SPE-171371-MS

Improved Artificial Lift Design for Solvent Assisted SAGD Processes

Babatunde Babayemi, Xingru Wu, and Suresh Sharma, University of Oklahoma

Copyright 2014, Society of Petroleum Engineers

This paper was prepared for presentation at the SPE Artificial Lift Conference & Exhibition-North America held in Houston, Texas, USA, 68 October 2014.

This paper was selected for presentation by an SPE program committee followingreview of information contained in an abstract submitted by the author(s). Contentsof the paper have not been reviewed by the Society of Petroleum Engineers and are subject to correction by the author(s). The material does not necessarily reflectany position of the Society of Petroleum Engineers, its officers, or members. Electronic reproduction, distribution, or storage of any part of this paper without the writtenconsent of the Society of Petroleum Engineers is prohibited. Permission to reproduce in print is restricted to an abstract of not more than 300 words; illustrations maynot be copied. The abstract must contain conspicuous acknowledgment of SPE copyright.

Abstract

Artificial lift design for heavy oil systems is a continual changing process in which evolving technologicaladvancements coupled with constant learning experience has led to production capabilities that were notfeasible in the past. There has been an increase in installation of temperature tolerant Electric Submersiblepumps (ESP) with the primary aim of improving deliverability from mature heavy oil fields. Similarly, inthe heavy oil industry, the desire to attain energy efficiency has birthed variations of the Steam AssistedGravity Drainage (SAGD) process that use solvent additives in further reducing bitumen viscosity. Theheavy oil emulsion formed in these systems exhibit rheological characteristics and outflow behaviorsdifferent to conventional SAGD systems. Tubing hydraulic performance, steam trap requirements andflow assurance behaviors associated with production from these systems are investigated in this work.

This paper presents dynamic multiphase simulations detailing fluid flow regimes, mass and heattransfer mechanisms and pressure/temperature changes as the reservoir fluid flows out of the reservoir,through an ESP and up the production tubing to the surface. The reservoir is a 3D fully coupledreservoir/wellbore model with properties similar to the Athabasca bitumen reservoir. The simulation wasconducted considering all periods in the lifecycle of production i.e. pre and post ramp up. This researchfinds that a detailed understanding of fluid phase behavior and reservoir operating parameters during thedifferent periods can dramatically improve operating efficiency and impact on ESP design.

Furthermore, results generated from this research can be used as a yardstick for SAGD productionengineers in designing artificial lift systems for solvent assisted processes

IntroductionOver the years, there have been improvements in bitumen exploitation strategies with Steam AssistedGravity Drainage (SAGD) operations recording significant success. SAGD is a heavy oil recovery processthat uses steam injection for heavy oil production as resources are becoming too deep for extractionthrough mining operations. The SAGD process involves drilling two horizontal wells with the injectoroften located a few meters above the producer. The steam injected heats up the reservoir leading toreduction in heavy oil viscosity, enhanced mobility and eventually initiates drainage to the producer.Figure 1 illustrates the cross section of the SAGD process and shows the process of steam injection andemulsion drainage to the producer.

The SAGD process was initially modelled by Butler et al. (1981) and is as follows:

The equation shows that reducing the kinematic viscosity of oil in place, v, contributes significantly tooil production. Hence, the ability to effectively reduce the viscosity of bitumen at every point duringproduction is an important attribute that impacts reservoir deliverability. This has led to continuous effortand research aimed at developing SAGD variants that would achieve this goal.

Solvent-assisted SAGD (SASAGD) processes were developed to meet such requirements and involvesolvent and steam injection strategies aimed at enhancing viscosity reduction of heavy oil. This leads toa corresponding decrease in energy use as less steam is generated to achieve the same productionthroughput. SAVEX (Gutek et al. 2003), SAS (Zhao 2007), SAP (Gupta et al. 2004) and ES-SAGD (Nasret al. 2003) are some of the more successful SASAGD processes that are being tested in field operations.In this work we focus primarily on the Expanding Solvent Steam Assisted Gravity Drainage (ES-SAGD)process because of its wide acceptance as a successful SASAGD strategy.

The rheology of the bitumen emulsion tremendously impacts the dynamics of production from SAGDsystems. It influences the interaction between inflow variables and outflow variables and therefore affectsengineering design for sustaining stable production. Emulsions are formed during SAGD operations.Emulsions are thermodynamically unstable mixtures of two immiscible fluids made stable by altering theinterfacial tension existing between each fluid. The large quantity of asphaltene molecules in heavy oilinitiates formation of oil-water emulsions as the asphaltene molecule acts as a natural emulsifier thatstabilizes the emulsion formed. Besides, the significant quantity of solid particles in place with heavy oilassist in stabilizing the oil-water emulsion, these type of emulsions are known as Pickering emulsions(Spiecker et al. 2003). The co-injection of solvents impacts the chemical properties of the asphaltenemolecules and as such alters the behavior of the heavy oil emulsion formed. An important propertyimpacted and which forms the subject of interest in this paper is the rheological behavior of the emulsion.

