Modeling of Severe Slugand Slug Control With OLGA

Mehrdad P. Fard, Hydro, Oil & Energy Brage Petroleum Technology; John-Morten Godhavn, Statoil ASA;and Svein Ivar Sagatun, SPE, Hydro, Oil & Energy Troll Petroleum Technology

SummaryWith the recent developments in the oil and gas industry, in whichthe number of satellite fields are increasing, the industry facesmajor challenges. A mixture of gas and oil is transported in mul-tiphase pipelines along the bottom of the sea from the wells and upthrough the riser to the oil rig. The development of unstable flowin multiphase pipelines is a major and expensive problem. Theirregular flow results in poor oil/water separation, limits the pro-duction capacity, and causes flaring. Handling slug flow thus hasbecome very important with regard to increasing the productionrate and avoiding possible equipment damages.

In this paper, we have developed a model of a physical processthat generates slugs. The model has been verified against experi-mental data. Several control strategies have been tested on themodel, and the simulation results are presented. It is shown bysimulation of multiphase flow that the unstable slug flow can bestabilized by feedback control. In addition, interesting and impor-tant phenomena such as inverse response of the pressure at the topand asymmetric response of the pressure at the bottom have beenrevealed and explained.

IntroductionMultiphase pipelines connecting remote wellhead platforms andsubsea wells are a common feature of offshore oil production.Recent developments indicate that this trend will be followed inthe future. In addition, the feasibility of using long-distance tie-back pipelines to connect subsea processing units directly to on-shore processing plants makes it likely that these will also bedeployed in the future. These trends in offshore oil production facea major challenge.

The slug phenomenon in multiphase flow lines with both gasand liquid hydrocarbons has gained increased interest in recentyears. Slugging can be characterized as either hydrodynamic orterrain slugging. Hydrodynamic slugs are built in horizontal partsof the pipeline when liquid and gas velocities are different. Theseslugs are usually short with higher frequency. The inlet separatorwill in most cases handle these slugs with fewer difficulties, be-cause the amount of liquid is small compared to the volume of theseparator. On the contrary, a terrain slug can contain a lot of liquidand represents a great challenge to the downstream processingsystem. Level and pressure variations in the first-stage separatorcan be propagated in the downstream process. Variations in theseparator level may cause poor separation and occasionally fluidflooding. Oscillations in the separator pressure can result in in-creased flaring, which is environmentally benign and is also costlybecause of CO2 taxes.

The degree of slugging depends on various factors. The mostimportant are pipeline pressure, pipeline topology, and productionrate. Terrain slugging is most likely to occur at low rates, with alow pipeline pressure toward the end of a well’s lifetime.

Considerable operational problems and equipment damagescaused by slugs have been reported. Some of the typical problemsencountered at various platforms include the following:

• Large disturbances in the separator train, causing� Poor separation with high water-in-oil percentage in theoil outlet.� Varying water quality at the separator water outlets, which inturn causes difficulties for downstream water-treatment system.

• Large and rapidly varying compressor operations caused bypressure variations.

• Large pressure fluctuations, which cause reductions in liftingcapacity of the wells.

Various attempts have been made to avoid the problems asso-ciated with slug flow (Hollenberg et al. 1995; Havre and Dalsmo2001; Skofteland and Godhavn 2003; Havre and Stray 1999; Cour-bot 1996). Some of these are design related, such as installing slugcatchers onshore or increasing the size of the first-stage separatorsto provide the necessary buffer capacity. However, such solutionsare not appropriate because of high costs and the need for largerspace capacity.

This paper describes the OLGA 2000* simulation model of aphysical process and compares the simulation results with experi-mental data. Furthermore, this paper shows how the OLGA modelcan be used in the control-design phase before testing.

Phenomena such as inverse response of the pressure at the topand asymmetric response of the pressure at the bottom have beenrevealed and explained. These phenomena play crucial rules in thedesign of controller strategies for stabilizing the unstable slugflow. This paper is organized as follows. The modeled test loop ispresented in the Loop Description section, the OLGA modeling ofthe Tiller Loop section is devoted to modeling and verificationissues, and some interesting and important observations are dis-cussed in the Observed Phenomena section. In the Slug ControlWith OLGA Metlab Toolbox section, slug suppression by meansof active control has been discussed and several control strategiesare tested on the developed model. And finally, some concludingremarks are made in the Conclusions section.

