Xiangfeng Zhang The Model Test of Deep Water S-Lay Stinger

Xiangfeng Zhang1

State Key Laboratory of Structural Analysis

and Industrial Equipment,

Dalian University of Technology,

Dalian 116024, China;

Geophysical Department Subsea Engineering

and Technical Center,

China Oilfield Services Limited,

Tianjin 300452, China

e-mail: [email protected]

Qianjin YueState Key Laboratory of Structural Analysis


Dean of School of Ocean

Science and Technology,


Dalian 116024, China


Wenshou ZhangState Key Laboratory of Structural Analysis



Dalian 116024, China


The Model Test of Deep WaterS-Lay Stinger Using DynamicalSubstructure MethodDeep water pipeline installations by S-lay present many challenges, especially in theoverbend section. The S-lay requires a long curved and stiff stinger to support thepipeline weight. The interaction between the overbend pipe and stinger is complicatedsuch that the numerical structure analysis could not sufficiently predict the mechanicalbehavior of the installing process. A dynamic substructure model test method with 1:20length scale for 2000 m water depth is addressed in this paper, where the large scalemodel structure can be tested to simulate the vessel movements during installation. Theroller forces influenced by the stinger stiffness and vessel movements are discussed basedon the test platform. [DOI: 10.1115/1.4028879]

Keywords: deep water pipelaying, stinger design, substructure test method, roller load,S-lay

1 Introduction

In S-lay technology, as outlined in Fig. 1, the pipeline could bedivided into the overbend and sagbend. During shallow waterinstallation, previous research focused on the sagbend section,where the minimum bending radius occurs at the touch downpoint, but as the water depth and pipe diameter increase, the prob-lems will move to the overbend section [1,2], where both the pipe-line and stinger will reach to their limiting state.

In the shallow water case, the stinger could be simplified as arigid structure [3–9]. With the water depth increasing, the stingermust be designed longer to support the heavy weight of the pipe-line, which causes the stiffness to become an important factor thatinfluences the dynamic interaction of the pipeline and rollers. Inprevious studies [10,11], the stinger was regarded as a rigid curve,but rollers were considered to be springs. One end of the springwas connected to the pipeline, and the other end was located on arigid arc. This approach can determine the interaction forcesbetween the pipeline and rollers and also determine the pressuredistribution along the overbend section. Furthermore, previousresearch studies [12,13] introduced a substructure method inwhich the stinger structural rigidity was translated to the boundarypoint located at rollers, and the stinger was simplified as a deform-able cure to simulate the contact statically between the pipelineand stinger. Based on this method, the static roller loads and pipestrain distributions were analyzed. The impact of the stinger’sstructural rigidity on roller loads was also discussed. However, inthe actual pipelaying process, both the pipeline and stinger struc-ture are in a marine environment and are dragged by a dynamicfloating body. The stinger’s performance as a deformable bodyunder the dynamic excitation was difficult to predict, and although

a number of research studies [14–16] have analyzed pipelinestress dynamically, but few works considered the stiffness of thestinger in this case [17,18].

This oversight is primarily attributable to the strong nonlinearissues between the stinger and the pipeline. During the laying oflarge diameter pipelines in ultra deep water, the overbend sectionof the pipeline material enters a nonlinear stage. Due to the largedeformation of the pipe, numerical methods produce nonlinearproblems. Moreover, taking the stiffness of the stinger intoaccount, uncertainties related to the pipeline and stinger contactstatus caused by hull motions will cause difficulties for theresearchers and designers in obtaining accurate roller loads. Thus,the stinger’s structural analysis and design is mostly based on theEngineer’s experience instead of a scientific design basis.

Fig. 1 S-lay

1Corresponding author.Contributed by the Ocean, Offshore, and Arctic Engineering Division of ASME

for publication in the JOURNAL OF OFFSHORE MECHANICS AND ARCTIC ENGINEERING.Manuscript received August 13, 2013; final manuscript received October 15, 2014;published online November 17, 2014. Assoc. Editor: Longbin Tao.