Reservoir depletion and the deep location of the reservoir make artificial lift systems necessary.Besides, the drive toward achieving an energy efficient process requires operating at lower pressureswhich implies that artificial lift might be required even before maturity is reached. The Electro Submers-

Figure 1Cross section of a SAGD operation

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ible Pump (ESP) is widely preferred because of itsflexibility in operation as well as newer units beingmore temperature tolerant. The change in fluid rhe-ology and the corresponding impact on engineeringdesign for solvent SAGD processes infers that ESPdesign would be equally affected. This paper inves-tigates this and presents results toward this effect.

SAGD production and EmulsionbehaviorSAGD production differs from conventional hydro-carbon production. The major variables influencingSAGD performance as pointed out by Yuan andNugent (2013) include the liquid head above theproducer, steam trap efficiency and the effectiveviscosity of the emulsion located between the injec-tor and producer.

The liquid head is the pressure exerted by the fluid that accumulates above the producer as a result ofthe downward flow of the heavy oil emulsion. Steam trap, also referred to as subcool, is the temperaturedifference between the measured temperature at a point and the saturated steam temperature at that point.In SAGD operations the subcool concept is used to achieve thermal efficiency by ensuring the temperatureof the production fluid at the bottomhole is lower than the saturation temperature of steam at the operatingbottomhole pressure. As shown in Figure 2 low subcool values indicates the steam chamber being closerto the production well, while high subcool values imply the steam chamber is further away from theproducer. Thus, it is expected that the subcool value strongly influences the viscosity of the emulsionlocated between the injector and producer.

High subcool values correlate with high emulsion viscosities between the injector and producer andvice versa. Although ESPs are being designed to handle steam (Noonan et al. 2013), steam influx isgenerally unwanted. Therefore, there is an optimal subcool value and corresponding viscosity valuerequired to attain the necessary flow rates during the different stages of production. The addition ofsolvents as an injection additive would alter the subcool-viscosity-productivity behavior that exists whencompared to conventional SAGD operations.

Proper characterization of the emulsion is an important step for engineers designing SAGD productionsystems as emulsions differ significantly in their rheological properties. There are two major forms ofemulsions typical to heavy oil production. These are Water in Oil (WIO) emulsions and Oil in Water(OIW) emulsions; although complex variations occur which include Water in Oil in Water (W/O/W) andOil in Water in Oil (O/W/O) emulsions.

As far as subsurface emulsion characterization is concerned, contrary views have been presented. In hisstudy (Bosch et al. 2004) explicitly states that OIW emulsions are produced at bottomhole conditionswhile another study (Chung and Butler 1988) has stated that WIO emulsions dominate during subsurfacebottomhole production. Nevertheless, it has been pointed out that WIO emulsions are received at surfacefacilities and require additional processing to make them suitable for pipeline transportation (McGowan1990). The rheology of the Oil-Water emulsion is an important factor that determines design parametersused in sizing subsurface and surface production systems.

This was taken into cognizance while carrying out this study and significantly contributed towards theconclusions reached.

Figure 2Illustration of low subcool vs high subcool

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Methodology

Variables that impact most on sustaining stable production from an inflow perspective are the subcoolvalue (determines effective viscosity), bottomhole pressures and liquid accumulation rate from thereservoir bulk. These variables vary significantly for different stages during production. Hence, it isimperative to analyze dominant stages and present results specific for each stage. The major stagesinvestigated are the Pre ramp-up stage, Post ramp-up stage and Wind-down stage of SAGD production.Production for SAGD operations is such that operating constraints are designed with the intent ofmaximizing recovery while ensuring stable production within operable limits. Figures 3 present a briefoverview of these periods and illustrate features characterizing each.Steps taken in the study are as follows:

1. Reservoir simulation of SAGD and ES-SAGD operations; component volumes, mass fractions andphase distributions for both operations are documented

2. Numerical simulation investigating relationship between subcool, bottomhole pressure and liquidhead above producer for the SAGD and ES-SAGD operations

3. Emulsion design/characterization based on inflow volumes and mass composition4. Dynamic simulation of outflow behavior using inflow parameters supporting stable productioncalculated in steps 1 to 3 above for both SAGD and ES-SAGD operations

5. ESP design and sizing for both operations

Figure 3The major stages of production investigated for both SAGD and ES-SAGD operations

