily limiting conditions that, if exceeded, would require cessation of platform operations. The DNV rules [2] permit allowable stress design but recommend the semi-probabilistic limit state design method, which the NPD rules also require [4]. BS6235 permits both methods but the design equations it gives are for the allowable stress method [1]. API-RP2A is very specific in recommending not to apply limit state methods. According to the DNV and the NPD rules for limit state design, four limit states must be checked: 1.
Ultimate limit state For this limit state the following two loading combinations must be used: Ordinary: 1,3 P + 1,3 L + 1,0 D + 0,7 E, and Extreme : 1,0 P + 1,0 L + 1,0 D + 1,3 E where P, L, D and E stand for Permanent (dead), Operating (live), Deformation (e.g., temperature, differential settlement) and Environmental loads respectively. For well controlled dead and live loads during fabrication and installation, the load factor 1,3 may be reduced to 1,2. Furthermore, for structures that are unmanned during storm conditions and which are not used for storage of oil and gas, the 1,3 load factor for environmental loads - except earthquakes - may be reduced to 1,15.
2.
Fatigue limit state All load factors are to be taken as 1,0.
3.
Progressive Collapse limit state All load factors are to be taken as 1,0.
4.
Serviceability limit state
All load factors are to be taken as 1,0. The so-called characteristic values of the loads used in the above combinations and limit states are summarized in Table 2, taken from the NPD rules.
6. CONCLUDING SUMMARY
In addition to environmental loads, an offshore structure must be designed for dead and live loads, fabrication and installation loads as well as accidental loads. Widely accepted rules of practice, listed in the references, are usually followed for specifying such loads. The type and magnitude of fabrication, transportation and installation loads depend upon the methods and sequences used for the corresponding phases. Dynamic and impact effects are normally taken into account by means of appropriate dynamic load factors. Accidental loads are not well defined with respect to intensity and probability of occurrence. They will typically require special protective measures. Load combinations and load factors depend upon the design method to be used. API-RP2A is based on allowable stress design and recommends against limit state design, BSI favours allowable stress design, while DNV and NPD recommend limit state design.
7. REFERENCES [1] BS6235, "Code of Practice for Fixed Offshore Structures", British Standards Institution, London, 1982. [2] "Rules for the Design, Construction and Inspection of Offshore Structures", Det Norske Veritas (DNV), Oslo, 1977 (with corrections 1982). [3] API-RP2A, "Recommended Practice for Planning, Designing and Constructing Fixed Offshore Platforms", American Petroleum Institute, Washington, D.C., 18th ed., 1989. [4] "Regulation for Structural Design of Load-bearing Structures Intended for Exploitation of Petroleum Resources", Norwegian Petroleum Directorate (NPD), 1985. [5] DOE-OG, "Offshore Installation: Guidance on Design and Construction", U.K. Department of Energy, London 1985. 8. ADDITIONAL READING 1. 2. 3. 4.
OCS, "Requirements for Verifying the Structural Integrity of OCS Platforms"., United States Geologic Survey, National Centre, Reston, Virginia, 1980. Hsu, H.T., "Applied Offshore Structural Engineering", Gulf Publishing Co., Houston, 1981. Graff, W.G., "Introduction to Offshore Structures", Gulf Publishing Co., Houston, 1981. Gerwick, B.C. Jr., "Construction of Offshore Structures", John Wiley, New York, 1986.
Table 1 Minimum design live load specification Loads to be taken into account (kN/m2)
For portions of the structure
For the structure as a whole
Zone considered
Flooring and joists
Other components
(3)
Process zone (around wells and largescale machines)
5 (1)
5 (1)
2.5
Drilling zone
5 (1)
5 (1)
2.5
Catwalks and walking platforms (except emergency exits)
3
2.5
1
Stairways (except emergency exits)
4
3
0
Module roofing
2
1.5
1
Emergency exits
5
5
0
Storage floors - heavy
18
12
8 (2)
Storage floors - light
9
6
4 (2)
Delivery zone
10
10
5
Non-attributed area
6
4
3
STORAGE
(1) Accumulated with a point load equal to the weight of the heaviest part likely to be removed, with a minimum value of 5 kN. Point loads are assumed as being applied to a 0,3m ´ 0,3m surface. (2) Applied on the entirety of the flooring surface (including traffic). (3) This column gives the loads to be taken into account for the structure's overall calculation. These values are the input for the computer runs. Table 2 Characteristic Loads according to NPD [4] LOAD TYPE
LIMIT STATES FOR TEMPORARY PHASES
LIMIT STATES FOR
Progressive Collapse Serviceability
Fatigue
Ultimate
Abnormal effects
Damage condition
Serviceability
DEAD
EXPECTED VALUE
LIVE
SPECIFIED VALUE
DEFORMATION ENVIRONMENTAL
ACCIDENTAL
Fatigue
Ult
EXPECTED EXTREME VALUE Dependent on operational requirements
Expected load history
NOT APPLICABLE
Value dependent on measures taken
Dependent on operational requirements
Dependent on operational requirements
Expected Annua load history excee probab
NOT APPLICABLE
A.4 - Analysis I OBJECTIVE/SCOPE To present the main analysis procedures for offshore structures. PREREQUISITES Lecture 15A.1: Offshore Structures: General Introduction Lecture 15A.2: Loads I: Introduction and Environmental Loads Lecture 15A.3: Loads II: Other Loads RELATED LECTURES Lecture 15A.5: Analysis II SUMMARY Analytical models used in offshore engineering are briefly described. Acceptance criteria for the verification of offshore structures are presented. Simple rules for preliminary member sizing are given and procedures for static in-place and dynamic analysis are described. 1. ANALYTICAL MODEL The analysis of an offshore structure is an extensive task, embracing consideration of the different stages, i.e. execution, installation, and in-service stages, during its life. Many disciplines, e.g. structural, geotechnical, naval architecture, metallurgy are involved. This lecture and Lecture 15A.5 are purposely limited to presenting an overview of available analysis procedures and providing benchmarks for the reader to appreciate the validity of his assumptions and results. They primarily address jackets, which are more unusual structures compared to decks and modules, and which more closely resemble onshore petrochemical plants. 2. ANALYTICAL MODEL The analytical models used in offshore engineering are in some respects similar to those adopted for other types of steel structures. Only the salient features of offshore models are presented here. The same model is used throughout the analysis process with only minor adjustments being made to suit the specific conditions, e.g. at supports in particular, relating to each analysis. 2.1 Stick Models Stick models (beam elements assembled in frames) are used extensively for tubular structures (jackets, bridges, flare booms) and lattice trusses (modules, decks). 2.1.1 Joints Each member is normally rigidly fixed at its ends to other elements in the model.
If more accuracy is required, particularly for the assessment of natural vibration modes, local flexibility of the connections may be represented by a joint stiffness matrix. 2.1.2 Members In addition to its geometrical and material properties, each member is characterised by hydrodynamic coefficients, e.g. relating to drag, inertia, and marine growth, to allow wave forces to be automatically generated. 2.2 Plate Models Integrated decks and hulls of floating platforms involving large bulkheads are described by plate elements. The characteristics assumed for the plate elements depend on the principal state of stress which they are subjected to. Membrane stresses are taken when the element is subjected merely to axial load and shear. Plate stresses are adopted when bending and lateral pressure are to be taken into account. 3. ACCEPTANCE CRITERIA 3.1 Code Checks The verification of an element consists of comparing its characteristic resistance(s) to a design force or stress. It includes:
a strength check, where the characteristic resistance is related to the yield strength of the element, a stability check for elements in compression where the characteristic resistance relates to the buckling limit of the element.
An element (member or plate) is checked at typical sections (at least both ends and midspan) against resistance and buckling. This verification also includes the effect of water pressure for deepwater structures. Tubular joints are checked against punching under various load patterns. These checks may indicate the need for local reinforcement of the chord using overthickness or internal ring-stiffeners. Elements should also be verified against fatigue, corrosion, temperature or durability wherever relevant. 3.2 Allowable Stress Method This method is presently specified by American codes (API, AISC). The loads remain unfactored and a unique coefficient is applied to the characteristic resistance to obtain an allowable stress as follows: Condition
Axial
Strong axis bending
Weak axis bending
Normal
0,60
0,66
0,75
Extreme
0,80
0,88
1,00
"Normal" and "Extreme" respectively represent the most severe conditions:
under which the plant is to operate without shut-down. the platform is to endure over its lifetime.
3.3 Limit State Method
This method is enforced by European and Norwegian Authorities and has now been adopted by API as it offers a more uniform reliability. Partial factors are applied to the loads and to the characteristic resistance of the element, reflecting the amount of confidence placed in the design value of each parameter and the degree of risk accepted under a limit state, i.e:
Ultimate Limit State (ULS):
corresponds to an ultimate event considering the structural resistance with appropriate reserve.
Fatigue Limit State (FLS):
relates to the possibility of failure under cyclic loading.
Progressive Collapse Limit State (PLS):
reflects the ability of the structure to resist collapse under accidental or abnormal conditions.
Service Limit State (SLS):
corresponds to criteria for normal use or durability (often specified by the plant operator). 3.3.1 Load factors Norwegian Authorities (2, 4) specify the following sets of load factors: Limit State
Load Categories P
L
D
E
A
ULS (normal)
1,3
1,3
1,0
0,7
0,0
ULS (extreme)
1,0
1,0
1,0
1,3
0,0
FLS
0,0
0,0
0,0
1,0
0,0
PLS (accidental)
1,0
1,0
1,0
1,0
1,0
PLS (post-damage)
1,0
1,0
1,0
1,0
0,0
SLS
1,0
1,0
1,0
1,0
0,0
where the respective load categories are: P are permanent loads (structural weight, dry equipments, ballast, hydrostatic pressure). L are live loads (storage, personnel, liquids). D are deformations (out-of-level supports, subsidence). E are environmental loads (wave, current, wind, earthquake).
A are accidental loads (dropped object, ship impact, blast, fire). 3.3.2 Material factors The material partial factors for steel is normally taken equal to 1,15 for ULS and 1,00 for PLS and SLS design. 3.3.3 Classification of Design Conditions Guidance for classifying typical conditions into typical limit states is given in the following table: Condition
Loadings P/L
Design E
D
A
Criterion
Construction
P
ULS,SLS
Load-Out
P
reduced wind
Transport
P
transport wind and wave
Tow-out (accidental)
P
Launch
P
ULS
Lifting
P
ULS
In-Place (normal)
P+L
wind, wave & snow
actual
ULS,SLS
In-Place (extreme)
P+L
wind & 100 year wave
actual
ULS
support disp
ULS ULS
flooded compart
PLS
SLS In-Place (exceptional)
P+L
wind & 10000 year wave
Earthquake
P+L
10-2 quake
ULS
Rare Earthquake
P+L
10-4 quake
PLS
Explosion
P+L
blast
PLS
Fire
P+L
fire
PLS
Dropped Object
P+L
drill collar
PLS
Boat Collision
P+L
boat impact
PLS
Damaged Structure
P + reduced L
reduced wave & wind
actual
PLS
PLS
4. PRELIMINARY MEMBER SIZING The analysis of a structure is an iterative process which requires progressive adjustment of the member sizes with respect to the forces they transmit, until a safe and economical design is achieved. It is therefore of the utmost importance to start the main analysis from a model which is close to the final optimized one. The simple rules given below provide an easy way of selecting realistic sizes for the main elements of offshore structures in moderate water depth (up to 80m) where dynamic effects are negligible. 4.1 Jacket Pile Sizes
calculate the vertical resultant (dead weight, live loads, buoyancy), the overall shear and the overturning moment (environmental forces) at the mudline. assuming that the jacket behaves as a rigid body, derive the maximum axial and shear force at the top of the pile. select a pile diameter in accordance with the expected leg diameter and the capacity of pile driving equipment. derive the penetration from the shaft friction and tip bearing diagrams. assuming an equivalent soil subgrade modulus and full fixity at the base of the jacket, calculate the maximum moment in the pile and derive its wall thickness.
4.2 Deck Leg Sizes
adapt the diameter of the leg to that of the pile. determine the effective length from the degree of fixity of the leg into the deck (depending upon the height of the cellar deck). calculate the moment caused by wind loads on topsides and derive the appropriate thickness.
4.3 Jacket Bracings
select the diameter in order to obtain a span/diameter ratio between 30 and 40. calculate the axial force in the brace from the overall shear and the local bending caused by the wave assuming partial or total end restraint. derive the thickness such that the diameter/thickness ratio lies between 20 and 70 and eliminate any hydrostatic buckle tendency by imposing D/t<170/3ÖH (H is the depth of member below the free surface).
4.4 Deck Framing
select a spacing between stiffeners (typically 500 to 800mm). derive the plate thickness from formulae accounting for local plastification under the wheel footprint of the design forklift truck. determine by straight beam formulae the sizes of the main girders under "blanket" live loads and/or the respective weight of the heaviest equipments.
5. STATIC IN-PLACE ANALYSIS The static in-place analysis is the basic and generally the simplest of all analyses. The structure is modelled as it stands during its operational life, and subjected to pseudo-static loads. This analysis is always carried at the very early stage of the project, often from a simplified model, to size the main elements of the structure. 5.1 Structural Model 5.1.1 Main Model
The main model should account for eccentricities and local reinforcements at the joints. Typical models for North Sea jackets may feature over 800 nodes and 4000 members. 5.1.2 Appurtenances The contribution of appurtenances (risers, J-tubes, caissons, conductors, boat-fenders, etc.) to the overall stiffness of the structure is normally neglected. They are therefore analysed separately and their reactions applied as loads at the interfaces with the main structure. 5.1.3 Foundation Model Since their behaviour is non-linear, foundations are often analysed separately from the structural model. They are represented by an equivalent load-dependent secant stiffness matrix; coefficients are determined by an iterative process where the forces and displacements at the common boundaries of structural and foundation models are equated. This matrix may need to be adjusted to the mean reaction corresponding to each loading condition. 5.2 Loadings This Section is a reminder of the main types of loads, which are described in more detail in Lectures 15A.2 and 15A.3. 5.2.1 Gravity Loads Gravity loads consist of:
dead weight of structure and equipments. live loads (equipments, fluids, personnel).
Depending on the area of structure under scrutiny, live loads must be positioned to produce the most severe configuration (compression or tension); this may occur for instance when positioning the drilling rig. 5.2.2 Environmental Loads Environmental loads consist of wave, current and wind loads assumed to act simultaneously in the same direction. In general eight wave incidences are selected; for each the position of the crest relative to the platform must be established such that the maximum overturning moment and/or shear are produced at the mudline. 5.3 Loading Combinations The static in-place analysis is performed under different conditions where the loads are approximated by their pseudo-static equivalent. The basic loads relevant to a given condition are multiplied by the appropriate load factors and combined to produce the most severe effect in each individual element of the structure. 6. DYNAMIC ANALYSIS A dynamic analysis is normally mandatory for every offshore structure, but can be restricted to the main modes in the case of stiff structures.
6.1 Dynamic Model The dynamic model of the structure is derived from the main static model. Some simplifications may however take place:
local joint reinforcements and eccentricities may be disregarded. masses are lumped at the member ends. the foundation model may be derived from cyclic soil behaviour.
6.2 Equations of Motion The governing dynamic equations of multi-degrees-of-freedom systems can be expressed in the matrix form: MX'' + CX' + KX = P(t) where M is the mass matrix C is the damping matrix K is the stiffness matrix X, X', X'' are the displacement, velocity and acceleration vectors (function of time). P(t) is the time dependent force vector; in the most general case it may depend on the displacements of the structure also (i.e. relative motion of the structure with respect to the wave velocity in Morison equation). 6.2.1 Mass The mass matrix represents the distribution of masses over the structure. Masses include that of the structure itself, the appurtenances, liquids trapped in legs or tanks, the added mass of water (mass of water displaced by the member and determined from potential flow theory) and the mass of marine growth. Masses are generally lumped at discrete points of the model. The mass matrix consequently becomes diagonal but local modes of vibration of single members are ignored (these modes may be important for certain members subjected to an earthquake). The selection of lumping points may significantly affect the ensuing solution. As a further simplification to larger models involving considerable degrees-of-freedom, the system can be condensed to a few freedoms while still retaining its basic energy distribution. 6.2.2 Damping Damping is the most difficult to estimate among all parameters governing the dynamic response of a structure. It may consist of structural and hydrodynamic damping. Structural Damping Structural damping is associated with the loss of energy by internal friction in the material.
It increases with the order of the mode, being roughly proportional to the strain energy involved in each. Hydrodynamic Damping Damping provided by the water surrounding the structure is commonly added to the former, but may alternatively be accounted as part of the forcing function when vibrations are close to resonance (vortex-shedding in particular). Representation of Damping Viscous damping represents the most common and simple form of damping. It may have one of the following representations:
modal damping: a specific damping ratio z expressing the percentage to critical associated with each mode (typically z = 0,5% structural; z = 1,5% hydrodynamic) proportional damping: defined as a linear combination of stiffness and mass matrices.
All other types of non-viscous damping should preferably be expressed as an equivalent viscous damping matrix. 6.2.3 Stiffness The stiffness matrix is in all aspects similar to the one used in static analyses. 6.3 Free Vibration Mode Shapes and Frequencies The first step in a dynamic analysis consists of determining the principal natural vibration mode shapes and frequencies of the undamped, multi-degree-of-freedom structure up to a given order (30th to 50th). This consists in solving the eigenvalue problem: KX = l MX For rigid structures having a fundamental vibration period well below the range of wave periods (typically less than 3 s), the dynamic behaviour is simply accounted for by multiplying the time-dependent loads by a dynamic amplification factor (DAF):
DAF = where b = TN/T is the ratio of the period of the structure to the wave period. 6.4 Modal Superposition Method A convenient technique consists of uncoupling the equations through the normal modes of the system. This method is only applicable if:
each mass, stiffness and damping matrix is time-independent. non-linear forces are linearized beforehand (drag).
The total response is obtained by summing the responses of the individual single-degree-of-freedom oscillators associated to each normal mode of the structure. This method offers the advantage that the eigen modes provide substantial insight into the problem, and can be re-used for as many subsequent response calculations as needed at later stages.
It may however prove time-consuming when a large number of modes is required to represent the response accurately. Therefore:
the simple superposition method (mode-displacement) is applied to a truncated number of lowest modes for predicting earthquake response. it must be corrected by the static contribution of the higher modes (mode-acceleration method) for wave loadings.
6.4.1 Frequency Domain Analysis Such analysis is most appropriate for evaluating the steady-state response of a system subjected to cyclic loadings, as the transient part of the response vanishes rapidly under the effect of damping. The loading function is developed in Fourier series up to an order h:
p(t) = The plot of the amplitudes pj versus the circular frequencies wj is called the amplitude power spectra of the loading. Usually, significant values of pj only occur within a narrow range of frequencies and the analysis can be restricted to it. The relationship between response and force vectors is expressed by the transfer matrix H, such as: H = [-M w2 + i x C w + K] the elements of which represent:
Hj,k = The spectral density of response in freedom j versus force is then:
The fast Fourier transform (FFT) is the most efficient algorithm associated with this kind of analysis. 6.4.2 Time Domain Analysis The response of the i-th mode may alternatively be determined by resorting to Duhamel's integral: Xj(t) = The overall response is then obtained by summing at each time step the individual responses over all significant modes. 6.5 Direct Integration Methods Direct step-by-step integration of the equations of motion is the most general method and is applicable to:
non-linear problems involving special forms of damping and response-dependent loadings. responses involving many vibration modes to be determined over a short time interval.
The dynamic equilibrium at an instant t is governed by the same type of equations, where all matrices (mass, damping, stiffness, load) are simultaneously dependent on the time and structural response as well. All available integration techniques are characterised by their stability (i.e. the tendency for uncontrolled divergence of amplitude to occur with increasing time steps). Unconditionally stable methods are always to be preferred (for instance Newmark-beta with b = 1/4 or Wilson-theta with q = 1,4). 7. CONCLUDING SUMMARY
The analysis of offshore structures is an extensive task. The analytical models used in offshore engineering are in some respects similar to those used for other types of steel structures. The same model is used throughout the analysis process. The verification of an element consists of comparing its characteristic resistance(s) to a design force or stress. Several methods are available. Simple rules are available for preliminary member sizing. Static in-plane analysis is always carried out at the early stage of a project to size the main elements of the structure. A dynamic analysis is normally mandatory for every offshore structure.
A.5 - Analysis II OBJECTIVE/SCOPE To present the analysis procedures for offshore structures relating to fatigue, abnormal and accident conditions, load-out and transportation, installation and local design. PREREQUISITES Lecture 15A.1: Offshore Structures: General Introduction Lecture 15A.2: Loads I: Introduction and Environmental Loads Lecture 15A.3: Loads II: Other Loads RELATED LECTURES Lecture 15A.4: Analysis I SUMMARY Methods of fatigue analysis are described including the fatigue model (structural, hydrodynamic loading, and joint stress models) and the methods of fatigue damage assessment. Abnormal and accidental conditions are considered relating to earthquake, impact and progressive collapse. Analyses required for load-out and transportation and for installation are outlined. Local analysis for specific parts of the structure which are better treated by dedicated models outside of the global analysis are identified. 1. FATIGUE ANALYSIS A fatigue analysis is performed for those structures sensitive to the action of cyclic loadings such as:
wave (jackets, floating structures). wind (flare booms, stair towers). structures under rotating equipments.
1.1 Fatigue Model 1.1.1 Structural Model The in-place model is used for the fatigue analysis. Quasi-static analysis is often chosen; it permits all local stresses to be comprehensively represented. The dynamic effects are accounted for by factoring the loads by the relevant DAF. Modal analysis may be used instead; it offers computational efficiency, but may also overlook important local response modes, particularly near the waterline where direct wave action causes high out-of-plane bending (see Section 5.2). The mode - acceleration method may overcome this problem. 1.1.2 Hydrodynamic Loading Model A very large number of computer runs may be necessary to evaluate the stress range at the joints. The wave is repeatedly generated for:
different blocks of wave heights (typically from 2 to 28m in steps of 2m), each associated with a characteristic wave and zero-upcrossing period. different incidences (typically eight). different phases to determine the stress range for a given wave at each joint.
1.1.3 Joint Stress Model Nominal joint stresses are calculated for eight points around the circumference of the brace. The maximum local (hot spot) stress is obtained by multiplying the former by a stress concentration factor (SCF) given by parametric formulae which are functions of the joint geometry and the load pattern (balanced/unbalanced). 1.1.4 Fatigue Damage Model The fatigue failure of joints in offshore structures primarily depends on the stress ranges and their number of occurrences, formulated by S-N curves: log Ni = log a + mlog Dsi The number of cycles to failure Ni corresponds to a stress range. The effect of the constant stresses, mainly welding residual stresses, is implicitly accounted for in this formulation. The cumulative damage caused by ni cycles of stress Dsi, over the operational life of the platform (30 to 50 years) is obtained by the Palmgren-Miner rule:
D= The limit of this ratio depends on the position of the joint with respect to the splash zone (typically +/-4m on either side of the mean sea level). The ratio should normally not exceed:
1,0 above, 0,1 within, 0,3 below the splash zone.
1.1.5 Closed Form Expression The damage may alternatively be expressed in closed form:
D= where a, m are coefficients of the selected S-N curve. Ds is the stress range exceeded once in N cycles. k is a long-term distribution parameter, depending on the position of the joint in the structure. N is the total number of cycles. 1.2 Deterministic Analysis
This analysis consists of time-domain analysis of the structure. The main advantage of this representation is that non-linear effects (drag, high order wave theories) are handled explicitly. A minimum of four regular waves described in terms of height and associated period are considered for each heading angle. 1.3 Spectral Analysis Waves of a given height are not characterised by a unique frequency, but rather by a range of frequencies. If this range corresponds to a peak in the structural response, the fatigue life predicted by the deterministic method can be seriously distorted. This problem is overcome by using a scatter diagram, in which the joint occurrence of wave height and period is quantified. Wave directionality may also be accounted for. Eventually the most thorough representation of a sea state consists of:
the frequency spectrum constructed from the significant wave heights and mean zero-crossing periods. the directionality function derived from the mean direction and associated spreading function.