Loop DescriptionThe process to be modeled is an experimental setup at the SINTEFPetroleum Research Multiphase Flow Laboratory (Hollenberg et al.1995) at Tiller, outside of Trondheim. The gas SF6 and the liquidExxsol** D80 are mixed in a 231-m, 3-in. closed loop with avertical riser for circulation of oil and gas. The process equipmentsbetween the inlet and outlet of the loop are shown in Fig. 1. Thefirst 100 m have a −0.1° declination and then a 180° horizontalU-turn. The diameter of the U-turn is 3.5 m, corresponding to alength of 70 diameters. After the U-turn, the pipe is declined −0.7°for approximately 105 m and finally there is a 15-m vertical riser.The riser ends in a double bend, where the flow is directed downinto an 8-in. vertical-drop leg. The vertical-drop leg ends in agas/liquid separator, where the gas is drawn into a demister toremove droplets and then into the compressor. The oil is drained tothe horizontal separator and recycled through the oil pump. At theinlet section, the gas is mixed with the oil through a 45° down-ward-inclined pipe. The separated phases are fed to the single-phase velocity-measurement stations and routed through flowme-ters by manually operated valves. The oil and gas then passthrough a 7-m-long flexible (rubber) pipe section and onto theinitial −0.1° section. The scientific instrumentations are located

Copyright © 2006 Society of Petroleum Engineers

This paper (SPE 84685) was first presented at the 2004 SPE International Symposium andExhibition on Formation Damage Control, Lafayette, Louisiana, 18–20 February, and re-vised for publication. Original manuscript received for review 28 March 2003. Revisedmanuscript received 28 February 2005. Paper peer approved 02 March 2005.

* OLGA 2000 is a trademarked product of Scandpower, Oslo, Norway.** Exxsol is a registered trademark of the ExxonMobil Corp., Irving, Texas (February 2004).

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along the entire loop but mostly along the last 100-m section andin the riser. The holdup is measured by means of seven single-energy narrow-beam gamma densitometers distributed along thepipe, with the two last densitometers in the riser.

The frequency-density responses presented in this paper aregenerated using the fast Fourier transformation function of Matlab.

OLGA Modeling of Tiller LoopThe main objective of this section is to develop an OLGA modelthat captures the main physical mechanism of the slugs generatedin the Tiller loop. The model developed in this section shows somediscrepancies such as the offset in valve openings and the over-calculation of the main frequency components in the top pressure.The model is intended to be used in the control-design stage of thestudy; therefore, it is important that the flow- and pressure-oscillation frequencies and amplitudes are close to the real data.This point has been confirmed in our verification study. The terrainslug mechanism of the slug-tracking module of OLGA 2000 (ver-sion 2.03) is used in the simulations. The default value of delayconstant is used.

Loop Topology. The loop has been modeled with the follow-ing topology:

• 100-m pipe inclined −0.1°.• 10-m flexible pipe for U-turn.• 105-m pipe inclined −0.7°.• 15-m vertical riser.The x-y coordinates of the pipelines are shown in Table 1. The

topology of the pipeline is shown in Fig. 2.The pipeline has been modeled as 3-in. pipe with an inner

diameter of 69 mm. An internal roughness of 3e-5 was assumed.The pipeline is divided into pipe elements of constant cross-

sectional areas with various lengths. Each pipe is subdivided intosections. The number and the length of sections vary for each pipeelement. The pipeline model consists of seven pipe elements with50 sections.

Boundary Conditions. The pressure at the inlet separator is as-sumed to be constant at 1.2 bar. Therefore, the outlet boundarycondition is modeled as a constant. To avoid numerical distur-bances introduced in the model, an extra horizontal pipe is addedto the top of the riser.

The pipeline inlet is modeled as a source with a specified totalmass flow rate. This module ensures a constant mass flow rate intothe loop. The simulations were performed by a constant mass flowrate of 0.8 kg/s. The gas mass fraction is set to 0.0023.