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Reservoir Simulation ModelThe reservoir model was built with thermal simula-tion software CMG STARS. The reservoir is a 3Dfully coupled model with properties similar to bitu-men from the McMurray formation of the Atha-basca area. The injector is 6 m above the producerwhich is 1 m above the bottom of the reservoir. Boththe producer and the injector run across the entirelength of the reservoir. Figure 4 illustates the 3Dstructure and details of the reservoir. Figure 5 showsits relative permeability curves. Table 1 presentsinformation on the input data used in building themodel and Table 2 gives details of solvent charac-terization data for ES-SAGD input generated using

Figure 43D reservoir illustrating IK and JK cross sections

Figure 5Relative permeability curves of reservoir model

Table 1Summary of the reservoir and fluid properties used inbuilding reservoir model

Reservoir Properties

Parameter Value

Length (m) 300

Width (m) 46

Thickness (m) 50

Grid blocks in x,y,z directions 26,3,26

Initial Temperature (C) 48

Reference Depth (m) 760

Reservoir Permeability (mD) 1468

Reservoir Porosity 0.25

Formation Compressibilty (1/kPa) 1.80E-05

Rock Heat capacity (kJ/m3-C) 1500

Thermal conductivity of Rock (kJ/m-day-K) 1.50E02

Heat loss to Over/Under burden (kJ/m-day-K) 1.00E03

SPE-171371-MS 5

phase behavior software WINPROP. The solvent is a pseudo-solvent containing a mixture of fractionsranging from C4-C10. K-value data were simulated across relevant temperature and pressure rangesapplicable to each operation. Viscosity data was tuned to match viscosity data calculated using Putang-uttas correlation (Puttagunata et al. 1993) with a bitumen viscosity of 20480 mPa.s at 25C

The maximum injection pressure of the injector was initially 11000 kPa and then switched to 5500 kPaafter circulation and depressurization stages which took five months. Production was allowed to continueuntil rates were less than 5 m3/d. For the solvent injection model 2% solvent mole fraction was co-injectedcontinously once production commenced. The steam quality was 0.9 for all cases.

Subcool, Bottomhole pressure and liquid head studyThe relationship between the subcool (SC) values, bottomhole pressures (BHP) and the effective liquidhead were investigated for both operations by using the reservoir model to conduct necessary sensitivityanalysis. The goal here is to reach a scenario for each operation and at the 3 stages of production such thatthe viscosity value at grid blocks beneath the injector all the way down to the grid block located at theproducer were of the same numerical value. It was assumed that the concept of having equal viscositiesat both locations implies production into the wellbore balances accumulation from the reservoir bulk forthe set operating condition. Hence, it can be stated that optimal pseudosteady state of production isreached in the model when no build up of emulsion allows a viscosity gradient to develop between theinjector and producer.

This was carried out at subcool values ranging from 1C to 35C for the selected periods in both cases.At 35C, production was unstable in all cases while 1C was assumed to be the onset of steam influx andis not accounted for in the production analysis. The flow rates and equivalent BHP are documented as astable combination for each case that allows stable inflow into the bottomhole.

Emulsion design/CharacterizationThe average mass fraction of oil, water and solvent (for the ES-SAGD case) produced into the bottomholefor both operations at each stage of production are presented in Table 3. These values are used for the PVTsimulation of the fluid with the emulsion characterization software. Taylors equation (Taylor 1932) wasused in calculating the viscosity of the emulsion.

For the solvated bitumen mixture, viscosity was calculated using viscosity correlation developed byShu et al (1984) before applying Taylors equation. The viscosity data in all cases were calculated overall pressure and temperature conditions attainable at bottomhole conditions. Figure 6 presents details ofthe viscosity variations for both the SAGD and ES-SAGD at 5500 kPa. The WIO emulsion was calculated

Table 2Summary of pseudo-solvent and heavy oil properties used for solvent characterization

Molecular Weight (kg/gmole) Mass Density (kg/m3) Critical molar volume(l/mol) Critical Temp (K) Critical Pres (atm)

Pseudo Solvent 97.35 722.5 0.399 546.9 29.62

Heavies 533.9 1013.2 1.557 1037.2 10.26

Table 3Mass fraction distribution of components at bottomhole location for both SAGD and ES-SAGD

SAGD ES-SAGD

Oil Water Oil Solv Water

Pre Ramp-Up 12.13% 87.87% 13.19% 0.14% 86.67%

Post Ramp-Up 12.58% 87.42% 8.64% 0.07% 91.29%

Wind Down 2.70% 97.30% 2.68% 0.02% 97.30%

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with water as the dispersed phase and oil as the continuous phase while the OIW emulsion was calculatedwith water as the continuous phase and oil as the dispersed phase.For both operations the WIO viscosityvalues were input into the software and used to tune the model. The inversion point was set at 0.7

Outflow model and ESP designDynamic multiphase simulator OLGA was used to build the outflow model for analyzing the flowbehavior of the fluid exiting the reservoir. Figure 7 presents the tubing configuration of the outflow model.The ESP is located 30 m from the heel of the lateral which connects to a 760 m vertical tubing string andexits at the wellhead after flowing through a 30 m flowline at the surface.

