Experimental and Simulation Study of the Steam–Foam Process

Article pubs.acs.org/EF

Experimental and Simulation Study of the Steam−Foam Process S. Reza Bagheri* Shell Global Solutions Canada Incorporated, Calgary Technology Centre, 3655 36 Street Northwest, Calgary, Alberta T2L 1Y8, Canada ABSTRACT: The steam−foam process is an enhanced oil recovery (EOR) method which improves the performance of a traditional steam drive using a foaming surfactant. In this study, the surfactant solution was coinjected with steam at a certain quality into a core holder filled with a sand pack. The core holder was kept inside an oven at 250 °C to mimic the near-wellbore temperature in a steam flood. By measuring the pressure drop along the core with and without the surfactant, the mobility reduction factor (MRF) of the generated foam was measured. Two different surfactants were used, and the effect of different parameters such as pressure, steam quality, and superficial gas velocity on the foam strength was studied. Some mechanisms have been suggested to explain the foam generation delay and the foam front retardation in the core. In the modeling section of this paper, the current STARS model for foam generation was studied and its shortcomings identified based on the experimental observations. A modified foam model has been proposed and used to simulate the core-flooding results.

1. INTRODUCTION Steam injection is the most widely used method to recover heavy oil.1 However, due to the low density and low viscosity of steam, it may be difficult to achieve a good macroscopic sweep efficiency in the reservoir.2 Gravity override and steam channeling are two types of reservoir problems that can reduce the effectiveness of steam-flooding applications. In both cases, the steam preferentially sweeps sections of the reservoir; however, a significant fraction of the oil-in-place can remain bypassed in the deeper regions of the reservoir (gravity override) or in the lower permeability zones (channeling). Steam−foam is a promising EOR method in which a surfactant is injected with steam to generate an in situ foam which can divert steam into the bypassed zones.3 The use of foam can reduce the gas mobility and thereby improve the sweep of steam.4 To create and propagate stable foam in the reservoir, a surfactant must be added to the liquid phase of the steam. In addition, a small amount of noncondensable gas may be added to the steam. The role of this noncondensable gas is to stabilize the collapse of bubbles of steam by condensation.4 There are several categories of models that have been proposed for modeling foam in a porous medium. These models, in general, can be texture-implicit local-equilibrium or texture-explicit population balance. Texture-implicit localequilibrium models, which are also called empirical or semiempirical models in the literature, do not attempt to explicitly incorporate the foam texture into the model. Instead, the effect of the presence of foam on the gas mobility is taken into account. These models modify gas relative permeability or gas viscosity, primarily on the basis of experimental observations and hypotheses.5 The commercial reservoir simulator STARS, developed by the Computer Modeling Group (CMG), uses such an texture-implicit local-equilibrium model.6 The aim of this paper is to study the mechanisms that determine the efficiency of a steam−foam flood. Most of the previous studies on steam−foam have used a noncondensable gas such as nitrogen to mimic the vapor part of the steam. The © XXXX American Chemical Society

use of noncondensable gas can make it easier to design and control the experiments; however, as is shown in this study, it will mask the effects of steam condensation on foam behavior. In this study, the effects of some parameters like steam quality, surfactant concentration, and superficial gas velocity on foam strength have been investigated. The STARS foam model is presented, and its shortcomings to model the results of this study are identified. An improved version of the STARS foam model has been proposed and implemented in Shell’s proprietary simulator MoReS to capture these results.

2. EXPERIMENTAL SECTION 2.1. Core-Flooding Setup. A schematic diagram of the experimental setup is depicted in Figure 1. The injection system was comprised of three high-pressure stream lines. The core holder was placed inside an oven at 250 °C. Deionized (DI) water was injected through two pumps to produce the steam at the desired quality. One pump provided the water for the steam generator, and the other provided the liquid water to be mixed with the generated superheated steam. The superheated steam and the liquid water were mixed together in the oven to generate saturated steam at a desired quality which was injected into the core. The superheated steam temperature (Ts) and the liquid water line temperature (Tw) were used along with the cold water equivalent (CWE) rates of the two steam lines to calculate the steam quality after the mixing point (point 3 in Figure 1). The steam pressure at this point was measured using a pressure transducer. From thermodynamic calculations, the quality of steam as well as its temperature was determined at point 3 before the injection of steam into the core. An energy balance was used to calculate the steam quality. The enthalpy of the superheated steam at point 1 (Figure 1) was estimated from the steam table by measuring its temperature (Ts), flow rate, and pressure (points 1 and 2 are just a few centimeters away from point 3, so the pressure measured at point 3 was used for the calculations). Similarly, the water line enthalpy at point 2 was estimated by measuring its temperature (Tw), flow rate, and pressure. These enthalpies were added together to estimate the Received: September 15, 2016 Revised: November 8, 2016 Published: December 1, 2016 A