Journal of Offshore Mechanics and Arctic Engineering FEBRUARY 2015, Vol. 137 / 011701-1Copyright VC 2015 by ASME

Downloaded From: http://asmedigitalcollection.asme.org/ on 11/26/2014 Terms of Use: http://asme.org/terms

A model test may be an alternative approach. However, deep-water pipelaying has its own characteristics where the pipelinediameter and stinger size is small compared to the water depth. Ina conventional tank test, if a larger scale is selected to meet thespecific pipeline and stinger model design and processing require-ments, then a very deep pool should be provided in the laboratory,which is not feasible. Conversely, a smaller scale meets the labrequirements but will be not suitable for the pipeline and stingermodel designs, which makes it difficult to satisfy the model andthe lab requirements simultaneously. This difficulty may be thereason that relevant studies testing the deep-water S-lay modelhave not been published to date.

To address this issue, a substructure model test method wasproposed in this paper. As illustrated in Fig. 2, the pipelaying ves-sel, the overbend section and stinger, and the sagbend sectionwere considered as substructure 1, 2, and 3, respectively. One ofthe advantages of this method is that there are no limits to thescale of the laying system model. A large scale model test wasproposed to analyze the effects of hull motions on the dynamicpressure acting on rollers. The test results indicate that the pro-posed model-based test method can address a variety of nonlinearproblems and uncertainties in the laying process and can be usedto determine dynamic roller loads.

2 Dynamical Substructure Model of Deep

Water Stinger

As shown in Fig. 2, the entire system involved in the pipelayingprocess has been divided into three substructures, the pipelayingvessel, the overbend section and stinger, and the sagbend section.Substructure 1 provides the power input for substructure 2, andsubstructure 3 acts as a bound on substructure 2. The stinger insubstructure 2 can be analyzed separately, and its kinetic equationmay be expressed as follows:

M€x tð Þ þ C _x tð Þ þ K tð Þx tð Þ ¼ FW tð Þ þ FV tð Þ þ FR tð Þ (1)

where M, K(t), and C are the mass, stiffness, and damping matrixof the stinger, respectively. Because the shape and conditions ofthe contact between the pipeline and stinger change with time,this stiffness matrix K(t) is time varying. x(t) is the motion vectorof the pipelay vessel. The external load Fw(t) is the environmentalload, which is mainly the wave load and is expressed as follows:

FW tð Þ ¼ 1

2qDCD u

*tð Þ

�� u* tð Þ þ p4

qD2CM_u*

tð Þ (2)

where Fw(t) is the fluid load that acts on the stinger, u*

and_u*

arethe fluid velocity and acceleration, respectively, and CD and CM

are the drag and inertia coefficients.

Fv(t) is the inertial force caused by vessel motions. FR(t) is theroller loads

FR tð Þ ¼Xn

i¼1

fri tð Þ (3)

where n is the number of the rollers and fri(t) is the load acting onthe ith roller, as depicted in Fig. 3.

In Eq. (1), Fw(t) and Fv(t) can be solved numerically given thevessel motions and the pipelaying environment. The roller loadFR(t) could not be obtained through numerical computationbecause it is difficult to determine whether the roller is in contactwith the pipeline in dynamic cases.

2.1 Static Roller Load. The static roller load, the basic forcefor stinger structural strength analysis, is controlled by the geo-metric parameters and the structural stiffness of stinger by giventhe pipe and tensioner. R and K are assumed as the radius and thestructural stiffness of the stinger, respectively, and define h and Kv

by the following equations:

h ¼ aR

K ¼Xn

i¼1

ksi

Kv ¼ �K

(4)

Fig. 2 Substructures and the overbend section

Fig. 3 Mechanical model of the stinger in substructure 2

Fig. 4 Finite element model of stinger

011701-2 / Vol. 137, FEBRUARY 2015 Transactions of the ASME


where h is defined as the height beyond the ideal roller position ofarc in radial direction, a and � are dimensionless factors, Ksi is theith the local rigidity of stinger, K is the rigidity of stinger struc-ture, and Kv is used to analyze the contact behavior between pipeand stinger given different rigidity.

