analysis of damage of jacket with barge

Engineering Structures 27 (2005) 1317–1326

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Evaluation of damage to offshore platform structures due to collisionlarge barge

Wei-liang Jina,∗, Jian Songa, Shun-feng Gonga, Yong Lub

aInstitute of Structural Engineering, Zhejiang University, Hangzhou 310027, PR ChinabSchool of Civil and Environmental Engineering, Nanyang Technological University, Singapore 639798, Singapore

Received 31 March 2004; received in revised form 21 February 2005; accepted 21 February 2005Available online 19 May 2005

Abstract

An offshore jacket platform in the South China Sea was impacted by a large derrick and lay barge during installation. This papera non-linear dynamical analysis procedure for firstly determining the impact action based on the forensic evidence from the dcomponents, and then evaluating the overall damage effects on the platform structure. The impact action of the barge is simua triangle impulse load with different collision contact times. The curves relating the indentation deformations of the damagedwith different collision contact times are simulated using an estimated velocity of the impacting ship. On the basis of these cuthe actual detected dent damages, the contact time and the maximum impact load on the platform are determined. Taking into aforce–deflection relationship of the local indentation of the damaged cross-diagonal brace, the transmission of the impact load to thstructure is simulated by a non-linear spring. The added mass coefficient with hydrodynamic effects and the pile–soil-structure inare considered in the computational model of the non-linear dynamic response of the platform structure. Subsequently, the dynamic respoof the offshore jacket structure is computed and the critical stress and deformation of the tubular joints are obtained as indicatodamage effects. The results are useful for choosing a feasible and reasonable repairing and strengthening scheme for the damagThe procedure presented in this paper is generally applicable for the evaluation of typical offshore platform structures in the case ofor collision.© 2005 Elsevier Ltd. All rights reserved.

Keywords: Offshore jacket platform; Collision dynamic response; Damage simulation; Ship/barge impact

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1. Introduction

Offshore jacket platforms have been widely usedoffshore oil and gas exploitation with complicated oceanenvironments. Besides the normal operational loads,platforms are subjected to other loads, such as wind, wcurrent and ice loads [1]. At the same time, they are alsoexposed to unexpected incidents inducing sudden loadsexample, collision of a vessel with the platform, or impafrom a heavy object dropping from the top of the platforThese may result in crooking or buckling of some membethus reducing their load bearing capacity and potentiaaffecting the safety and the integrity of the whole platfor

∗ Corresponding author.

0141-0296/$ - see front matter © 2005 Elsevier Ltd. All rights reserved.doi:10.1016/j.engstruct.2005.02.010

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structure. To effectively repair the damaged members arestore the desired state of the structure requires a gassessment of the condition of the structural system afteraccidental event. For this reason, how to analyze and assthe damage to the platform structures due to collision, athe influences of such damage on the integrity, load bearcapacity and the fatigue lifetime of the platform, havbecome important topics in offshore platform risk studies.

The concerns for ship collision are reflected in varioudesign codes [2,3]. For a general assessment of a damagplatform impacted by a ship, it may be possible to turn thdynamic problem of collision into a normal statics problemwith equivalent static loads [4]. Such analysis can be usefufor understanding the general effects of the collision adetermining the residual strength of the affected membe

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1318 W.-l. Jin et al. / Engineering Structures 27 (2005) 1317–1326

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Fig. 1. Flow diagram of platform structure damage analysis.

However, collision is actually a dynamic process, andinvolves more complicated dynamic factors which couaffect the structural response,e.g. the way that the collisionhappened between the barge and the platform structurecontact time of collision, the pile–soil-structure interactioduring the dynamic response of the platform structure.the course of a collision, one important problem is enerabsorption and dissipation. Jorben Amadahl [5] analyzedthe impacts between supply vessels and offshore structuin particular two areas were studied, energy dissipation inthe ship’s bow and stern structures and the deformatibehavior of tubular bracings. Various mechanisms of enedissipation in a ship structure subjected to collision loawere identified and described; design curves were propofor bow and stern impacts with supply vessels. The differemodes of energy dissipation were described, for assesthe load carrying capacity in the beam mode of deformataccounting for the detrimental effect of local indentation.