DOI: 10.1021/acs.energyfuels.6b02341 Energy Fuels XXXX, XXX, XXX−XXX

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Energy & Fuels

Figure 1. Steam−foam experimental setup. total enthalpy at point 3 (since points 1 and 2 are very close to point 3, it was assumed that the heat transfer between the lines and the oven when fluids travel from point 1 and 2 to 3 is negligible). By measuring the pressure of point 3 and knowing its enthalpy, the steam quality can be estimated from the steam table. During the surfactant flooding, the liquid water vessel was replaced with a surfactant solution vessel. Nitrogen was the third line that met the two other injection lines at the same junction. It was injected in low concentration to stabilize the foam (0.1 mol % of total steam rate). Thus, the volumetric flow rate of nitrogen was negligible compared to that of the vapor part of steam, and it was not included in the superficial gas velocity calculations. The main part of the experimental setup was the core holder, which had a cylindrical lead sleeve inside. The core was 37.46 cm long with a diameter of 3.81 cm. The thickness of the core holder was 2.29 mm. The core was mounted vertically in these experiments to give a uniform frontal flow. The lead sleeve was packed with unconsolidated Ottawa sand to have a permeability of about 3 D. The lead sleeve had two pressure taps along the core such that the core was divided into three evenly spaced sections. Two absolute pressure transducers recorded the pressure at the two end sides of the core. Three differential pressure transducers read the differential pressures in each of the one-thirds of the core. These transducers are displayed in Figure 1, and their pressures were identified as Pin and Pout for absolute and ΔP1, ΔP2, and ΔP3 for differential pressures. A thermocouple at the beginning of the core measured the core temperature near the first absolute pressure transducer. The precision in temperature and pressure measurements was approximately 0.1 °C and 0.25%, respectively. A heat-traced line came out of the core holder to a heat-exchanger bore. The temperature of the heat exchanger was controlled by a water bath. This temperature was set to 90 °C such that, in the tests with oil, the oil viscosity out of the oven stayed low so that the oil would flow. The heat exchanger was connected to a highpressure Jurgusen visual cell which again was heat traced to the same temperature. The gas and aqueous effluent from the core was directed to this cell; thus, it played the role of a separator. We were able to extract samples for analysis from the bottom of the Jurgusen cell. From the top of the visual cell and through another line, the gas that was separated in the separator was used for feeding the back-pressure regulator (BPR). 2.2. Surfactants. In this study, two candidate surfactants were used for the core-flooding tests. Both of the surfactants were alkyl benzenesulfonate (ABS) anionic surfactants, identified as surfactants A and B in this paper. These surfactants were screened by conducting foam height tests and thermal degradation analysis of several commercial surfactants in the market to find a surfactant with high foamability and thermal stability at high temperatures. It should be

noted that when the weight percent of the surfactant is reported for a test, it means the weight percent of the surfactant in the liquid part of the steam. 2.3. Experimental Procedure. During steam flooding, the back pressure, temperature, and rates of water and superheated steam were set to get the desired steam quality at the inlet of the core. The surfactant was added to the water line with no change in the back pressure, water and superheated steam rates, and temperatures. However, after the start of foam generation, the pressure in the core should increase significantly since the foam has a strong ability to reduce the gas mobility in porous media. This pressure increase is an important criterion for assessing the foam strength. The mobility reduction factor (MRF), which is usually reported in the steam−foam literature, is the ratio of pressure drop across the core in the presence and absence of foam7

MRF =

ΔPf ΔPnf

(1)

where ΔPf is the steady-state pressure drop across the core with foam flow and ΔPnf is the steady-state pressure drop across the core with gas/water flow at rates equivalent to those that would be seen in the presence of foam. As is shown in Figure 1, three differential pressure transducers read the differential pressures in each one-third of the core. In each run, the pressure drops across each section of the core (ΔP1, ΔP2, and ΔP3 in Figure 1) were measured in a baseline steam flood in order to determine the values of ΔPg for that section. After the start of surfactant injection, the new pressure drops were measured at each time interval to have the value of ΔPf for each third of the core. Then the MRF value for each section was calculated using eq 1. If no foam was generated, the pressure drop would not increase and the MRF would be equal to 1. An MRF of larger than 1 is an indication of foam generation. The larger the MRF is, the stronger the foam. The same sand pack was used for all runs, but it was cleaned before the start of each run.