This approach requires that the physical process be approximately linear (or properly linearised) and stationary. Transfer functions TF are determined from time-domain analyses involving various wave heights, each with different period and incidence:
The response has normally a narrow-banded spectrum and can be described by a Rayleigh distribution. The zero-upcrossing frequency of stress cycles is then approximated by:
Tz = where mn is the nth order moment of the response. The significant stress range is readily obtained for each sea state as:
ssig = where S(w,q) is the directional wave energy spectrum. 1.4 Wind Fatigue 1.4.1 Wind Gusts The fatigue damage caused by the fluctuating part of wind (gusts) on slender structures like flare booms and bridges is usually predicted by spectral methods. The main feature of such analysis is the introduction of coherence functions accounting for the spanwise correlation of forces. 1.4.2 Vortex Shedding
Vortex induced failure occurs for tubes subjected to a uniform or oscillating flow of fluid. Within a specific range of fluid velocities, eddies are shed at a frequency close to the resonant frequency of the member. This phenomenon involves forced displacements, which can be determined by models such as those suggested in [1]. 2. ABNORMAL AND ACCIDENTAL CONDITIONS This type of analysis addresses conditions which may considerably affect the integrity of the structure, but only have a limited risk of occurrence. Typically all events with a probability level less than the 10-4 threshold are disregarded. 2.1 Earthquake Analysis 2.1.1 Model Particular attention shall be paid to:
foundations: the near field (i.e. the soil mass in the direct vicinity of the structure) shall accurately represent loaddeflection behaviour. As a general rule the lateral foundation behaviour is essentially controlled by horizontal ground motions of shallow soil layers. modal damping (in general taken as 5% and 7% of critical for ULS and PLS analyses respectively).
2.1.2 Ductility Requirements The seismic forces in a structure are highly dependent on its dynamic characteristics. Design recommendations are given by API to determine an efficient geometry. The recommendations call for:
providing sufficient redundancy and symmetry in the structure. favouring X-bracings instead of K-bracings. avoiding abrupt changes in stiffness. improving the post-buckling behaviour of bracings.
2.1.3 Analysis Method Earthquake analyses can be carried out according to the general methods presented in Lecture 15A.4. However their distinctive feature is that they represent essentially a base motion problem and that the seismic loads are therefore dependent on the dynamic characteristics of the structure. Modal spectral response analysis is normally used. It consists of a superposition of maximum mode response and forms a response spectrum curve characteristic of the input motion. This spectrum is the result of time-histories of a SDOF system for different natural periods of vibration and damping. Direct time integration can be used instead for specific accelerograms adapted to the site. 2.2 Impact The analysis of impact loads on structures is carried out locally using simple plastic models [2]. Should a more sophisticated analysis be required, it can be accomplished using time-domain techniques presented in Section 6 of Lecture 15A.4.
The whole energy must be absorbed within acceptable deformations. 2.2.1 Dropped Object/Boat Impact When a wellhead protection cover is hit by a drill collar, or a tube (jacket leg, fender) is crushed by a supply boat, two load/deformation mechanisms occur simultaneously:
local punch-through (cover) or denting (tube). global deformation along plastic hinges with possible appearance of membrane forces.
2.2.2 Blast and Fire Owing to the current lack of definitive guidance regarding explosions and fire, the behaviour of structures in such events has so far been only predicted by simple models based on:
equivalent static overpressure and plastic deformation of plates for blast analysis. the reduction of material strength and elastic modulus under temperature increase.
In the aftermath of recent mishaps however, more accurate analyses may become mandatory, based on a better understanding of the pressure-time histories and the effective resistance and response of structures to explosions and fire. 2.3 Progressive Collapse Some elements of the structure (legs, bracings, bulkheads) may partially or completely loose their strength as a result of accidental damage. The purpose of such analysis is to ensure that the spare resistance of the remaining structure is sufficient to allow the loads to redistribute. Since such a configuration is only temporary (mobilisation period prior to repairs) and that operations will also be restricted around the damaged area, reduced live and environmental loads are generally accepted. In this analysis, the damaged elements are removed from the model. Their residual strength may be represented by forces applied at the boundary nodes with the intact structure. 3. LOAD OUT & TRANSPORTATION 3.1 Load-Out The load-out procedure consists in moving the jacket or module from its construction site to the transportation barge by skidding, or by using trailers underneath it. The barge may be floating and is continuously deballasted as the package progresses onto it, or grounded on the bottom of the harbour. 3.1.1 Skidding The most severe configuration during skidding occurs when the part of the structure is cantilevering out:
from the quayside before it touches the barge. from the barge just after it has left the quay.
The analysis should also investigate the possibility of high local reactions being the result of settlement of the skidway or errors in the ballasting procedure.
3.1.2 Load-Out by Trailers As the reaction on each trailer can be kept constant, analysis of load-out by trailers only requires a single step to determine the optimal distribution of trailers. 3.2 Transportation 3.2.1 Naval Architectural Model The model consists of the rigid-body assembly of the barge and the structure. Barges are in general characterised by a low length/beam ratio and a high beam/draught ratio, as well as sharp corners which introduce heavy viscous damping. For jacket transport, particular care shall be taken in the representation of overhanging parts (legs, buoyancy tanks) which contribute significantly to the righting moment. Dry-transported decks and modules may be simply represented by their mass and moments of inertia. This analysis shall provide the linear and angular accelerations and displacements of the structure to be entered in the structural model as inertia forces, and also the partition and intensity of buoyancy and slamming forces. 3.2.2 Structural Model The jacket model is a simplified version of the in-place model, from which eccentricities and local reinforcements may be omitted. The barge is modelled as a plane grid, with members having the equivalent properties of the longitudinal and transversal bulkheads. As the barge passes over a wave trough or a crest, a portion only of the barge is supported by buoyancy (long barges may be spanning over a whole trough or be half-cantilevered). The model therefore represents the jacket and the barge as two structures coupled together by the seafastening members. 4. INSTALLATION 4.1 Launching 4.1.1 Naval Architectural Model A three dimensional analysis is carried out to evaluate the global forces acting on the jacket at various time steps during the launch sequence. At each time step, the jacket/barge rigid body system is repositioned to equilibrate the internal and external forces produced by:
jacket weight, inertia, buoyancy and drag forces. barge weight, buoyancy and ballast forces. vertical reactions and friction forces between jacket and barge.
The maximum reaction on the rocker arm is normally obtained when the jacket just starts rotating about the rocker hinge. 4.1.2 Structural Model
The structural model is in all aspects identical to the one used for the transportation analysis, with possibly a finer representation of the launch legs. The rocker arm is also represented as a vertical beam hinged approximately at midspan. Interface loads obtained by the rigid body analysis are input at boundary conditions on the launch legs. All interface members must remain in compression, otherwise they are inactivated and the analysis restarted for that step. Once the tilting phase has begun, the jacket is analysed at least for each main leg node being at the vertical of the rocker arm pivot. 4.2 Upending No dedicated structural analysis is required for this phase, which is essentially a naval architecture problem. A local analysis of the lugs is performed for crane-assisted upendings. 4.3 Docking Docking of a jacket onto a pre-installed template requires guides to be analysed for local impact. The same requirement applied for bumpers to aid the installation of modules. 4.4 Unpiled Stability The condition where the jacket may for a while stand unpiled on the seafloor is analysed for the design installation wave. The stability of the jacket as a whole (overturning tendency) is investigated, together with the resistance of the mudmats against soil pressure. 4.5 Piling The piles are checked during driving for the dynamic stresses caused by the impact wave of the hammer blow. The maximum cantilevered (stick-up) length of pile must be established for the self-weight of the pile and hammer combined, accounting for first and second order moments arising from the pile batter. Hydrodynamic actions are added for underwater driving. Elements in the vicinity of the piles (guides, sleeves) shall also be checked, see Section 5.1. 4.6 Lifting 4.6.1 Model The model used for the lift analysis of a structure consists of the in-place model plus the representation of the rigging arrangement (slings, spreader frames). For single lifts the slings converge towards the hook joint, which is the sole vertical support in the model and shall be located exactly on the vertical through the centre of gracity (CoG) of the model. For heavier dual-crane lifts, the CoG shall be contained in the vertical plane defined by the two hook joints. The mathematical instability of the model with respect to horizontal forces is avoided by using soft horizontal springs at the padeyes. The force and elongation in these springs should always remain small. 4.6.2 Design Factors Different factors are applied to the basic sling forces to account for specific effects during lifting operations.
4.6.2.1 Skew Load Factor (SKL) This factor represents the effect of fabrication tolerances and lack-of-fit of the slings on the load repartition in a statically undetermined rigging arrangement (4 slings or more). Skew factors may either be directly computed by applying to a pair of opposite slings a temperature difference such that their elongation/shortening corresponds to the mismatch, or determined arbitrarily (typically 1/3 - 2/3 repartition). 4.6.2.2 Dynamic Amplification Factor (DAF) This factor accounts for global dynamic effects normally experienced during lifting operations. DnV [24] recommends minimum values as follows: Lifted Weight W (tonnes)
up to 100 t
100 t to 1000t
1000 t to 2500t
more than 2500 t
DAF offshore
1,30
1,20
1,15
1,10
DAF inshore
1,15
1,10
1,05
1,05
4.6.2.3 Tilt Effect Factor (TEF) This factor accounts for additional sling loading caused by the rotation of the lifted object about a horizontal axis and by the longitudinal deviation of the hooks from their theoretical position in the case of a multi-hook lift. It shall normally be based on 5° and 3° tilt respectively depending on whether cranes are on different vessels or not. 4.6.2.4 Yaw Effect Factor (YEF) This factor accounts for the rotation of the lifted object about a vertical axis (equal to 1,05 typically). 4.6.3 Consequence Factors Forces in elements checked under lift conditions are multiplied by a factor reflecting the consequence a failure of that specific element would have on the integrity of the overall structure:
1,30 for spreader frames, lifting points (padeyes) and their attachment to the structure. 1,15 for all members transferring the load to the lifting points. 1,00 for other elements.
5. LOCAL ANALYSES AND DESIGN Local analyses address specific parts of the structure which are better treated by dedicated models outside the global analysis. The list of analyses below is not exhaustive and more information can be found in [1-24] which provide a complete design procedure in each particular case. 5.1 Pile/Sleeve Connections Underwater pile/sleeve connection is usually achieved by grouting the annulus between the outside of the pile and the inner sleeve. The main verifications address:
the shear stresses in the concrete.
the fatigue damage in the shear plates and the attachment welds to the main jacket accumulated during pile driving and throughout the life of the platform.
5.2 Members within the Splash Zone Horizontal members (conductor guide frames in particular) located within the splash zone (+/-5m on either side of the meansea-level approximately) shall be analysed for fatigue caused by repeated wave slamming. A slamming coefficient Cs=3,5 is often selected. 5.3 Straightened Nodes Typical straightened nodes (ring-stiffened nodes, bottle legs nodes with diaphragms) are analysed by finite-elements models, from which parametric envelope formulae are drawn and applied to all nodes representative of the same class. 5.4 Appurtenances 5.4.1 Risers, Caissons & J-Tubes Static In-Place and Fatigue Risers, caissons and J-tubes are verified either by structural or piping programs for the action of environmental forces, internal pressure and temperature. Particular attention is paid to the bends not always satisfactorily represented by structural programs and the location of the touch-down point now known a-priori. A fatigue analysis is also performed to assess the fatigue damage to the clamps and the attachments to the jacket. Pull-In J-tubes are empty ducts continuously guiding a post-installed riser pulled inside. They are verified by empirical plastic models against the forces generated during pull-in by the friction of the cable and the deformation of the pull head, see [22]. 5.4.2 Conductors Conductors are analysed in-place as beam columns on discrete simple supports, these being provided by the horizontal framing of the jacket (typically 20 to 25 m span). The installation sequence of the different casings must be considered to assess the distribution of stresses in the different tubes forming the overall composite section. Also the portion of compression force in the conductor caused by the hanging casings is regarded as an internal force (similar to prestressing) which therefore does not induce any buckling tendency, see [23]. 5.5 Helidecks The helideck is normally designed to resist an impact load equal to 2,5 times the take-off weight of the heaviest helicopter factored by a DAF of 1,30. Plastic theories are applicable for designing the plate and stiffeners, while the main framing is analysed elastically. 5.6 Flare Booms Analyses of flare booms particularly consider:
variable positions during installation (horizontal pick-up from the barge, lift upright). reduced material characteristics due to high temperature in the vicinity of the tip during operation. dynamic response under gusty winds. local excitation of diagonals by wind vortex-shedding.
6. CONCLUDING SUMMARY
With the trend to ever deeper and more slender offshore structures in yet harsher environments, more elaborate theories are necessary to analyse complex situations. There is a risk for the Engineer having increasingly to rely on the sole results of computer analyses at the expense of sound design practice. To retain enough control of the process of analysis, the following recommendations are given:
× check the interfaces between the different analyses and ensure the consistency of the input/output. × verify the validity of the data resulting from a complex analysis against a simplified model, which can also be used to assess the influence of a particular parameter. × make full use of "good engineering judgement" to criticise the unexpected results of an analysis. 7. REFERENCES [1] Skop R.A. & Griffin O.M., An Heuristic Model for Determining Flow-Induced Vibrations of Offshore Structures/OTC paper 1843, May 1973. [2] De Oliveira J.G., The Behaviour of Steel Offshore Structures under Accidental Collisions/OTC paper 4136, May 1981. [3] API-RP2A, Recommended Practice for Planning, Designing and Constructing Fixed Offshore Platforms/18th edition, September1989. [4] DnV, Rules for the Classification of Fixed Offshore Structures, September 1989. [5] DnV, Standard for Insurance Warranty Surveys in Marine Operations, June 1985. [6] NPD, Regulation for Structural Design of Loadbearing Structures Intended for Exploitation of Petroleum Resources, October1984 and Veiledning om Utforming, Beregning og Dimensjonering av Stalkonstruksjoner i Petroleumsvirksomheten, December1989. [7] DoE, Offshore Installations: Guidance on Design and Construction/London, April 1984. [8] McClelland B. & Reifel M.D., Planning and Design of Fixed Offshore Platforms/Van Nostrand Reinhold, 1986. [9] UEG, Node Flexibility and its Effect on Jacket Structures/CIRIA Report UR22, 1984. [10] Hallam M.G., Heaf N.J. & Wootton L.R., Dynamics of Marine Structures/ CIRIA Report UR8 (2nd edition), October 1978. [11] Wilson J.F., Dynamics of Offshore Structures/Wiley Interscience, 1984. [12] Clough R.W. & Penzien J., Dynamics of Structures/McGraw-Hill, New York, 1975. [13] Newland D.E., Random Vibrations and Spectral Analysis/Longman Scientific (2nd edition), 1984. [14] Zienkiewicz O.C., Lewis R.W. & Stagg K.G., Numerical Methods in Offshore Engineering/Wiley Interscience, 1978. [15] Davenport A.G., The Response of Slender Line-Like Structures to a Gusty Wind/ICE Vol.23, 1962.
[16] Williams A.K. & Rhinne J.E., Fatigue Analysis of Steel Offshore Structures/ICE Vol.60, November 1976. [17] Anagnostopoulos S.A., Wave and Earthquake Response of Offshore Structures: Evaluation of Modal Solutions/ASCE J. of the Structural Div., vol. 108, No ST10, October 1982. [18] Chianis J.W. & Mangiavacchi A., A Critical Review of Transportation Analysis Procedures/OTC paper 4617, May1983. [19] Kaplan P. Jiang C.W. & Bentson J, Hydrodynamic Analysis of Barge-Platform Systems in Waves/Royal Inst. of Naval Architects, London, April 1982. [20] Hambro L., Jacket Launching Simulation by Differentiation of Constraints/ Applied Ocean Research, Vol.4 No.3, 1982. [21] Bunce J.W. & Wyatt T.A., Development of Unified Design Criteria for Heavy Lift Operations Offshore/OTC paper 4192, May 1982. [22] Walker A.C. & Davies P., A Design Basis for the J-Tube Method of Riser Installation/J. of Energy Resources Technology, pp. 263-270, September 1983. [23] Stahl B. & Baur M.P., Design Methodology for Offshore Platform Conductors/J. of Petroleum Technology, November 1983. [24] DnV - Rules for the Classification of Steel Ships, January 1989.
A.6: Foundations OBJECTIVE\SCOPE
to classify different types of piles to understand main design methods to cover various methods of installation
PREREQUISITES Lecture 1B.2.2: Limit State Design Philosophy and Partial Safety Factors Lectures 10.6: Shear Connection Lectures 12.4: Fatigue Behaviour of Hollow Section Joints Lecture 15A.12: Connections in Offshore Deck Structures Lecture 17.5: Requirements and Verifications of Seismic Resistant Structures A general knowledge of design in offshore structures and an understanding of offshore installation are also required. SUMMARY In this lecture piled foundations for offshore structures are presented. The lecture starts with the classification of soil. The main steps in the design of piles are then explained. The different kinds of piles and hammers are described. The three main execution phases are briefly discussed: fabrication, transport and installation. 1. INTRODUCTION 1.1 Classification of Soils The stratigraphy of the sea bed results from a complex geological process during which various materials were deposited, remoulded and pressed together. Soil texture consists of small mineral or organic particles basically characterized by their grain size and mutual interaction (friction, cohesion). The properties of a specific soil depend mainly on the following factors:
density. water content. over consolidation ratio.
For design purposes the influence of these factors on soil behaviour is expressed in terms of two fundamental parameters:
friction angle. undrained shear strength Cu.
Since the least significant of either of these parameters is often neglected, soils can be classified within "ideal" categories:
granular soils. cohesive soils.
1.2 Granular Soils Granular soils are non-plastic soils with negligible cohesion between particles. They include:
sands : characterized by large to medium particle sizes (1mm to 0,05mm) offering a high permeability, silts : characterized by particle sizes between 0,05 and 0,02mm; they are generally over-consolidated; they may exhibit some cohesion.
1.3 Cohesive Soils Clays are plastic soils with particle sizes less than 0,002mm which tend to stick together; their permeability is low. 1.4 Multi-Layered Strata The nature and characteristics of the soil surrounding a pile generally vary with the depth. For analysis purposes, the soil is divided into several layers, each having constant properties throughout. The number of layers depends on the precision required of the analysis. 2. DESIGN Steel offshore platforms are usually founded on piles, driven deep into the soil (Figure 1). The piles have to transfer the loads acting on the jacket into the sea bed. In this section theoretical aspects of the design of piles are presented. Checking of the pile itself is described in detail in the Worked Example.
2.1 Design Loads These loads are those transferred from the jacket to the foundation. They are calculated at the mudline. 2.1.1 Gravity loads
Gravity loads (platform dead load and live loads) are distributed as axial compression forces on the piles depending upon their respective eccentricity. 2.1.2 Environmental loads Environmental loads due to waves, current, wind, earthquake, etc. are basically horizontal. Their resultant at mudline consists of:
shear distributed as horizontal forces on the piles. overturning moment on the jacket, equilibrated by axial tension/ compression in symmetrically disposed piles (upstream/downstream).
2.1.3 Load combinations The basic gravity and environmental loads multiplied by relevant load factors are combined in order to produce the most severe effect(s) at mudline, resulting in:
vertical compression or pullout force, and lateral shear force plus bending.
2.2 Static Axial Pile Resistance The overall resistance of the pile against axial force is the sum of shaft friction and end bearing. 2.2.1 Lateral friction along the shaft (shaft friction) Skin friction is mobilized along the shaft of the tubular pile (and possibly also along the inner wall when the soil plug is not removed). The unit shaft friction:
for sands: is proportional to the overburden pressure, for clays: is calculated by the "alpha" or "lambda" method and is a constant equal to the shear strength C u at great depth.
Lateral friction is integrated along the whole penetration of the pile. 2.2.2 End bearing End bearing is the resultant of bearing pressure over the gross end area of the pile, i.e. with or without the area of plug if relevant. The bearing pressure:
for clays: is equal to 9 ´ Cu. for sands: is proportional to the overburden pressure as explained in Section 6.4.2 of API-RP2A [1].
2.2.3 Pile penetration The pile penetration shall be sufficient to generate enough friction and bearing resistance against the maximum design compression multiplied by the appropriate factor of safety. No bearing resistance can be mobilized against pull-out: the friction available must be equated to the pull out force multiplied by the appropriate factor of safety. 2.3 Lateral Pile Resistance
The shear at the mudline caused by environmental loads is resisted by lateral bearing of the pile on the soil. This action may generate large deformations and high bending moments in the part of the pile directly below the mudline, particularly in soft soils. 2.3.1 P-y curves P-y curves represent the lateral soil resistance versus deflection. The shape of these curves varies with the depth and the type of soil at the considered elevation. The general shape of the curves for increasing displacement features:
elastic (linear) behaviour for small deflections, elastic/plastic behaviour for medium deflections, constant resistance for large deflections or loss of resistance when the soil skeleton deteriorates (clay under cyclic load in particular).
2.3.2 Lateral pile analysis For analysis purposes, the soil is modelled as lumped non-linear springs distributed along the pile. The fourth order differential equation which expresses the pile deformation is integrated by successive iterations, the secant stiffness of the soil springs being updated at each step. For large deformations, the second order contribution of the axial compression to the bending moment (P-Delta effect) shall be taken into account. 2.4 Pile Driving Piles installed by driving are forced into the soil by a ram hitting the top. The impact is transmitted along the pile in the form of a wave, which reflects on the pile tip. The energy is progressively lost by plastic friction on the sides and bearing at the tip of the pile. 2.4.1 Empirical formulae A considerable number of empirical formulae exist to predict pile driveability. Each formula is generally limited to a particular type of soil and hammer. 2.4.2 Wave equation This method of analysing the driving process consists of representing the ensemble of pile/soil/hammer as a onedimensional assembly of masses, springs and dashpots:
the pile is modelled as a discrete assembly of masses and elastic springs. the soil is idealized as a massless medium characterized by elastic-perfectly-plastic springs and linear dashpots. the hammer is modelled as a mass falling with an initial velocity. the cushion is represented by a weightless spring (see Figure 3). the pile cap is represented by a mass of infinite rigidity.
The energy of the ram hitting the top of the pile generates a stress wave in the pile, which dissipates progressively by friction between the pile and the soil and by reflection at the extremities of the pile. The plastic displacement of the tip relative to the soil is the set achieved by the blow. Curves can be drawn to represent the number of blows per unit length required to drive the pile at different penetrations. The wave equation, though representing the most rigorous assessment to date of the driving process, still suffers a lack of accuracy, mostly caused by the inaccuracies in the soil model. 3. DIFFERENT KINDS OF PILES Driven piles are the most popular and cost-efficient type of foundation for offshore structures. As shown in Figure 2, the following alternatives may be chosen when driving proves impractical:
insert piles. drilled and grouted piles. belled piles.
3.1 Driven Piles Piles are usually made up in segments. After placing and driving the first long segment, extension segments called add-ons are set on piece by piece as driving proceeds until the overall design length is achieved. In recent years one-piece piles have been widely used in the North Sea since the offshore work is considerably reduced. Wall thickness may vary. A thicker wall is sometimes required:
in sections from mudline down to a specified depth within which bending stresses are especially high, at the pile tip (driving shoe) to resist local bearing stresses while driving.
Uniform wall thickness is however preferable thus avoiding construction and installation problems. 3.2 Insert Piles Insert piles are smaller diameter piles driven through the main pile from which the soil plug has been previously drilled out. They are therefore not subjected to skin friction over the length of the main pile and can reach substantial additional penetration. The insert pile is welded to the main pile at the top of the jacket and the annular space between the tubes is grouted. This type of pile is used:
in a preplanned situation: performance is good although material and installation costs are higher than for normal driven piles. as an emergency procedure: when scheduled piles cannot be driven to the required penetration, resulting therefore in one of the following drawbacks.
× a thicker wall section of the main pile will be within the jacket height instead of below the mudline. × reduced friction area and end bearing pressure, × difficulties often noted for the setting-in of all the required volume of grouting, i.e. the concern is the leakage of grout or the impossibility to fill with the calculated volume of grout. 3.3 Drilled and Grouted Piles This procedure is the only means of installing piles with tension resistance in hard soils or soft rocks; it resembles that for drilling a conductor well. An oversized hole is initially drilled to the proposed pile penetration depth. The pile is then lowered down, sometimes centred in the hole by spacers and the annular space between the pile shaft and the surrounding soil is grouted. Design uncertainty results because:
hard soil formation softens when exposed to the water or mud used during drilling and exhibits lower skin friction resistance. in case of calcareous sand, external grouting just crushes the sand, slightly extending the effective pile diameter but not increasing the friction significantly.