This paper reports a comprehensive evaluation ofdamage to an offshore platform structure, which was acdentally impacted by a large derrick and lay barge durithe installation. In this study, the impact load is determinon the basis of the forensic damage evidence detectedthe offshore structure was impacted by the barge. The addemass coefficient for the hydrodynamic effect is considerin the evaluation of the collision effect on the platformstructure, while the pile–soil-structure interaction is consered in the development of the computational model forstructural system. The force–deflection relationship of thlocal indentation for the damaged cross-diagonal bracesimulated by a non-linear spring. The dynamic response

he

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ysdtg

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ei-

er

d

-e

isf

Fig. 2. Photograph of the platform jacket structure (center) impacted by alarge derrick and lay barge (right).

Fig. 3. Layout of skirt piles.

the structure is then analyzed, and the critical stressdeformation responses of the tubular joints in the offshojacket structure are obtained to assess the damage effFig. 1gives an outline of the general analysis procedure.

2. Description of the collision case

An offshore platform was impacted by a ship during iinstallation.Fig. 2provides a snapshot view of the collisioncaptured at the time when it took place. The platform issteel jacket deep-water platform with four legs and eigskirt piles. Fig. 3 shows the layout of the skirt piles. Thewater depth of the seabed is 117.2 m. The penetration deof the piles is 91 m, and the diameter of the piles is 1829 m

The large derrick and lay barge was anchored besideoffshore platform. At the time of collision, the barge ran ouof control so that it flew onto the platform jacket structudue to sea wind and current and collided with the jackstructure. The damage was reported on the diagonal bracbetweenleg A2–B2 (Fig. 4, location I). The damaged areastarted at 4 m from the top weld of the node and endedabout 5.40 m from the node connecting to leg A2. Inspect

W.-l. Jin et al. / Engineering Structures 27 (2005) 1317–1326 1319

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Fig. 4. The damage areas due to the collision.

revealed that the diagonal member was squashed for a150 mm and cut open over an area of about 130× 130×110 mm. On the opposite side of the member there wa furrow about 40 mm deep. Removal of paint and soscratching were also evident on this diagonal member.

In addition, as shown inFig. 4, leg B2 was squashed in atthe level of the first riser clamp in the water. The supportthe riser clamp was partially cut open along the weld onreinforcement plate, while the weld between the leg and threinforcement plate was partially cut open.

3. Mechanics model of colliding system

3.1. Collision mechanics

When a platform is collided with by a ship, it may bassumed that the time of collision is far smaller than tmotion period of the ship. After collision, the ship woulmovetogether with the platform structure. Before settingthe mechanics model of the colliding system, the followiconsiderations [6] are given:

• The hull of the ship is assumed to be a rigid body wicertain speed and mass for the calculation of the colliseffect on the platform structure, and the deformationthe hull structure is neglected.

• The collision effect is evaluated in accordance with tlaws of momentum conservation and conservationenergy.

• The mass of the ship includes its self-mass,ms, andthe additional mass due to the hydrodynamic interactibetween sea water and the ship. It can be expressed bfollowing formula:

m1 = ms + kms (1)

ut

s

he

wherek is the additional mass coefficient. For collisionalong the side direction (lateral),k = 0.4; in the casewhere thebow or the stern bumps against the platformk = 0.1.

• The materials within the collision contact area arperfectly elasto-plastic.

According to the law of object momentum conservation,

m1v1 + m2v2 = (m1 + m2)v12. (2)

Hence the common speed after colliding is

v12 = (m1v1 + m2v2)/(m1 + m2) (3)

wherem2 is the mass of the platform;v1 is the velocityof ship movement before colliding;v2 is the velocity ofthe platform movement before colliding;v12 is the commonspeed of the ship and platform after colliding.