Model Verification. A simulation is run to verify the modelagainst the experimental data. The simulation is run with a fullyopened valve, as in the experiment. The comparison of thepressures at the top and the bottom of the riser is shown inFigs. 3 and 4. It is shown that the frequency and the amplitudematch very well in the case of pressure at the bottom of the riser.However, there is some discrepancy between the simulated andmeasured pressure at the top of the riser. From the frequencydensity of the upper plot in Fig. 4, we can see that the mainfrequency component of top-pressure response computed byOLGA is almost twice the measured top pressure.

For further verification of the developed model for dynamicresponses, a 11⁄2-hour scenario is simulated. The same scenario isrun as an experiment to collect data for verification of the model(Table 2).

The controller parameters for both simulation and experimentare given in Table 3.

Note that the valve opening in the OLGA 2000 varies betweenthe 0 and 1 intervals, while in the experiment, the valve openingcan vary between the 0 and 100 intervals. This is the reason whythe gains are different.

Fig. 2—The topology of the pipeline.

Fig. 1—Diagram of the Tiller loop.

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The responses are compared and shown in Figs. 5 and 6. Thevalve openings are shown in Fig. 7. It can be seen that there is adiscrepancy in the valve opening. The mean value of the valveopening from measurements is approximately 24% with automaticcontrol. However, the mean value of the simulated valve openingis approximately 46%.

Observed PhenomenaThe following observations in the simulations are in accordancewith what have been reported from corresponding experiments inGodhavn et al. (2005). These behaviors may have a more signifi-cant impact on the responses in a larger-scale system where thelength of the riser is significantly larger.

Asymmetric Pressure Response. An interesting observation wasmade with regard to the step response of the pressure. A smallincrease in the valve opening during stable flow (Fig. 8) results ina much faster pressure reduction in the bottom of the riser than thepressure increase obtained with a similar small decrease in thevalve opening (Fig. 9). A decreased valve opening increases theamount of liquid in the riser. The pressure buildup is limited by theliquid flow rate into the riser and is, therefore, slow. An increasedvalve opening reduces the amount of liquid in the riser. The liquidis replaced by (compressed) gas in the pipeline that will expand inthe riser and is, therefore, fast. The asymmetry cannot be handledproperly with linear controllers. The controller must be tuned withrespect to the fastest response. This means, for example, that apressure controller must be evaluated when the pressure referenceis decreased. The performance with an increased pressure refer-ence will then be slow.

Inverse Response at Top Pressure. An alternative to using bot-tom-pressure measurements is to use pressure measurements up-stream of the topside choke (top pressure). This alternative has theadvantage of being independent of underwater instrumentation.Bottom-pressure measurements are, in most cases, less availableand more expensive than topside measurements. However, the toppressure behaves differently from the bottom pressure. The toppressure is given by the choke downstream pressure (separatorpressure) and the differential pressure over the choke. The latter isagain given by the rate flowing through the valve, while the bottompressure is given by the top pressure plus hydrostatic pressuregiven by the amount of liquid in the riser.

We observed that the top pressure has an inverse responsecompared with the bottom pressure on steps in the valve opening(open-loop experiment—no feedback). A small opening of thechoke during stable flow (almost constant pressure both at the topand at the bottom of the riser) gives a pressure fall in the bottom

Fig. 4—Frequency density and time response of the pressure atthe bottom of the riser.

Fig. 3—Frequency density and time response for the pressureat the top.

Fig. 5—Comparison between the simulated and measured pres-sures at the bottom of the riser.

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of the riser, as expected. However, we observed a temporary pres-sure increase in the top of the riser before the top pressure wasdecreased to a new stationary value (Fig. 8). Similarly, a smallclosure of the valve gives the expected pressure increase in thebottom and a more surprising pressure decrease at the top, beforethe expected stationary increase (Fig. 9).

A decreased valve opening gives a reduced mass rate throughthe valve and out of the riser. And, with a constant rate flowinginto the riser, liquid accumulates in the riser and gives an increasedbottom pressure. The relative velocity between the gas and liquidphases (slip velocity) in the top of the riser increases, resulting inless liquid (mass) through the valve. A reduced mass rate throughthe valve results in a reduced differential pressure across the valve.The same reduction in the differential pressure across the valvealso will take place at the top pressure, because the separatorpressure is constant. Liquid is now accumulated in the riser be-cause less liquid flows out, and the flow into the riser is constant.The liquid buildup and bottom-pressure increase constitute the firstpart of a pressure oscillation toward a new equilibrium. This canresult in a riser slug if the valve-opening change is sufficientlylarge. Similarly, an increased valve opening gives the following:

• An increased mass rate out of the riser.• Reduced relative velocity between gas and liquid.• Increased differential pressure across the valve.• Increased top pressure.