3.4 Belled Piles
While belled piles, on land, are used to decrease the bearing stress under a pile, offshore belled piles provide a large bearing area to increase tip uplift resistance. The main pile, normally driven, serves here as a casing through which a rig drills a slightly oversized hole ahead. A belling tool (underreamer) then enlarges the socket to a conical bell with a base diameter a few times that of the main pile. A heavy reinforcement cage is lowered inside the bell which is subsequently filled with concrete made using fine aggregate (maximum size 10mm). 4. FABRICATION AND INSTALLATION 4.1 Fabrication The piles are usually made up of "cans" - cylinders of rolled plate with a longitudinal seam. Single cans are typically 1,5m long or more. Longitudinal seams of two adjacent segments are rotated 90° apart at least. Bevelling is mandatory should the wall thickness difference exceed 3mm between adjacent cans. Maximum deviation from straightness is specified (0.1% in length). Commonly used steel grade is X52 or X60. The outside surface of grouted piles should be free of mill scale and varnished. In certain instances, steel piles are protected underwater by sacrificial anodes or by impressed current. In the splash zone additional thickness to allow for corrosion (3mm for example) and epoxy or rubberized coating, monel or copper-nickel sheeting are provided. 4.2 Transportation 4.2.1 Barge transportation Pile segments are choked and fastened to the barge to prevent them from falling overboard under severe seastates. Pile plate should be thick enough to prevent any deformation caused by stacking. 4.2.2 Self floating mode This method is attractive where long segments of pile are to be lifted and set in guides far below the sea surface (skirt piles for example). The ends of the piles are sealed by steel closure plates or rubber diaphragms which should be able to resist wave slamming during the tow. 4.2.3 Transport within the jacket The piles are pre-set inside the main legs or in the guides/sleeves, generating additional weight and possibly buoyancy (if closed). They are held in place by shims which prevent them from escaping from their guides during launch and uprighting of the jacket. Several piles are driven immediately after the jacket has touched down, providing initial stability against the action of waves and current. 4.3 Hammers Piles are positioned:
either inside the jacket legs, extending the full height of the jacket,
or encased in sleeves protruding at the bottom of the jacket, running vertical or parallel to the legs (typical batter 1/12 to 1/6).
Piles can then be driven using any type of hammer (or a combination of types). Hammers are illustrated in Figure 3. 4.3.1 Steam hammers Steam hammers are widely used for offshore installation of jackets. They are generally single acting with rates of up to 40 blows/minute. Energies of current hammers range from 60 000 to 1 250 000 ft lb/blow. (82KNm to 1725KNm per blow). During driving, the hammer with attached driving head rides the pile rather than being supported by leads. The hammer line from the crane boom is slackened so as to prevent transmission of impact and vibration into the boom. 4.3.2 Diesel hammers Diesel hammers are much used at offshore terminals. They are lighter to handle and less energy consuming than steam hammers, but their effective energy is limited. 4.3.3 Hydraulic hammers Hydraulic hammers are dedicated to underwater driving (skirt piles terminating far below the sea surface). Menck hydraulic hammers are widely used. They utilize a solid steel ram and a flexible steel pile cap to limit impact forces. They are double acting. Hydraulic fluid under high pressure is used to force a piston or set of pistons, and in turn, the ram up and down. Properties of some hammers used offshore are shown in Table 1. A selection of large offshore pile driving hammers driving on heavy piles is also shown in Table 2. 4.3.4 Selection of hammer size Selection of hammer size is based on:
experience of similar situations (see Quality Control: Section 4.6), numerical modelling of driving for each particular site (see Pile Driving: Section 2.4)
Typical values of pile sizes, wall thicknesses, and hammer energies for steam hammers are shown in Table 3. 4.4 Installation 4.4.1 Pile handling and positioning Figure 4 shows the different ways of providing lifting points for positioning pile sections. Padeyes are generally used (welded in the fabrication yard; their design should take into account the changes in load direction during lifting). Padeyes are then carefully cut before lowering the next pile section.
Sketch E shows the different steps for the positioning of pile sections:
pile or add-on lifted from the barge deck. rotation of the crane to position add-on. installing and lowering of the pile add-on.
4.4.2 Pile connections Different solutions for connecting pile segments back-to-back are used:
either by welding, Shielded Metal Arc Welding (SMAW) or flux-cored, segments held temporarily by internal or external stabbing guides as shown in Figure 4. Welding time depends upon:
- pile wall thickness: 3 hours for 1in. thick (25,4mm); 16 hours for 3in. thick, (76,2mm) (typical). - number and qualification of the welders. - environmental conditions.
or by mechanical connectors (as shown in Figure 4):
- breech block (twisting method). - lug type (hydraulic method). 4.4.3 Hammer placement Figure 5 shows the different steps of this routine operation:
lifting from the barge deck.
positioning over pile by booming out or in (the bell of the hammer acts as a stabling guide... very helpful in rough weather). alignment of the pile cap. lowering leads after hammer positionment.
Each add-on should be designed to prevent bending or buckling failure during installation and in-place conditions. 4.4.4 Driving Some penetration under the self weight of the pile is normal. For soft soil conditions, particular measures are taken to avoid an uncontrolled run. Piles are then driven or drilled until pile refusal. Pile refusal is defined as the minimum rate of penetration beyond which further advancement of the pile is no longer achievable because of the time required and the possible damage to the pile or to the hammer. A widely accepted rate for defining refusal is 300 blows/foot (980 blows/metre). 4.5 Pile-to-Jacket Connections 4.5.1 Welded shims The shims are inserted at the top of the pile within the annulus between the pile and jacket leg (see Figure 6) and welded afterwards.
4.5.2 Mechanical locking system
This metal-to-metal connection is achieved by a hydraulic swaging tool lowered inside the pile and expanding it into machined grooves provided in the sleeves at two or three elevations as shown on Figure 7.
This type of connection is most popular for subsea templates. It offers immediate strength and the possibility to re-enter the connection should swaging prove incomplete. 4.5.3 Grouting This hybrid connection is the most commonly used for connecting piles to the main structure (in the mudline area). Forces are transmitted by shear through the grout. Figure 8 shows the two types of packers commonly used. The expansive, non-shrinking grout must fill completely the annulus between the pile and leg (or sleeve).
Bonding should be excellent; it is improved by shear connectors (shear keys, strips or weld beads disposed on the surface of the sleeve and pile in contact with the grout). The width of the annulus between pile and sleeve should be maintained constant by use of centralizers and be limited to:
1,5in. minimum, (38,1mm) about 4in. (101,6mm) maximum (to avoid destruction of the tensile strength of the grout by internal microcracking).
Packers are used to confine the grout and prevent it from escaping at the base of the sleeve. Packers are often damaged during piling and are therefore:
installed in a double set. attached to the base of the sleeve to protect them during pile entry and driving.
Thorough filling should be checked by suitable devices, e.g. electrical resistance gauges, radioactive tracers, well-logging devices or overflow pipes checked by divers. 4.6 Quality Control Quality control shall:
confirm the adequacy of the foundation with respect to the design. provide a record of pile installation for reference to subsequent driving of nearby piles and future modifications to the platform.
The installation report shall mention:
pile identification (diameter and thickness). measured lengths of add-ons and cut-offs. self penetration of pile (under its own weight and under static weight of the hammer). blowcount throughout driving with identification of hammer used and energy, as shown in Figure 9. record of incidents and abnormalities:
- unexpected behaviour of the pile and/or hammer. - interruptions of driving (with set-up time and blowcount subsequently required to break the pile loose). - pile damage if any.
elevations of soil plug and internal water surface after driving. information about the pile/structure connection:
- equipment and procedure employed. - overall volume of grout and quality. - record of interruptions and delays.
4.7 Contingency Plan Contingency documents should provide back-up solutions in case "unforeseen" events occur such as:
impossibility to get the required pile penetration. mechanical breakdown of the hammer. grout pipe blockage.
5. CONCLUDING SUMMARY This lecture has described:
the difficult aspects of foundations in a variety of soils.
the multiplicity of solutions and the different kind of piles and hammers. the complexity of the process from design to installation.
6. REFERENCES [1] API-RP2A, "Recommended Practice for Planning, Designing and Constructing Fixed Offshore Platforms", American Petroleum Institute, Washington, D.C., 18th ed., 1989. 7. ADDITIONAL READING 1.
McClelland, B. and Reifel, M. D., Planning and design of fixed offshore platforms, Von Mostrand Reinhold Company (1982). 2. Bowles, J. E., Foundation analysis and design, MacGraw Hill Book Company (4th edition 1988). 3. Bowles, J. E., Analytical and computer methods in Foundation Engineering, MacGraw Hill Book Company (1983). 4. Poulos, H. G. and Davis, E. H., Pile foundation analysis and design, John Wiley and Sons (1980). 5. Graff, W. J., Introduction to offshore structures, Gulf Publishing Company (1981). 6. Le Tirant, P., Reconnaissance des sols en mer pour l'implantation des ouvrages Pétroliens, Technip (1976) 7. Pieux dans les formatines carbonates - Technip ARGEMA (1988). 8. Capacité patante des pieux - Technip ARGEMA (1988). 9. Dawson, T. H., Offshore Structural Engineering, Prentice Hall Inc (1983). 10. Gerwick, Ben C., Construction of Offshore Structures, John Wiley and Sons (1986). A. Air/Steam Hammers Make
Model
Rated
Ram
Max.
Energy
Weight
Stroke
Std. Pilecap Weight
(ft-lbs)
(kips)
Hammer Weight
Rated Operating
Steam
Air
Hose
Rated
Consumption
Consumption
ST/F
BPM
(lbs ht)
(lbs ht)
.....
Pressure
(m) (kips)
Conmaco
Typical
(w/leads) (kips)
(psi)
6850
510.000
85
72
57,5
312
180
31.500
5650
325.000
65
60
59,0
262
160
5300
150.000
30
60
12,7
92
160
8.064
300
90.000
30
36
12,7
86
150
200
60.000
20
36
12,7
74
Menck
12500
1.582.220
275,58
69
154,32
(MRBS)
8800
954.750
194,01
59
8000
867.960
176,37
7000
632.885
5000 4600
7.500
2@4
40
3@4
45
1.711
4
46
6.944
1.471
3
54
120
5.563
1.195
3
59
853
171
53.910
26.500
2@6
36
103,62
600
150
32.400
16.700
8
36
59
85,98
564
142
30.860
15.900
8
38
154
49
92,4
583
156
30.800
14.830
4@4
35
542.470
110,23
59
66,14
335
150
20.940
10.400
6
40
499.070
101,41
59
52,91
313
142
19.840
9.900
6
42
3000
325.480
66,14
59
33,07
205
142
12.130
6.000
5
42
1800
189.850
38,58
59
22,05
125
142
7.060
3.700
4
44
850
93.340
18,96
50
11,5
64
142
3.530
1.950
3
45
OS60
18.000
60
36
120.000
40
36
60.000
20
36
38,65
150
MKT
OS40 OS20
C. Hydraulic Hammers Make
Model
Rated Energy
Ram Weight
Standard
Hammer Weight
Pilecap Weight (kips)
HMB
Menck
(ft-lb)
(kips)
(kips)
4000
1.200.000
205
490
3000A
800.000
152
414
3000
725.000
139
33
1500
290.000
55
17,6
900
170.000
30,8
500
72.000
9,5
1,1
27,5
MRBU
760.000
132
84
415
MHU 1700
1.230.000
207
77
617
MHU 900
650.000
110
MH 195
141.000
22,0
6,0
59
MH 165
119.000
19,0
6,0
51
MH 145
105.000
16,5
6,0
46
MH 120
87.000
13,9
6,0
40
MH 96
69.000
11,0
1,9
27
172 88
386
MH 80
58.000
9,3
1,9
24
TABLE 1 Properties of some hammers used offshore Hamme r
Rated Striking Energy
Expected Net Energy (ftlb x 1000)
Type
Blows per Minute
Weight including Offshore Cage, if any (metric tons)
Vulcan 3250
Single-acting steam
60
300
750
1040
673
600
HBM 3000
Hydraulic underwater
50-60
175
1034
1430
542
542
HBM 3000 A
Hydraulic underwater
40-70
190
1100
1520
796
796
HBM 3000 P
Slender hydraulic underwater
40-70
170
1120
1550
800
800
Menck MHU 900
Slender hydraulic underwater
48-65
135
-
-
651
618
Menck MRBS 8000
Single-acting steam
38
280
868
1200
715
629
Vulcan 4250
Single-acting steam
53
337
1000
1380
901
800
HBM 4000
Hydraulic underwater
40-70
222
1700
2350
1157
1157
Vulcan 6300
Single-acting steam
37
380
1800
2490
1697
1440
Menck MRBS 12500
Single-acting steam
38
385
1582
2190
1384
1147
Menck MHU 1700
Slender hydraulic underwater
32-65
235
-
-
1230
1169
IHC S300
Slender hydraulic underwater
40
30
220
300
-
-
IHC S800
Slender hydraulic underwater
40
80
580
800
-
-
(ft-lb x 1000)
KNm
On Anvil
On Pile
IHC S1600
Slender hydraulic underwater
30
160
1160
1600
-
-
IHC S2000
Slender hydraulic underwater
-
260
1449
2000
-
-
IHC S2300
Slender hydraulic underwater
-
-
1566
2300
-
-
TABLE 2 Large pile driving hammers Pile Outer Diameter
Wall Thickness
Hammer Energy
(in.)
(mm)
(in.)
(mm)
(ft-lb)
(kN-m)
24
600
5/8 - 7/8
15-21
50.000 - 120.000
70 - 168
30
750
¾
19
50.000 - 120.000
70 - 168
36
900
7/8 - 1
21-25
50.000 - 180.000
70 - 252
42
1.050
1 - 1¼
25-32
60.000 - 300.000
84 - 120
48
1.200
17- 1¾
28-44
90.000 - 500.000
126 - 700
60
1.500
17 - 1¾
28-44
90.000 - 500.000
126 - 700
72
1.800
1¼ - 2
32-50
120.000 - 700.000
168 - 980
84
2.100
1¼ - 2
32-50
180.000 - 1.000.000
252 - 1.400
96
2.400
1¼ - 2
32-50
180.000 - 1.000.000
252 - 1.400
108
2.700
1½ - 2½
37-62
300.000 - 1.000.000
420 - 1.400
120
3.000
1½ - 2½
37-62
300.000 - 1.000.000
420 - 1.400
Note 1: With the heavier hammers in the range given, the wall thicknesses must be near the upper range of those listed in order to prevent overstress (yielding) in the pile under hard driving. Note 2: With diesel hammers, the effective hammer energy is from one-half to two-thirds the values generally listed by the manufacturers and the above table must be adjusted accordingly. Diesel hammers would normally only be used on 36-in. or less diameter piles. Note 3: Hydraulic hammers have a more sustained blow, and hence the above table can be modified to fit the stress wave pattern. TABLE 3 Typical values of pile sizes, wall thickness and hammer energies
A.7: Tubular Joints in Offshore Structures OBJECTIVE/SCOPE To present methods for the design of large tubular joints typically found on offshore structures. PREREQUISITES Lecture 15A.1: Offshore Structures: General Introduction RELATED LECTURES Lecture 15A.8 : Fabrication Lecture 15A.12: Connections in Offshore Deck Structures SUMMARY The lecture defines the principle terms and ratios used in tubular joint design. It presents the classifications for T, Y, X, N, K and KT joints and discusses the significance of gaps, overlaps, multiplanar joints and the details of joint arrangements. It describes design methods for static and fatigue strength, presenting some detailed information on stress concentration factors. 1. INTRODUCTION The main structure of a topside consists of either an integrated deck or a module support frame and modules. Commonly tubular lattice frames are present, however a significant amount of rolled and built up sections are also used. This lecture refers to the design of tubular joints. These are used extensively offshore, particularly for jacket structures. Connection of I-shape sections or boxed beams whether rolled or built up, are basically similar to those used for onshore structures. Refer to the corresponding lectures for appropriate design guidance. Two main calculations need to be performed in order to adequately design a tubular joint. These are: 1. 2.
Static strength considerations Fatigue behaviour considerations
The question of fatigue behaviour always has to be addressed, even where simple assessment of fatigue behaviour shows this will not be a problem. The joint designer must therefore always be "fatigue minded". 2. DEFINITIONS The following definitions are universally acknowledged [1]: (refer to Figure 1 for clarification):
The CHORD is the main member, receiving the other components. It is necessarily a through member. The other tubulars are welded to it, without piercing through the chord at the intersection. Other tubulars belonging to the joint assembly may be as large as the chord, but they can never be larger. The CAN is the section of the chord reinforced with an increased wall thickness, or stiffeners. The BRACES are the structural members which are welded to the chord. They physically terminate on the chord skin. The STUB is the extremity of the brace, locally reinforced with an increased wall thickness. Different positions have to be identified along the brace - chord intersection line:
CROWN position is located where the brace to chord intersection crosses the plane containing the brace and chord. SADDLE position is located where the brace to chord intersection crosses the plane perpendicular to the plane containing the brace and chord, which also contains the brace axis.
2.1 Geometrical definitions Refer to Figure 1 L is the length of the chord can D is the chord outside diameter T is the chord wall thickness d is the brace outside diameter t is the brace wall thickness (where there are several braces, a subscript identifies the brace) g is the theoretical gap between weld toes e is the eccentricity × Positive when opposite to the brace side, Negative when on the brace side q is the angle between brace and chord axis. 2.2 Geometrical ratios
a=
b=
Can slenderness ratio
Brace to chord diameter ratio (always £ 1)
g=
Chord slenderness ratio
t=
Brace to chord thickness ratio
z=
Relative gap
These are non-dimensional variables for use in parametrical equations. 3. CLASSIFICATION Load paths within a joint are very different, according to the joint geometry. The following classification is used, see Figure 2.
3.1 T and Y Joints These are joints made up of a single brace, perpendicular to the chord (T joint) or inclined to it (Y joints). In a T joint, the axial force acting in the brace is reacted by bending in the chord. In a Y joint, the axial force is reacted by bending and axial force in the chord. 3.2 X Joints X joints include two coaxial braces on either side of the chord. Axial forces are balanced in the braces, which in an ideal X joint have the same diameter and thickness. In fact, other considerations such as brace length, which can be very different on each side of the chord, may lead to two slightly different braces. Angles may be slightly different as well. The important point to note is the balance of forces in the braces. If the axial force in one brace is far higher than the one in the other brace, the joint may be classified as a Y (or a T) joint rather than an X joint. 3.3 N and K Joints These joints include two braces. One of them may be perpendicular to the chord (N joint) or both inclined (K joint). The ideal load pattern of these joints is reached when axial forces are balanced in the braces, i.e. net force into chord member is low. 3.4 KT Joints These joints include three braces. The load pattern for these joints is more complex. Ideally axial forces should be balanced within the braces, i.e. net force into chord member is low. 3.5 Limitations For a joint to be able to be fabricated and to be effective, the geometrical ratios given in Section 2.2 have limitations. Table 3.1 shows these limits and their typical ranges. Parameter
q
Typical range
Limitations min
max
0,4 - 0,8
0,2
1
12 - 20
10
30
0,3 - 0,7
0,2
1 (2)
40° - 90°
30° (3)
90° (1)
(1) Physical limitation (2) Brace shall be less or equal to chord thickness (see punching shear) (3) Angle limitation to get a correct workmanship of welds. Table 3.1 Geometrical Limits and Typical Ranges 3.6 How to classify a joint This classification deals only with braces located in one plane. It must always be remembered that this classification is based on load pattern as well as the geometry. Engineering judgement must therefore be used to classify a joint. For example a geometrical K joint may be classified as.
a K joint when forces are balanced within braces. a Y joint when the force in one brace is reacted predominantly by the chord, rather than by the second brace.
4. GAP AND OVERLAP 4.1 Definitions The GAP is the distance along the chord between the weld toes of the braces (Figure 3).
The theoretical gap is the shortest distance between the outer surfaces of two braces, measured on the line where they cross the chord outer surface. The real gap is the one measured at the corresponding location, between actual weld toes.
A brace OVERLAPS another brace when one brace is welded to the other brace. The overlapping brace is always the thinner brace. The overlapped brace is always completely welded to the chord. 4.2 Limitations The minimum gap allowed is 50mm. This limitation is set to avoid two welds clashing. This is important because the gap is a highly stressed zone. 4.3 Multiplanar Joints The same definitions and limitations apply to multiplanar joints. 5. JOINT ARRANGEMENT As a rule, welds in a joint have to be kept away from zones of high stress concentration. The following practice, see Figure 4, should be followed: 1. 2. 3. 4. 5.
The chord circumferential welds are to be located at either 300mm or a quarter of the chord diameter, whichever is the greater, from the nearest point of a brace-chord connection. The brace circumferential welds are to be located at either 600mm or a brace diameter, whichever is the greatest, from the nearest point of the brace-chord connection. The actual gap shall not be less than 50mm. To achieve this, most designers use a 70 or 75mm theoretical gap. Eccentricity and offset are to be kept within a quarter of the chord diameter. When higher values can not be avoided, secondary moments have to be introduced in the structural analysis by introducing extra nodes. Thickness transitions are smoothed to a 1 in 4 slope, by tapering the thicker wall.
6. STATIC STRENGTH 6.1 Loads taken into account The loads considered in a joint static strength design are the axial force, the in-plane bending moment and the out-of-plane bending moment for each brace. The other components (transverse shear and brace torsion moment) are usually neglected since unlike the preceding loads, these loads do not induce bending in the chord wall. Nevertheless, their presence must never be forgotten and in some specific cases, their effects must be assessed. The axial load, in-plane and out-of-plane bending moments are normally the dimensioning criterion for tubular joints. 6.2 Punching shear 6.2.1 Acting punching shear The acting punching shear is the shear stress developed in the chord by the brace load. The acting punching stress vp is written as: vp = t f sin q where f is the nominal axial, in-plane bending or out-of-plane bending stress in the brace (punching shear for each kept separate), see Figure 5.
6.2.2 Allowable punching shear
Allowable punching shear values in the chord wall are determined from test results carried out on full scale or on reduced scale models. Tests are performed on experimental rigs such as the one shown in Figure 6. They are performed for a single load-case (axial force, in-plane bending, or out-of-plane bending).
The ultimate static strength obtained through these tests can then be expressed in terms of punching shear, as defined above. Statistical treatment of results allow formulae to be defined for the allowable punching shear stress. 6.2.3 The API method Several offshore design regulations are based on the punching shear concept [1,2]. The following method is presented in API RP2A [2]: A. Principle
This method applies to a single brace without overlap, for a non-stiffened joint. When the joint includes several braces, each brace connection is checked independently. Punching shear for each load component (axial force, in-plane bending, and out of plane bending) is calculated and compared to the allowable punching shear stress for the appropriate load and geometry. Interaction formulae are given for combined loading, combining the three punching shear ratio calculated for each component.
B. Allowable punching shear stress The allowable punching shear stress for each load component is:
Vpa = Qq Qf where: Fyc is the yield strength of the chord member Qq is to account for the effects of type of loading and geometry, see Table 6.1. Qf is a factor to account for the nominal longitudinal stress in the chord
Qf = 1 - l g fAX, fIPB, fOPB are the nominal axial, in-plane bending and out of plane bending stresses in the chord Value for l and Qq are given in Table 6.1 Load component
Axial load
In-plane bending
Out of plane bending
Stress in brace
fax
fby
fbz
Acting punching shear
Vpx = t fax sin q
Vp = t fby sin q
Vp = t fbz sin q
0,045
0,021
Qq
K joints
T & Y Joints
w/o diaphragm
Tension
Compression
X
w diaphragm l
0,030
Table 6.1 Values of Qq for allowable punching shear stress from APIRP2A
Qg = 1,8 - 0,1
for g £ 20
Qg = 1,4 - 4 g/D for g > 20
but Qg must be ³ 1,0
Qb =
for b > 0,6
QB = 1,0 for b £ 0,6 C. Loading Combination For combined loadings involving more than one load component, the following equations shall be satisfied:
where: IPB refers to in-plane bending component OPB refers to out-of-plane bending component AX refers to axial force component and
ax
where: arc sin term is in radians. 6.3 Overlapping joints The parametric formulae discussed in Section 6.2 were specifically established for non-overlapping joints with no internal reinforcement. These formulae cannot be used for overlapping joints. In an overlapping joint, part of the load is transferred directly from one brace to the other through the overlapping section, without that part of the load transferring through the chord. The static strength of an overlapping joint is higher than a similar joint without an overlap. API RP2A, [2] allows the static shear strength of the overlapping weld section to be added to the punching shear capacity of the brace-chord connection, see Figure 7.
The allowable axial load component perpendicular to the chord, P^ (in Newtons) should be taken to be: P^ = (vpa T l1) + (2vwa tw l2) where: vpa is the allowable punching shear stress (MPa) for axial stress. l1 is the circumference for that portion of the brace which contacts the chord (mm), see Figure 7. vwa is the allowable shear stress for weld between braces (MPa). tw is the lesser of the weld throat thickness or the thickness t of the inner brace (mm). l2 is the projected chord length (one side) of the overlapping weld, measured perpendicular to the chord (mm), see Figure 7. 6.4 Reinforced joints 6.4.1 Definition Large chord wall thickness may be reduced by stiffening the chord. The most usual reinforcement consists of ring stiffening inside the chord. Some joints may require more complex stiffening. This is the case for large diameter chords which would otherwise require an un-economic chord wall thickness. There are very many different stiffening solutions for a large diameter chord. Therefore there are no parametric formulae available for these designs. Specific analyses must therefore be carried out for an accurate solution. This may involve finite element analysis. 6.4.2 Ring Stiffening Ring stiffening consists of ring plates welded in the chord can prior to welding the braces to it. The punching shear capacity of the chord still may be taken into account when calculating the forces acting on the stiffeners. Ring stiffeners can be justified through parametric formulae available in various publications, the best known being published by Roark [3]. 7. STRESS CONCENTRATION As in any mechanical body presenting discontinuities, stresses are not uniform along the connecting surface of a brace and chord. Figure 8 shows an example of the stress distribution in a joint with local discontinuities at and in the vicinity of the brace chord intersection.