The kinetic energy before colliding is partially absorbeby the plastic distortion of the ship and platform structureso the conservation of energy can be expressed as

1

2m1v

21 + 1

2m2v

22 = 1

2(m1 + m2)v

212 + Es + Ep (4)

where Es is the energy absorbed by the ship;Ep isthe energy absorbed by the platform. According to therigid-body assumption for the ship,Es may be neglected.Consequently, the energy absorbed by the platform canconservatively written as

Ep = 1

2m1v

21/(1 + m1/m2). (5)

Generally, the impact energy absorbed by an offshorejacket structure from a ship involvesthe following energyabsorption processes:

• local denting or crushing of the tubular member section• elastic beam bending;• plastic bending/hinge formation;• global structural deformation (elastic and plastic).

In the particular case under investigation, the total maof the jacket platform is 1.889× 106 kg and the mass of thebarge is about 4.2 × 107 kg. The total mass of the jacketplatform structure is far less than the mass of the barge. Tadditional mass factor for the barge is assumed to be 0.4take into account the installment equipment in the vessel.

3.2. Local dent of tubular member

To study the global structural response due to collisiothe local dent of the tubular member under impact loamust be discussed first so that the transmission ofimpact load can be established. The shape and area of thlocal dent depends on the collision modes. Because ofcomplexity of the impact problem, it is difficult to find asimple analytical model to establish the relationship betweethe local denting of the tubular memberδ and the impact


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Fig. 5. Non-linear spring force–deformation relation.

load P. Bai [7] proposed an empirical formula for relatingthe elastic deformation,δE, to the impact loadPE:

PE = 0.1116(D/t)3E LCδE (6)

whereE is the Young’s modulus,t is the tube wall thickness,D is the tubediameter, andLC is the axial characteristiclength of the contact area.LC depends on the tube diameteD, the tube length, and the shape of the dent. On the baof a series of denting experiments [8] and linear shell finiteelement analyses [9], it was proposed thatLC = 1.9D.

When the impact forceP is greater than a critical valueP0, a permanent dent deformation would be produced on thtube wall. The critical valueP0 can be derived from the rigidplastic finite element analysis [9], as

P0 = 2Fyt2LC/D (7)

whereFy is the yield stress of the material.An empirical formula relating the permanent dent

deformation,δP, to the impact load,PP, can be obtainedaccording to API RP 2A-WSD [2], as

PP = 40Fyt2(δP/D)0.5. (8)

The total dent depthδ will include the elastic dent depthδE and the permanent dent deformationδP when the impactforce P is greater than the critical valueP0, i.e.

δ = δE + δP. (9)

Using Eqs. (6)–(8), the P–δ relationship can be obtainedA non-linear spring, as shown inFig. 5, having this P–δ

relationship is introduced into the computational modto represent the force–deflection relationship of the locindentation for the damaged cross-diagonal brace. The nlinear spring is effective under unilateral compression, athe final deformation of the spring represents the dent deof the tubular member under the impact load.

3.3. Motion equations of collision

Transient dynamic analysis can be used to explain tdeformation, strain, stress, and force with time under steaload, transient load and simple harmonic load. The basicequation of motion can be written as

[M]{u} + [C]{u} + [K ]{u} = {P(t)} (10)

-

Fig. 6. Definition of the impact load.

where[M] is the mass matrix,[C] is the damping matrix,[K ] is the stiffness matrix,{u} is the nodal accelerationvector, {u} is the nodal velocity vector, and{u} is thenodal displacement vector. For the impact problem undconsideration, the impact action from the ship,P(t), may besimplified into an isosceles triangle impulse load, as showin Fig. 6. It can be expressed as

P(t) =

2Ft/t0 0 ≤ t ≤ t0/22F(1 − t/t0) t0/2 ≤ t ≤ t0

0 t0 ≤ t

. (11)