3. FOAM MODELING The STARS simulator developed by the Computer Modeling Group uses an implicit-texture foam model based on the concept of the mobility-reduction factor.8 In this model, it is assumed that the gas-phase mobility is reduced by a dimensionless interpolation factor, FM, which is defined as FM = B

k rgf k rgnf

(2) DOI: 10.1021/acs.energyfuels.6b02341 Energy Fuels XXXX, XXX, XXX−XXX

Article

Energy & Fuels where kfrg is the relative permeability of gas in the presence of foam and knf rg is the relative permeability of gas in the absence of foam. Thus, by knowing the relative permeability of gas in the absence of foam and calculating FM, the actual gas relative permeability with the addition of foam can be determined. FM is unity without foam and decreases with increasing foam strength. In STARS, FM is generally defined as5,6 FM =

1 1 + FMMOB. F1. F2. F3. F4...

(3)

where the FMMOB is a constant, called the reference mobility reduction factor. Each of the Fi functions is unity when the foam is at maximum strength with respect to that dependent property. For example, the function F1 describes the effect of surfactant concentration on the foam strength and is defined as ⎛ Csurf ⎞EPSURF ⎟ F1 = ⎜ ⎝ FMSURF ⎠

Figure 2. Pressure drop profile in run 19 with 0.5 wt % of surfactant A, steam quality 40%, and superficial gas velocity of 847 m/d. (4)

absolute pressure increases. This pressure increase has a negligible effect on steam quality in the gas phase; however, it can significantly change the steam quality inside of an individual foam bubble. The pressure difference between the inside and the outside of a bubble depends upon the surface tension and the radius of the bubble. The Young−Laplace equation describes the equilibrium pressure difference sustained across the interface between two static fluids, such as water and air, due to the phenomenon of surface tension. For a spherical bubble the equation can be written as 2γ Pi − Po = (6) r

Here FMSURF is the critical surfactant concentration and EPSURF is an exponential parameter. There are also some other functions in STARS which are not mentioned here, as they were not used in the modeling section; these can be found elsewhere.6 In this study, the STARS model is used as a backbone for foam modeling; however, as it was not possible to describe all of the experimental results using STARS, a revised model was scripted and implemented in Shell’s proprietary simulator MoReS. The original model was improved by adding more parameters to capture the experimental observations discussed later in this paper, and using the equations described herein, these improvements could be made in another simulator other than MoReS. One difference is that the gas mobility reduction was achieved by applying FM to gas viscosity instead of gas relative permeability. Thus, here FM =

where Pi is the pressure inside the bubble, Po is the pressure outside the bubble in the continuous phase, γ is the surface tension, and r is the bubble radius. If the surface tension is very low or the bubble size is large, the pressure difference can be negligible. However, at high temperatures, the surfactant’s ability to decrease the gas−water surface tension probably decreases as it would under lower temperatures. As a result, this pressure difference can be important for smaller bubbles. The higher pressure inside the smaller bubbles can decrease the steam quality significantly. The result of this decreased steam quality can be the immediate collapse of the bubble or a decrease in the bubble volume. If the bubble volume decreases then according to eq 6 the internal pressure will increase and the steam quality will decrease again. Thus, the bubble will shrink again until it collapses. Even in the absence of this pressure effect, the bubbles are unstable and finally collapse because of drainage. However, it can be assumed that below a critical radius (rc) the rate of bubble collapse because of this capillary effect should be much faster than drainage, and these unstable bubbles cannot contribute to the mobility reduction. Thus, for each pressure (Po), there is an allowed range of bubble size which can change the mobility of the gas phase (see Figure 3). Bubbles smaller than the critical size either are extremely unstable or will not form at all. When the foam generation starts initially, there is a critical radius below which no bubbles can be generated at the local pressure. Initially, this radius can be very small, and most of the bubbles generated in the core will have a radius higher than the critical size. After the start of foam generation, the local pressure (Po in eq 6) will increase, and this results in collapse of the smaller bubbles. In other words, the critical radius increases with increasing pressure, and this narrows down the window of

μgnf μgf

(5)

where μfg is the gas viscosity in the presence of foam and μnf g is the gas viscosity in the absence of foam.