7.1 Stress concentration factor The stress concentration factor (SCF) is defined as the ratio of the highest stress in the connection (or hot spot stress f HS) to the nominal brace stress fNOM: SCF = fHS/fNOM 7.2 Kellog equation This approximate formula can be used for rapidly assessing SCF, for preliminary analyses. fHS/vp = 1,8 √g vp being the punching shear. 7.3 Parametric formulae SCF parametric formulae have been determined based on a large number of finite element analyses and cross-checked with either full scale or model tests. They are based on many man years of work by numerous research teams. A large number of parametric formulae have been published [4]. Sections 7.3.1 to 7.3.3 give, as an example, the most commonly used and acknowledged formulae. In using any set of formulae, care should be taken in classifying the situation and ascertaining any limitations that apply. The only alternatives to these formulae are to perform model tests (full size or at reduced scale) or finite element analyses. No parametric formulae are presently available for stiffened joints. The only ones published to date concern non-stiffened, non overlapping joints. 7.3.1 Kuang equations for T/Y joints [4] Axial load SCFCHORD = 1,981 g0,808 t1,333 exp(-1,2b3 a0,057 sin1,694 q SCFBRACE = 3,751 g0,55 texp(-1,35b 3) a0,12 sin1,94 q Out-of-plane bending
SCFCHORD = 1,024 g1,014 t0,889 b0,787 sin1,557 q
0,3 £ b £ 0,55
SCFCHORD = 0,462 g1,014 t0,889 b(-0,619) sin1,557 q
0,55 £ b £ 0,75
SCFBRACE = 1,522 g0,852 t0,543 b0,801 sin2,033 q
0,3 £ b £ 0,55
SCFBRACE = 0,796 g0,852 t0,543 b(-0,281) sin2,033 q
0,55 £ b £ 0,75
In-plane bending SCFCHORD = 0,702 g0,60 t0,86 b(-0,04) sin0,57 q SCFBRACE = 1,301 g0,23 t0,38 b(-0,38) sin0,21 q 7.3.2 Kuang equations for K joints [4] Balanced axial load SCFCHORD = 1,506 g0,666 t1,104 b(-0,059) (g/D)0,067 sin1,521 q SCFBRACE = 0,92 g0,157 t0,56 b(-0,441) (g/D)0,058 Exp(1,448 sin q ) In-plane bending (bending moment applied to one brace only) SCFCHORD = 1,822 g0,38 t0,94 b0,06 sin0,9 q SCFBRACE = 2,827 t0,35 b-0,35 sin0,5 q 7.3.3 Kuang equations for KT joints [4] Balanced axial load Outer braces only loaded SCFCHORD = 1,83 g0,54 t1,068 b0,12 sin q
0° < q £ 90°
SCFBRACE = 6,06 g0,1 t0,68 b-0,36 {(g1+g2)/D}0,126 sin0,5 q
0° < q ³ 45°
SCFBRACE = 13,8 g0,1 t0,68 b-0,36 {(g1+g2)/D}0,126 sin2,88 q
45° £ q ³ 90°
SCFBRACE = 4,89 g0,123 t0,672 b-0,396 {(g1+g2)/D}0,159 sin2,267 q In-plane bending - as for K joint Validity range The above equation for T/Y, K and KT joints are generally valid for joint parameters within the following limits: 8,333 £ g £ 33,3 0,20 £ t £ 0,8 0,3 £ b £ 0,8 unless stated otherwise 6,667 £ a £ 40 unless stated otherwise
0° £ s £ 90° unless stated otherwise. 8. FATIGUE ANALYSIS A fatigue analysis of a joint consists of the following steps: 1. 2. 3.
Calculation of nominal stress ranges in the brace and the chords Calculation of hot-spot stress range Calculation of joint fatigue lives using S-N curves for tubular members at joints.
8.1 Nominal stress range Nominal stress ranges in braces and chords are calculated by a global stress analyses. 8.1.1 Wave histogram A wave histogram has to be obtained for each direction around the platform. A simple form of a wave histogram is as follows: Wave height (metres)
Average number per year
0-1,5
3 100 000
1,5-3
410 000
3-4,5
730 000
4,5-6
5 000
6-8
800
8-10
20
8.2.2 Nominal stress ranges Nominal stress ranges can be calculated by following the steps below: 1. 2. 3.
Wave heights are grouped in "blocks", for which just one stress range will be calculated. Different wave directions need to be considered with a minimum of three "blocks" per wave direction. For each block one representative wave is chosen, whose action is supposed to represent the action of the whole block. The highest wave of the block is normally chosen. Nominal stresses for each joint component are then calculated for different phase angles of the chosen wave, for one complete cycle (360°). The nominal stress range for the joint component is defined as the difference between the highest and the lowest stress obtained for a full wave cycle. Four to twelve phase angles per wave are usually considered.
8.2 Hot spot stress ranges Hot spot stress ranges are then evaluated for each chosen joint location by applying parametric formulae [4] (or by applying the SCF calculated from a detailed analysis). When using parametric formulae, stress components (axial, in plane bending and out of plane bending) have to be distinct throughout the calculations, as the SCF formulae apply individually for each load component.
Where a chord and brace intersect, four to eight locations are usually chosen around the intersection line. For each of these locations the stress response for each sea state should be computed, giving adequate consideration to both global and local stress effects. 8.3 S-N Curves S-N curves to be used for offshore structures are given by statutory regulations [1,2]. APIRP2A uses the curves shown in Figure 9.
The X and X1 curves should be used with hot spot stress ranges based on suitable stress concentration factors. The permissible number of cycles is obtained from the S-N curve by taking the hot spot stress range, and entering the graph. It should be noted that Curve X presumes welds which merge smoothly with the adjoining base metal. For weld without such profile control, the X¢ curve is applicable. 8.4 Cumulative Fatigue Damage Ratio
The stress responses should be combined into the long term stress distribution, which should then be used to calculate the cumulative fatigue damage ratio, D, given by:
D= where n is the number of cycles applied at a given stress range N is the number of cycles to cause failure for the given stress range (obtained from appropriate S-N curve). In general the design fatigue life of each joint and member should be at least twice the intended service life of the structure, i.e. a safety factor of 2,0. For critical elements whose sole failure would be catastrophic, use of a larger safety factor should be considered. 9. CONCLUDING SUMMARY
Terminology, geometric ratios and joint classifications are now standardised for tubular joints. The presence of gaps and overlaps significantly influence joint behaviour. Determination of static strength is generally based on the concept of punching shear, with the allowance of overlapping joints. Special analysis are required for reinforced joints. Stress concentration factors (SCF) are defined for most commonly occurring joints. Determination of fatigue strength is based on nominal stress range multiplied by appropriate SCF.
10. REFERENCES [1] Offshore Installations: Guidance on Design, Construction and Certification. Fourth Edition, HMSO, 1990. [2] Recommended Practice for Planning, Designing and Constructing Fixed Offshore Platforms, API RP2A Nineteenth Edition. [3] Young, Warren C, Roark's Formulae for Stress and Strain. Sixth Edition, McGraw-Hill. [4] Stress Concentration Factors for Simple Tubular Joints, 1989, Volumes 1 to 5, Lloyds Register of Shipping-Offshore Division.
A.8: Fabrication OBJECTIVE To describe the general methods of jacket fabrication. To discuss the various stages of operation from material selection, through erection, including construction practices and equipment. To indicate the calculations normally involved. PREREQUISITES Lecture 15A.1: Offshore Structures: General Introduction Lectures 3.1: General Fabrication of Steel Structures Lectures 3.2: Erection Lectures 3.3: Principles of Welding Lectures 3.4: Welding Processes RELATED LECTURES Lectures 15: Structural Systems: Offshore SUMMARY The construction philosophy and definition of the construction phases of the fabrication of offshore structures are described. The overall execution plan and the contractor's organisation for its implementation are introduced and constructability i.e. the more general aspects of design - the size and transportability of components, welding access considerations, construction tolerance, is discussed. The fabrication of nodes and reinforced tubulars, including the fabrication procedure for a typical node is described together with jacket assembly and erection and the procedures for "big lift". 1. INTRODUCTION 1.1 Construction Phases Jacket construction involves the following work phases: Procurement The technical and commercial activities required to supply material and specialised products to enable the execution of construction activities. Fabrication The processes normally carried out in a fabrication shop to produce relatively small units. Thus fabrication includes processes such as cutting, rolling, pressing, fitting, welding, stress relieving on such items as welded tubulars, beams, nodes, girders, cones, supports, clamps, etc. Assembly
The processes normally performed outside the fabrication shop but at ground level in order to assemble groups of shop fabricated items into an (assembled) unit for subsequent erection in accordance with a construction sequence. Erection The processes required to install assembled and shop fabricated items together in their final configuration. These processes include fitting and welding. However the emphasis is on the transportation and lifting of heavy assemblies. 1.2 Construction Philosophy The design of a jacket, i.e. a lifted, launched or self-floating jacket, is determined primarily by the offshore installation equipment available and the intended water depth. In general the preference is to lift the jacket in place. The size of such jackets has being increasing as offshore lifting capacity has grown. With modern lifting capacity now up to 14,000 tonnes, jackets approaching this order of magnitude are now candidates for lifting into position. For jackets destined for shallow water, where the height is of the same order as the plan dimensions, erection is usually carried out vertically, i.e. in the same attitude as the final installation. Such jackets may be lifted or skidded onto the barge. Jackets destined for deeper water are usually erected on their side. Such jackets are loaded by skidding out onto a barge. Historically most large jackets have been barge launched. This method of construction usually involves additional flotation tanks and extensive pipework and valving to enable the legs to be flooded for ballasting the jacket into the vertical position on site. This method of construction is currently applicable for jackets up to 25,000 tonnes. Very large jackets, in excess of this, have been constructed as self-floaters in a graving dock and towed offshore subsequent to flooding the dock. In considering the construction philosophy and contract strategy, the objectives of achieving quality requirements and efficiency are of fundamental importance. An offshore jacket goes through a series of very distinct stages as it moves from fabrication to load-out. These stages range from operations which are almost totally automatic under very controlled conditions, e.g. steel production, automatic welding, to operations which are almost totally manual in very variable conditions, e.g. yard erection, offshore activities. Thus decreasing efficiency occurs as progress through these operations advances. In addition, the stable conditions in repetitive processes of the early operations are more conducive to the maintenance of high quality. A third basic consideration is that risk increases with each progressive stage. These general trends during construction are shown in Table 1. It is clear therefore that, as a general principle, as much work as possible should be undertaken in the earlier more productive, higher quality, less risky phases of the project. Some of the principles which reduce the time and cost of construction are:
Subdivision into as large components and modules as it is possible to fabricate and assemble. Concurrent fabrication of major components in the most favourable location and under the most favourable conditions applicable to each component. Planning the flow of components to their assembly site. Providing adequate facilities and equipment for assembly, including such items as synchrolifts, and heavy-lift cranes. Simplification of configurations and standardisation of details, grades and sizes. Avoidance of excessively tight tolerances. Selection of structural systems that utilise skills and trades on a relatively continuous and uniform basis. Avoidance of procedures that are overly sensitive to weather conditions; ensuring that processes which are weather sensitive are completed during shop fabrication, e.g. protective coating.
Quality management is a vital and integral component of all aspects of offshore fabrication. Essentially it involves ensuring that what is produced is what is needed. The requirements for documentation, hold point, audits, reviews and corrective actions are part of the quality assurance process. They are crucial tools for controlling the project execution and providing verifiable evidence of the fabricator's competence.
Quality control, inspection and testing should be performed during all phases of construction to ensure that specified requirements are being met. The most effective quality scheme is one which prevents the introduction of defective materials and workmanship into a structure, rather than finding problems after they occur. A general note on Quality Assurance for Offshore Construction is included in Appendix 1. It is applicable to this lecture and also to Lecture 15A.9: Installation. 2. ENGINEERING OF EXECUTION Engineering of execution, 'construction engineering', entails the work required during each phase of execution to ensure that the design requirements are fulfilled. A general method of execution is envisaged at the jacket design stage. Since the shape of the jacket, its form and properties require quite specific methods of load-out, offshore transportation and installation (which are construction activities executed under contractor responsibility), there is considerable interfacing of engineering requirements in these phases. In the earlier phases, i.e. procurement through assembly and erection, the contractor, while being limited by design specification requirements, has freedom of choice with regard to the exact method of execution adopted. However, in all phases the contractor is required to demonstrate that the methods which he adopts are compatible with the specification requirements and do not affect the integrity of the structure. Each phase of execution has its own specific engineering requirements which are determined by the processes executed during that phase. These processes range from those which are largely repetitive early in execution to one-off activities in the latter phases. Accordingly the engineering which supports procurement and shop fabrication is voluminous but repetitive, e.g. material take-offs, shop drawings, cutting plans, etc. The assembly and erection phases are supported by a mix of repetitive engineering, e.g. scaffolding, and specific studies for limited series of activities. The volume of contractor construction engineering on a large jacket is typically 130,000/150,000 hours. The typical organisation of a contractor's technical documents is shown in Table 2. When designing larger components consideration must be given to their subdivision into elements which will not distort when fabricated and which can be relatively easily assembled without welding/dimensional problems. For instance, nodes are categorised as either complex or simple from the execution viewpoint based on the number of separate fitting-weldingNDT (non-destructive testing) cycles required during fabrication and the possibility of automatic welding between the node can and the tubular during sub-assembly. The number of fitting-welding-NDT cycles depends on the existence of ring stiffeners and the number and disposition of stubs. For reasons relating to weld distortion and to allow automatic welding, it is almost essential that ring stiffeners be installed prior to fitting/welding of stubs. This adds an extra cycle to the fabrication of the node. Thus ring stiffeners are best avoided. Where this is not possible, care should be taken to define them at an early stage on critical nodes. Node stubs can be classified as simple or overlapping. Overlapping stubs add at least one complete cycle to node fabrication and should therefore be avoided where possible. The minimum separation between the weld toes of adjacent simple stubs is typically specified as 50mm, see API RP2A, Fig. 4.3.1-2 [1]. However this distance is too small to allow simultaneous welding of adjacent stubs - 150mm is a more practical distance. 3. FABRICATION 3.1 Fabrication Processes The specifications for fabrication of offshore jackets are determined by the designer. They are usually based on one or more of the well known codes, with additional requirements dictated by the specific design, client standards, statutory rules, etc. Two recognised codes which are used extensively for establishing general requirements are the API RP2A Recommended Practice for Planning, Designing and Constructing Fixed Offshore Platforms, [1] and AISC Specification for the Design, Fabrication and Erection of Structural Steel for Buildings [2]. For larger jackets, the nodes tend to be fabricated separately under highly controlled shop conditions. Alternatively cast steel nodes may be used in order to eliminate critical welding details. Recent experience, both in the laboratory and as a result of in-service inspection, has prompted increasingly greater attention to the welding aspects of fabrication. In particular greater attention has been focused on the importance of
complete joint penetration groove welds, elimination of "notch effects" at the root and especially the cap of node welds, and achieving the required weld profile. Welds which are critical for fatigue endurance may be required to be ground to a smooth curve. This process reduces the probability of brittle failure. However it also implies increasingly sophisticated and stringent fabrication and quality assurance/quality control (QA/QC) requirements. Typical welding details from API RP2A [1], showing tubular members framing into or overlapping another member with access from one side only, are shown in Figure 1. However, a lot of emphasis is placed on designing stubs which can be welded from both sides. For instance, in the weld details for the Bouri jacket, Figure 2, most stubs are accessible from both sides.
Welding procedures are required, detailing steel grades, joint design, welding consumables, etc. Welds are typically subject to 100% visual, magnetic particle inspection (MPI) and ultrasonic test (UT) inspection. The weld acceptance criteria, e.g. maximum weld undercut length (t/2 or 10mm), and maximum depth (t/20 or 0,25mm), imply an exceptionally high quality of welding. In addition all welders should be qualified for the type of work assigned to them and certified accordingly.
The location and orientation of circumferential and longitudinal welds during construction is based on minimising interferences and ensuring the minimum distance between circumferential welds. Special attention is required on items such as pile sleeve shear plates, launch runners, mudmats, etc. where planned avoidance of weld interference is critical. All temporary plates and fittings should be subjected to the same requirements for weld testing as the member to which they are being affixed. There is also an overriding necessity to ensure that such attachments are located at a safe distance from main structural welds in order to minimise the risk of defect propagation. This is not unduly conservative - the "Alexander Kielland" capsized due to a fatigue crack initiated at the attachment of a sonar device to a principal structural member. Temporary cut-outs should be of sufficient size to allow sound replacement. Corners should be rounded to minimise stress concentrations. Where welds are found to be defective, they should be rectified by grinding, machining or welding as required. Welds of insufficient strength, ductility or notch toughness should be completely removed prior to repair. In general, sub-assemblies are executed so that at least one of the two edges which will mate during subsequent assembly/erection has a cut-off allowance. This procedure provides flexibility in that the sub-assemblies can be sent to the field with the cut-off allowance in place and cut to fit on location. Alternatively they can be cut to exact dimensions during sub-assembly where the as-built dimension has already been determined. 3.2 Node Fabrication The primary structure nodes are frequently geometrically complex. Accordingly their fabrication presents particular problems, especially from the points of view of welding and dimensional control. On a complex jacket the designer may specify the node cans, or the whole node including stubs and ring stiffeners, in material with specified through-thickness properties. This requirement is introduced because of tearing or punching effects likely to be sustained by these elements during their design life and indeed during fabrication. The designer may also "thicken" or reinforce the cans to withstand local stresses. Finally, in an effort to ensure that node welds contain minimal levels of residual stress due to fabrication, thermal stress relieving or post-weld heat treatment (PWHT) of the heavier more restrained welds may be prescribed. This is frequently a requirement for thicker walled North Sea jackets. API RP2A [1] provides specific tolerances for final fabrication. The contractor must work within tolerances which preserve dimensional compatibility and observe weight control requirements at each phase of construction. Bearing these requirements in mind, node fabrication tolerances are tight, e.g. typical working points within 6mm of theoretical, stub angle within 1 minute, all braces within 12mm of the design dimension. The typical fabrication process for a conventional node, assuming that the can (with or without ring stiffeners) has already been fabricated, commences with profiling of stubs and terminates with UT inspection of the finished node after PWHT. The intermediate stages can be performed in several different ways, some of which depend on the specific geometry of the node and many of which depend on fabricator preference. Some fabricators prefer to orient the can upright, maintaining that it enables more stubs to be fitted simultaneously. However the majority of fabricators tend to fit the stubs to a can placed on horizontal rollers. The sequential steps in the fabrication of a typical node are as follows:
Trace generators, working points, etc. onto the can. Cut and profile the stubs. Touch up bevels and trace generators onto the stubs. Trace node locations onto the can surface and grind or blast areas. UT the cleaned areas to ensure that the steel is free from laminations. Particular care is required where shrinkage strains in the through-thickness direction may lead to lamellar tearing in highly restrained joints. Assemble one or two adjacent stubs in the same plane on the can. Tack-weld in position. Verify dimensional control and weld preparations around stub. Weld according to predetermined sequence to limit deformation. The welding processes used are usually shielded metal arc welding (SMAW) or flux cored arc welding (FCAW), see Lecture 3.4 Welding Processes. If the weld is double-sided, after 3 or 4 passes, back-grind and clean weld roots from opposite side. Perform MPI test on ground roots. Complete weld body. Deposit weld bead for cap profiling. Toe-grind profiles if required. Grind weld beads at base metal to remove undercut. Allow welds to cool. Visually inspect finished welds. MPI and UT finished welds. Repeat previous steps for successive stubs.
When all stubs have been fitted and welded out perform post weld heat treatment (PWHT) as required, blast or grind welds and perform NDT re-test on all welds. Cut any required off-cuts on cans or stubs. Perform final dimensional control of node.
3.3 Jacket Sub-assemblies Sub-assembly can be considered as an intermediate stage between standard shop fabrication, i.e. nodes, tubulars, beams etc. and assembly or erection. The emphasis is on performing the maximum number of welds in the shop. This ensures the highest weld quality since many node and tubular welds can be double-sided and/or automatic when performed in the fabrication shop. When defining sub-assemblies, the principal factors to be borne in mind are the following:
Size/Weight/Dimensions: these are largely governed by considerations of transportability. Welding Sequence: sub-assemblies should not imply a difficult welding sequence causing distortion or induced stresses during sub-assembly welding or the subsequent assembly or erection. Constructability: certain sub-assemblies may have specific construction difficulties associated with them, e.g. short, large diameter infillings are difficult to erect vertically and are best included in sub-assemblies, if possible.
3.4 Dimensional Control Of all the areas of quality control (QC) which require attention, that of dimensional control, as emphasised in the code and specifications, tends to be exaggerated. However, it is clear that attention must be paid to the dimensions which have structural significance, e.g. the straightness of elements, ovality of tubulars, eccentricities at node joints, etc. It is also clear that on a jacket the global alignment/verticality of items such as pile sleeves, conductor guides, launch runners, etc. are also important. Finally dimensional control of items which are intended for "mating" or "removal" offshore, for example piles/pile sleeves, jacket top/MSF base, buoyancy tank/supports, etc. is vital to the efficiency of offshore installation. There are therefore, many aspects where the attention to dimensional control is justified even if the overall design might occasionally benefit if the designer did not always require that everything fitted so tightly. The principal reason for requiring such accurate dimensional control of nodes and tubulars during fabrication is not because of the structural consequences of out-of-tolerance but rather because the parts may not fit together in the yard. It is one of the most vexing incongruities of the tubular steel jacket concept that the theoretical tolerances on node stub eccentricity are generous from the structural viewpoint while the actual tolerances are very tight because of considerations regarding the fitting together of components during subsequent phases of construction. The dimensional control of node fabrication, in particular, involves potentially intricate calculations in the shop. However, the most successful systems simply involve the inclusion on the shop drawings of several additional "checking" measurements and the correct marking of the node can and stub generators and offsets. 4. JACKET ASSEMBLY AND ERECTION 4.1 Jacket Assembly Shop fabricated sub-assemblies and loose items are incorporated into assemblies which constitute the major lifts of the erection sequence. Thus for a large jacket, the assemblies are typically of four types:
Jacket levels incorporating conductor guide frames Top frames Jacket rows i.e. bents or partial bents Pile sleeve clusters.
The assembly and erection phases are based on the following objectives:
Maximise on-the-ground assembly (as opposed to erection) and maximise access around the jacket during execution.
Minimise erection joints in principal structural elements, such as jacket legs, launch runners, rows, levels. Align critical areas such as conductor guides, pile sleeves, launch runners. Sub-assemble principal structural elements of jacket such as jacket legs, rows, levels. Sub-assemble, and where possible pre-test, systems such as grouting, ballasting. Include maximum quantity of secondary items such as anodes, risers, J-tubes, caissons. Coat or paint required areas (top of jacket, risers) prior to erection. Minimise the use of temporary items which require subsequent removal, such as scaffolding, walkways, lifting aids, etc. and pre-install such aids where they are necessary.
The assembly of a jacket frame, often having a spread at the base of 50m or more, places severe demands on field layout and survey and on temporary support and adjustment bracing. Such large dimensions mean that the thermal changes can be significant. Temperature differences may be as great as 30°C between dawn and afternoon and as much as 15°C between various parts of the structure, resulting in several centimetres distortion. However, the practice of 'using the sun' to fit elements which are not dimensionally in-tolerance is common in the field. This procedure in itself tends to induce residual stresses in the structure. Because of the difficulty associated with thermal distortion, it is normal to "correct" all measurements to a standard temperature, e.g. 20° C. Elastic deflections are also a source of difficulty in maintaining tolerances in the location of nodes. Foundation displacements under the skid beams and temporary erection skids must be carefully calculated and monitored. The overall assembly sequence and programme requires that each assembly be completed prior to lifting. It is normal to determine the exact location, orientation and attitude, i.e. face-up or face-down, of each assembly in the field in anticipation of its lifting procedure. Assembly layout drawings are usually prepared showing central co-ordinates for each assembly. The central co-ordinates are then used as local bench marks with the object of defining the assembly, the sub-assemblies, loose items, appurtenances and temporary attachments which comprise, field welds, overall dimensions, weight, reference drawings, etc. Dimensional control of the assembly both prior to and after welding, can be by means of a series of self-checking measurements on the structure itself. Provided cross checks are adequate, the time consuming exercise of referring measurements to an external bench mark can be avoided. Normally the assembly is tacked in position to theoretical dimensions using allowable positive tolerances to compensate for weld shrinkage. Perhaps the most fundamental rule in fitting is the avoidance of "force-fitting" of members prior to welding or to force stresses into unwelded members through the welding sequence since such conditions cannot have been foreseen by the designer. An outline sequence of events which apply to all types of assembly is as follows:
Preparation of assembly support and staging Rough setting of assembly main structure and position tacking. Dimensional control of assembly main structure. Infilling of secondary structure and position tacking. Dimensional control of assembly and secondary structure. Preweld inspection. Weldout of structure subject to continuous inspection and according to approved sequence. Installation of appurtenances (e.g. anodes, supports, walkways, risers, J-tubes, caissons, grouting and ballasting) and scaffolding, lifting, aids, erection guides, temporary attachments. Test (e.g. hydrotest) if required. Overall NDT, dimensional control. Blasting and painting or touch up. Removal of temporary assembly supports and staging. Preparation for transport/lift/erection.