At any arbitrary time,t , the motion equations expressed inEq. (10) may be considered as a series of static equilibriuequations with inertia force([M]{u}) and damping force([C]{u}). There are typically two methods for solving thesequations. One is the modal superposition method, andother is step-by-step integration. The natural mode shaand frequencies of the structure must be solved first whthe modal superposition method is adopted. Becauseimpulse type load may excite numerous vibration modeit is necessary to consider a sufficient number of modes inorder to obtain a satisfactory solution. Moreover, the modsuperposition method is only applicable for linear structursystems. For the problem under consideration, non-lineresponses may take place at some critical regions ofjacket structure, as well as at the impact regions (represenby a non-linear spring) and in the pile–soil interactionTherefore, direct integration using the Newmark methodadopted in this study for solving the motion equations undthe impact load.

4. Impact load identification

4.1. Non-linear numerical simulation analysis

Numerical simulation analysis is carried out to inverseidentify the impactload characteristics according to thedetected damage on the impacted member. In this procedureforward non-linear finite element analysis is performeto calculate the dent damage for various possible loconditions, so that the relationship between the dent deand the loading parameters canbe established. From there


Fig. 7. Finite element of the damaged member.

Fig. 8. Distribution of stress anddeformation for the element under 2.45× 106 N.

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the actual load parameterscan be determined according tthe measureddent information.

Both an equivalent static load and a more realistic impload as expressed in Eq. (11) are determined through thisnumerical analysis procedure. The equivalent static loadbe used to perform a quick static analysis to assessmagnitude of the structural response, while the impact locan be used for a more detailed and more accurate dynaresponse analysis of the platform jacket structure.

4.2. Equivalent static load identification based on tdamagedmember

The equivalent static load analysis is first performon the damaged cross-diagonal bracing member assumtwo fixed ends. The finite element model for the damagmember between the two adjacent nodes was createshown inFig. 7. The diameter of the member is 914 mmthe thickness is 19 mm and the length is 12 400 mThree-node and four-node shell elements are used to mthe model, with denser mesh arranged in regions near thconnections and the loading area. The shell element, whichas plasticity, creep, stress stiffening, large deflection a

t

ne

dic

ngdas

.sh

small strain capability, is suitable for the non-linear problemunder consideration. The member was meshed with 1shell elements and 2016 nodes. The ideal elastic–placonstitutive relationship is adopted for the steel, with elasmodulusE = 2.0× 105 MPa, yield strengthFy = 345 MPaand Poisson’s ratio= 0.3.

The loads are distributedin the damaged area of themember in a triangle form, with maximum loading densiat the center of the damaged area, as shown inFig. 7. Aseries ofanalyses with different total load values are carriout.Fig. 8shows a typical local dent damage scenario fromthe numerical simulation. On the basis of the results, itfound that when the total load value is 2.45 × 106 N, themaximum radial deflection is 150.886 mm, which is closethe actually detecteddent depth of 150 mm.

4.3. Equivalent static analysis of cross-diagonal brace

To better simulate the boundary conditions of thdamaged member, another round of FE analysis are carout on a substructure containing the damaged cross-diagbraces. The substructure is depicted inFig. 9. At the top righttubular joint, all members connecting to the damaged brac


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Fig. 9. Model of the cross-diagonal brace substructure.

are included in the model for a length equal to three timestheir respective diameters. On the basis of the results frthe analysis described in the previous section, three total lovaluesequal to 2.4×106 N, 2.45×106 N and 2.5×106 N areapplied. The simulated damage under the 2.45× 106 N loadis depicted inFig. 10. The corresponding maximum radiadeflection is 155.3 mm, which does not differ much fromthe analysis result for the damaged member alone and is aclose to the measured dent depth of 150 mm.

It is therefore concluded that the equivalent static loaof the collision with respect to the local dent damageabout 2.45 × 106 N. This static load may be used in anapproximate static procedure to assess the general effecthe collision on the overall platform structure. The presestudy, however, will be based on the dynamic impact loadescribed in what follows.