Fluid flow behaviors investigated include superficial velocities, flow regimes, liquid hold up, steamflash dynamics and changes in rheology. Their impact on pressure and temperature variations and heattransfer mechanisms as the fluid flows through the system was investigated. Outflow dependent variablesincluding total pressure head, frictional pressure losses and stable flow rates at the surface coupled withinflow variables which include pump intake pressure (PIP), optimal subcool and stable rates at thesandface were used in sizing the ESP. Table 4 provides relevant information for design of the outflow

Figure 6Viscosity variations with temperature for dead oil case and solvated case

Figure 7Illustration of the production outflow model used in analyzing tubing hydraulic performance

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tubing model with details of tubing dimensions and heat exchange parameters for the tubing andformation.

Results and Discussion

Reservoir Simulations SAGD vs ES-SAGD production variationThe effect of solvent co-injection in comparison to pure steam injection is observed by referring to Figure8 and Figure 9. Figure 8 illustrates the mass flow rate over the entire production period for 9 years whileFigure 9 illustrates its corresponding cumulative production profile as well as the Cumulative Steam OilRatio (CSOR) in each case. CSOR is the yardstick used to measure the thermal efficiency of a SAGDproject, the higher the CSOR the less efficient and vice versa. In Figure 8 the ES-SAGD case shows thatproduction during the pre ramp-up phase is higher compared to the SAGD case. After reaching the peakrate (post ramp-up), production rates decline rapidly and are lower for the ES-SAGD case when comparedaverage SAGD rates.

The solvent injection in combination with steam leads to a rapid increase in production because of theadditional viscosity reduction effect caused by the presence of the solvent. The drop in production ratesfor the ES-SAGD case post ramp-up is because of a larger fraction of bitumen produced pre ramp-up. Thedilution and dissolution effect of the injected solvent allows the steam chamber to expand further outbefore reaching the top of the formation.

Table 4Relevant information required for design of outflow model

Pipe Dimensions Pipe Heat Transfer Formation Heat Transfer

Vertical Height (m) 760 Hmean (W/m2-C) 5 rformation (kg/m

3) 2243

Tubing Diameter (m) 0.102 Csteel (J/Kg-C) 500 Cformation (J/Kg-C) 1256

Roughness (m) 4.50E05 Ksteel (W/m-C) 50 Kformation (W/m-C) 1.59

Figure 8Illustration of production and injection flow rates over the lifecycle of production

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The mass rates of water as illustrated in Figure 8 represents the steam injection rate for both cases andis lower in the ES-SAGD case for all periods although it seemed to increase slightly during the post rampup period. The cumulative bitumen produced in both cases is illustrated in Figure 9 and shows that theoil recovery is higher in the ES-SAGD case than in the SAGD case. The corresponding increase incumulative oil recovery combined with less steam injection at the expense of the solvent injected is thecause for the reduction in CSOR for the ES-SAGD case.The variation in CSOR value over the entirelifecycle of the project is presented in Figure 9. The initial increase is because of the increase in steaminjection volumes and a corresponding low bitumen production rate during the circulation and depres-surization phases.

Numerical Simulations Subcool, Bottomhole pressures and Stable ratesIn understanding the relationship that exists between the subcool value, bottomhole pressure andproduction inflow to the wellbore, comparisons were made on a stage by stage basis for both operations.

SAGD vs ES-SAGD Pre Ramp-Up Figure 10 presents the average stable rates attained at standardconditions for SAGD and ES-SAGD cases. The rates vary from 214 m3/d at 30C SC to 244 m3/d at 5CSC for the SAGD case. The increasing flow rates as subcool reduces is because of the effect of lessviscous drag between liquid layers which resist the effect of gravity on downward drainage. Hence,gravity drainage is more effective at lower viscosity values (lower subcool). ES-SAGD flow rate valuesrange from 221.4 m3/d at 30C SC to 254.6 m3/d at 5C SC. The flow rate values are larger in theES-SAGD case because of a larger delivery from the reservoir.

