Lecture 15A.1
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.
None.
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 aspects is presented. Finally terms are introduced through a glossary.
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.
Page 1 of 32
Lecture 15A.1
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.
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Lecture 15A.1
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Lecture 15A.1
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
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Lecture 15A.1
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
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).
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.
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Lecture 15A.1
Page 12 of 32
Lecture 15A.1
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].
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.
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Lecture 15A.1
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).
The design of offshore structures has to consider various requirements of construction relating to: 1. fabrication. 2. weight. 3. load-out. 4. sea transport. 5. offshore installation. 6. module installation. 7. hook-up. 8. 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.
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.
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Lecture 15A.1
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.
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).
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Lecture 15A.1
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.
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Lecture 15A.1
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.
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.
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.
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Lecture 15A.1
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.
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.
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Lecture 15A.1
For lifting, steel wire ropes in a four-sling arrangement are used which directly rest in the fourpoint 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.
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Lecture 15A.1
Operator
Name
Mode
Thor
Monohull
Odin
Type
Lifting capacity (Tonnes)
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
Monohull
Heerema Hermod
Balder
DB50
DB100
Semisub
Semisub
Monohull
Semisub
McDermott
DB101
Semisub
DB102
Semisub
Rev
6000 + 6000 = 12000
Micoperi
M7000
Semisub
Rev
7000 + 7000 = 14000
ETPM
DLB1601 Monohull
Rev.
1600
Notes: 1. Rated lifting capacity in metric tonnes. 2. 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.
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Lecture 15A.1
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.
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Lecture 15A.1
For load-out three basic methods are applied: •
skidding
•
platform trailers
•
shearlegs.
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.
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.
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Lecture 15A.1
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.
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Lecture 15A.1
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.
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.
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.
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.
The major phases of construction when structural integrity may be endangered are: •
Load-out
•
Sea transport
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Lecture 15A.1
•
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.
Page 27 of 32
Lecture 15A.1
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 t opside level. SEA-FASTENING The structure to keep the object rigidly connected to the barge during transport. SHACKLES Connecting element (bow + pin) between slings and pade yes. 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.
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Lecture 15A.1
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.
[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 erectio n 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.
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Lecture 15A.1
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.
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.
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Lecture 15A.1
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 Hydrimechanics; 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.
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Lecture 15A.1
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.
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Lecture 15A.2
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 highorder wave theories can be found in the literature, e.g. see [9].
Page 5 of 16
Lecture 15A.2
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.
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Lecture 15A.2
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,
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Lecture 15A.2
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.
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 = Cif c A ......................................... (7) Where, A is the exposed area of structure, f c 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].
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.
Page 12 of 16
Lecture 15A.2
[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 P ublishing Co., Houston, 1981. [13] Gerwick, B.C. Jr., "Construction of Offshore Structures", John Wiley, New York, 1986.
Phase θ = kx - ω t Relative water depth d/L
Deep water
Finite water depth
d/L ≥ 0,5
d/L < 0,5
Velocity potential θ Surface elevation z
ζa cos θ
ζa cos θ
ρ gζa ekz cos θ
Dynamic pressure pdyn =
Water particle velocities ζa ω ekz cos θ
horizontal u =
vertical w =
ζa ω ekz sin θ
Water particle accelerations ζa ω2 ekz sin θ
horizontal u' =
vertical w' =
-ζa ω2 ekz cos θ
Page 15 of 16
Lecture 15A.2
Wave celerity c =
Group velocity cgr =
Circular frequency ω =
Wave length L =
Wave number k =
co =
c=
cgr =
cgr =
ω =
ω =
Lo =
L=
ko =
kd tanh kd =
Water particle displacements horizontal ξ
-ζa ekz sin θ
vertical ζ
ζa ekz cos θ
Particle trajectories
Circular orbits
Elliptical orbits
Where ζ a =
Page 16 of 16
Lecture 15A.3
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
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Lecture 15A.3
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. Previous experience along the tow route 2. Exposure time and reliability of predicted "weather windows" 3. Accessibility of safe havens 4. Seasonal weather system 5. 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.