4.4. Maximum percussive force analysis

In order to more realistically reproduce the impaceffect, it is necessary to identify the dynamic impact loaparameters. To completely define the impact load accordingto Eq. (11), it is necessary to determine the maximumimpact force,F , and theduration (contact time),t0. The totalimpulse is equal to the momentum of the platform after thcollision,

m2v12 = 1

2Ft0 (12)

wherev12, as shown in Eq. (3), is dependent on the velocityof the impacting ship prior to collision(v1). Therefore,once the ship velocity is known, the impulse will beconstant, and the actual values oft0 andF can be identifiedby numerical trial analysis to reproduce the measured ddamage.

According to the documentation of the accident, thplatform was laterally impacted by the barge with a veloci

f

d

o

of

t

between 1.0 and 2.0 m/s. The analysis is performedassuming the two limit speed values. A detailed finielement analysis [10] has been carried out to simulatethe dent damage on the damaged member for the abovetwo ship velocities for various combinations of the impacontact time and peak impact load.Table 1 summarizesthe numerical simulation results.Figs. 11 and 12 depictthe relationship between thecalculated dent depth of thedamaged memberδP and the collision contact timet0 forthe ship velocity(v1) equal to 1 m/s and 2 m/s, respectively.The fit curves can be expressed as

δP = 20.15t30 − 27.64t2

0 − 133.08t0 + 246.30

(v1 = 1 m/s) (13)

δP = −35.87t30 + 243.42t2

0 − 622.01t0 + 718.45

(v1 = 2 m/s). (14)

According to the above expressions, for the actual detecdent depth equal to 150 mm, the contact time is found to0.682 s when the impact velocity of barge is 1.0 m/s, andit is 2.15 s if the impact velocity of the barge is 2.0 m/s.Correspondingly, the maximum percussive force wasF =5.367× 106 N andF = 3.405× 106 N, respectively.

Table 1Dent depth of damaged member for different collision contact times

Contact time Maximum Dent deptht (s) percussive force δ (mm)

F (MN)

(a) Ship impact velocity= 1.0 m/s

0.2 18.6722 220.2910.4 9.3361 184.4410.6 6.2241 167.3430.8 4.6681 130.4811.0 3.7344 104.4691.2 3.1121 82.321

(b) Ship impact velocity= 2.0 m/s

0.4 18.6755 512.5560.8 9.3377 335.9271.2 6.2252 287.3451.6 4.6689 190.6092.0 3.6155 156.6282.4 3.1126 134.563

The above two combinations of the impact load will bconsidered in the subsequent dynamic response analysthe platform jacket structure.

5. Dynamic response analysis

With the determination of the impact load as describin the preceding section and the derivation of the non-linearspring representing the transmission of the impact loadthe platform structure (Section 3.2), the dynamic responseanalysis of the platform jacket structure can be performeTo take into account the interaction between the piles an


Fig. 10.Z-directional displacement distribution of the node of the cross-diagonal brace under 2.45× 106 N impact load.

s

s

l

n,dda

thee

seared

anly,ofn

tednt

ring

in

Fig. 11. Fit curve of the dent depth ofthe damaged tubular member versucollision contact time, for barge velocity= 1.0 m/s.

Fig. 12. Fit curve of the dent depth ofthe damaged tubular member versucollision contact time, for barge velocity= 2.0 m/s.

the soil, the computational model of the platform structurewill consist of an appropriate pile model and the FE modefor the jacket structure supported by the piles.

5.1. Pile model

In order to simulate the structure-pile–soil interactiothe bearing capacity of a single pile needs to be analyzefirst. The diameter of the piles is uniformly 1829 mm, anthe length is 91 m. In the finite element (FE) model,3-D beam element is used to simulate the pile. Because thepile penetrates through several layers of different soils,discretization of the piles along the vertical direction is madsuch that within each layer of thesoils the portion of the pileis divided into an integer number of elements. This allowan easier calculation of the parameters of the non-linsprings representing the soil reaction. Each pile is dividinto approximately1.5 m long elements along the length.