The bottomhole pressure for each case is equal to the pressure value higher than the saturation pressureat the producer grid blocks as defined by the subcool value. Hence the higher the subcool (SC) value, thehigher the bottomhole pressure (BHP). In most cases a positive pressure difference is required betweenthe operating BHP and BHP values that intiates steam influx. Going forward we assume all operations arecarried out at 10C SC and steam influx occurs at 1C.

In the SAGD case the BHP at 10C SC is 5469.3 kPa and is 5472 kPa for the ES-SAGD case whileat 1C, the BHP for the SAGD case is 5468 kPa while for the ES-SAGD case the BHP is 5467 kPa. Thisimplies the ES-SAGD case has a larger leverage of drawdown prior to steam influx; 5 kPa for the

Figure 9Illustration of cumulative production and CSOR values for both the SAGD and ES-SAGD operation

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ES-SAGD case and 1.3 kPa for the SAGD case. This probably occurs because the solvent creates a solventvapor chamber below the base of the steam chamber acting as a buffer where concurrent condensation andvaporization leads to cooling of bitumen at the steam chamber-bitumen interphase. This effect becomesvisible as subcool values become lower and aids in delaying steam breakthrough

SAGD vs ES-SAGD Post Ramp-Up Figures 11 shows the average stable rates at standard conditionsand corresponding BHP at various subcool values for the post ramp up stage of the SAGD and ES-SAGDoperations. The rates vary from 182.16 m3/d at 30C SC to 199.72 m3/d at 5C SC for the SAGD caseand from 155.56 m3/d at 30C SC to 176.62 m3/d at 5C SC for the ES-SAGD case. The rates for bothcases are less than their pre ramp-up equivalent however; the reduction in the ES-SAGD cases iscomparatively higher and is equally lower than in the SAGD post ramp-up case as most of the oil wasproduced in the pre ramp-up stage.

At 1C SC the BHP for the SAGD case is 5492.3 kPa and at 10C SC it is 5492.2 kPa, for theES-SAGD case at 1C SC the BHP is 5495.3 kPa while at 10C SC it is 5496.8 kPa. The drawdown is1.5 kPa for the ES-SAGD and for the SAGD case and is 0.1 kPa for the SAGD case. The drawdowntrend is still higher in the ES-SAGD case compared to the SAGD case as seen in the pre ramp-up stage.

Figure 10SAGD vs ES-SAGD relationship between subcool, flowrate and bottomhole pressure during pre ramp-up stage

Figure 11SAGD vs ES-SAGD relationship between subcool, flowrate and bottom hole pressure during Post ramp-up stage

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The less bitumen in place during this phase of production is because of a thinner bitumen column meaningthe solvent effect on steam breakthrough is not as pronounced.

SAGD vs ES-SAGD Wind-down Figure 12 shows the average stable rates at standard conditions andthe corresponding BHP at various subcool values for the wind-down stage of the SAGD and ES-SAGDoperations. The rates are significantly less for both cases with flow rates ranging from 19.1 m3/d at 30CSC to 31.8 m3/d at 5C SC for the SAGD case and 13.37 m3/d at 30C SC to 29.9 m3/d at 5C SC forthe ES-SAGD case.

For the SAGD case the BHP at 10C SC is 5507.8 kPa and is 5500.9 kPa at 1C while for theES-SAGD case the BHP is 5501.3 kPa at 10C and is 5500.6 kPa at 10C. Here the operating BHP at 10CSC is significantly less in the ES-SAGD case in comparison to the SAGD case as less bitumen is beingproduced with more solvent being recovered, hence a relatively less hydraulic pressure potential on thewellbore pressure.

Outflow modelDiscussions were presented with reference to flow through a production tubing. 0 m is the inlet of theproduction tubing i.e. at the exit of the ESP and it increases to 820 m at the surface where it exits at thewellhead. The Pump intake pressure (PIP) was assumed to be 500 kPa higher than the saturation pressureof the production fluid at the sandface for both operations.

Fluid RheologyFigure 13 illustrates the rheological behavior of the emulsion produced for both SAGD and ES-SAGDoperations during transport through the production tubing to the surface. Results show that the fluidflowing out of the ESP into the vertical is an emulsion of the OIW type which forms because thetemperature extremes at the bottomhole neutralize the natural tendency of the asphaltene molecule to aidformation of a WIO emulsion. Moreover, it would be difficult for a WIO emulsion to exist stably with thehigh water to oil mass fraction as shown in laboratory experiments (Bennion et al. 1993).