For a given stinger, as shown in Fig. 4, effects on the roller loadfor different a and � are shown in Figs. 5 and 6. Both of the fig-ures were obtained by the substructure method from a previousstudy [12]. In Fig. 5, the x-axis expresses the roller number andthe y-axis is the pressure acting on each roller. This figure wasused to show the pressure variation when the height of one roller(roller 5) changed, described as a percentage of the stinger radiushigher than the ideal location. From Fig. 5, it is clear that theroller load is sensitive to roller height, even if h is 0.1% of thestinger radius, the roller load will exceed the pipeline design pres-sure limit. However, if one roller height changes, it only impactsthe adjacent two rollers loads and has little impact on rollers far

away, which is important for the practical application of thestinger to transform roller height.

Figure 6 expresses the influence on roller loads from stingerstiffness. The x-axis is the value of v, as v increases, the stinger ismore rigid, and the y-axis is the force acting on the rollers. Thisfigure was to analysis roller loads when the stinger structure rigid-ity was taken into account. As you can see from this figure, whichis different from the roller height, the stinger stiffness that isdenoted by v has different effects on the rollers in different loca-tions. This difference is observed primarily because when thestinger stiffness decreases, its deformation will increase under thepipeline pressure. The point that the pipeline leaves the stingermoves near the vessel, this makes the active length of the stingerdecrease, when v decreases to a certain extent, certain rollers willseparate from the pipeline and their loads will be 0. At the sametime, certain rollers loads will increase correspondingly, such asroller 4 and roller 1.

Static roller loads could be obtained and analyzed by numericalmethod, but the dynamic roller loads caused by hull motions aremore important.

2.2 Dynamic Roller Loads. Because the stiffness matrix K(t)in Eq. (1) is highly nonlinear and time depended, it is difficult tosolve this equation and also Eq. (3). Thus, the contact statebetween the stinger and the pipe is not easily determined; there-fore, the roller loads cannot be readily obtained. A model experi-ment is an effective approach for solving these problems. In thisexperiment, the overbend and stinger is regarded as a separatesubstructure. Based on appropriate assumptions, experimentalanalyses for this substructure can be conducted separately. Thedynamic substructure method can be performed with a large scalesuch that the laboratory space can be fully utilized by constructingmodels to test this substructure. In this paper, a 1:20 large scalestinger model was tested. Using this model, the effects of themovements of the pipelaying vessel were considered and thedynamic pipeline pressure between the stinger and pipeline(the dynamic roller load) was analyzed.

Fig. 5 Concentrated roller loads

Fig. 6 Variations of roller forces with m Fig. 7 The principle of the substructure simulation platform

Journal of Offshore Mechanics and Arctic Engineering FEBRUARY 2015, Vol. 137 / 011701-3


3 Test Setup

In deep and ultra deep water pipelaying, the flexibility of thesagbend pipeline will be close to the catenary, and during thedeep water pipelaying process, the inflection point of the pipelinevaries within a small range. By assuming that the scope of thisvariation is negligible relative to the water depth in which the pipeis laid, substructure 3 can be regarded as a spring–mass system. Inaddition, dynamic roller loads are primarily caused by vesselmotions. Due to the installation are typical selecting a very calmsea environment, hence the roller loads caused by the marine envi-ronment are ignored. In this paper, the response of the stinger andthe overbend under vessel motions are primarily analyzed. Sub-structure 1, the pipelay vessel, is considered to be rigid. The hullmotions are achieved by a motion platform, and the effect of thepipeline and stinger acting on the hull is not considered, therebyallowing the focus to remain on the contact state of substructure 2,as shown in Fig. 7. Compared with other tests of deep water struc-ture models, a notable feature of this test is that the stinger and theoverbend section were simplified as a substructure, and a large

scale of 1:20 was used for simulating pipelaying in a water depthof 2000 m.