4.4 Jacket Erection In this phase assembled, sub-assembled and fabricated structures, together with loose items, are incorporated into the final structure according to the sequence outlined in Figure 3a - 3e.
Jacket frames are typically laid out flat and then rolled using multiple crawler cranes. Co-ordinating such a rigging and lifting operation requires thoroughly developed three-dimensional layouts, firm and level foundations for the cranes and experienced, well rehearsed operators. Twenty four cranes were involved in the two major side frame lifts during the erection of platform Cerveza, which was 300m long. For the Magnus platform and Bouri DP3, a different procedure known as "toast rack" was used. Here the jacket horizontal levels were fabricated, erected in place and tied in to complete the jacket. For the Bullwinkle jacket, one of the world's largest, sections of the jacket were fabricated in Japan, transported by barge to Texas and assembled using jacking towers which rolled up the sections to heights as great as 140m. For jackets destined for shallow water erection is usually carried out vertically, i.e. in the same attitude as the final installation. Such jackets may be lifted onto the barge or skidded out. In this latter case, adequate temporary pads and braces must be provided under the columns to distribute the loads for skidding. The structural analysis associated with the erection procedure for a given assembly usually involves a computer model with all relevant structural characteristics. The assembly is analysed for a number of load cases which correspond (approximately) to the support conditions of the assembly at its presumed critical attitudes, i.e. the locations of the cranes, bogies, saddles, etc. when the panel is being transported and when it is in horizontal and vertical attitudes. The structural analysis for lift/transport identifies the worst cases from the point of view of structural response. These cases are then analysed to determine the maximum stresses and displacements. The calculations should show that global and local stresses are within allowable limits according to API/AISC codes. Frequently, a structural analysis computer programme is used for this purpose. The analysis will indicate where bending stresses are high and/or crane, bogie or support loads inadmissible. Thus modifications can be made to redistribute
structural stresses and loads at "supports" to optimise both and ensure that neither the cranes nor the structure can be overloaded during erection. An outline sequence for the erection of all major components would be:
Technical appraisal of lift methods. Calculations for crane configuration, rigging accessories, etc. Preparation of cranes for lift. Preparation for rigging. Transport assembly to lift location. Roll-up into position with scaffolding and staging in position, if possible. Preparation of fixing system and wind bracing (usually done by means of guy wires and turnbuckles). Weldout at least sufficient to allow crane release. Crane release. Removal of rigging and temporary attachments.
Jacket structural completion is followed by a short phase during which all the jacket systems, both permanent and those required during installation, are completed and rendered functional. The load-out operations are covered in Lecture 15A.9: Installation. 5. CONCLUDING SUMMARY
The design of a jacket is determined primarily by the offshore installation equipment available and the intended water depth. In general, the preference is to lift the jacket in place. Jackets destined for deeper water are usually erected on their side. As a general principle, as much of the execution as possible should be undertaken in the early phases of fabrication. Each phase of execution has its own engineering requirements which are determined by the processes executed during that phase. The specifications for fabrication of offshore jackets are determined by the designer and are usually based on one or more of the well known codes. Shop fabricated sub-assemblies and hose items are incorporated into assemblies which constitute the major lifts of the erection sequence. Assembled, sub-assembled and fabricated structures, together with loose items, are incorporated into the final structure in a sequence which takes account of structural analyses of bending stresses, and crane, bogie and support loads.
6. REFERENCES [1] API RP2A, Recommended Practice for Planning, Designing and Construction of Fixed Offshore Installations, latest edition. Engineering design principles and practices that have evolved during the development of offshore oil resources. [2] AISC Specification for the Design, Fabrication and Erection of Structural Steel for Buildings, latest edition. API code refers to this specification for calculations of basic allowable stresses of all jacket members. 7. ADDITIONAL READING [1] Det Norkse Veritas Marine Operations Recommended Practice RP5 - Lifting (June 1985). Principles and good practice for offshore heavy lifts. [2] AWS Structural Welding Code AWS D1.1-88. All jacket welding and weld procedure qualifications are required by the API code to be undertaken in accordance with this code.
[3] Det Norske Veritas, Rules for the Design, Construction and Inspection of Offshore Structures, 1977. Rules for construction and installation of steel jackets as required by DNV. [4] Lloyd's Register of Shipping, Rules and Regulations for the Classification of Fixed Offshore Installations, 1989. Based on Lloyd's experience from certification of over 500 platforms world-wide. APPENDIX 1 Quality Assurance and Quality Control It is becoming increasingly common for operators to specify that the quality of construction for offshore structures be controlled by a recognised quality system management standard. ISO 9000/EN 29000, Standard for Quality Systems Management, is recognised as the accepted standard in such situations. These standards set down the requirements that a soundly based quality management system must fulfil if it is to assist in properly defining and controlling product quality. Because the standards deal with the quality system, and are not product standards, they are applicable to many sectors of industry including offshore construction. They apply in any situation where management wish to adopt a clearly defined policy and an orderly approach to providing a quality product. All aspects of a company's activities are covered in the standards including:-
Design
Product Traceability
Contract Review
Process Control
Documentation Control
Inspection/Testing
Management Responsibility
Calibration of Equipment
Purchasing
Control of Non-Conformances
Corrective Action
Handling/Storage/Delivery
Quality Records
Training
Management Review/Audits
Etc.
QA Management Complexity The overall programme for a jacket construction, shows that there are a very considerable number of offshore activities in many different locations within a very short period of time. The evaluation of the performance of such a range of activities and at a number of centres is a major QA/QC undertaking. It is difficult to fully appreciate the scope of documentation on a jacket construction project. Consider the documentation which is expected to flow from one location to another in respect to a single node. From the time the plate is manufactured until it is located in the final structure, a dossier must be compiled. This documentation could commence with copies of certificates from the steel plate manufacturer and progress through several welding, NDT, dimensional control phases at a number of successive locations, culminating in the issue of a Release Note at the node fabrication shop.
Clearly this is necessary on some items, e.g. steel, welds, NDT Certificates for the jacket primary structure, risers, etc. These documents may be useful during maintenance of the platform enabling many in-service problems to be traced to abnormalities which occurred during construction. Construction of a large jacket typically involves thousands of steel plates. Each plate inevitably becomes an individual as it is allocated a unique number corresponding to a Material Utilisation Schedule or Cutting Plan. The individual number of pieces of plate could be in excess of 20,000 items. The primary object of material control is to ensure that, at any stage of construction, the origin of each and every item can be traced back to a material certificate which in turn corresponds to a set of test/chemical composition etc. as contained in the Data Dossier. However, voluminous as this documentation may be, it constitutes less than half of the total documentation produced for a complete jacket. Consider for instance the number of welds in a complex buoyancy tank, the walkways on top of the jacket, the anodes, launch runners, grout lines, etc. Each of these must be welded, several must be individually inspected. However the requirement to produce sophisticated documentation in respect of each is questionable. For this reason it is important that agreement is reached at an early stage as to the individual items which require identification, that these be kept to a minimum and that the identification system be simple. In actual practice it has proven to be very difficult to make all materials really traceable. Much more could be done to structure such documentation in such a way that it would really be of help throughout the platform life. Procedures and Specifications Within the Contractor's organisation QA/QC procedures must be developed for the project, many of which will be specifically for jacket construction. These are divided into Management Procedures (e.g. Management of Non Conformities, Management of Jacket Completion Onshore, etc.) and Control Procedures (e.g. Procedure for Ultrasonic Testing of welds at Jacket Yard, Dimensional Control Procedure for Node Fabrication at factory etc.). Construction Procedures/Specifications are also required (e.g. Jacket Assembly and Erection Procedure, Pile Installation Procedure, etc.) in addition to a vast number of weld procedure specifications and qualifications, welder qualifications and inspection plans. Even if the number of specific procedures required from each subcontractor is minimised, a fabrication subcontractor will still be required to develop procedure and specifications for the following typical functions/activities: subcontract organisation, material control, fabrication method/sequence, procedures for cutting, forming, pre-heating, post-weld heat treatment along with the more obvious welding and NDT procedures and Inspection Plans. Typically hundreds of procedures/specifications must be developed by jacket subcontractors. Certification On most offshore projects, the underwriters normally agree to insure the plant during its operating life provided it is designed, constructed and maintained to predetermined standards and certified as such. This certification is also almost invariably required by the state authorities in whose waters the plant is installed. It is normally performed by one of the traditional ship classification societies known as the Certifying Authority (CA). In the widest sense, certification requires that the CA carry out independent surveillance to ensure that the standards chosen for the project are satisfactory and that the project is performed in accordance with them. Formerly this meant that the CA inspected every activity likely to influence the adequacy of the final product - an enormous task. More recently with the advent of QA, the certification function can mean audits of the construction so that, rather than inspect everything, the CA satisfies himself that the manner in which the construction is being managed and performed (based on incomplete but comprehensive inspection) is likely to lead to a satisfactory product. Phase
Work Centre
Efficiency
Quality Variability
Risk
Engineering
Office
Decreasing
Increasing
Increasing
Procurement
Factory
Decreasing
Increasing
Increasing
Fabrication
Fabrication Shop
Decreasing
Increasing
Increasing
Assembly and Erection
Yard Site
Decreasing
Increasing
Increasing
Loadout and
Transition
Decreasing
Increasing
Increasing
Seafastening Transport and Installation
Offshore Site
Decreasing
Increasing
Increasing
Table 1 Jacket Construction Phases and Characteristics No.
Document Series
Group or Individual Subject Title
1
Shop Drawings, Cutting Plan
Welding standards, nodes, tubulars, piles, pile sleeves, clusters, conductor guide frames, launch runners, buoyancy tanks, cathodic protection system, protective coating systems, risers, j-tubes, caissons, boat landings, boat bumpers, walkways, grouting systems, ballasting system, installation aids, as-built drawings.
2
Method and Temporary Works Drawings
Subassemblies, assemblies, supports, access, scaffolding, lifting and transportation onshore, test and commission, identification. Onshore construction accessories. Offshore installation (preparation, transportation, lifting, launching, anchor patterns etc.). Offshore installation accessories (tools, guides, access, handling etc.).
3
Quality Assurance Procedures
Documentation identification, distribution and approval, witness and hold points, technical modifications and non-conformance management, material control, material identification and traceability, procurement and subcontracting, weld parameter control, management of specific problem areas.
4
Quality Control Procedures
NDT methods (visual, UT, x-ray, dye penetrant, MPI), dimensional control, destructive testing methods, NDT operator training and qualification, calibration of inspection equipment, pressure testing, miscellaneous testing.
5
Manuals
Testing, commissioning and preparation of jacket for tow. Load-out manuals - jacket piles, topsides. Installation manuals - jacket, piles, topsides.
6
Weld Procedures
For each location - weld procedures - repair procedures.
7
Design Reports, Reviews and Specifications
Quay design, skidway design, mooring system design, soil improvement spec., skidding system spec., dredging spec., transportation of jacket and piles, buoyancy tanks, jacket launching and emplacement, on-bottom stability, pile driveability, jacket levelling study.
8
Engineering Meetings
Normally held at critical phases of construction at the various construction locations.
9
Fabrication, Assembly and Erection
Fabrication/welding sequence (for principal items), forming, bending, stress relieving, coating, assembly and erection, temporary and secondary attachments, lifting and transporting, jackdown, weight control, settlement control, jacket weighing.
10
Inspection plan
Steel supply (at each supplier). Fabrication of typical jacket and pile components (at relevant centres). Assembly and erection.
11
Technical Proposals and NonConformance Resolutions
Technical Clarification Requests ) Technical Relaxation Requests ) Possible at every Major Non-Conformance Reports ) Phase of location Minor Non-Conformance Reports ) of the Project.
Table 2 Jacket Construction Engineering: Typical Organisation of Contractor's Technical Documents
A.9: Installation OBJECTIVE To describe the general methods of jacket installation. To discuss the various stages of operation from loadout through offshore positioning and installation, including construction practices and equipment. To indicate the calculations normally involved. PREREQUISITES Lectures 15A.1: Offshore Structures : General Introduction RELATED LECTURES Lectures 3.1: General Fabrication of Steel Structures Lectures 3.2: Erection Lecture 3.3: Principles of Welding Lecture 3.4: Welding Processes Lectures 15A: Structural Systems: Offshore SUMMARY The phases of installation of a steel jacket - loadout, seafastening, offshore transportation and installation - are described and the associated analyses are indicated. 1. INTRODUCTION 1.1 Project Phases A steel jacket installation usually consists of the following project phases: Loadout - Comprises the movement of the completed structure onto the barge which will transport it offshore. Seafastening - Comprises fitting and welding sufficient structure between the structure and the barge to prevent the jacket shifting during transit to the offshore site. Offshore Transportation - Comprises the tow to the location offshore and arrival of the barge at the offshore site with the seafastened structure. Installation - Comprises the series of activities required to place the structure in the final offshore location. These activities include jacket lift and upending, positioning, pile installation, jacket levelling and grouting, together with support services for these activities. 1.2 Construction Philosophy In deciding how best to fabricate (i.e. vertical or on its side) and install (i.e. lifted, launched or self-floating) a given jacket, the options are principally determined by the installation equipment available and the jacket's intended water depth. In general, the preference is to lift the jacket into location. The motivation for this installation method, rather than the more traditional barge-launching, is a reluctance to spend money on jacket steelwork which will only be used during the temporary installation phase. The size of such lifted jackets has been increasing as offshore lifting capacity has grown. With modern
lifting capacity now up to 14000 tonnes (see Table 1), jackets approaching this order of magnitude are now candidates for lifting into position. Figure 1 shows how the 6000 tonne jacket for the Kittiwake field in the North Sea was lifted from the barge into the water and up-ended in a continuous operation, ending with the jacket on the seabed ready for piling. The advantage of this approach is that the jacket, being lowered into the water, does not require the launch frames necessary for launching from a barge. Also, since the weight of the jacket is taken by the cranes throughout, there is no need for special buoyancy tanks and deballasting systems.
Jackets destined for deeper water are heavier and are usually erected on their side and launched from a barge (Figure 2). This method of construction is currently applicable for jackets up to 25000 tonnes. A launched jacket usually requires additional buoyancy tanks with extensive pipework and valving to enable the legs and tanks to be flooded in order to ballast the jacket into the vertical position on site. For instance, in the case of the Brae 'B' jacket (a large 19000 tonne jacket installed in 100m water depth in the North Sea) it was necessary to provide 11000 tonnes of additional buoyancy. This
buoyancy was primarily to limit the jacket trajectory through launch (i.e. to stop it hitting the sea bed) but was also essential for maintaining bottom clearance during up-ending. The additional buoyancy took the form of two 'saddle' tanks, two pairs of twin 'piggy-bank tanks' and twelve 'cigar' tubes installed down the pile guides (Figure 3). Altogether the auxiliary buoyancy added about 3,300 tonnes additional weight to the jacket.
Very large jackets, in excess of launch capacity, have been constructed as self-floaters in a graving dock, towed offshore subsequent to flooding the dock, and installed on location by means of controlled flooding of the legs (see Figure 4).
1.3 Installation Planning The installation of a jacket consists of loading out, seafastening and transporting the structure to the installation site, positioning the jacket on the site and achieving a stable structure in accordance with the design drawings and specifications, in anticipation of installation of the platform topsides. An important aspect is the avoidance of unacceptable risk during offshore activities from loadout through to platform completion. It is recognised that the potential cost to the project associated with failure to successfully execute marine activities is particularly high. Normally therefore the contractor is obliged to produce procedures for these activities which demonstrate that the risk of failure has been reduced to acceptable levels. He is also required to demonstrate that, prior to the commencement of an activity, all the necessary preparations have been completed. An installation plan must be prepared for each installation. The plan will include the method and procedures developed for the loadout, seafastening and transportation and for the complete installation of the jacket, piles, superstructure and equipment. Depending on the complexity of the installation, detailed procedures and instructions may be needed for special items such as grouting, diving, welding inspections, etc. Limitations on the various operations due to factors such as environmental conditions, barge stability, lifting capacity, etc. must be defined. The installation plan is normally subdivided into phases, e.g., loadout, seafastening, transportation and installation. Installation drawings, specifications and procedures must be prepared showing all the pertinent information necessary for construction of the total facility on location at sea. These drawings typically include details of all inspection aids such as lifting eyes, launch runners or trusses, jacking brackets, stabbing points, etc. For jackets installed by flotation or launching, drawings showing launching, up-ending and flotation procedures must be prepared. In addition, details are also provided for piping, valving and controls of the flotation system, etc. as well as barge arrangements and tie-down details. The engineering input into an offshore installation project also involves the design of all temporary bracing, seafastenings, rigging, slings, shackles and installation aids, etc. These must be designed in accordance with an approved offshore design code, e.g. API RP2A [1].
Quality management is a vital and integral component of all offshore installation projects. A general note on Quality Assurance for Offshore Construction is appended toLecture 15A.8 : Fabrication. It is equally applicable to an offshore installation project. 2. LOADOUT AND SEAFASTENING Loadout entails the movement of the completed structure onto the barge which will transport it offshore. Seafastening entails fitting and welding sufficient ties between the jacket and the barge to prevent the jacket shifting while in transit to the offshore site. Jackets which have been fabricated on their side are usually loaded by skidding the entire structure onto a cargo or launch barge. During loadout, the jacket is supported on the skid ways, usually on two inner legs of the jacket, see Fig. 9 of Lecture 15A.1. The legs function as the bottom chord of a large truss, which can span between points of support, especially when part of the jacket is on the barge and part still on the skid ways. Where jackets are fabricated in the vertical, i.e. in the same attitude as the final installation, they may be lifted onto the barge or skidded out. In this latter case, adequate temporary pads and braces must be provided under the columns to distribute the loads during skidding. Initial friction of the jacket on the skid ways may be as high as 15 per cent, especially if the jacket has been erected with its weight bearing continuously on the skid way. In some cases the jacket is initially fabricated slightly above the skid ways using hydraulic or sand jacks. Then, at the time of loadout, the jacket is lowered onto the skid ways. To reduce the sliding friction, grease on hardwood, or heavy lubricating oil on steel, or even fibre-filled Teflon faced pads, are used. Values of sliding friction as low as 3 per cent are usually attained. The barge should be of adequate size and structural strength to ensure that the stability and static and dynamic stresses in the barge and seafastenings due to the loading operation and during transportation remain within acceptable limits. The barge must also have the capability to launch the jacket, if this is required, without the use of a derrick barge. For a barge which floats during the loadout, the ballast system must be capable of compensating the changes in tide and loading. It is usual in this case to load out on a rising tide so that the tide assists the ballast system. In the case of a barge which will be grounded during loadout, the barge must have sufficient structural strength to distribute the concentrated deck loads to the supporting foundation material. The jacket must be loaded in such a manner that the barge is in a balanced and stable condition. Barge stability can be determined in accordance with regulations such as those published by Noble Denton, The American Bureau of Shipping, or the US Coast Guard. Allowable static and dynamic stresses in the barge hull and framing due to loadout, transportation and launching must not be exceeded. A simplified check list for the operations relating to jacket loadout might be: 1. 2. 3. 4.
5.
Is the jacket complete? Has the structure been analysed for loadout stresses on the basis of the actual structure as fabricated at the time of loadout? Is the launch barge securely moored to the loadout dock, so that it won't move out during the loading? Is the barge properly moored against sideways movement? If compression struts are used between the barge skid ways and those on shore, are they accurately aligned and supported so they won't kick out during loadout? Have the pull lines, shackles, and pad eyes been inspected to ensure they are properly installed and can't foul during loadout? Can the barge be properly ballasted? If the tide will vary during loadout, are ballasting arrangements adequate? Will ballast be adjusted as the weight of the jacket goes onto the barge? Are there proper controls? Is there an adequate standby ballast system? Are there back-up systems to pull the jacket back to shore if anything goes wrong during loadout? If the ballast correction is to be made iteratively, step-by-step as the jacket is loaded, are there clear paint marks so that each step can be clearly identified? Have clear lines of supervision and control been established? Are the voice radio channels checked? Have the marine surveyors been notified so that they can be present? Owner's representatives? Certifying Authority? Have their approvals been received?
Once the jacket is on the barge, the barge must be ballasted for transportation. During loadout, many tanks will be partially full, in order to control deck elevation and trim. However, with the jacket fully supported on the barge, these considerations are no longer relevant and the tanks can be ballasted to suit the demands of the sea voyage. Ballast tanks should normally either be full or completely empty, to eliminate free surface and sloshing effects. The draft and freeboard will have been carefully selected to maximise stability, and especially to minimise submersion of projecting members of the jacket during the tow and the consequent slamming, buoyancy and collapse forces. Large jacket launch and cargo barges are relatively flexible structures in that the jacket structure is normally (much) stiffer. Therefore, ballasting the barge to obtain the required draft and trim should preferably be done at the dock side before seafastenings are attached. If one scheme of ballasting is to be used for a sheltered channel tow and another for the open sea, the seafastenings should be freed during the reballasting to avoid imposing undue stresses on the jacket legs or, alternatively, calculations should be performed to demonstrate that freeing is not required by the reballasting procedure. Seafastenings are installed after loadout and must be completed prior to sailaway. They are major structural systems, subjected to both static and dynamic loads. When the barge is on the high seas it must be assumed that it can encounter conditions which are "as bad as could have been statistically foreseen". Accordingly, the gravity and inertial forces involved must be calculated for all anticipated barge accelerations and angles of roll and pitch during the design sea conditions adopted for the tow (usually the 10-year return storm for that season and location). In determining this criteria, the reliability of the short term weather forecast should be considered. Since the loads are dynamic, impact must be minimised. Seafastenings should be attached to the jacket only at locations approved by the designer. They should be attached to the barge at locations which are capable of distributing the load to the barge internal framing. They should also be designed to facilitate easy removal on location. Seafastenings are normally subject to the same code requirements for fabrication as the jacket. 3. OFFSHORE TRANSPORTATION The transportation of heavy components from a fabrication yard to the offshore site is a critical activity. It is especially so in the case of the jacket since the behaviour of the unit usually influences the verification of barge strength, the design of seafastenings, and indeed the design of the jacket itself. Also there are the practical aspects of tug selection, tow route, etc. to be considered. The size and power requirements of the towing vessels and the design of the towing arrangement must be calculated or determined from past experience. Tug selection involves such considerations as length of tow route, proximity of safe harbours and the anticipated weather conditions and sea states. As a minimum the tugs should be capable of maintaining station in a 15 metre/second wind with accompanying waves. However, this criterion depends on the location. For instance, the requirement in the Mediterranean is typically that the main tug should maintain station against a 20 metre/second wind, 5,0m significant sea-state and 0,5 metre/second current, acting simultaneously. Weather forecasting is provided throughout the tow so that, if exceptional weather threatens, a pre-arranged port of refuge may be sought. Experience has shown that the first phase of transportation is the most treacherous. There are several reasons for this. In the harbour area a big tug can normally exercise very little control even with a shortened towline. With a short towline between two considerable masses, the large tug and the much larger barge/jacket, the risk of snapping is high. Thus it is standard practice to lengthen the towline once out of the port. Also, because of the nature of many ports, close control is essential in order to avoid the possibility of running aground. Normally, therefore, the harbour tugs take the barge out under the guidance of a pilot who knows the port. When the barge is out of the port the problems are not totally solved since it must be assumed that the worst can happen, i.e. the towline may break. The tug must have sufficient time to pick up the emergency towline and control the barge before it drifts into shallow water. Thus the departure is normally subject to strict weather forecast conditions for a period which assumes that the speed of the tow is between 1 and 2 knots for the first 100 nautical miles from the coast. Consequently, as a minimum, a favourable 48hour weather forecast is required, e.g. Force 5 and decreasing. Once the tow is under way, trim will be adjusted to optimise tow speed and give directional stability during tow. Usually the barge will be trimmed down by the stern.