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Lecture 15A.3
LIMIT STATES FOR TEMPORARY PHASES LOAD Progressive Collapse
TYPE Service
Fatigue
Ultimate
ability
Serviceability Abnormal
Damage
effects
condition
DEAD
EXP
LIVE
SPE
DEFORM
EXPECTE
ATION Depen dent on ENVIRON
operati
MENTAL
onal require
Expect
Value
ed load
dependent on
history
measures taken
Dependent on operational requirements
ments
ACCIDEN TAL
Dependent on NOT APPLICABLE
operational requirements
NOT APPLIC
LIMIT STATES FOR NORMAL OPERATIONS
Progressive Collapse Fatigue
Ultimate Abnormal effects
Damage condition
ECTED VALUE CIFIED VALUE
D EXTREME VALUE
Annual Expect
Annual
exceedanc
ed load
exceedance
e
history
probability 10-2
probability
Annual exceedance probability 10-2
10-4
Annual exceedanc ABLE
e
NOT APPLICABLE
probability 10-4
Page 13 of 13
Lecture 15A.4
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).
The material partial factors for steel is normally taken equal to 1,15 for ULS and 1,00 for PLS and SLS design.
Guidance for classifying typical conditions into typical limit states is given in the following table:
Construction
P
Load-Out
P
ULS,SLS reduced wind
support
ULS
disp Transport
P
transport wind
ULS
and wave Tow-out
P
flooded compart
PLS
(accidental) Launch
P
ULS
Lifting
P
ULS
In-Place
P+L
(normal)
wind, wave &
actual
ULS,SLS
snow
Page 5 of 14
Lecture 15A.4
In-Place
P+L
(extreme)
wind & 100
actual
ULS
year wave SLS
In-Place
P+L
(exceptional)
wind & 10000
actual
PLS
year wave
Earthquake
P+L
10-2 quake
ULS
Rare
P+L
10-4 quake
PLS
Earthquake Explosion
P+L
blast
PLS
Fire
P+L
fire
PLS
Dropped
P+L
drill collar
PLS
P+L
boat impact
PLS
Object Boat Collision Damaged
P + reduced L
Structure
reduced wave
PLS
& wind
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.
Page 6 of 14
Lecture 15A.4
p(t) = The plot of the amplitudes p j versus the circular frequencies ω j is called the amplitude power spectra of the loading. Usually, significant values of p j 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 ω2 + i x C ω + K] the elements of which represent:
H j,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.
The response of the i-th mode may alternatively be determined by resorting to Duhamel's integral:
X j(t) = The overall response is then obtained by summing at each time step the individual responses over all significant modes.
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.
Page 13 of 14
Lecture 15A.4
•
responses involving many vibration modes to be determined over a short time interval.
The dynamic equilibrium at an instant τ 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 characterized 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 β = 1/4 or Wilsontheta with θ = 1,4).
•
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.
Page 14 of 14
Lecture 15A.5
[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.
Page 15 of 16
Lecture 15A.5
[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.
Page 16 of 16
Lecture 15A.6
These loads are those transferred from the jacket to the foundation. They are calculated at the mudline.
Gravity loads (platform dead load and live loads) are distributed as axial compression forces on the piles depending upon their respective eccentricity.
Page 3 of 25
Lecture 15A.6
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).
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.
The overall resistance of the pile against axial force is the sum of shaft friction and end bearing.
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 Cu at great depth.
Lateral friction is integrated along the whole penetration of the pile.
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.
Page 4 of 25
Lecture 15A.6
Page 7 of 25
Lecture 15A.6
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.
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.
Page 8 of 25
Lecture 15A.6
account the changes in load direction during lifting). Padeyes are then carefully cut before lowering the next pile section.
Page 13 of 25
Lecture 15A.6
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.
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 i n 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).
Figure 5 shows the different steps of this routine operation:
Page 14 of 25
Lecture 15A.6
•
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 positioning.
Each add-on should be designed to prevent bending or buckling failure during installation and in-place conditions.
Page 15 of 25
Lecture 15A.6
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/feet (980 blows/meters).
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.
Page 16 of 25
Lecture 15A.6
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.
Page 17 of 25
Lecture 15A.6
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.
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).
Page 18 of 25
Lecture 15A.6
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.
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.
Page 19 of 25
Lecture 15A.6
•
information about the pile/structure connection:
- equipment and procedure employed. - overall volume of grout and quality. - record of interruptions and delays.