In dynamic cases, the dynamic similarity principle wasemployed. The frequency ratio between the model and the proto-type of the stinger structure and vessel motions is the same.According to geometric scale, kL¼ 1:20 and the constructedstinger model, the frequency scale is kx¼ 4.47. All three substruc-tures were designed based on the scale, detailed parameters of thestinger and pipeline are provided in Table 1. Substructure 1 wasdesigned to simulate the hull motions acting on substructure 2(Fig. 8). The primary geometric parameters of substructure 2including the overbend and stinger were designed based on thegeometric scale, especially the outer diameter and the wall thick-ness of the pipe, the length and the radius of curvature of thestinger, the space and the length of rollers. The stinger model hadthree sections, connected by hinges at the bottom and a link beamat the top. The length of the link beam could be changed to fit dif-ferent pipeline laying (Fig. 9). For substructure 3, the sagbendweight and the departure angle at the top of the sagbend sectionwas calculated. Substructure 1 was controlled by the control sys-tem to simulate the vessel motions during laying. Substructures 2and 3 were driven by substructure 1 to simulate the real pipelay-ing, and data were collected by sensors and a measuring system,as shown in Fig. 10.

The roller numbering scheme is presented in Fig. 7. To deter-mine real laying conditions, static and dynamic model tests wereperformed in the laboratory.

4 Test Results and Analysis

4.1 Static Tests. The static model tests were intended to ver-ify the reliability of the experimental approach and ensure that anappropriate roller height was used.

The roller height should be adjusted several times beforeobtaining an ideal roller height to have a more evenly distributedroller loading among all rollers. Cases 1 and 2 are two cases of

Table 1 Characteristics of the pipelaying model

Prototype Model

Stinger Length (m) 88 4.4Radius (m) 73 3.65Weight (T) 500 0.25Fundamental frequency (Hz) 1.42 6.26

Roller Quantity 10 10Length (m) 9 0.45Spacing (m) 3 0.15

Pipeline Thickness (m) 0.0381 0.002Outer diameter (m) 0.3239 0.0162Departure angle (deg) 83.7 83.7Sagbend weight (T) 334.886 0.04186

Fig. 8 Substructure models: (a) stinger and overbend model, (b) sagbend model, and (c)pipelay vessel model



improper height arrangements, and case 3 is an appropriate one.Cases 1–3 show the static roller load distributions during rollerheight adjustment in lab, as displayed in Fig. 11. This figurereveals that under extreme laying conditions, the pipeline pressuredistribution and the roller distribution state are related. If the roll-ers are not aligned along the same arc, the roller load distributionis uneven, thereby causing excessive concentration of this load, asshown in Fig. 5. As the rollers become increasingly aligned alongthe same arc, the pipeline pressure distributions become increas-ingly uniform. However, even though the rollers were on thesame arc, pressures acting on the rollers in various locations haddifferent values, while roller 8 always had the greatest load. Thisproperty is observed primarily because the last roller and pipelinegenerally cannot be in contact with each other during deep-waterlaying. In this situation, roller 9 did not come in contact with thepipeline completely, thereby leading to a concentrated load onroller 8. Finally, the test results in case 3 were compared withresults calculated by a previous study [12] and the commercial

program Offpipe, as shown in Fig. 11, from which we can see theresults are similar. This finding verified the feasibility of the sub-structure test method.

4.2 Dynamic Tests. Based on the static test results, a suitablespatial location for the rollers was obtained. Next, dynamic modeltests of substructure 2 were performed. In the dynamic case, ves-sel motions are needed. Pipelay vessel motions with 0 deg, 45 deg,and 90 deg wave approach angles were provided by ShanghaiJiaotong University, and partial cycles of hull motions in the timedomain are presented in Fig. 12. These motions were used by sub-structure 1 and applied to substructure 2. Dynamic roller load testsin the constructed model were conducted for different hullmotions and wave approach angles. The roller load responses areillustrated in Figs. 13 and 14 below.