The behaviour of the jacket seafastened to the barge must be satisfactory both from the point of view of static and dynamic stability. Both are verified by means of numerical analyses. However, particularly for larger structures, the sensitivity of the dynamic analysis will usually warrant verification by model testing. The intact static stability criteria usually adopted is that the righting arm be positive throughout a range of 36° about any axis. The so-called dynamic stability of wind overturning criteria simply ensure that for a given wind, the energy which tends to overturn the barge is at least 40% less than that which is available due to the inherent righting stability of the barge. In considering the motions of the jacket and barge it is intuitively plausible that roll will be the most problematical motion (from the point of view of body accelerations) and that the largest roll will be caused by a beam sea. It may be less obvious, but nevertheless true, that if the barge width and, to a lesser extent the length, are reduced, the roll will diminish and if the barge is set at a (much) deeper draft, the roll will also diminish. All of these considerations reflect static properties of the jacket and barge. Improvements can occasionally be made by choosing a narrower barge (although obviously stability will suffer) or increasing the draft (although in this case stability may again suffer and parts of the structure which were previously 'dry' may now be subjected to 'slamming'). Incorrect "balancing" of these aspects can have very serious cost/risk implications in overall project terms. Thus, for a large jacket, the barge selection process is normally performed at a very early stage of the design process. 4. OFFSHORE INSTALLATION This section is concerned with the stages of jacket installation commencing with removal of the jacket from the barge to its placing on the sea bed and temporary on-bottom stability. Lecture 15A.6: Foundations covers the subject of pile installation. 4.1 Removal of Jacket from Barge Unless a jacket is a self floater, it must first be removed from the transportation barge. There are two basic methods used:
launch lift
4.1.1 Launch The launch site is normally at or near the installation location. With heavy jackets in shallow water it may be necessary to launch the jacket in deep water at some distance from the installation location and tow the jacket to site. Immediately prior to launch, the seafastening securing the jacket to the barge is cut. The jacket is then pulled along the barge skid ways (which were used for loadout) by winches. As the jacket moves towards the stern of the barge, the barge start to tilt and a point is reached when the jacket is self sliding. An initial tilt to the barge may have been provided by ballasting immediately prior to launch. A stern trim of approximately 5° is usually aimed for. The skid ways terminate in rocker arms at the stern of the barge. As the jacket moves along the skid ways its centre of gravity reaches a point where it is vertically above the rocker arm pivot. Further movement causes the rocker arm and jacket to rotate. The jacket will then slide under its own self weight into the water. Various stages in the launch of a jacket are shown in Figure 1a to 1d. Once in the water the self floating jacket is brought under control with lines from tugs and/or the installation vessel. The jacket must be designed and fabricated to withstand the stresses caused by the launch. This can be achieved either by strengthening those members which might be over-stressed by the launching operation, or designing into the jacket a special truss, commonly referred to as a launch truss. Spacing between jacket members or launch trusses will be dictated by the spacing between launch skid ways. Thus a jacket will generally be designed from the outset for installation by a specific barge. Once launched the jacket must float with a reserve of buoyancy in order to stop the downward momentum of the jacket. This requires the jacket to be water tight. It is common practice to gain additional buoyancy by sealing jacket legs and pile sleeves with removable rubber diaphragms. However, there is frequently a need for even more buoyancy. This is achieved
by adding buoyancy tanks. These need to be removable and are located where they give most benefit. Buoyancy tanks from previous launches are often used. The launch of a jacket is clearly a critical phase in the life of the jacket. Considerable design effort is required in order to ensure that the launch sequence is feasible. A jacket launch naval analysis is required in order to:
ensure that an adequate sliding velocity is maintained during the rocker arm rotation; verify that the trajectory followed has a safe seabed clearance; determine the jacket behaviour during launch; define operational requirements during launch, including ballast configuration; check the stability of the jacket during launch and when free floating.
The plots shown in Figures 1a to 1d are extracted from such an analysis. The jacket weight was 14,000 tonnes and was being installed in 105 metres of water. The analysis showed that it should take approximately 2 minutes between start of self sliding (Figure 1a) and the jacket reaching its final floating position (Figure 1d). 4.1.2 Lift An increasing number of jackets are being installed by direct lift. This trend has been encouraged by the availability of large crane vessels such as the Micoperi 7000. Curves showing load capacity against lifting radius are shown in Figure 2. Another factor tending to increase direct lift jackets are savings in weight that are being achieved in jacket design. In a direct lift the jacket is lifted off the barge completely in air. A second form of lift is the buoyancy assisted lift. In this case the barge is flooded and hence submerged. This results in part of the jacket being buoyant, reducing hook loads. Buoyancy tanks may be added to the jacket if required. Shallow water jackets may be lifted in the vertical position. In this case no up-ending is required and installation is straight forward. Deep water jackets will in general be lifted on their side. Two cranes will normally be used, noting that large derrick barges such as the Micoperi 7000 are fitted with two cranes as standard. When considering a tandem lift it should be noted that it is unlikely that both hooks will carry the same load, and that the maximum permissible jacket weight will be less than the sum of the two crane capacities. It should also be noted that cranes are frequently guyed back to give maximum lift capacity and carry less load if they are revolving. This can further reduce the apparent lift capacity. Finally, the weight of lifting slings need to be considered, these contributing as much as 7% of the lift weight. When the jacket is to be removed from the transportation barge by lifting, it is normal for the installation vessel to be correctly moored and positioned so that up-ending and set-down may proceed as one integral lift operation. The selection of a suitable installation vessel is clearly essential. In addition to lift capacity, it is also necessary to consider stability and motion response characteristics. In the harsh North Sea environments installation vessels are usually semisubmersibles such as the Micoperi 7000. In more moderate waters they are often flat bottomed barges. In intermediate environments, e.g., the Gulf of Mexico, ship-shaped vessels may be used. The large semi-submersible crane vessels used in the North Sea have full dynamic positioning systems for locating themselves on site. They also have sophisticated computer controlled ballast systems to keep the vessel level during lifting operations. During a lift the ballast system is also used to counteract heel and increase hoisting and lowering speeds during the crucial lift-off and set-down operations. The natural period of large installation vessels in roll, pitch and heave tend to be close to the typical peak periods of the sea spectra encountered offshore. These motions therefore predominate. Normally this means that beam seas should be avoided since this excites roll which is the most disruptive motion. However, the "best attitude" is not always possible since it depends on the work that the vessel is required to perform. Accordingly vessel operators perform extensive studies to determine permissible sea states for specific operations and vessel captains invariably "experiment" with different headings in a particular sea in order to minimise motions and maximise workability. The first stages in lifting a jacket from the transportation barge involve positioning the barge and connecting the slings to the hook. The barge will normally be controlled by tugs. Once everything is ready for lift to proceed the seafastenings will be cut.
The next stage is to transfer the weight of the jacket from the barge to the crane. The general requirement here is to lift as rapidly as possible. However, careful control and phasing with barge and crane vessel motions is required in order to ensure that once the jacket is lifted clear of the barge it does not hit the barge as a subsequent wave passes through. The same lift procedure is adopted in both a direct and buoyancy assisted lift. Once the jacket is lifted clear of the barge, the barge is removed by tugs. Up-ending of the jacket will then normally proceed directly. 4.2 Jacket Up-ending and Set-down Unless a jacket is transported and lifted in its upright position, it will be necessary to up-end the jacket at the installation location. Up-ending may be achieved by controlled flooding of buoyancy tanks, by using a crane vessel or by a combination of both. 4.2.1 Up-ending by Ballast control and Flooding A large crane vessel will not normally be required for either a launched or self-floating jacket. Upending is therefore achieved by controlled flooding. A small installation vessel will usually be required for the installation of piles once the jacket has been set-down, so this is used as the platform from which to control the various flooding operations. This installation vessel will also be used to help position the jacket. Figure 3 shows a sketch of the Brae 'B' jacket showing the auxiliary buoyancy tanks. In this case the flooding system involved 42 primary and 22 contingency subsea valves under direct hydraulic control. The nitrogen power source and associated control panels were contained in watertight capsules. Figure 4 shows a sequence of sketches indicating how a self floating jacket is upended. In step 1 the waterline compartments at one end of the jacket are flooded. More water line tanks are flooded in step 2 until by step 3 the upper frame of the jacket reaches waterline and may also be flooded. The jacket is then allowed to rotate until all legs are equally flooded as in step 4. The jackets natural position will then be floating upright as in step 5. Further flooding of the jacket as in step 6 will enable the jacket to be lowered onto the sea bed in a controlled manner. The up-ending of a launched jacket will be similar to that shown in Figure 4. The main difference is that there may be less excess buoyancy with which to control the operation. In this case a combination of flooding and lift, as shown in Figure 5, may be used to up-end and set-down the jacket.
The crane and ballasting operations need to be clearly defined before the operation begins. This involves careful naval analysis of the free floating position of the jacket at various stages during the up-ending procedure. A feature of these analyses is the need to consider what happens in the event of buoyancy tanks being accidentally flooded, or of flooding valves failing to operate. Contingency procedures and equipment must be provided. 4.2.2 Up-ending using the crane vessel Figure 5 shows the most simple use of a crane to up-end and set-down a jacket. This is acceptable for jackets that are launched. For horizontally oriented jackets that are lifted directly the procedure is more involved. A horizontally lifted jacket may be upended in one of two ways. Perhaps the most straight forward is to lower the jacket into the water so that it floats. Slings can then be removed and new slings attached at the top of the jacket. The jacket may then be up-ended as shown in Figure 5. This may require closures to legs and some additional buoyancy. A second method is to up-end directly, as shown in Figure 6. This requires special padears so that the necessary rotation between slings and jacket can occur. Careful naval analysis is also required in order to carefully determine hook loads and to ensure that the jacket remains stable.
Once up-ended the jacket can be set-down on the sea-bed. Since the lifting points are submerged divers may be required to disconnect the slings from the jacket. Although two crane hooks are shown in Figure 6, it should be noted that for light weight jackets it is possible to up-end using a single crane. In this case the main and auxiliary hooks are used together, for example the main hook taking the weight of the jacket with the auxiliary hook providing the upending force.
An increasing trend is to install a jacket over an existing well or wells. A pre-drilling template will have been used to position the wells, the same template being used to position the jacket. It is necessary to ensure that the well heads are protected from damage due to accidental contact with the jacket. Once set-down the jacket should be positioned at or near grade and levelled within the tolerances specified in the installation plan. Once level, care should be exercised to maintain grade and levelness of the jacket during subsequent operations. Levelling the jacket after all piles have been installed should be avoided if at all possible as it is costly and frequently ineffective. If necessary, levelling should take place after a minimum number of piles have been driven by jacking or lifting. In this instance procedures should be used to minimise bending stresses in the piles. 4.3 On-bottom Stability Once set-down on the sea bed, it is normal for piling to proceed as rapidly as possible. However, this far into the installation procedure the weather and hence sea conditions may be detioriating. This is a result of long term weather forecasting being less reliable than short term forecasting. It should also be noted that any problems encountered during the installation procedure will result in delay and that it may be some time before the jacket is adequately fixed to the sea bed by piling. It is necessary for the jacket to be stable and level during piling. A separate on-bottom stability analysis is therefore carried out. Three conditions need to be met: (1) vertical resistance to jacket weight and piling loads; (2) stability against sliding under wave/current loading; (3) stability against overturning under wave/current loading. In carrying out the above analyses it is necessary to use an appropriate sea-state to generate hydro-dynamic loading. This should be the maximum statistical wave which may occur prior to piling being completed. Assuming installation to occur in the summer months, a typical criteria may be a 1 year summer storm wave. The provisions that need to be made to ensure on-bottom stability vary greatly depending on jacket location, height and on sea-bed soil conditions. For example, with good soil conditions the jacket may be able to be supported directly on existing jacket steel with no extra provision made. However, with poor soil conditions large 'mudmats' may be required in order to spread the load. These can influence launch and installation dynamics. For many jackets it is not possible to achieve stability against sliding and overturning using flat mudmats. In these circumstances mudmats with skirts may be used. Skirts considerably improve the resistance to sliding, and in silty or clay soils can allow nominal tension loading to resist overturning. Another option frequently used is to stab a number of piles as soon as the jacket is set-down. These will penetrate some distance under self weight providing additional sliding resistance. Since most piles are inclined, the piles also provide a degree of resistance to over turning. 5. CONCLUDING SUMMARY
There are broadly four phases to the installation of a steel jacket - loadout, seafastening, offshore transportation and installation offshore. In deciding how best to fabricate and install a given jacket, the options are principally determined by the installation equipment available and the jacket's intended water depth. An installation plan must be prepared for each installation. Loadout entails the movement of the completed structure onto the barge which will transport it offshore. Seafastening entails fitting and welding sufficient ties between the jacket and the barge to prevent shifting while in transit to the offshore site. The transportation of heavy components from a fabrication yard to the offshore site is a critical activity requiring careful calculation and planning. Removal of the jacket from the barge is accomplished either by direct lifting with a derrick barge and lowering into position, or by launching. A number of engineering studies are required for jacket launch and set-down.
6. REFERENCES [1] API RP2A, Recommended Practice for Planning, Designing and Construction of Fixed Offshore Installations, latest edition. Engineering design principles and practices that have evolved during the development of offshore oil resources. 7. ADDITIONAL READING 1. 2. 3. 4. 5.
Det Norske Veritas Marine Operations Recommended Practice RP5 - Lifting (June 1985). Principles and good practice for offshore heavy lifts. AISC Specification for the Design, Fabrication and Erection of Structural Steel for Buildings, latest edition. API code refers to this specification for calculation of basic allowable stresses of all jacket members. AWS Structural Welding Code AWS D1.1-88. All jacket welding and weld procedure qualifications are required by the API code to be undertaken in accordance with this code. Det Norske Veritas, Rules for the Design, Construction and Inspection of Offshore Structures, 1977. Rules for construction and installation of steel jackets as required by DNV. Lloyds Register of Shipping, Rules and Regulations for the Classification of Fixed Offshore Installations, 1989. Based on Lloyd's experience from certification of over 500 platforms world-wide.
Operator
Name
Type
Mode
Lifting Capacity
Heerema
Thor
Monohull
Fix
2720
Rev
1820
Fix
2720
Rev
2450
Fix
4536 + 3628 = 8164
Rev
3630 + 2720 = 6350
Fix
3630 + 2720 = 6350
Rev
3000 + 2000 = 5000
Fix
4000
Rev
3800
Fix
1820
Rev
1450
Fix
3360
Rev
2450
Odin
Hermod
Balder
McDermott
DB50
DB100
DB101
Micoperi
Monohull
Semisub
Semisub
Monohull
Semisub
Semisub
DB102
Semisub
Rev
6000 + 6000 = 12000
M7000
Semisub
Rev
7000 + 7000 = 14000
Notes: 1. 2.
Rated lifting capacity in metric tonnes When the crane vessels are provided with two cranes, these are situated at the vessels stern at approximately 60m distance ctc.
Table 1 Major Offshore Crane Vessels
A.10: Superstructures I OBJECTIVE/SCOPE To introduce the functional requirements; to identify major interfaces with the process, equipment, logistics, and safety; to introduce the structural concepts for jacket and gravity based structure (GBS) topsides; to elaborate on structural design for deck floors. PREREQUISITES Lectures 1A & 1B: Steel Construction Lecture 2.4: Steel Grades and Qualities Lecture 2.5: Selection of Steel Quality Lectures 3.1: General Fabrication of Steel Structures Lecture 6.3: Elastic Instability Modes Lecture 7.6: Built-up Columns Lectures 8.4: Plate Girder Behaviour & Design Lectures 11.2: Welded Connections Lecture 12.2: Advanced Introduction to Fatigue Lectures 15A: Structural Systems - Offshore SUMMARY The topside lay-out is discussed, referring to API-RP2G [1], and to general aspects of interface control and weight control. The different types of topside structures (relevant to the type of substructure, jacket or GBS) are introduced and described. These types are: 1. 2. 3.
integrated deck. module support frame. modules.
Floor concepts are presented and several aspects of the plate floor design are addressed. 1. INTRODUCTION This lecture deals with the overall aspects of the design of offshore topsides. The topside of an offshore structure accommodates the equipment and supports modules and accessories such as living quarters, helideck, flare stack or flare boom, microwave tower, and crane pedestals. The structural concept for the deck is influenced greatly by the type of substructure (jacket or GBS) and the method of construction, see Figures 1 and 2.
Heavy decks, over 10,000 tons, are provided with a module support frame onto which a number of modules are placed. Smaller decks, such as those located in the southern North Sea, are nowadays installed complete with all equipment in one lift to minimize offshore hook-up. Most of this lecture refers to this type of integrated deck such as is shown in Figures 3 and 4.
The selection of the concept for the structural deck is made in close cooperation with the other disciplines. 2. BASIC ASPECTS OF DESIGN 2.1 Space and Elevations The first step in developing a new design concept is to consider all the requirements for the deck structure. The design requirements and their impact on the structural system are discussed below. The lay-out of the deck is influenced by the type of hydrocarbon processing to be undertaken. The area required for the equipment, piping and cable routings, the vertical clearance as well as the access/egress requirements determine the deck area and deck elevations. The elevation on the lowest decks depends on the environmental conditions. The elevation of the cellar deck, i.e. the lowest deck, is based on the maximum elevation of the design wave crest, including tide and storm sway, plus an air gap of 1,5m minimum. The vertical distance between the decks of the topside is generally in the range 6 - 9m in the North Sea. Consideration of the prevailing wind direction is very important in determining the position of various components on the platform, such as the vent of the flare, cranes, helideck; and the logistic and safety provisions. 2.2 Lay-out Requirements
The requirements for the various topside components are briefly described below, based on API-RP2G [1]. Wells: the position of the wells depends heavily on whether the wells will be drilled and worked with a separate cantilevered jack-up rig or with a platform-based rig. In the first case the wells must be close to the platform edge and require significant deck area above free of obstacles. In the second case a pair of heavy beams to support the drill rig must be provided. Equipment, piping and cable-supports: all devices to treat the oil or gas shall comply with the requirements of API-RP2G [1]. Living quarters and helideck: the helideck should be in the vicinity of the living quarters to enable fast evacuation. Usually the helideck is located in the obstacle free area on top of the living quarters. Gas compressor module: the pressure in gas reservoirs declines due to production. Future compression may be needed in order to achieve acceptable gas flow through the export pipeline. Water or gas injection module: oil production declines after some years of operation. The reservoir then requires stimulation by, for example, injection of water. Deck crane: the location of the crane should be selected so as to obtain maximum deck coverage and to enable the crane operator to keep eye contact with the lifted object and the supply vessel. The location of the deck crane should be outside the obstacle free area of the helideck and should not interfere with future facilities. Vent/flare boom or stack: a vent discharges gaseous products in the air without burning them; a flare discharges and burns these products. Both vents and flares should be located outside hazardous areas and away from the helideck. The tip shall exceed the elevation of the helideck by at least 100 feet. Heat radiation shall be checked. Microwave tower: A high mounting is required to provide obstacle free support for microwave antennae. A stiff support is required in order to comply with the stringent deflection criteria. Survival capsules and man-overboard crane: the supporting structures for these items usually cantilever from the main structure. Shock loading and dynamic amplification increases the support reactions during operation. Walkways, ladders and stairs: these items should be kept obstacle free, be non-slippery and have sufficient width to allow evacuation of personnel on stretchers. Cladding, walls, doors and louvres: the type of cladding depends on the operational requirements and the preference of the oil company. For safety reasons, walls and doors may have to satisfy specified explosion and fire resistance requirements. Louvres may be used to allow natural ventilation, whilst preventing entry of rain, snow and birds. Lay down areas for equipment, spares and consumables: these areas are provided by cantilevering from the main structure in order to allow access to the lower deck levels by the deck crane, without providing hatches through the decks. Hatches: access to the lower decks within reach of the crane is required to enable maintenance, repair and platform modification. The hatches should be identified early in the design. Risers, caissons, sumps: the riser section of the pipeline rises from the seabed to the deck. It introduces vertical and horizontal loads (environmental and operational) in the deck structure. Caissons for pumps and sumps for discharge are hung from the cellar deck and introduce significant vertical and horizontal loads in the deck. Drainage provisions: provision is required for spillage in drip pans under the equipment and for collecting oil-polluted rainwater to prevent spilling into the sea. Deck penetrations: pipes connecting process-items on different decks and, vessels, cable routings, etc. can require significant areas to be clear of structural members. The major penetrations should be identified early in the design and coordinated with major structural members.
Other provisions: items such as monorails and inspection gangways may also be required. 2.3 Loads In Lecture 15A.3 the different types of loads have been identified and partly quantified. Dead weight, tankful live load and wind load are discussed here. Dead weight includes the weight of structure, equipment, piping, cables, machinery, and architectural outfitting. Tankful live load covers weight of potable water, diesel fuel, helifuel, glycol, methanol, well-kill mud, lubrication oil, waste, etc. Live load also covers all sorts of miscellaneous loads such as bagged or palletized consumables, spare parts, maintenance equipment, etc. The application of live load is typical for topsides. For design considerable engineering judgement is required concerning:
the magnitude of the load to be applied to the various structural items:
- direct loaded deck stringer - deck beam - deck truss - deck leg - jacket - pile - pile bearing resistance
the area to which live load is to be applied. This area is described in the code as the non-occupied area.
For local strength, the walkways, escape routes, etc. are considered as non-occupied by equipment and are thus loaded by live load. For overall strength, the walkways, escape routes, etc. are considered as occupied (kept clear for evacuation) and consequently no live load is applied.
the arrangement of loads that generates maximum stress. A policy on this item should be prepared for each project, stating both variation of loads over one deck, and variation over various decks.
Wind loads should be properly assessed. For overall structural integrity, the complex shape of the platform creates problems in assessing the effective area for wind load. Special elements such as communication towers and flare booms require consideration as wind sensitive structures. To control the design process, weight engineering as explained in Section 2.5 below, shall be performed by the project management staff. Any structural analysis must be linked to the latest available information in the weight report. This requires that the load file for the structural analysis and the weight report are compatible with respect to total weight, weight distribution and centres of gravity. 2.4 Interface Control The many functions of the topside result in the involvement of many disciplines in the design. Due to the high cost of providing platform space, the facility must be designed to be very compact. This requirement leads to several major areas of interdisciplinary control.
Space allocation: the structure should not use space allocated for equipment or access routes. Overhead clearance between piping, cable routes, equipment and the floor overhead should be respected. Direct interface control; pumps, vessels and piping require support by the steel structure.
Interface between drilling and workover operations. Interface between platform crane and helideck, deckhouse, drilling rig, and flare stack. Interface with the export riser. Interface between the deck and modules. Interface between the topside and bridge from adjacent platform. Interface with the substructure.
2.5 Weight Engineering The weight of the overall facility as well as its major components is critical. Lack of weight control can lead to costly design changes as well as to major provisions in order to keep within the limits of the construction strategy. Weight engineering consists of:
weight prognosis weight reporting weight control weighing
Weight prognosis is the methodology which applies an uncertainty surcharge as high as +30% in the conceptual design phase, to +5% in the final fabrication phase, see Figure 5.
3. STRUCTURAL SYSTEMS 3.1 Selection of Topside for a Main Jacket-Based Structure
The selection of the concept for the topside structure is the second step in the development of a structural system. The two possible basic alternatives: a truss type (Figure 4) or a portal-frame type without braces (Figure 3), are compared in Table 1. Table 1 Comparison of concepts for main jacket-based structures No.
Item
Truss type
Frame type
1.
Discipline non-interference
-
++
2.
Flexibility during construction
-
++
3.
Flexibility during operation
-
++
4.
Automated fabrication
-
++
5.
Construction depth
++
0
6.
Inspection
-
++
7.
Maintenance
-
+
8.
Weight of structure
+
0
9.
Strength reserve
+
++
10.
Potential for high strength steel
+
++
11.
Structural CAPEX
+
+
12.
Platform CAPEX
+
++
Note: ++ denotes greater benefit -- denotes greater disadvantage The selection of the topside main structure concept, truss or portal frame, is linked with the decision of the position of the longitudinal structure in the cross-section. In a 20-25m wide deck, trusses will generally be arranged in longitudinal rows: centre line and both outer walls (Figure 6).
In such decks, however, portal frames will be arranged in 2 longitudinal rows, approximately 14-16m apart, allowing floor cantilevers of approximately 5m (Figure 3). 3.2 Selection of Topsides for Gravity Based Structures Topsides of gravity based concrete structures (GBS) are quite different from the jacket based topsides, see Lecture 15A.1. The topside structure is an important element in the overall portal-type system. Gravity based substructures have been built with one to four shafts. A rectangular or a T-arrangement of four shafts has been adopted. The basic form is a modularized topside with a grid of heavy box girders. A few elements only of the GBS-topside structural design are indicated below:
due to portal frame action, the deck is subject to fatigue; a design case difficult to control in topside design. equipment lay out optimization, piping and cable routes, logistic and emergency routes require many big openings and perforations of plate walls, thus creating stress concentrations. attachment of secondary structures and of equipment/pipe/cable supports to the main structure must be strictly controlled, to avoid fatigue problems. the connection area with the concrete shaft must provide the transition from circular (shaft) to square (deck). It accommodates high strength anchor bars, temporary crushing devices for deckmating, and requires tolerances on deck and substructure dimensions. inspection and repair options must be planned carefully, especially as fatigue may occur.