Page 20 of 25
Lecture 15A.6
A. Air/Steam Hammers
Rated
Ram
Max.
Std.
Typical
Pilecap
Rated Operating
Steam
Air
Hose Rated
Hammer Make
Model
Energy
(ft-lbs)
Weight
(kips)
Stroke
(m)
Weight
(kips)
Weight
(w/leads) (kips)
Pressure
(psi)
Consumption
Consumption
ST/F BPM
(lbs ht)
(lbs ht)
.....
2@ 6850
510.000
85
72
57,5
312
180
5650
325.000
65
60
59,0
262
160
31.500
7.500
4
40
3@
45
4 Conmaco
5300
150.000
30
60
12,7
92
160
8.064
1.711
46 4
300
90.000
30
36
12,7
86
150
6.944
1.471
54 3
200
60.000
20
36
12,7
74
120
5.563
1.195
59 3
2@ 12500
1.582.220
275,58
69
154,32
853
171
53.910
26.500
6
36
8800
954.750
194,01
59
103,62
600
150
32.400
16.700
8
36
8000
867.960
176,37
59
85,98
564
142
30.860
15.900
8
38
7000
632.885
154
49
92,4
583
156
30.800
14.830
4@
35
4
Menck 5000
542.470
110,23
59
66,14
335
150
20.940
10.400
40 6
(MRBS) 4600
499.070
101,41
59
52,91
313
142
19.840
9.900
42 6
3000
325.480
66,14
59
33,07
205
142
12.130
6.000
42 5
1800
189.850
38,58
59
22,05
125
142
7.060
3.700
44 4
850
93.340
18,96
50
11,5
64
142
3.530
1.950
45 3
MKT
OS-60
18.000
60
36
OS-40
120.000
40
36
OS-20
60.000
20
36
38,65
150
3
Page 23 of 25
60
Lecture 15A.6
C. Hydraulic Hammers
Rated Energy
Make
Ram Weight
Standard
Hammer
Typical
Weight
Operating
Rated
Pilecap
Model
Rated
Weight
Pressure
Oil Flow BPM
(ft-lb)
(kips)
4000
1.200.000
205
490
3000A
800.000
152
414
3000
725.000
139
(kips)
(kips)
(psi)
(gal. min)
33
HMB
40-70 1500
290.000
55
17,6
172
900
170.000
30,8
500
72.000
9,5
1,1
27,5
1700
760.000
132
84
415
3400
845
50-80
MHU
1.230.000
207
77
617
3400
845
32-65
650.000
110
386
3100
580
48-65
195
141.000
22,0
6,0
59
3550
98
38
MH
119.000
19,0
6,0
51
3190
103
42
105.000
16,5
6,0
46
2755
102
42
145
87.000
13,9
6,0
40
2320
103
44
MH
69.000
11,0
1,9
27
2830
75
48
58.000
9,3
1,9
24
2465
75
48
88
MRBU
MHU
900
MH
Menck
165
MH
120
MH 96
MH 80
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.
Page 24 of 25
Lecture 15A.6
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.
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
Page 25 of 25
Lecture 15A.7
To present methods for the design of large tubular joints typically found on offshore structures.
Lecture 15A.1: Offshore Structures: General I ntroduction
Lecture 15A.8 : Fabrication Lecture 15A.12: Connections in Offshore Deck Structures
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.
The main structure of 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. Connections 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. Static strength considerations 2. Fatigue behaviour considerations
Page 1 of 23
Lecture 15A.7
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".
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.
Page 2 of 23
Lecture 15A.7
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.
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 θ is the angle between brace and chord axis.
Page 3 of 23
Lecture 15A.7
α=
β=
γ=
Can slenderness ratio Brace to chord diameter ratio (always ≤ 1) Chord slenderness ratio
τ=
Brace to chord thickness ratio
ζ=
Relative gap
These are non-dimensional variables for use in parametrical equations.
Load paths within a joint are very different, according to the joint geometry. The following classification is used, see Figure 2.
Page 4 of 23
Lecture 15A.7
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.
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.
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.
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.
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.
Page 5 of 23
Lecture 15A.7
θ
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.
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.
Page 6 of 23
Lecture 15A.7
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.
Page 7 of 23
Lecture 15A.7
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.