The SAGD produced fluid stays as an OIW emulsion until a switch to a WIO emulsion occurs whichis recognized because of the sudden increase in viscosity. This occurs 228m from the bottom for thewind-down case and approximately 390m from the bottom for both the pre ramp-up and post ramp-upcases. The formation of a lubricating layer of continuous aqueous phase (water in this case) along thewalls of the pipe (Nez et al. 1996) leads to a reduction in the water to bitumen ratio. As the water tobitumen ratio reduces it gets closer to the inversion point, once the water to bitumen fraction reaches 70%by volume the flip to a WIO emulsion occurs.

Figure 12SAGD vs ES-SAGD relationship between subcool, flowrate and bottomhole pressure during Wind-Down stage

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The less extreme temperature as the emulsion is transported towards the surface favors reactivation ofthe more dominant lipophilic bonds in the asphaltene molecule. The flip occurs more readily in thewind-down case because of a lesser water fraction relative to the surface area of the pipe and as such theinversion point is reached more readily. The spike in viscosity as seen in all profiles indicates the onsetof water flashing to steam as it approaches the surface. This occurs as the system pressure drops belowthe saturation pressure of the emulsion in the pipe.

For the ES-SAGD case, the injection of the solvent promotes precipitation of some asphaltenemolecules in the reservoir and alters the chemical composition of the remaining asphaltene molecules tobe more hydrophilic. This stabilizes the OIW emulsion that is formed and no switch is seen in theES-SAGD case. This is highest for the wind-down case because of its lesser water fraction relative to thesurface area of the pipe when compared to the other stages of production.

Figure 14liquid hold up variation on production tubing for each of the phases for SAGD and ES-SAGD production

Figure 13Rheological variations along production profile for both SAGD and ES-SAGD emulsions.

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Liquid holdup and Fluid RegimesFigure 14 illustrates the variation in liquid hold up for both SAGD and ES-SAGD operations across theentire length of the production tubing. In the SAGD case the drop in hold up as the fluid approaches thesurface indicates water flashing to steam on pressure reduction below equivalent saturation pressure forthe temperature reached at that point. The relatively less volume of emulsion formed in the wind-downstage leads to constant emulsion instability and hence large free water flashes much lower down theproduction tubing during transport to the surface. Steam starts flashing 250 m from the base of the verticalfor the wind-down case and occurs 680 m from the base of the vertical for both the pre ramp-up and postramp-up cases. The liquid hold-up has an effect on the viscosity profile of the system as illustrated inFigures 15. As the emulsions water phase flashes into steam it assists in reducing the viscosity of theemulsion further.

In the ES-SAGD production case, the larger volume of liquid flashed is because of the presence of thesolvent. The solvent has saturation properties close to water and as such the solvent flashes as conditionsare reached that initiates water flash. The corresponding reduction in solvent volume leads to further

Figure 16Superficial velocities of liquid and gas phases for SAGD pre ramp-up (represents other stages for both operations)

Figure 15Viscosity and liquid holdup relationship for SAGD Pre ramp up stage.

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alteration in chemical composition of the emulsion and a sudden drop out of large volumes of free water.In both operations the volume of steam flash is highest in the wind-down case because of the larger waterto bitumen ratio compared to other stages of production as illustrated in Table 3

Figures 16 illustrates the superficial velocity of liquid and gas phases for the SAGD and ES-SAGDoperations respectively and variations that occur within the tubing string. In both cases the liquidsuperficial velocity exceeds the gas superficial velocity. It can thus be inferred that bubble flow dominatesvertical transport and stratified flow dominates the horizontal segments, this was verified with thesimulator.

Temperature Gradient and Heat transfer mechanismFigures 17 present a graphical illustration of the heat transferred from the production fluid to thesurrounding formation for the ES-SAGD and SAGD operations respectively. The temperature of the fluidexiting the reservoir is high and as such drives heat loss to the production tubing and then to the formationduring transport to the surface. Two factors determine the quantity of heat loss; the volumetric flow ratesof the fluid (either in liquid or gaseous phase) and the gas fraction present after vaporization. Analysis ispresented comparing SAGD production stages with corresponding ES-SAGD production stages.

The pre ramp-up stage for the ES-SAGD case has a higher flow rate leading to higher heat transferrates when compared to the SAGD case as heat transfer is more effective at high fluid velocities. Thehigher volume of liquid flashing to steam for the ES-SAGD case implies higher cooling effect on thetubing because of latent heat of vaporization. Water in vapor phase (steam) has a relatively high thermaldiffusivity value. The presence of a larger volumetric flow of steam corresponds to a larger heatconduction potential as the production fluid approaches the surface. Hence, the heat transfer effect ishighest in the ES-SAGD case compared to the SAGD case for the pre ramp-up stage of production.