The effects of the soil reactions on each pile element cbe simplified into three kinds of non-linear springs, namea lateral spring representing the lateral bearing capacitythe soil, a vertical spring representing the vertical frictioforce on the pile surface, and a torsion spring representingthe circumferential friction force on the pile surface. Theacting points of the lateral and torsional springs are locaat the mid-height of the element, while the acting poiof the vertical spring is locatedtowards thebottom of theelement to simulate the verticalpile–soil friction force. Foreach pile as a whole, there is also an end support spwhich represents the end-bearing capacity of the pile.Fig. 13depicts the arrangement of soil springs.

The spring parameters arecalculated according to thesite investigation and pile testing data [11]. A typicalQ–Z (vertical force–displacement) datasheet is shown


65.85.85.8

65.85.85.8

65.85.85.85.8

65.8

65.865.865.865.8

Table 2Q–Z (vertical load–vertical displacement) datasheet for a single pile for a steel-pipe pile of diameter 1829 mm

Penetration Q(1) Z(1) Q(2) Z(2) Q(3) Z(3) Q(4) Z(4) Q(5) Z(5) Q(6) Z(6) Q(7) Z(7) Q(8) Z(8)

depth (m)

24.00 0. 0 .71 3.7 1.42 23.8 2.13 76.8 2.55 133.5 2.84 182.9 2.84 274.4 2.84 326.70 0. 0. .33 3.7 .65 23.8 .98 76.8 1.17 133.5 1.30 182.9 1.30 274.4 1.30 3628.20 0. 0. .33 3.7 .65 23.8 .98 76.8 1.17 133.5 1.30 182.9 1.30 274.4 1.30 3633.70 0. 0. 2.34 3.7 4.68 23.8 7.02 76.8 8.42 133.5 9.35 182.9 9.35 274.4 9.35 337.90 0. 0. .35 3.7 .71 23.8 1.06 76.8 1.28 133.5 1.42 182.9 1.42 274.4 1.42 3640.80 0. 0. .35 3.7 .71 23.8 1.06 76.8 1.28 133.5 1.42 182.9 1.42 274.4 1.42 3641.40 0. 0. .75 3.7 1.50 23.8 2.25 76.8 2.70 133.5 3.00 182.9 3.00 274.4 3.00 342.70 0. 0. .27 3.7 .54 23.8 .81 76.8 .97 133.5 1.08 182.9 1.08 182.9 1.08 3648.50 0. 0. .35 3.7 .71 23.8 1.06 76.8 1.28 133.5 1.42 182.9 1.42 274.4 1.42 3667.80 0. 0. .44 3.7 .89 23.8 1.33 76.8 1.60 133.5 1.77 182.9 1.77 274.4 1.77 3687.20 0. 0. .53 3.7 1.06 23.8 1.60 76.8 1.92 133.5 2.13 182.9 2.13 274.4 2.13 391.90 0. 0. 2.76 3.7 5.52 23.8 8.28 76.8 9.93 133.5 11.03 182.9 11.03 274.4 11.03 365.896.00 0. 0. 1.89 3.7 3.77 23.8 5.66 76.8 6.79 133.5 7.54 182.9 7.54 274.4 7.54 3

102.40 0. 0. 1.89 3.7 3.77 23.8 5.66 76.8 6.79 133.5 7.54 182.9 7.54 274.4 7.54 3107.90 0. 0. .89 3.7 1.77 23.8 2.66 76.8 3.19 133.5 3.55 182.9 3.55 274.4 3.55 3120.80 0. 0. .89 3.7 1.77 23.8 2.66 76.8 3.19 133.5 3.55 182.9 3.55 274.4 3.55 3

Notes:Q = Load in meganewtons,Z = displacement in millimeters; the penetration depth is in meters.

todipndipd

gsThngntF

ar

rmhe

g,dthe

Fig. 13. Schematic illustration of the pile model.