The same definitions and limitations apply to multiplanar joints.
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. 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. 2. 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. 3. The actual gap shall not be less than 50mm. To achieve this, most designers use a 70 or 75mm theoretical gap. 4. 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. 5. Thickness transitions are smoothed to a 1 in 4 slope, by tapering the thicker wall.
Page 8 of 23
Lecture 15A.7
Page 9 of 23
Lecture 15A.7
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.
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 = τ f sin θ 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.
Page 10 of 23
Lecture 15A.7
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 treatments of results allow formulae to be defined for the allowable punching shear stress.
Page 11 of 23
Lecture 15A.7
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 - λ γ f AX, f IPB, f OPB are the nominal axial, in-plane bending and out of plane bending stresses in the chord
Page 12 of 23
Lecture 15A.7
Value for λ and Qq are given in Table 6.1
Stress in brace
f ax
f by
f bz
Acting punching shear
Vpx = τ f ax sin θ
Vp = τ f by sin θ
Vp = τ f bz sin θ
0,045
0,021
K joints
T & Y Joints
Tension
Compression
w/o diaphragm Qq X
w diaphragm
0,030
λ
Qg = 1,8 - 0,1
for γ ≤ 20
Qg = 1,4 - 4 g/D for γ > 20 but Qg must be ≥ 1,0
Qβ =
for β > 0,6
QB = 1,0 for β ≤ 0,6
Page 13 of 23
Lecture 15A.7
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.
The parametric formulae discussed in Section 6.2 were specifically established for nonoverlapping 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 si milar 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.
Page 14 of 23
Lecture 15A.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 t hat portion of the brace which conta cts 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.
Large chord wall thickness may be reduced by stiffening the chord. The most usual reinforcement consists of ring stiffening inside the chord.
Page 15 of 23
Lecture 15A.7
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.
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].
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.
Page 16 of 23
Lecture 15A.7
- as for K joint
The above equation for T/Y, K and KT joints are generally valid for joint parameters within the following limits: 8,333 ≤ γ ≤ 33,3 0,20 ≤ τ ≤ 0,8 0,3 ≤ β ≤ 0,8 unless stated otherwise 6,667 ≤ α ≤ 40 unless stated otherwise 0° ≤ σ ≤ 90° unless stated otherwise.
A fatigue analysis of a joint consists of the following steps: 1. Calculation of nominal stress ranges in the brace and the chords 2. Calculation of hot-spot stress range 3. Calculation of joint fatigue lives using S-N curves for tubular members at joints.
Nominal stress ranges in braces and chords are calculated by a global stress analyses.
Page 19 of 23
Lecture 15A.7
A wave histogram has to be obtained for each direction around the platform. A simple form of a wave histogram is as follows:
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
Nominal stress ranges can be calculated by following the steps below: 1. 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. 2. 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. 3. 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.
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).
Page 20 of 23
Lecture 15A.7
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.
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.
Page 21 of 23
Lecture 15A.7
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.
•
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.
Page 22 of 23
Lecture 15A.7
[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.
Page 23 of 23
Lecture 15A.8
Page 13 of 22
Lecture 15A.8
Page 14 of 22
Lecture 15A.9
Page 3 of 19
Lecture 15A.9
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.
Page 4 of 19
Lecture 15A.9
Page 5 of 19
Lecture 15A.9
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).
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.
Page 6 of 19
Lecture 15A.9
5. 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
Thor
Odin
Type
Mode
Lifting Capacity
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
Monohull
Monohull
Heerema
Hermod
Balder
DB50
DB100
Semisub
Semisub
Monohull
Semisub
McDermott
DB101
Micoperi
Semisub
DB102
Semisub
Rev
6000 + 6000 = 12000
M7000
Semisub
Rev
7000 + 7000 = 14000
Notes: 1. Rated lifting capacity in metric tonnes 2. When the crane vessels are provided with two cranes, these are situated at the vessels stern at approximately 60m distance etc.
Page 19 of 19
Lecture 15A.10
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 desi gn for deck floors.
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
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.
integrated deck.
2.
module support frame.
3.
modules.
Floor concepts are presented and several aspects of the plate floor design are addressed.
Page 1 of 21
Lecture 15A.10
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.