Comparing post ramp-up production stages for each case, the SAGD and ES-SAGD cases show similarheat transfer profiles across all sections of the tubing. The higher fluid flow rate in the SAGD casepartially offsets the higher steam fraction of the ES-SAGD case and as such the magnitude of differenceseen in the pre ramp-up stage is not imminent here. Comparing for the wind-down stage, the flow ratesare similar for both SAGD and ES-SAGD cases, the higher steam fraction in the ES-SAGD case howeverwould determine the heat transfer rate and leads to a higher heat transfer rate for the ES-SAGD case.

The heat transfer rates impact the temperature gradient for each case and the final temperature at thesurface as shown in Figures 18. Note that in making comparison among the different stages of productionfor each operation (either SAGD or ES-SAGD) the increasing flow rate is the dominating factor thatdetermines the heat transfer rate.

Figure 17Illustration of heat transfer profile across production tubing for SAGD and ES-SAGD production

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Pressure gradientThe pressure gradient is the sum of the potential pressure gradient and frictional pressure gradientassuming acceleration effects are neglected. The potential pressure gradient is primarily dependent on themixture density as well as the height of the liquid column. The density of the emulsion is affected by therelative volume of the water phase and this varies at the different production stages for each operation.This is beyond the scope of our work and is not investigated. However, Figure 19 presents the potentialpressure gradient variation with depth for various stages of production in both operations.

The frictional pressure gradient is expected to differ more significantly because of the varyingrheological behaviors for each operation. Figure 20 presents the frictional pressure gradient profile forSAGD and ES-SAGD operations. Locations dominated by OIW emulsions have less frictional gradientvalues because of lower viscosity values while locations dominated by WIO emulsions have largerfrictional gradient values because of higher viscosity values. The roughness value of the pipe presentedin Table 4 implies the pipe inner surface is not smooth and the velocity distribution of fluid flowinfluences the frictional pressure drop. The increase in volumetric flow rate of steam implies that theReynolds number would be easily exceeded in comparison to the liquid volumetric flow rate andturbulence would be easily reached. This magnifies the frictional pressure drop effect.

Figure 18Temperature gradient for each stage of production for both operations

Figure 19Potential pressure gradient for each stage of production for both operations

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For the SAGD case, the larger steam throughput during the post ramp-up stage when compared to thepre ramp-up stage is responsible for the higher frictional pressure gradient as illustrated in Figure 20. Thepeaks correspond with points that have maximum steam velocity in each case. In the wind down case thevolume of steam is high, however the velocity is not high enough to reach turbulence therefore, frictionalpressure drops are not imminent.

The ES-SAGD case has a constant frictional pressure gradient and this is minimal compared to theSAGD case. This is because of the stable OIW emulsion being transported from the bottomhole to thesurface and its lower viscosity value when compared to the WIO emulsion. The faster the flow the higherthe frictional gradient, so it is expected that the pre ramp-up period would have the highest frictionalgradient and then post ramp-up and then wind-down in that order. Figures 21 illustrates the variation ofaverage frictional pressure drop variation with depth for both the SAGD and ES-SAGD cases to give afeel of the difference in magnitude on production for both operations.

ESP DesignThe ESP design was carried out analytically using results attained from the reservoir simulations,sensitivity analysis on inflow variables and outflow variables obtained from the wellbore model. Tables

Figure 20Frictional pressure gradient profile for SAGD and ES-SAGD case across production tubing

Figure 21Frctional pressure drop variation with depth for SAGD and ES-SAGD operations

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5 present the model parameters used in sizing the pump. Figure 22 gives details of the pump curves usedin sizing the pre ramp-up, post ramp-up and wind-down stages for both the SAGD and ES-SAGDoperations. The curves were chosen as the rated flow rate of the pump for optimal efficiency fell withinthe operating flow rate ranges for each operation.

From Table 5, it is realized that the potential pressure head is within the same range for all cases,however the frictional pressure head is significantly higher in the SAGD case and as such the overall

Table 5Outflow, Inflow and ESP variables

SAGD ESSAGD

PRERAMPUP

POSTRAMPUP WINDDOWN

PRERAMPUP

POSTRAMPUP WINDDOWN

INFLOW BHP (bara) 54.69 54.92 55.07 54.71 54.96 55.01

PIP (bara) 51.47 51.68 51.81 51.49 53.44 52.36

Psat (bara) 46.47 46.68 46.81 46.49 48.44 47.36

Dpmax (bara) 3.22 3.24 3.26 3.22 1.52 2.64

Surface Rate (m3/d) 241.4 200.5 28.71 278.4 174.9 28.8

Bottomhole Rate (m3/d) 300.3 249.0 35.57 345.3 217 35.64

OUTFLOW Mixture density (kg/m) 966.9 972.3 971.5 982.0 995.5 997.7

Potential head (bara) 74.82 74.79 71.17 74.92 75.42 76.20

Frictional head (bara) 0.41 0.67 0.04 0.010 0.008 0.001

Total head (bara) 75.23 75.46 71.21 74.93 75.43 76.20

ESP DESIGN PUMP HEAD (m) 253.2 251.8 205.5 245.7 227.2 245.9

PUMP SPEED (rpm) 2917 2917 2917 2917 2917 2917

No of stages 51 46 34 55 38 41

ESP duty (HP) 17.72 13.73 2.74 19.11 11.36 3.27

Figure 22Pump curves used in sizing ESP for the different phases of production.