Table 2. The spring parameters were input into pile modeldetermine the overall bearing capacity of the single pile anthe load (at the top of pile) versus displacement relationshFig. 14(a) shows the relationship between the axial load aaxial displacement of a single pile, while the relationshbetweenthe pile cap torsion and the angle of twist is depictein Fig. 14(b).

5.2. Structural model for the platform jacket

Fig. 15shows the structural model for the platform jacket.The model is constructed according to the design drawinPipe elements are used to mesh the platform structure.damaged member is meshed with pipe elements 0.4 m loThe other members are meshed with 2 m long pipe elemeThere are in total 2082 elements and 1907 nodes in the

.

.e.

s.E

(a) Axial load versus axial displacement.

(b) Cap torsion versus angle of twist.

Fig. 14. Load versus displacement relationship of a single pile.

model. The pile–soil interaction is simulated by non-linesprings.

The damaged diagonal bracing isφ914× 19 mm, and itsyield stress is 345 MPa. The superstructure of the platfoitself is not included in the structural model because tcollision happened during the installation.

The force–deformation curve of the non-linear sprinwhich is used to simulate the collision of the platform anthe barge and located at the point of the dent damage on


e

ea

e

rae

e

c

oc

sa

the

the

edsingers

mthe

tictheted

ce-ngtoa-

rio

Fig. 15. The platform jacket structure model.

tubular member, is obtained using Eqs. (6)–(8). The commonvelocity v12 of the platform and barge after collision can bdetermined according to the law of momentum conservationas described inSection 2. It is assumed that no secondcollision between the barge and the platform happenand the impact action of the barge is transformed intoisosceles triangle impulse load.

5.3. Dynamic response analysis for stress and displacemof the jacket structure

The dynamic response of the platform jacket structuis calculated to examine the magnitude of the stressesdisplacements at critical regions and the nodal points. Thanalysis is carried out for twoimpact loading scenarios,namely, (a) impactwith the barge velocity equal to 1.0 m/s;as described inSection 4.4, the corresponding contact timis t0 = 0.682 s and the peak impact force isF = 5.367×106 N; and (b) impact at barge velocity equal to 2.0 m/s, forwhich t0 = 2.15 s andF = 3.405× 106 N.

Fig. 16 shows the computed displacements at the impaloading point (the loading side of the non-linear collisionspring, denoted as UX1852) and at the damage locationthe diagonal brace (denoted as UX1837), for the impaloading scenario (a).Fig. 17 shows the corresponding strestime histories at several critical locations on the diagon

d,n

nt

end

t

ft

l

Fig. 16. Response time histories of displacements at the joints ofdamaged member, for barge velocity= 1 m/s (collision time = 0.682 s,peak loadF = 5.367× 106 N).

Fig. 17. Response time histories of stresses at critical regions ofdamaged member, for barge velocity= 1 m/s (collision time = 0.682 s,peak loadF = 5.367× 106 N).

tubular member, namely at its connection to leg A2 (denotas S12), at the impact location (S1837), and near the crospoint between the two cross-diagonal brace memb(S1765). The von Mises strength criteria are adopted.

It can be observed from these plots that the maximustresses of the damaged diagonal brace have reachedyield strength at the impactlocation and at its connectionto leg A2. The element stressat the crossing point betweenthe two cross-diagonal members remains within the elasrange. The element stresses in the remaining part ofjacket structure are rather small and they are not presenin detail here.

The computed results also show that the total displament of the overall jacket structure under the impact loadiis small. It is expected that they would be able restoretheir original positions, and no specific rehabilitation mesures are required concerning the displacement of the plat-form structure.

The computed responses for the impact loading scena(b), as shown inFigs. 18and 19, are generally similar to


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f

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ore

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se

89,

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ite.