• Less liquid in the riser.• Reduced bottom pressure.This behavior of the top pressure imposes limitations on the

performance of the system from a control point of view. This issuewill be addressed later in the paper.

Slug Control With OLGA/Matlab ToolboxThis section is dedicated to the control issues of the model devel-oped in the OLGA Modeling of the Tiller Loop section. Scand-power, in cooperation with Norsk Hydro ASA, has developed anOLGA/Matlab toolbox (Fard 2001). This toolbox establishes aclient/server connection between the Matlab and the OLGA 2000environment through the OLGA server, so that the simulationresults from OLGA 2000 become available for Matlab online. Thistoolbox allows for advanced process simulation and control. In thefollowing sections, all the control schemes are implemented in theMatlab environment, and the OLGA model is controlled andsteered from a Matlab application.

Single-Loop Bottom-Pressure Control. This control scheme usesfeedback from pressure at the bottom of the riser. The bottom riserpressure will increase when the amount of liquid in the riser in-creases. The idea is to stabilize the amount of liquid in the riser byusing the topside choke and, in this way, avoid slugs from buildingup. The control structure is shown in Fig. 10.

Fig. 6—Comparison between the simulated and measured pres-sures at the top of the riser.

Fig. 7—Comparison between the simulated and measuredvalve opening.

Fig. 8—Inverse response at the top pressure with increasingvalve opening.

Fig. 9—Inverse response at the top pressure with decreasingvalve opening.

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The controller parameters are shown in Table 4. The simula-tion starts with a fully opened valve. The slug period is approxi-mately 204 seconds. The controller is then switched on after 350seconds. The set point is reached after almost 10 minutes. The topand bottom pressures are shown in Fig. 11, while the volumetricflow rate and the valve opening are shown in Fig. 12.

Cascade Control of Bottom Pressure and Flow Rate. The cas-cade-control scheme used here is shown schematically in Fig. 13.The master (i.e., outer) controller is a pressure controller, while theslave (i.e., inner) controller is a flow controller. Both controllersare of standard proportional and integral type. The controller pa-rameters are shown in Table 5. The simulation starts with a fullyopened valve. The slug period is 200 seconds. The controller isswitched on 350 seconds after start. The slug is immediately sup-pressed, and the set point is reached after 4 minutes. The pressures,flow rate, and valve opening are given in Figs. 14 and 15.

Cascade Control of Top Pressure and Flow Rate. A disadvan-tage of former control schemes is the need for underwater mea-surement of pressure. This is a costly and sometimes unreliablemeasurement. The control scheme in this section removes thisdisadvantage by using the pressure upstream of the choke as acontrol variable in the master loop. This control structure is shownin Fig. 16.

It is clear from simulation results that the top pressure showssome sort of nonminimum-phase behavior. This is a well-known

fact that the bandwidth of such a system is upperbounded. Theinverse response is a constraint that cannot be removed with linearfeedback control. The controller must, therefore, have low propor-tional gain and high integration time. The issue of nonminimum-phase behavior of top pressure has been discussed in more detail inthe Inverse Response at Top Pressure section.

This controller was difficult to tune. As is often seen in prac-tical cases in the fields, first the valve was choked manually to30% opening; after a short period of time, the controller took over.The responses are shown in Figs. 17 and 18. It is clear fromFig. 18 that, with the controller being active, we have been able toachieve a lower top pressure, which in turn results in a higherproduction rate.

ConclusionsA dynamic OLGA multiphase-simulation model of the Tiller loophas been developed and verified against test data. The model hascaptured the physical mechanisms of the slugs generated in theTiller loop. Important phenomena such as inverse response of thetop pressure and asymmetric step response of the bottom pressurehave been reported.