Both Figs. 13 and 14 are in the time domain with the y axisshowing the pressure acting on the rollers, as obtained by the datacollection system in the lab. Among the ten rollers used in theexperimental tests, three rollers were selected, rollers 2, 5, and 8,located in the middle of each stinger section. The dynamic load ofthese rollers was assessed. Fig. 13 was used to contrast dynamicroller loads in different locations on the stinger, and Fig. 14showed the analysis of the loads with different wave approachangles.

Fig. 9 The pup piece and the roller arrangement

Fig. 10 Pipelay models Fig. 11 Static roller loads



The comparison of Figs. 13 and 14 revealed that the amplitudeof the dynamic loads acting on the roller, changing in a periodicmanner similar to the movement cycle of the hull, was propor-tional to the distance between rollers and the hull. Thus, furtherdistances from the hull were associated with greater amplitudes.Moreover, the mean value of the dynamic roller loads and thestatic load exhibited different relationships at different positions.As the distance decreased, the mean dynamic loads became lessthan the static load, and for rollers distant from the hull, the meandynamic loads were slightly greater than the static load. For roll-ers in intermediate locations, the mean dynamic load was similarto the static load. This phenomenon relates to the structural rigid-ity distribution of the stinger and is consistent with conclusions inthe extant literature [12].

Meanwhile, the changes of the dynamic roller loads caused byhull motions are more complex. The primary forms of hull motionthat affect the roller load are pitch and heave. Heave producesvertical accelerations. However, pitch generates tangentialaccelerations that are distributed differently at each roller, therebyproducing more complicated effects on the dynamic loads. Thus,among hull motions, pitch has the most dramatic effect on

Fig. 12 Partial cycles of hull motions: (a) Heave, (b) roll, and(c) pitch

Fig. 13 Dynamic roller loads for different roller locations: (a)roller 2, (b) roller 5, and (c) roller 8



pipelaying (Fig. 13), especially when the wave approach anglesare 0 (Fig. 14).

4.3 Discussion. One difference between the numerical analy-sis and the model test is that the model test considers the effectsof the dynamic rigidity of the stinger on the roller load.

4.3.1 Ideal Rigid Model. If the stinger is assumed to be a rigidbody, the dynamic load of the rollers caused by hull movementscan be expressed as follows [19]:

DRf ¼ dDT (5)

where DRf represents the dynamic load of the rollers caused bythe inertial dynamic load of substructure 3. d¼ S/q is a dimen-sionless quantity, where S is the distance between rollers and q isthe radius of curvature of the stinger. DT is the inertial load of thesagbend section of the pipeline caused by hull motions. In thisideal model, the stinger is regarded as a rigid body, and thedynamic loads acting on the rollers will be the same, irrelevant totheir locations.

4.3.2 Elastomer Model in Static Case. As a deformable body,the stinger will experience radial deformation due to static pres-sure from the pipeline. In this case, compared with the ideal rigidmodel, changes in the contacts between the stinger and the pipewill occur, as depicted in Fig. 15. From Fig. 15 we can see thatunder the huge pressure, compared with the rigid stinger, theelastomer stinger had a little deformation. This effect caused anumber of rollers to miss contact with the pipe, while certain roll-ers were pressed tighter and this was also verified by Figs. 6and 11 (cases 1–3).

Furthermore, assuming that the uplift point position is fixed,this point divides the stinger into two sections. The section fromthe uplift point to the stinger tail may be regarded as a cantilever,this section exhibits minimal rigidity but has maximal deforma-tion. The section of the stinger between the main hinge point andthe uplift point can be viewed as a simply supported beam. Rela-tive to the cantilever, the latter section will be more rigid andhave less deformation, which means that the contact status is dif-ferent along the deformable stinger, and the rollers on differentsections will have various loads, as shown in Fig. 11 (cases 1–3).This finding is different from the ideal rigid model.