4. RESULTS AND DISCUSSION 4.1. Pressure Drop Profile. Most of the previous studies have used a noncondensable gas such as nitrogen for steam− foam core floods instead of real steam, due to the practical difficulties involved in experimentation. When nitrogen is used, the pressure drop along each core section will increase after foam generation and propagation in that section. Once the foam completely fills the section, however, the pressure drop reaches its maximum value and forms a plateau.9 However, the pressure profile when using steam is completely different. Figure 2 shows typical pressure drop profiles of a steam−foam core flood with surfactant A. In the first and second sections of the core, the pressure drop initially increases and reaches a maximum value but then decreases to a much lower value and remains constant at later times. Once the pressure drop in each section reaches its maximum value, the pressure drop in the following section starts increasing. The same thing happens for each section until foam fills the core, and foam breakthrough is observed at the core outlet. This phenomenon can be attributed to the pressure increase after foam generation. Once the foam starts to form, the C

DOI: 10.1021/acs.energyfuels.6b02341 Energy Fuels XXXX, XXX, XXX−XXX

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Energy & Fuels

generation and result in a decrease in lamellae density and foam strength. As the pressure increases, this critical size will increase too, fewer pores contribute to foam regeneration, and the foam strength keeps decreasing. Thus, the result will be very similar to that of the previous case. The shape of the MRF (or simply pressure drop) curves in Figure 2 can be explained by assuming a nonhomogeneous foam generation along the core. The foam, which is initially generated at the core inlet, forms a strong front, which starts propagating along the core (Figure 4). The generation of this foam front will increase the local pressure, which results in collapse of the smaller bubbles and a decrease in the foam strength. Thus, after some time, the core inlet starts generating a weaker foam which follows the strong front as a weaker tail. The pressure drop peak observed in each section is the result of the strong foam front passing through that section. Once the foam front leaves a section and enters the next section, the pressure drop in that section starts decreasing and the pressure drop in the next section starts increasing. When the foam front enters the last section, it will be weakened due to the capillary end effects which will be discussed in section 4.5. Thus, the DP3 peak is a little smaller. However, the viscous foam creates a large pressure drop along the corer, so the local pressure at the core outlet is much lower than that of the core inlet. This lower pressure results in a stronger foam generation at the last section. Thus, the tail of the pressure drop curve at the last section (DP3 in Figure 4) has a higher value compared to other sections. If nitrogen is used instead of steam for foam flooding, the increase in pressure does not result in the condensation of nitrogen, so the foam texture and strength does not change because of that, and a pressure drop response like that of Figure 4 will not be observed. 4.2. Foam Frontal Velocity. The foam frontal velocity can be measured in the experimental setup by monitoring the pressure drop response. As mentioned before, there are three pressure transducers which are roughly equally spaced along the core, which make it possible to measure the pressure drop in the first, second, and third sections of the core (see Figure 1). The length of each section is ∼1/3 of the total core length. Using these transducers, it is possible to estimate the average foam frontal velocity at each section. Table 1 summarizes the foam frontal velocity in the first and second one-third sections of the core for several runs. The results suggest that, under similar conditions for surfactant A, a stronger foam with higher MRF always moves faster than a weaker foam with a lower

Figure 3. Relation between the bubble pressure and its radius.

bubble radius for foam generation. By increasing the critical radius, the foam texture becomes coarser and foam will be much weaker (Figure 3). The previous argument assumed that the foam bubble size is smaller than the pore size. It is also possible that the bubble size is bigger than the pore size. In this case lamellae exist individually, and each lamella separates two bubbles of gas. To generate an individual lamella vapor should enter a pore throat and overcome the entry pressure of the pore. The smaller the pore throat, the higher the entry pressure. Again, by increasing the pore pressure after foam generation, the pore pressure plus the entry pressure may be too high for the vapor to enter the pore without condensation. As a result, by increasing the pressure after foam generation, the steam vapor may not be able to generate foam in pore throats smaller than a critical size. Thus, these smaller pores will not contribute to foam

Figure 4. Relation between the foam front movement and the pressure drop response. D

DOI: 10.1021/acs.energyfuels.6b02341 Energy Fuels XXXX, XXX, XXX−XXX

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In run 29, the surfactant injection initially started at lower gas velocities with no foam generation, which is why the breakthrough time is much longer compared to those of the other runs.