Page 2 of 21
Lecture 15A.10
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 hookup. Most of this lecture refers to this type of integrated deck such as is shown in Figures 3 and 4.
Page 3 of 21
Lecture 15A.10
The selection of the concept for the structural deck is made in close cooperation with the other disciplines.
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 d eck 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.
Page 4 of 21
Lecture 15A.10
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 portalframe type without braces (Figure 3), are compared in Table 1.
Page 9 of 21
Lecture 15A.10
Note:
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
+
++
++ 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).
Page 10 of 21
Lecture 15A.10
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.
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.
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. sea fastening
•
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.
Page 13 of 21
Lecture 15A.10
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
Page 14 of 21
Lecture 15A.10
•
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.
Deck beams supporting the floor panels or providing direct support to major equipment are generally provided as HE 800-1000 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.
Page 17 of 21
Lecture 15A.10
Figures 8 and 9 illustrate the problems.
Page 18 of 21
Lecture 15A.10
[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.
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 17-44.
Good background to theory of plated structures.
Page 21 of 21
Lecture 15A.11
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).
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
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.
fabrication
2.
weighing
3.
load out
4.
sea transport
5.
offshore installation especially deckmating
6.
module installation
Page 1 of 22
Lecture 15A.11
7.
hook-up
8.
commissioning
A brief discussion on inspection and repair and on platform removal concludes this lecture.
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)
} Discussed in
Horizontal bracing
} Lecture 15A.10
Deck beams
}
Primary girders
}
Vertical trusses or bracing
} Discussed in
Deck legs
} this lecture
Page 2 of 22
Lecture 15A.11
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.
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.
Page 7 of 22
Lecture 15A.11
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.
Page 8 of 22
Lecture 15A.11
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.
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 postbuckling strength. The depth/thickness limits given above do not then apply.
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
Page 9 of 22
Lecture 15A.11
These joints are discussed in Lecture 15A.12.
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.
Page 10 of 22
Lecture 15A.11
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.
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.
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.
Page 13 of 22
Lecture 15A.11
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.
Page 14 of 22
Lecture 15A.11
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.
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
Page 17 of 22
Lecture 15A.11
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
Page 18 of 22
Lecture 15A.11
re-use of the facility is planned, then removal engineering should be developed early in the design.
•
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 t he lecture.
[1] API-RP2A: Recommended practice for planning, designing and constructing fixed platforms. American Petroleum Institute, 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 for structural code for topsides. [3] API-Bulletin 2V: Bulletin on design of flat plate structures. American Petroleum Institute, Institute, 1st ed., 1987. Valuable specialist addendum to API-RP2A.
Page 21 of 22
Lecture 15A.11
[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.
Page 22 of 22
Lecture 15A.12
•
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 practices. 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.
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.
Page 3 of 19
Lecture 15A.12
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 l ength ls for the stringer,
Page 4 of 19
Lecture 15A.12
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.
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.
Page 7 of 19
Lecture 15A.12
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 f y 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 = M cw/(h1 - t 1) (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 multi-planar 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
Page 8 of 19
Lecture 15A.12
amounting to 60% of the yield value, the strength of the connection has to be reduced by 20%.
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].
Page 9 of 19
Lecture 15A.12
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 f yo is the yield stress of joint "can" to is the wall thickness of joint "can" β 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 ≤ β ≤ 1,0 For bending moments in-plane, the axial force N is derived from N = Mcw/(h1 -t 1) 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)
Page 10 of 19
Lecture 15A.12
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.
Page 13 of 19
Lecture 15A.12
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 (b m1 tp + bm2 tw) For the brace intersection the effective width is given by: Aeff.b = 2 (b e1 + be2) tp The strength of the connection is thus given by: N2sin θ2 = Aeff.c f yo and N2sin θ2 = Aeff.b f yo 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.
Page 14 of 19
Lecture 15A.12
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.
Page 15 of 19
Lecture 15A.12
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 t E) f y N2 = 1,08 (c t L + (D - d) t E) f y N3 = 0,87 d (tL + 2 tE) f y
Page 16 of 19
Lecture 15A.12
[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
1. Marshall, P.W. "Design of Welded Tubular Connections: Basis and Use of AWS Provisions". Elsvier, 1991 2. 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 3. 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
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