The material used at present is high strength steel typically of 355 MPa yield stress. There is a trend to use higher strength steel (420-460 MPa). 3.3 Floor Systems The concept for the floor-system in offshore structures is conventional: hot rolled beams, typically at 1000-1200mm centres, are covered by a chequered or flat steel plate 6-10mm thick. The options are:
conventional steel floor steel grating (bar-type or plate type) aluminium floor system orthotropic deck in steel corrugated steel plate
The conventional steel floor contributes approximately half of the weight of the steel structure of an offshore deck. Steel gratings, especially with the plate type, could gain increased application as their weight per sq.m. is attractive. Aluminium has attracted much interest recently; current development in Norway will show its real potential. Orthotropic decks in steel have found application in helidecks. Further study is required to assess their actual feasibility for floors of offshore modules. Corrugated steel plate (approximately 1-3mm thick) as sub-flooring has been used in living quarters. In summary, the floor concept used for a typical floor of an offshore deck of a module is a conventional steel floor or steel grating. 3.4 Floor Panel Concept for Conventional Steel Floor The floor panel, defined as the assembly of the floor plate and the stringer, can be connected to the overall structure in two ways:
stacked: stringer over the top of deckbeams. flush: stringer welded in between deck beam, with top flange in one plane. It is practically impossible to change from the flush to stacked arrangement in a later phase of the design.
All elevations and overhead clearances are involved in the choice of arrangement. Clearances are very important for equipment height, pipe routing, pipe stress, cable routing, etc. The single most important structural aspect is the amount of prefabrication that can be carried out away from the main fabrication yard. The cost is also a very important factor. 3.5 Floor Stabilization Concept The deck structure requires lateral stabilization of each floor with respect to:
lateral instability of beams horizontal forces, e.g. wind, pipe reactions, sea transport horizontal components of permanent braces horizontal components of temporary braces, e.g. seafastening horizontal components of sling forces module skewing during installation.
There are essentially two options for floor stabilization:
provision of separate underfloor horizontal bracings allocate the stabilization function to the floor plate.
There is a clear preference for the stabilization by the floor plate. Where underfloor bracing is adopted, there are two configuration options (see Figure 11). The rhomboid solution should be chosen for the upper deck, due to congestion at the column by the padeyes for lifting. The underfloor bracing under a plate floor does create a very unclear structural situation. The bracing is assumed completely to perform the stabilizing function, but in practice the floor plate is much too stiff to allow that. It is common practice in the structural analysis for underfloor bracing to neglect completely the floor plate.
4. DECK FLOORING DESIGN 4.1 Introduction The selection of the main deck dimensions have been considered above in relation to lay-out requirements. The interactive process of conceptual design of the jacket and deck yields the spacing of the columns. In the Dutch sector of the North Sea, transverse column spacings are typically 9m for a wellhead platform to 15m for a production platform. Longitudinally spacings are typically 15m. Next decisions are made on:
floor system: plate versus grating main structure: truss versus portal frame
floor panel concept: stacked versus flush floor stabilization underfloor bracing versus plate.
The structural concept is then complete. A principle for economic design of steel structures is that the load-paths should be short. For a floor design of a production deck typical dimensions are: Structural item Typical span 1. 2. 3. 4. 5.
floor plate 1m stringer (longitudinal) 5m deck beam (transversal) 15m main structure (longitudinal) 15m column
These components are identified in Figure 6. 4.2 Floor Plate Design Options are to choose between flat plate, chequered plate or tear plate. Another option for providing slip resistance is to coat with a sand finish. The floor plate thickness is usually 8-10mm and 6mm for lighter loaded floors, although welding distortion may rule out the 6 mm thickness. In practice the floor plate acts as horizontal bracing between the columns. Special attention is required to ensure that all welds between the floor plate and the underlying structure do not form brittle points. Failure of such welds could lead to crack initiation in the rest of the structure. The same attention applies to the buckling of the floor plate by stresses which are picked up unintentionally. Strength of Floor Plate The strength of the floor plate is very high both for uniform as well as concentrated loads. Elastic, small-deflection theory provides uneconomical conservative results. API-RP2A (2) does not specify live loads. They are specified by the operator. For main decks generally accepted figures are: p = 20kN/sqm, or F = 10-25 kN on a 0,3 x 0,3m load area Det Norske Veritas [3] presents an expression for the required plate thickness t, which incorporates membrane effects and is of special interest for design for local loads. Equipment and containers are regularly offloaded by the crane in some deck areas, such as lay-down areas and food container platforms. An increased plate thickness may be required in these areas due to larger concentrated loads (1).
4.3 Stringers The typical stringer for a production platform is an IPE 240-270 or HE 240-280A profile positioned at approximately 1m centres and spanning 5m. It is important to choose, especially for stacked floor panels, a profile which allows selection of heavier sections with practically identical depth to accommodate local heavy equipment. Designers should avoid choosing deeper sections or reinforcing them to accommodate late extra load requirements by welding another section underneath. Interference with small diameter, hard piping or with cable trays then is quite likely. Joining floor plate and stringers requires welding. Intermittent welding is generally not accepted. A continuous thin weld (a = 4 mm) is usually specified. The shear in this weld is generally quite low. The joint between the stringer and the deck beam differs with the floor panel concept chosen.
stacked floors have a continuous fillet weld around the flange contact area and generally do not have web stiffening of the stringers.
If the top of the deck beam becomes inaccessible for maintenance, some operators will require seal plates to be welded between the deck beam and the floor plate. This is quite expensive. A typical joint is depicted in Figure 7.
The decision on the type of stringer joint should preferably be made prior to material ordering.
flush floors. Welding the floor between deck beams requires removal of the top-flange of the stringer near its end and perfect fit between the deck beams and floor. Deck beam prefabrication is also required.
4.4 Deck Beams Deck beams supporting the floor panels or providing direct support to major equipment are generally provided as HE 8001000 beams, though HL 1000 (400mm wide) or HX 1000 (450mm wide) are also used for heavier loads or greater spans. The major joint in the deck beam is that with the main structure. The joint configuration is strongly determined by the prefabrication concept and elevation of the flanges. It is different for the stacked and for the flush concept. Stacked Floor Concept Figures 8 and 9 illustrate the problems.
For the full stacked concept (Figure 9), where both transverse and longitudinal main beam are positioned lower, welding of the top flanges is straightforward. The lower flange, typically 40mm thick, can only be welded to the web, typically 20-25mm thick, if alignment of both flanges is ensured. The lower flange of the main structure should be at least 250mm underneath, to enable back welding of the root. For the less suitable partially stacked concept (Figure 8), where only the transverse main beam is positioned lower, connection for the top flange of the transverse deck beam is more difficult. Direct welding of the top flange of the deck beam to the web should be rejected. Options are shown in Figure 9 with detail (a) haunching and detail (b) slotting the top flange through the web. Again it is apparent that a decision on joint configuration is required prior to material ordering. Flush Floor Concept Detailing is dependent of the prefabrication policy. If the deck panel is prefabricated as an assembly of plate, stringer and deck beam, the detail shown in Figure 10a is the more appropriate.
To allow top flange welding a strip of the floor plate is fitted and welded last. If the deck panel is fabricated as an assembly of plate and stringer only, the detail Figure 10b will be the most feasible. 4.5 Horizontal Bracing In Section 3.5 the preference for the floor plate to act as horizontal bracing was indicated. If however separate bracing members are required, the elevation must be chosen carefully. The bracing members have to pass with sufficient clearance under the stringers, penetrate the web of the deck beams at sufficient distance from the lower flange. They also require good access for welding of the joint. These requirements generate the elevation and the maximum feasible diameter of the brace (Figure 11). Horizontal bracing can easily clash with vertical piping and major hatches. Assembly of the braces is generally quite cumbersome. 5. CONCLUDING SUMMARY
The topside lay-out was discussed, referring to API-RP2G, together with general aspects of interface control and weight control. Based on the type of substructure, jacket and GBS, the different types of topside structure were introduced and described. These types are:
integrated deck. module support frame. modules.
Floor concepts were described. Several aspects of the plate floor design were addressed.
6. REFERENCES [1] API-RP2G: Production facilities on offshore structures. American Petroleum Institute 1 ed. 1974. Presents the basic requirements. [2] API-RP2A: Recommended practice for planning, designing and constructing fixed platforms. American Petroleum Institute, 18th ed., 1989. The structural offshore code governs the majority of platforms. [3] DNV: Rules for the classification of steel ships. Part 5, Chapter 2.4.C, Permanent decks for wheel loading. Det Norske Veritas. Practical approach for economic floor plate design under static load. 7. ADDITIONAL READING 1.
M. Langseth & c.s.: Dropped objects, plugging capacity of steel plates. BOSS Conference 1988 Trondheim, pp 1001-1014. Floor and roof plate behaviour under accidental loading.
2.
D. v.d. Zee & A.G.J. Berkelder: Placid K12BP biggest Dutch production platform. IRO Journal, nr. 38, 1987, pp 3-9. Presents a recent example for a portal-framed topside.
3.
P. Gjerde et al: Design of steel deckstructures for deepwater multishaft gravity concrete platform. 9th. OMAE conference Houston 1990, paper 90-335. Most recent presentation on GBS topside structure.
4.
P. Dubas & c.s.: Behaviour and design of steel plated structures, IABSE Surveys S 31/1985, August 1985, pp 1744.
Good background to theory of plated structures.
A.11 - Superstructures II OBJECTIVE/SCOPE To elaborate on structural steel concepts for integrated decks, module support frames, and modules. To show principles and methods of construction (from yard to offshore site). PRE-REQUISITES Lectures 1A & 1B: Steel Construction Lecture 2.4: Steel Grades and Qualities Lecture 2.5: Selection of Steel Quality Lectures 3.1: General Fabrication of Steel Structures Lecture 6.3: Elastic Instability Modes Lecture 7.6: Built-up Columns Lectures 8.4: Plate Girder Behaviour and Design Lecture 11.2: Welded Connections Lecture 12.2: Advanced Introduction to Fatigue Lecture 15A: Offshore Structures SUMMARY Structural systems for each type of topside structure are introduced, i.e. truss, portal frame, box girder, and stressed skin. Some special topics of design are addressed and the different construction phases are presented in more detail, i.e.: 1. 2. 3. 4. 5. 6. 7. 8.
fabrication weighing load out sea transport offshore installation especially deckmating module installation hook-up commissioning.
A brief discussion on inspection and repair and on platform removal concludes this lecture. 1. INTRODUCTION This lecture deals with the structural design of jacket-based offshore deck structures, following the introduction in Lecture 15A.10. Heavy decks, over 10.000 tons, are provided with a module support frame onto which a number of modules are placed, see Lecture 15A.1, Figs. 4 and 5. Smaller decks, such as those located in the southern North Sea, are nowadays installed
complete with all equipment in one lift to minimize offshore hook-up. Most of this lecture refers to this type of integrated deck as described in Lecture 15A.10. The selection of the concept for the structural deck is made in close cooperation with the other disciplines. For the design of the deck structure, the in-place condition has to be considered, together with the various previous stages such as fabrication, load-out, transport and installation. A structural system for a deck structure comprises several of the following elements: Floors (steel plate or grating) Deck stringer (H beams, bulbs or troughs) Horizontal bracing Deck beams
} } Discussed in } Lecture 15A.10 }
Primary girders Vertical trusses or bracing Deck legs
} } Discussed in } this lecture
2. MAIN STRUCTURE DESIGN 2.1 Introduction Some major topics in topside structural design are reviewed below. 2.2 Main Structure-Portal Frame Design A portal frame design has been used in recent major projects in the Dutch sector such as Amoco P15, Placid K12 [5] and Penzoil L8. The main girder/column joint, as shown in Figure 1, is very important in determining the height. It is most practical to position the longitudinal and transverse main girder flanges at the same elevation.
Haunching of the transverse main girder , which is more lightly loaded-in-plane, however is not an option as these girders become highly loaded during transport. The severe restraint of welding a tubular in a diaphragm requires the selection of TTP steel for the column section. Due to the high importance of the diaphragm plates in the overall integrity of the structure and the welding constraints on the web plates in between, TTP-steel is chosen also for the diaphragm. Another option is to weld the girders directly onto the unstiffened can section of the column. The assessment of ultimate resistance as well as fatigue strength has been the subject of recent research (see Lecture 15A.12). Further improvement of the theoretical and experimental background is required. For lighter loaded truss structures, this non-stiffened type of joint has been used successfully. A third solution is to weld the girders directly to the can section of the column, which is internally stiffened by rings. Its most severe disadvantage is the difficulty of inspecting the column interior. The disadvantage of both direct girder-column joints is that the girder sizing is governed by the very high moments at the column/beam transition point.
Cast steel nodes form an alternative to the welded designs. Member selection for portal frame structures with increasing section moduli usually includes:
300 mm wide rolled beam. 400 mm wide rolled beam. 450 mm /460 mm wide rolled beam. castellated beams fabricated from rolled beams, giving a height 1,5 times the original beam height. built-up girders fabricated from rolled beam T-sections with a web plate welded-in-between. plate girder.
The plate girder of course provides the greatest flexibility for design, material selection and procurement, though its cost per tonne is approximately twice that of a rolled beam. 2.3 Main Structure-Truss Design Most offshore structures of moderate size have been provided with a truss-type structure. Typically such trusses consist of rolled beams as chords and tubulars as diagonals. Truss design requires several choices which affect the structural efficiency and have impact on other disciplines:
number and configuration of braces falling or rising braces intermediate load carrying of chords presence of external moments on joints braces: tubulars or H-rolled sections chords: rolled section or plate girders truss joints: locally reinforced chord or prefabricated node section.
Figure 2 shows different arrangements of braces (basically N or W-type) obtained by variation of the number of nodes. It should be kept in mind that all diagonals and verticals form obstructions for piping and cable routings of all kinds.
For the transverse trusses, transparency is even more important, especially near the well area. The number of members required should therefore be reduced to a minimum. Providing a W-truss with light verticals should be evaluated against choosing a heavier chord section. If a joint, e.g. at the top deck, is subject to severe moments due to lifting, ventstack, or crane pedestal for example, much of the bracing stress would result from unintended bending. Generally the deck leg restraint creates a similar problem in the lower deck. An evaluation should yield a preferred location therefore for the node of the end brace. The truss deflects under its vertical load which leads to restraint of the chord in the column and to bending of the chord. Both effects can quite severely effect the efficiency. The chord section should be kept compact therefore and not given too much height. Tubulars (circular, square or rectangular) or rolled sections can be chosen for the braces. The choice depends primarily on the loads and the chord width. A chord width of 300mm can accommodate a 10 in. brace only. Thus a wider chord flange is preferred. 2.4 Main Structure-Stressed Skin Design A third major structural option is the stressed skin concept, where full height plate walls take the function of the truss or the frame. Modules for living quarters are frequently built to this concept. Other types of modules have not been built with stressed skin since the obstruction they cause during construction is severe. For smaller stressed skin modules, trapezoid corrugated plate can be used to provide a wall in a frame of square hollow sections. For bigger modules, flat plate stiffened with through-stiffeners is used for the walls. The detailed design can only be made with a clear plan for assembling the module which shows the panels that must be prefabricated. 2.5 Non-Load Bearing Walls Blast or fire walls are provided in offshore platforms. Due to their function full welding to the main structure is often unavoidable, see Figure 3a.
Special attention is required concerning:
the capability of the walls to comply with the deformation of the main structure during load-out, sea transport, lifting and in-service. the strength of welds to the main structure being stronger than the plate to avoid rupture and potential crack initiation of the main structure.
One solution is to provide a flexible detail, see Figure 3b and 3c, with stiffeners falling short. 2.6 Crane Pedestals Crane pedestal, are discussed briefly below. It is structurally economical to put the crane pedestal on top of a main column. For a truss type the main structure will be close to the platform periphery so a moderate length of crane boom is sufficient. For a portal frame type with columns closer to the outer periphery, the pedestal requires a special column in order to avoid using a crane with large boom length. Figure 4 depicts such a solution.
The functions of the main structure with respect to the crane pedestal are:
to provide torsional support preferable at top deck level to provide lateral restraint at top deck level to provide lateral restraint at the lower end of the pedestal to provide vertical support, preferably at the lower end of the pedestal.
Bending restraint by deck beams and/or main structure girders is not required and should be reduced where possible. Torsion caused by slewing of the crane should preferably be resisted by the floor plate, the stiffest element. It has become practice to include the tapered top section of the pedestal in the supply package of the crane. The top section contains the large flange for the slewing bearing. Fatigue due to crane operations is a design criterion and requires careful detailing of the pedestal and the adjoining structure.
3. ANALYSIS OF DECK STRUCTURES 3.1 Introduction Although the analysis of deck structures is a standard task, several aspects require special attention:
Plate girder design Strength of joints Strength of the floor plate Lifting points Modelling of floor plates Support of modules.
3.2 Plate Girder Design Design of plate girders requires selection of many dimensional variables and of approaches for assessing load-carrying resistance. Lectures 8.4 deal in more detail with plate girder design. Web buckling due to bending, normal force and shear restricts the slenderness of the web which is expressed as the height of the web (h) divided by the web thickness (t). API-RP2A [2] refers to the AISC manual [3] which gives the figures below for material with yield-stress of 355 MPa: Allowable bending stress
0,66 Fy
0,60 Fy
Ratio web height h to thickness t
90
138
Ratio flange width b to thickness t
18
27
Instead of the above approach, more recent research, [3] and [6], allows use of the post-buckling strength. The depth/thickness limits given above do not then apply. 3.3 Strength of Joints The most important joints in a topside steel structure are:
the ring stiffened joint between rolled beams or plate girders with a circular column. the non-stiffened joint between rolled beams or plate girders with a circular column. the tubular brace joint to single web beams. the non-overlapped tubular joint.
These joints are discussed in Lecture 15A.12. 3.4 Lifting Points The effect of lifting points on deck design is considerable. For example the local forces that act on the lifting points (Figure 5) have to be transmitted safely through to the deck structure.
There are two types of lifting points, trunnions and padeye, Figure 6.
Trunnions, though favourable from other points of view, see Section 4, can generate considerable offset of the sling force with respect to the topdeck system points. Significant bending is generated which is transferred to the topdeck girders to the extent that they contribute to joint stiffness. It is most efficient to leave these bending moments in the column, by providing stiff columns. Padeyes generally provide a good opportunity to minimize or eliminate offset, as far as they can be situated on top of the column. The requirement of recessed padeyes (recessed padeyes are those which are positioned between the top and bottom flange elevation) or the presence of other structures on the top deck can lead to very eccentric positioning and resulting heavy moments. For this reason the lifting concept must be developed in the concept phase of the structural development. API-RP2A [1] requires larger load factors to be used for members direct-loaded by padeyes or trunnions. 3.5 Modelling of Floor Plates There are two points of major interest:
representation of the floorplate in the structural model true elevation
There are several ways to model the plate. The most direct is to choose a computer-program which allows selection of plate elements. A second option is to define representative members which model the plate stiffness by diagonals. The deck plate is often positioned in the model at the elevation of the centre line, i.e. the mid height of the main structure girders, in order to save nodes in the model. It should however be recognised that this "error" of elevation, amounting to 0,5 - 1m, can affect the results. A separate evaluation should then be performed on the effect to this deliberate "error" at least at some critical points. 3.6 Support of Modules Modules and deck structures interact structurally. API-RP2A [1] requires that modules are modelled as elastic structures for the analysis of the supporting deck. In the 1970's major difficulties arose in the decks for concrete gravity structures, because modules were represented as a set of loads for the different load cases, acting at the support points, and neglecting structural interaction. The basic phenomenon of this interaction is that the distribution of the support reactions of the module is quite unequal and varies with the load case. Dimensional control of the module as well as the support, with corrective measures, further provide control over the module - deck interaction. Some modules, such as living quarter modules, gas compressor and injection modules, are often placed on anti-vibration pads in order to isolate them from vibrations. 4. CONSTRUCTION 4.1 Introduction In Lecture 15A.1 the principal aspects of construction of offshore structures and their major equipment was introduced. For topsides more specific aspects are discussed below. 4.2 Fabrication 4.2.1 Operations The design should allow efficient prefabrication of major sections. Prefabrication will avoid congestion in one working area and it speeds up the whole construction process.
Prefabrication and assembly shall properly incorporate the aspects of installation of major and smaller mechanical equipment, as well as outfitting with piping, electrical and instrument cables and lines. It should be recognized that major mechanical and electrical equipment is often not available at the start of assembly and must be brought in during fabrication. 4.2.2 Design aspects Since the overhead space is well covered by extensive piping routes as well as cable trays during construction, "late" structural work should preferably not be positioned overhead in that underfloor area. Fabrication of offshore steel structures is principally assembly by welding. The prefabrication concept and joint detailing should maximize welding productivity with many horizontal welds preferably made using SMAW technology. Support to the topside during construction should be well controlled to avoid settlement and to keep within construction tolerances. Special consideration should be given to the selection of materials suitable for the fabrication. Where thick-walled elements are involved requiring Post Weld Heat Treatment (PWHT), the design should position such welding and the PWHT in the prefabrication phase. 4.3 Weight Engineering The topside must be kept under strict weight control, as explained in Lecture 15A.10. To that end the topside is usually weighed prior to load out. The basic design of a weighing system usually consists of a set of hydraulic jacks with electrical load cells on top, installed between the topside and the shop floor. The accuracy of such systems is typically 0,5-1%. Accuracy is necessary in order to check the actual position of the centre of gravity. Knowledge of the position is vital for the installation. The system for support of the topside should be similar to the anticipated method of load out. 4.4 Load Out 4.4.1 Operations The load out usually combines two operations:
moving the topside from the fabrication hall to the nearby quay. moving the topside from the quay onto the barge.
The short journey on land can be complicated when the track is not flat or curves have to be taken. The most preferred option for load out is therefore to use a platform trailer with individual suspended wheels, see Figure 7 and Slide 1.
Slide 1 : General arrangement of a load out through skidding
The trailer drives from the quay over a rocker flap resting on the quay and the barge and then slowly onto the barge. The barge is kept in right trim by ballast pumping. When it reached the right position, the topside is set down on the beam grid of the sea fastening. 4.4.2 Design aspects load out When using platform trailers the lower deck should be designed to meet three basic load-out requirements:
the bottom flange plates of the transverse beams should all be in one plane. the distance of transverse beams should not exceed approximately 7 m. the lower deck should be able to take an upward reaction typically in the range of 50-60 kN/m2 of ground area.
A uniform distribution of loads is assumed for platform trailers. Skid systems which are not provided with a proper load sharing system will lead to a non-uniform load distribution. Design for load-out requires coordination with sea fastening design. 4.5 Sea Transport and Sea Fastening 4.5.1 Operations Sea transport is a very critical operation, especially for topsides (see Slide 2).
Slide 2 : Seafastening of 105MN Brent C topside After completion of the load out and full fastening to the barge, the barge is ballasted to its target draft and cleared for the transport. The barge is towed by one or two tugs to the offshore location. There the barge is positioned close alongside the crane vessel.
Prior to lifting, the sea fastening is cut free. Planning the sea transport contains several steps:
identification of critical clearances, e.g. (harbour depth, width of bridges or locks, etc inshore) barge selection (a.o. stability, dynamic behaviour, location of bulkends). evaluation of sea route (weather, length of tow). assessment of barge motions in sea state. development of a sea fastening concept. assessment of deck/module integrity. assessment of barge integrity.
There is also the option with some crane vessels to transport the top side on board. Usually an extra take over is required as the draft of the crane vessel exceeds the depth at the fabricator's quay. The advantage however is that sea fastening requires less effort. Furthermore, the offshore operation is simpler and quicker, as the most critical and weather sensitive operation - lift off the barge- is avoided. 4.5.2 Design aspects of sea transport and sea fastening Several elements of the structure are dominated by the load condition during transport, see Lecture 15A.1. All equipment in or on the topside is also subject to heavy loads, e.g. control panels, generator skids, platform crane, during transport. Internal bracing of a topside for transport is not favoured since it creates obstacles and risk of damage or fire to cables, instruments, piping and equipment during subsequent removal. External bracing is also not without problems. The width of the topside requires an extra wide barge. It is difficult to find "strong" points on the topside exterior. The basic concept is therefore to fix the topside to the barge by its columns only. The designer should be aware that the bending stiffness of the topside often exceeds that of the barge. Considerable "composite" action can result when the barge deflects in heavy head-on seas. It is very important for any sea fastening concept to consider aspects of de-seafastening, i.e. cutting free, prior to lift off, and the need to remain safe in a moderate sea state. De-seafastening should not require any handling by cranes. Braces cut loose at one end should therefore remain stable and safe while fixed at one end only. Design of the sea fastening should not require any welding in the column joint, since the topside would not then be ready for immediate set down onto the jacket. When the tow is more than one or two days long, fatigue may have to be considered on critical nodes. 4.6 Installation 4.6.1 Operations Installation on the substructure can be:
deck mating with a deep submerged floating GBS (Slide 3) lifting onto an already installed jacket (Slide 4).