Fig. 18. Response time histories of displacements at the joints of tdamaged member for ship velocity= 2 m/s (collision time= 2.15 s, peakload F = 3.405× 106 N).

Fig. 19. Response time histories of stresses at critical regions of tdamaged member for ship velocity= 2 m/s (collision time= 2.15 s, peakload F = 3.405× 106 N).

those for scenario (a). However, due to an increase of timpact duration, the primary response duration is longer, athe overall response amplitudes also appear to be higher tfor scenario (a).

On the basis of the above results, it is suggested thfurther examination of the butt weld of the damagediagonal brace and the attachment weld of this diagonbrace to leg A2 be conducted by appropriate non-destructtesting. A proper repairing and strengthening scheme for thjacket platform structure around the damaged regions shobe determined accordingly.

6. Conclusions

A comprehensive evaluation procedure is presented fassessing the damage effects to an offshore platform strture due to collision by a large barge. The computationmodel involves the following aspects: (a) the determinatioof the maximum impact load and the impact duration; this isdone on the basis of the observed damage state, particul

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the damage dent deformation, in conjunction with non-lineFE analysis for thedamage in the bracing member; (b) thincorporation of a non-linear spring to model the transmsion of the impact load to the offshore platform structure ding the collision, and the derivation of the non-linear spriproperties from the analysis ofthe elastic and plastic dendeformation of thetubular member subjected to the impac(c) the modeling of the pile–soil interaction using three typof non-linear springs; and (d) the finite element modelingthe platform jacket structure. Subsequently, the dynamicsponse of the platform structure subject to the impact loadobtained. The response histories in terms of stress at critregions and the nodal displacements are obtained for the assessment of the integrity of the structural system. Forparticular case under investigation, it is found that yieing occurred only for the diagonal brace member arouits connections to the two legs, while the remaining part ofthe structure exhibited no inelastic response. Repairing astrengthening appears to be necessary only for the dianal member which was directly hit during the collision. Thgeneral procedure presented in this paper is applicablethe damage assessment of other offshore platform structin the case of accidental collisions.

Acknowledgment

The support of China Offshore Oil Research Centergratefully acknowledged.

References

[1] Jin W-l et al. Reliability based design of jacket platform under extremloads. China Ocean Engineering 1996;10(2):145–60.

[2] API. Recommended Practice 2A-WSD (RP 2A-WSD). 21st eAmerican Petroleum Institute; 2000.

[3] DnV. Recommended Practice RP-C203, DET NORSKE VERITA(DnV); May 2000.

[4] L i R-p, Chen W-g, Gu Y-n. Static analysis of collision strength ooffshore platform. Ocean Engineering 1995;13(2):14–21 [in Chinese].

[5] Jorgen A. Energy absorption in ship–platform impacts. Report ofNorwegian Institute of Technology, The University of TrondheimSeptember 1983.

[6] Petersen MJ, Pedersen PT. Collision between ships and offshplatform. In: Proceedings of13th annual offshore technologyconference. 1981, p. 163–71.

[7] Bai Y, Pedersen PT. Elastic–plastic behavior of offshore steestructures under impact loads. International Journal of ImpEngineering 1993;13(1):99–115.

[8] Smith CS. Assessment of damage in offshore steel platforms. In:Proceedings of international conference on marine safety. 19p. 279–305 [Paper 15].

[9] Ueda Y, Murakawa H, Xiang D. Classification of dynamics responof a tubular beam under collision. In: Proceedings of 8th internationalconference on offshore mechanics and arctic engineering. 19p. 645–52.

[10] Jin W-l, Gong S-f, Song J. Final report of damage assessment anafor WEN 13-1 jacket platform. Institute of Structural EngineeringZhejiang University, 2001. p. 12.

[11] Ocean Engineering Geologic Research Report of Platform SEngineering Survey Company, COGC, 1999. p. 4.

analysis of damage of jacket with barge

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