8.6/3.4 62.8/25.9 32.5 29.0

A/0.35 A/0.5 A/0.35 A/0.5 A/0.5 A/0.5 B/0.5 B/0.5 B/0.5 B/0.5 13 21 11 19 29 18 17 26 25 27

MRF. The surfactant breakthrough time was estimated to be much sooner than that of foam in these runs, so the foam front travels much more slowly than the surfactant front. The foam frontal velocities are much lower than those of the gas velocities. Some previous studies reported that, at high temperature, there was no retardation of the foam front relative to that expected from the volume of injected fluids.10 Friedmann et al.9 reported a lag of the foam front behind the surfactant front in high-temperature core floods and attributed it to the trapping of a fraction of the flowing bubbles in the smaller gas-saturated pores. 4.3. Foam Generation Delay. Table 2 shows the delay time for foam generation after the start of surfactant injection in Table 2. Delay Time for Foam Generation after the Start of Surfactant Injection for Different Runs run no.

surfactant/ concentration (wt %)

steam quality (%)

superficial gas velocity (m/d)

delay in foam generation: time(min)/CWE volume of injected steam (PV)

33 18 19 20 21

A/0.5 A/0.5 A/0.5 A/0.5 A/0.5

80 60 40 60 40

847 847 847 1301 1301

14.24/1.85 7.71/1.33 4.88/1.31 10.54/2.89 2.66/1.06

several of the experimental runs. It is obvious that by increasing the steam quality at the core inlet the time required for foam generation will increase. Zhong et al.11 studied foam injection into the vadose zone (the vadose zone, also termed the unsaturated zone, is the part of Earth between the land surface and the top of the phreatic zone, i.e., the position at which the groundwater (the water in the soil’s pores) is at atmospheric pressure). This is not an EOR application, but their results showed some similar observations to those of this study. The important feature of the vadose zone is that it is unsaturated. The results of Zhong et al.11 indicated that, under unsaturated conditions, a liquid front was formed ahead of the foam front. The accumulated liquid in this front was mainly composed of the water injected with the foam. When the foam was first injected into the column sediment (the sediment in the column was unconsolidated and of different particle sizes), no flow of bubbles could be seen at the column inlet. Instead, a liquid wetting front was observed. Thus, there was a lag time for foam generation in the porous medium. Visual observation indicated that the liquid content in the wetted sediment section increased with time until the minimum water saturation required for foam generation was reached and bubble formation was initiated at the upstream end of the wetted sediment section. The foam injection pressure was primarily attributed to the resistance to foam flow by the sediment but not to the movement of the water ahead of the foam front. As a result, this water front was almost invisible in the pressure profile of the core. The separation of the foam, liquid-wetting, and gas fronts is an important feature of foam transport in an unsaturated zone. The link between Zhong et al.’s study11 and hightemperature steam foam core flooding is that, at high temperature, a high steam quality (mass fraction of the vapor phase) results in a low volume fraction of the liquid phase. For example, a steam quality of 60% at 250 °C corresponds to a liquid volume fraction of less than 2%. The steam at this condition is probably too dry for the formation of lamellae. The water saturation inside the porous media can be high; however, it is important to note that the majority of this water is connate

a

9.0/3.6 8.5/3.4 16.8/3.5 13.3/3.6 123.3a/2.2 31.1/5.4 53.4/21.3 10.8/4.3 no BT 19.7/5.3 125.7a/3.7 60.7/10.5 8.8 54.2 N/A 25.6 144.9 8.8 3.1 21.3 N/A 26.1 159.9 19.7

surfactant/concentration (wt %) run no.

40 40 40 40 40 60 40 40 60 60

1301 1301 847 847 235 847 847 1301 684 1301

9.1 20.7 3.5 25.3 66.5 20.0 no foam 20.6 no foam no foam

12.8 79.1