Slide 3 : Deckmating of the 500MN Gullfaks-C topside
Slide 4 : Installation of 60MN K12-BP topside by floating crane Deck mating is a floating operation in a sheltered location, e.g. a Norwegian fjord or Scottish loch. Deck mating requires that the deck is temporarily supported with the final supports free. This requirement creates a very awkward load situation for the deck structure. Lifting is the usual installation method for jacket-based topsides. During development of a platform concept, the lift strategy should be defined as part of the overall construction strategy. The lifting capacity of crane vessels is defined by hook-load and reach. The required reach is determined mainly by the width of the topside and/or the transport barge. The major steps are:
review of the weight report. assessment of "critical" elevations.
assessment of feasible crane vessels. development of a lift concept. preliminary sizing of slings, shackles, trunnions, etc. concept design of guides and bumpers. analysis of deck or module structure for lift condition.
4.6.2 Design aspects of installation by lifting The lift concept consists of several elements:
the single or dual crane lift the sling configuration choice of topside pick-up points the necessity (or not) for spreader bars or even spreader frames the single, double or paired slings the choice of padeyes, or trunnions.
Crane vessels were listed in Lecture 15A.1. Slings are available up to over 400mm nominal diameter with safe working loads of 20-25 MN. A basic element in all elevations is the inevitable tolerance in sling length which leads to an unequal distribution of sling forces (typically 25%-75%) in a four sling lift. The unequal sling forces lead to significant stresses in the module (see Figure 8).
The use of spreader bars leads to a fully balanced lift without distorting the module. However the spreader bar is quite expensive and usually leads to a requirement for a higher hook elevation. The use of a spreader frame should only be considered in exceptional cases and does not prevent module distortion. The padeye/shackle option is limited by the safe working load (maximum 10MN) of the biggest shackle. The trunnion can accommodate higher loads. 4.7 Hook up Hook up is the completion of all joints and connections after installation. For economic reasons, the overall construction strategy should keep hook up work to a minimum. Critical hook up work is the work required immediately to secure the object in order to survive the next storm. 4.8 Commissioning Commissioning is not relevant to the structural design. 4.9 Inspection Maintenance and Repair (IMR) These activities are a major source of operational expenditure, OPEX, as introduced in Lecture 15A.1. Some requirements are:
inspection of the primary structure is a statutory, fully planned activity. inspection is only possible when proper access to the area or joint is provided. gaining access is costly and requires space to be left behind equipment. minimum provisions, e.g. small clamps under the deck, greatly speeds up scaffolding. crack growth through fatigue is slow. A crack is usually detectable before one quarter of its life is passed. dirt accumulation promotes corrosion damage. maximum use should be made of the results of inspection. Evaluation should lead to modification of the inspection programme where appropriate.
4.10 Removal Removal requirements are different from country to country. In some depths of water full removal is required in some countries from the mudline upward. Elsewhere only the structure 75 m or more above the mudline must be removed. Extensive engineering of removal is required to achieve a safe and effective operation. In the Gulf of Mexico removed structures are dumped in the form of reefs. It is very difficult and inefficient at present to include conceptual removal engineering in the design phase. When re-use of the facility is planned, then removal engineering should be developed early in the design. 5. CONCLUDING SUMMARY
Structural systems for each type of topside structure were introduced, i.e. truss, portal, box girder, and stressed skin systems. In the section on design some topics were addressed in more detail. In the section on construction the different phases were presented in more detail, i.e.
i. fabrication ii. weighing iii. load out iv. sea transport v. offshore installation especially deckmating
vi. module installation vii. hook-up viii. commissioning
A brief discussion on inspection and repair and on platform removal concluded the lecture.
6. REFERENCES [1] API-RP2A: Recommended practice for planning, designing and constructing fixed platforms. American Petroleum Institute, 18th ed., 1989. The structural offshore code, governs the majority of platforms. [2] AISC: Allowable stress design manual (ASD). 9th ed., American Institute of Steel Construction, 1989. Widely used structural code for topsides. [3] API-Bulletin 2V: Bulletin on design of flat plate structures. American Petroleum Institute, 1st ed., 1987. Valuable specialist addendum to API-RP2A. [4] API-Bulletin 2U: Bulletin on stability design of cylindrical shells. American Petroleum Institute, 1st ed., 1987. Valuable specialist addendum to API-RP2A. [5] D.v.d. Zee & A.G.J. Berkelder: Placid K12BP biggest Dutch production platform. IRO Journal, nr. 38, 1987, pp 3-9. Presents a recent example for a portal framed topside. [6] R. Narayanan: Plated structures/Stability and Strength. Applied Science Publishers, London, 1983. Good designers guide to plated structures design. [7] ANON: Gullfaks C platform deckmating. Ocean Industry, April 1989, pp 24. Good description of the actual mating of deck to GBS. [8] A.G.J. Berkelder: Seafastening 105 MN Brent C deck. Bouwen met Staal, nr.24 1979. Presentation of seafastening design for GBS topside.
A.12: Connections in Offshore Deck Structures OBJECTIVE/SCOPE To outline and explain the best methods for forming structural connections in offshore deck structures; to discuss the importance of a proper choice of connection type to achieve both the required strength and stiffness, and ease of fabrication. PREREQUISITES Lectures 11.2: Welded Connections Lectures 11.4: Analysis of Connections Lectures 13: Tubular Structures Lectures 15A: Structural Systems: Offshore RELATED LECTURES (covering specific items in greater detail) Lecture 2.4: Steel Grades and Qualities Lecture 2.5: Selection of Steel Quality Lectures 3.6: Inspection/Quality Assurance Lecture 4A.5: Corrosion Protection in Offshore Structures and Sheet Piling Lecture 11.5: Simple Connections for Buildings Lecture 12.2: Advanced Introduction to Fatigue Lectures 12.4: Fatigue Behaviour of Hollow Section Joints SUMMARY Various forms of structural connections in steel offshore deck structures are discussed; these cover the connections between deck stringers and main beams, between main beams themselves, between main beams and deck legs, truss connections and connections between columns and beams. The importance of designing and dimensioning to minimise fabrication and maintenance is emphasised. 1. INTRODUCTION Large offshore deck structures have traditionally been built up using modular components, see Lectures 15A.10 and 15A.11; a module support frame is built on top of the deck legs of the jacket structure on which the various modules are installed. With the high lifting capacities currently available, the topsides of light to medium offshore structures can now be installed in one lift. This development has had a considerable influence on the fabrication and design of deck structures, and has resulted in heavier modules, constructed of larger and heavier members, with consequences for the connections. Another aspect influencing fabrication, and thus the design, was the development of cleaner steels, with modified chemical compositions and good through-thickness properties. This so-called TTP steel (i.e. steel with through-thickness properties, see Lecture 2.4) has a low sulphur content to avoid lamellar tearing. Furthermore, if the carbon and carbon equivalent (CEV)
is low, the preheat temperature of the steel can be lowered, resulting in easier welding (without preheating) which again influences the connection design. The increase in lifting capacity, and the exploration for gas and oil in deeper water, have both resulted in larger structures, and have stimulated the use of higher strength steels, with yield strengths above 355 N/mm2. The joints have to be designed to withstand the various loading conditions (see Lectures 15A.2 and 15A.3) experienced during fabrication, load-out, transport, installation and the in-place condition (operation and storm). In order to allow redistribution of stresses it is important that the joints are stronger than the connected members; if this is not the case the joints themselves must have sufficient deformation/rotation capacity. The connection design should take account of all the aforementioned aspects and should be considered as an interactive procedure involving the choice of the structural layout, the fabrication sequence and the steel grades and qualities to be used. Other aspects such as inspection and corrosion protection requirements must also be considered. Since the fabrication costs are mainly governed by the costs of welding, the connections should be simple, and where possible, avoid the use of stiffeners. 2. CONNECTIONS IN OFFSHORE DECK MODULES The type of connections used in offshore deck modules depends directly on the type of structure involved:
truss types frame types stressed skin
As discussed in more detail in Lectures 15A.10 and 15A.11, the structural system for a deck includes several of the following elements:
floor (steel plate or grating) deck stringers (I-beams, bulb flats or troughs) deck beams main beams or girders (beams on main grid lines) vertical trusses or braces deck legs and columns
Depending on their function, loading, and availability of sections, these elements can be made of rolled I or H-sections, rolled circular or rectangular hollow sections, or welded sections; for the larger sizes, welded I or box plate girders, or welded tubular members are used. These elements have to be connected together; since the modules are generally fabricated under controlled conditions at the fabrication yard, welded connections are common practice. The main connection types are discussed more in detail below. Although it is common practice in offshore design to use the API-RP2A [1] or the AISC rules [2], the basic joint behaviour is discussed in this lecture without reference to the safety factors to be used. 3. CONNECTIONS BETWEEN DECK STRINGERS AND BEAMS The deck floor structure can be designed as a floor plate with stringers, or as an orthotropic plate. The floor plate with stringers is the most common type as it gives design flexibility for later changes (local loads, deck penetrations, etc). Orthotropic plate structures, are generally used in helidecks, see Lectures 15A.10 and 15A.11. The use of stacked stringers, as shown in Figure 1, facilitates fabrication and is, therefore to be preferred to the use of continuous connections, as shown in Figure 2.
For ease of fabrication, stiffeners should be avoided if possible. This means that the vertical loads have to be transmitted by the webs, as shown in Figure 1, over a length ls for the stringer, and lb for the deck beam; web crippling failure is also possible and should be checked. These are a common details which are dealt with in Eurocode 3 [3] and other codes. For the continuous connections, shown in Figure 2, the moment is assumed to be transferred by the flange connection and the shear by the web connection. The type of full penetration weld at the top flange for continuous connections depends on the fabrication sequence and should be decided by the fabricator. The bottom flange and web can generally be connected by fillet welds. A full penetration weld of the flange, without a 'mouse hole', is preferred because of corrosion protection although this results in a small weld defect at the neck between flange and web. However, even under fatigue loading such a defect can be accepted [4] the same is also valid for static loading. Only in cases where very high strength steel (fy > 500N/mm2) is used and a high yield to ultimate stress ratio, e.g. fy/fu > 0,9 occurs should this detail be evaluated rigorously. Since all loading cases are not always checked, the welds have to be designed to have at least the same strength as the connected parts, i.e. as the flange or web. It should be recognized that the shear stress distribution (Figure 2) for a detail with a 'mouse hole' is more severe than that without a 'mouse hole'. Special attention should be given to the unsupported upper side of the web in Figure 2b, as local buckling may be a problem, see Lecture 6.2 and [5]. 4. CONNECTIONS BETWEEN INTERMEDIATE AND MAIN DECK BEAMS The connection between the deck beams is most convenient if these beams have the same height, as shown in Figure 3b. Here the flanges are connected with full penetration welds, and the web by fillet welds or a full penetration weld depending on the thickness. Tolerance control is necessary to avoid differences at the deck floor level, between stringers. The shear loads are generally too high to allow a single or double sided notch as shown in Figure 3b since this results in a higher shear stress, see Figure 2. In case of equal heights, no TTP requirements are necessary for the beams. For the connection of beams with unequal heights, however, as shown in Figure 3a, the web of the main beam should have a TTP quality due to the loads being transferred through the web thickness. Furthermore, to satisfy the requirements for avoidance of cold cracking, etc., either the flange thickness of the intermediate beam should be less than 1,5 times the web thickness of the main beam, or the material should have a low carbon content (see Lecture 2.5).
As an expensive alternative solution, a plate connecting the flanges can be slotted through the web, as shown in Figure 4. Haunched alternatives are given in Lectures 15A.10 and 15A.11.
All welds should be designed to have the strength of the connected parts.
As a consequence the connection is as strong as the member; only in case of large 'mouse holes' the shear stress and possible local buckling of the unsupported web part [5] have to be checked. 5. BEAM TO DECK LEG CONNECTIONS The main beams, either rolled H sections or plate girders, must be connected to the deck legs, which are normally fabricated tubular members. For a frame type structure, this connection should be rigid and capable of transmitting the yield moment resistance of the connected beams. These connections, or nodes, are generally prefabricated, consisting of a tubular "can" with surrounding "diamond" (diaphragm) plates for the connection with the beams, as shown in Figure 5. This type of connection requires special material specifications and special welding procedures.
Stiffened Connections
The shear loads are transferred by the connection of the web plates to the tube walls. The moment is transmitted by the diamond plate in combination with an effective ring width of the tubular "can". The design resistance, for factored loading, is normally checked with the experimental Kamba formula, which is simplified by Kurobane [6] as follows:
NRd = where: NRd is the design resistance for the flange for factored loading fy is the yield stress of deck leg "can" b1 is the flange width of deck beam do is the outer diameter of tube to is the wall thickness of the deck leg "can" ts is the thickness of ring plate hs is the smallest width of the ring plate
bf¢ = Validity ranges:
The axial force in the flange N, is derived from N = Mcw/(h1 - t1) (see Figure 5). This formula is based on the test results for a ring-stiffened joint with two opposite loads; more detailed research is currently being carried out [7]. In the case of multiplanar loading, for four loads acting in the same sense, the joint strength will be greater. However, if the two loads in one direction are tensile and the two in the direction perpendicular to that are compressive, the joint strength may be decreased. Reference [7] reports that this decrease was found to be a maximum of 30%. Furthermore, if the deck leg is loaded by axial compressive stress amounting to 60% of the yield value, the strength of the connection has to be reduced by 20%. Non-Stiffened Connections For truss type frames, the beam to deck leg connection has to transfer mainly axial loading and an unstiffened connection, as shown in Figure 6, could be used; this is, however, not yet common practice. If sufficient deformation capacity exists, the secondary bending moments can be neglected for static loading. If fatigue loading has to be checked, however, care should be taken with these secondary bending moments, because the stress concentration factors at the flange to tubular connection are rather high. For practical cases these stress concentration factors can be in the order of 10 for
, see [8].
The static design resistance for factored load of the unstiffened connection is determined by the strength of the flange to tube connection, which can be based on Togo's ring model, see Lecture 13.2. The design resistance for flange loads in one direction (X-joint loading) is given by Eurocode 3 [3] and [9].
NRd = where: NRd is the design strength for the flange for factored loading fyo is the yield stress of joint "can" to is the wall thickness of joint "can" b is the flange width b1 to "can" diameter do ratio kp is the influence function for additional stress in the chord. Validity ranges:
0,4 £ b £ 1,0 For bending moments in-plane, the axial force N is derived from N = Mcw/(h1 -t1) as shown in Figure 5. For an axial loading the flange connections can interact such that the connection strength (I to tubular) is not twice the strength of one flange connection but:
NRd . Consequently the beam to deck leg connection has to be checked for:
NSd £ NRd Mipsd £ NR.d (h1 - t1) Currently, for multi-planar loading with loads and moments acting in the opposite sense, the same 30% reduction in joint strength as before is recommended, although initial investigations indicate that this may be conservative [10]. No reduction has to be applied if the loads are acting in the same sense. 6. CONNECTIONS BETWEEN BEAMS AND COLUMNS Columns between decks are necessary where external surfaces of the modules are clad, or where cantilevers or laydown areas are provided. The connection with the deck beams can be flexible in the longitudinal direction if these columns have only to withstand lateral loading. If, however, they are used to transfer loadings from cantilevers to both decks, the connections should have the same strength as the column or they should have sufficient deformation capacity. Figure 7 shows a possible full strength detail for columns connected to a plate girder, with possible connecting side beams and an extended cantilever. Here the web of the plate girder is ended before the flange to allow a tubular section to be welded between the flanges. I-beam sections, even with different depths, can be easily welded to this tubular section, and the columns can be welded to the flanges.
The "joint can" should have about the same diameter and thickness as the column. In Figure 7 longitudinal beams and a cantilever beam are also connected to this can. The bending moment resistance is here determined by the connection of the bottom flange to the tubular can, similarly as discussed in Section 5. 7. TRUSS CONNECTIONS Since the chords of the trusses are part of the deck floors, they are almost always made from an I or H-section; in exceptional cases, welded box sections are used. The diagonals are tubular, rectangular hollow sections, or H sections; all have their advantages and disadvantages with regard to material costs, maintenance and fabrication. Where these diagonals are connected to an I section chord, the chord should be stiffened to obtain a full strength connection; it should be kept in mind that intermediate beams may have to be connected to the chord at this location. The connection should be designed in such a way that fabrication and inspection will be easily possible. Figure 8 shows some connection details for lightly loaded trusses.
These connections generally do not develop a strength equal to or larger than the yield strength of the diagonals. Consequently the connection should have sufficient deformation capacity. However experimental evidence is only available for the connection according to Figure 8a. From a fabrication point of view, the connections with a gap between the braces are preferred. However the connections with overlapped braces as shown in Figures 8c and 8d are stronger. The connection strength may be governed by various criteria, depending on the geometry, i.e: a. chord web strength b. chord web crippling under a compression brace c. chord web shear between the diagonals of a gap joint d. chord web buckling e. brace (diagonal) effective width f. brace shear failure at the flange connection g. weld failure (to be avoided by full strength welds) h. lamellar tearing (to be avoided by TTP material for the flange).
For connections according to Figure 8a, Eurocode 3 [3] provides design strength formulae which can be used in a modified way for the connections of Figure 8b to 8d. Within the scope of this lecture it is not possible to deal with all connections in detail, however one example is given for a connection between tubular braces and an I-section chord as shown in Figure 9.
The strength of the connection for axial loads at the chord intersection (cross-section A) is governed by the effective width area:
Aeff.c = 2 (bm1 tp + bm2 tw) For the brace intersection the effective width is given by: Aeff.b = 2 (be1 + be2) tp The strength of the connection is thus given by: N2sin q2 = Aeff.c fyo and N2sin q2 = Aeff.b fyo The effective widths bm1, bm2, be1 and be2 are given in Eurocode 3 (6.6.8 and Appendix K, Table K.8.2). As an additional check the chord cross-section between the braces has to be checked for shear and shear in combination with axial loading and bending moments, see Table K.8.2 of Eurocode 3. The chord and braces have furthermore to satisfy the limits for d/t and h/t to avoid local buckling. Weld failure and lamellar tearing should always be avoided by choosing full strength welds and proper selection of the steel grade and quality. In these cases where the joint strength is lower than the brace member strength, sufficient rotation capacity should be available if the bending moments are neglected. Since it is difficult to show that sufficient deformation capacity exists due to a lack of research evidence, either the bending moments have to be incorporated in the strength assessment or the joint is stiffened to such an extent that the joint strength is larger than the brace member strength, e.g. as shown in Figure 10.
8. SPECIAL CONNECTIONS The previous sections dealt with the most common types of connection; however, depending on the platform layout, other types of connections may be necessary. Figure 11, for example, shows the connection between two panels of stiffened plates. Here both panels are made by (semi) automatic welding processes. Allowance is made for welding tolerances by welding the ends of the stringers after the fitting together of the panels. This procedure can be used for modules which are designed using the stressed skin method.
Special provisions are necessary for lifting the modules; padeyes or trunnions, for example, can be provided for this purpose, as shown in Figure 12; nowadays these devices are sometimes made of cast steel. It is important that these lifting devices are designed in such a way that they can be connected to the deck structure at a later stage when the precise location of the centre of gravity of the module, and the lifting method, are known.
Strength of padeyes is often assessed by means of "Lloyds" formulae, which are presented in the SWL (safe working load) format. The SWL is the least of the following values of Ni: N1 = 0,60 (a tL + 2 b tE) fy N2 = 1,08 (c tL + (D - d) tE) fy N3 = 0,87 d (tL + 2 tE) fy where the following limitations apply:
1,0 £
and if
£ 8,0 £ 1,0
then put tL + 2 tE = d in the above formulae.
tE not to exceed tL/2 dHOLE/dPIN £ 1,05
Tubular connections are not dealt with in this lecture since these are discussed in more detail in Lectures 13.2 and 13.3. For offshore deck structures, built up from stiffened plate panels, reference should be made to Lectures 8.3 and 8.4. For living quarters and helicopter decks, use can be made of the information in the previous sections. 9. CONCLUDING SUMMARY
The optimal design of offshore deck structures depends, to a large extent, on the coordination between the specialists for the various disciplines; for the layout, coordination between structural, mechanical, electrical, fabrication, load out and installation engineers is important. The structural designer has to consider the fabrication sequence; the conditions for welding and inspection (e.g. can it be welded properly?); the consequences of the choice of material grade and quality on the fabrication; and the various load conditions. In general, most connections can be designed with the basic formulae used for tubular connections and beam-tocolumn connections. Background information is given in [1, 2, 9, 11 - 15]. Recently a study has been carried out to investigate the use of RHS in deck structures [16]. This shows that the use of RHS, instead of beams, for deck trusses can be economical. However, due to restrictions in available sizes, economical solutions are mainly found for smaller platform sizes and for secondary steelwork such as staircase towers, access platforms and equipment supports.
10. REFERENCES [1] API-RP2A "Recommended Practice for Planning, Designing and Constructing Fixed Offshore Platforms". American Petroleum Institute, 18th Edition, 1989 [2] AISC "Specification for the Design, Fabrication and Erection of Structural Steel for Buildings". American Institute of Steel Construction, Chicago, 1980 [3] Eurocode 3: "Design of Steel Structures": ENV 1993-1-1: Part 1.1, General Rules and Rules for Buildings, CEN, 1992.
[4] Dijkstra, O.D., Wardenier, J. "The Fatigue Behaviour of Welded Splices with and without Mouseholes in IPE 400 and HEM 320 beams". Paper 14 Int. Conference Weld Failures, November 1988, London [5] Lindner, F. and Gietzeit, R. "Zur Tragfähigkeit ausgeklinkter Träger" Stahlbauwz. 1985. [6] Kurobane, Y. "New Developments and Practices in Tubular Joint Design". IIW doc. XV-488-81/XIII-1004-81, International Institute of Welding, 1981 [7] Rink, H.D., Wardenier, J. and Winkel, G.D. de "Numerical Investigation into the Static Strength of Stiffened I-Beam to Column Connections". Proceedings International Symposium on Tubular Structures, Delft, June 1991. Delft University Press. [8] Hertogs, A.A., Puthli, R.S. and Wardenier, J. "Stress Concentration Factors in Plate to Tube Connections". Proceedings ASME/OMAE Conference, March 1989, Vol. II, pp. 719-727 [9] Wardenier, J. "Hollow Section Joints". Delft University Press, Delft, 1982 [10] Broek, T.J. van der, Puthli, R.S. and Wardenier, J. "The Influence of Multiplanar Loading on the Strength and Stiffness of Plate to Tubular Column Connections". Proceedings International Conference "Welded Structures 90", London, UK, November 1990 [11] DNV "Rules for the Design, Construction and Inspection of Fixed offshore Structures" 1977 (with corrections 1982) [12] Lloyd's Register "Rules and Regulations for the Classification of Fixed Offshore Installation". London, July 1988 [13] IIW-XV-E "Design Recommendations for Hollow Section Joints - Predominantly Statically Loaded - 2nd edition". 1989, IIW doc XV-701-89 [14] UEG "Design of Tubular Joints for Offshore Structures". UEG, London, 1985 (3 volumes) [15] Voss, R.P. "Lasteinleitung in geschweisste Vollwandträger aus Stahl im Hinblick auf die Bemessung von Lagersteifen". Ph.D-Thesis, TU Berlin D83, 1983 [16] Guy, H.D. "Structural Hollow Sections for Topside Constructions". Steel Construction Today, 1990, 4 11. ADDITIONAL READING 1. 2. 3.
Marshall, P.W. "Design of Welded Tubular Connections: Basis and Use of AWS Provisions". Elsvier, 1991 Schaap, D., Pal, A.H.M. v.d., Vries, A. de., Dague. D. and Wardenier,J. "The Design of Amoco's 'Rijn' Production Platform". Proceedings of the International Conference on Steel and Aluminium Structures, Cardiff, UK, 8-10 July 1987, Vol. Steel Structures Paul, J.C., Valk, C.A.C. v.d., and Wardenier, J. "The Static Strength of Circular Multiplanar X-joints". Proceedings of the third IIW International Symposium on Tubular Structures, Lappeenranta, September 1989