Several control strategies have been tested on the developedmodel. It appears that a cascade-control strategy with feedbackfrom bottom pressure and flow rate at the top of the riser is bestcapable of suppressing the slugs. However, this control strategyrequires pressure measurements at the bottom of the riser. In ab-sence of such measurements, an alternative strategy must be imple-mented, as described in Cascade Control of Top Pressure and FlowRate. This control strategy has a weakness: Because of the non-minimum-phase behavior of the system, it suffers from limitedbandwidth and, therefore, slow control action. Simulation showedthat the controller was not capable of suppressing the slugs byitself. The choke was first choked back, and then the controllertook over. After that, the controller was capable of achieving andmaintaining various set points. However, this is not a major draw-back. In practice, a well is stabilized manually with a higher bot-tom pressure. Then, the controller takes over and decreases thebottom pressure gradually.

The controller with feedback from the bottom pressure is pref-erable because the controller is faster and reaches its pressure setpoint sooner than a simple bottom-pressure controller.

Fig. 10—Control structure with feedback from pressure atthe bottom.

Fig. 11—Simulated pressure response at the top and bottom ofthe riser with the single-pressure controller.

Fig. 12—Simulated flow response and valve opening with thesingle-pressure controller.

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NomenclatureKp � controller gainTj � controller integral time

ReferencesCourbot, A. 1996. Prevention of Severe Slugging in the Dunbar 16� Mul-

tiphase Pipeline. Paper OTC 8196 presented at the Offshore Technol-ogy Conference, Houston, 6–9 May.

Fard, M.P. 2001. The OLGA-Matlab Toolbox. Technical Report No. NH-00036701. Bergen, Norway: Norsk Hydro ASA.

Godhavn, J.M., Fard, M., and Fuchs, P.H. 2005. New Slug Control Strat-egies, Tuning and Experimental Results. J. of Process Control 15 (5):547–557.

Havre, K. and Dalsmo, M. 2001. Active Feedback Control as the Solutionto Severe Slugging. Paper SPE 71540 presented at the SPE AnnualTechnical Conference and Exhibition, New Orleans, 30 September–3 October.

Havre, K. and Stray, H. 1999. Stabilization of Terrain Induced Slug Flowin Multi Phase Pipelines. Trondheim, Norway: Servomøtet.

Hollenberg, J.F., Wolf, S., and Meiring, W.J. 1995. A Method to SuppressSevere Slugging in Flow Line Riser Systems. Proc. of the 7th BHRGroup Intl. Conference on Multiphase Production, Cannes, France.

Skofteland, G. and Godhavn, J.-M. 2003. Suppression of Slugs in Mul-tiphase Flow Lines by Active Use of Topside Choke—Field Experi-ence and Experimental Results. Proc. of the 11th BHR Group Intl.Conference on Multiphase Production, San Romeo, Italy, 11–13 June.

Mehrdad P. Fard is currently with Hydro, Oil & Energy BragePetroleum Technology. Before that, he was a teaching assis-tant at Norwegian U. of Technology and Science (NTNU),Trondheim, Norway. Fard holds an MSc degree in engineeringcybernetics from the U. of Technology, Trondheim, and a PhDdegree in engineering cybernetics from NTNU. John-MortenGodhavn is currently with Statoil ASA. Before that, he workedfor Kongsberg Seatex ASA. Godhavn holds an MSc degree inengineering cybernetics from NTH and a PhD degree in engi-neering cybernetics from NTNU. Svein Ivar Sagatun is currentlywith Hydro, Oil & Energy Troll Petroleum Technology. Beforethat, he worked for ABB and TTS. Sagatun holds an MSc degreein naval architecture from NTH and a PhD degree in engineer-ing cybernetics from NTNU.

Fig. 13—Cascade-control structure of bottom pressure andvolumetric flow.

Fig. 14—Simulated pressure response at the top and bottom ofthe riser with the cascade controller.

Fig. 15—Simulated flow rate and valve opening with the cas-cade controller.

Fig. 16—Cascade-control structure of top pressure and volu-metric flow.

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Fig. 17—Simulation: pressure responses with cascade control,feedback from top pressure, and volumetric flow rate.

Fig. 18—Simulation: flow rate and valve opening with feedbackfrom cascade control, top pressure, and volumetric flow rate.

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