4.3.3 Dynamic Roller Loads. The ratio between the dynamicload amplitude and static load of a roller is defined as the dynamicload factor, which can be expressed as follows:

DRd ¼ Ai=Ri (6)

where DRd is the dynamic load factor of the ith roller, Ai and Ri

are the dynamic load amplitude and static load of the ith roller,respectively.

At wave approach angles of 0 deg, by using Eq. (6) to analyzeexperimental data (Figs. 13 and 14(a)), the factor DRd and themean value of dynamic loads for each roller were obtained. Theenvelope of the increments was shown in Fig. 16. As indicated inthis figure, the dynamic load factors are different for each roller.These factors increase as the rollers move further away from the

Fig. 14 Dynamic roller loads for different wave approachangles: (a) 0 deg, (b) 45 deg, and (c) 90 deg

Fig. 15 The impact of stinger structural rigidity on rollercontact states



hull. For the rollers between the hull and the uplift point (rollers1–3), the dynamic load factors are small, and changes in these fac-tors are relatively flat. The dynamic load factors for the two rollersnear the uplift point are similar and are both slightly larger thanthe factors for rollers 1–3. As rollers become further away fromthe uplift point, their dynamic load factors gradually increase. Thedynamic load factor of roller 8, under the examined experimentalconditions reached 5.3%. This roller bears the maximum staticpressure among the rollers, and this large factor could be verydangerous during laying.

5 Conclusions

By treating the pipelay vessel, the overbend section and stingerand the sagbend as different substructures, this paper established asubstructure model for experimental testing. In this approach, alarge scale simulation of the stinger and overbend section of thepipeline can be used. This approach provides the research fordynamic roller load analyses and stinger design in deep waterS-lay and was conducted for a particular deep water stinger toassess dynamic roller loads and their changes.

Model tests revealed that dynamic roller load distributions areperiodic, similar to the periods of hull motions. Furthermore,changes and mean values of dynamic load amplitudes are propor-tional to the distance from the pipelaying vessel, rollers far fromthe stinger sustain greater changes and mean values. Meanwhile,the mean values near the vessel are slightly less than their staticloads. As the distance from the hull increases, the roller’s meandynamic load gradually increases and eventually exceeds theirstatic load and attains the maximum values at the end of thestinger. Near the uplift point, the mean dynamic loads and staticloads are similar. The magnitude of the incremental dynamicroller loads behaves in a similar way. The magnitude is also pro-portional to the spacing between the rollers and vessel, and withlonger distances, the dynamic roller load factors are greater.Under the tested experimental conditions, the roller with the great-est static load is also subjected to the greatest dynamic load. Thisphenomenon should be carefully considered by designers of

pipelaying and stinger structures. Furthermore, heave and pitchare the primary forms of hull motions that affect dynamic rollerloads. Pitch is maximized at 0 deg of wave approach angles andproduces the most severe changes in dynamic roller loads.

Acknowledgment

The financial supports for this research were provided by theNational High Technology Research and Development Programof China (No. 2006AA09A105-2). The first author also would liketo thank Professor Ren Ping at Shang Hai Jiao Tong Universityfor his kind help to offer the hydrodynamic vessel motion testresults.

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Fig. 16 The DRd envelope



http://dx.doi.org/10.1016/S0029-8018(96)00012-1

http://dx.doi.org/10.1016/S0141-1187(99)00020-6

http://dx.doi.org/10.3969/j.issn.1000-0879.2002.02.006

http://dx.doi.org/10.1115/1.4004628

http://dx.doi.org/10.4028/www.scientific.net/AMR.443-444.138

http://dx.doi.org/10.4028/www.scientific.net/AMR.443-444.138

http://dx.doi.org/10.7511/jslx201102023

http://dx.doi.org/10.1016/0045-7949(90)90339-4

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Xiangfeng Zhang The Model Test of Deep Water S-Lay Stinger