SPE-171371-MS 17

pressure head is reduced in the ES-SAGD cases. Figure 23 gives a graphical illustration of variations inpump duty and flow rates through the pump for both cases across the different phases of production.

From Figure 23 the high pump duty at the pre-ramp up stage is because of higher flow rates for theES-SAGD case compared to its SAGD equivalent. It can be inferred that for equal production throughputfor both operations a higher pump duty is required for the SAGD case. However, it should be noted thatproduction rates in this analysis are low compared to average field rates. At higher rates, the differencein frictional losses and corresponding pump duties would be magnified

ConclusionThe study investigates the significance of differing rheology of emulsions formed during solvent assistedSAGD and conventional SAGD production and analyzes its impact on tubing hydraulic performance andESP design.

Major conclusions reached are as follows:

The presence of solvents could aid in delaying steam breakthrough thereby increasing theoperating drawdown envelope relative to the producer sandface when compared to conventionalSAGD operations.

Solvent injection leads to precipitation of asphaltene molecules within the reservoir and alterationof the molecular make up of the asphaltene molecules such that it stabilizes the formation of anOIW emulsion.

OIW emulsions are formed at the bottomhole for both SAGD and ES-SAGD production. However,during transport to the surface a combination of factors including reduction in the water to bitumenvolume fraction and less extreme temperatures leads aids inversion to a WIO emulsion for theSAGD case. The stable OIW emulsion formed in the ES-SAGD case does not flip.

Larger volumes of steam flash for the ES-SAGD case is because of the presence of the solventwhich has saturation properties close to water. This equally leads to a large dropout of free waterbecause of the alteration in emulsion properties.

Figure 23Illustration of variation between pump duty and flow rates into ESP across the different phases

18 SPE-171371-MS

The combination of varying rheological properties and differing steam flash volumes influencesthe frictional pressure gradient and average pressure drop during production. The larger volume ofsteam flash for the ES-SAGD case however is not offset by the large rheological differencesbetween SAGD and ES-SAGD cases. Hence, the frictional pressure loss is higher in the SAGDcase.

ESP design for solvent assisted SAGD processes requires less pump duty to deliver the sameproduction through put when compared to conventional SAGD operations. Higher rates magnifythe difference in pump duty required to transport heavy oil when comparing conventional SAGDto solvent assisted SAGD operations.

AcknowledgementThe authors would like to thank the Oklahoma Geological Survey for their financial support. We alsoacknowledge the support of CMG Ltd and Schlumberger for providing the license to use their software.

Nomenclatures

SymbolC Specific Heat Capacity L2t2T1, J/Kg-Cg Gravitational acceleration Lt 2, m/s2

H Reservoir Thickness L, mk Permeability L2, DarcyK Thermal Conductivity mLt3T1, W/m-Cm Parameter used for viscosity functionS Oil saturationSC SubcoolT Temperature T, CU Overall Heat Transfer Coefficient mt3T1, W/m2-C

Greek Letters

Thermal diffusivity L2t1, m2/s Density mL3, Kg/m3

Porosityv Kinematic Viscosity L2t1,m2/s Difference

Subscripts

o Oilr Reservoir bulks Steamsteel Production tubing

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Improved Artificial Lift Design for Solvent Assisted SAGD ProcessesIntroductionSAGD production and Emulsion behaviorMethodologyReservoir Simulation ModelSubcool, Bottomhole pressure and liquid head studyEmulsion design/CharacterizationOutflow model and ESP designResults and DiscussionReservoir Simulations SAGD vs ES-SAGD production variationNumerical Simulations Subcool, Bottomhole pressures and Stable ratesSAGD vs ES-SAGD Pre Ramp-UpSAGD vs ES-SAGD Post Ramp-UpSAGD vs ES-SAGD Wind-down

Outflow modelFluid RheologyLiquid holdup and Fluid RegimesTemperature Gradient and Heat transfer mechanismPressure gradientESP Design

Conclusion

AcknowledgementReferences