1. Introduction
As the third-largest proven oil reserves globally [
1,
2], the Canadian oil sands are exploited by the thermal recovery method using steam-assisted gravity drainage (SAGD) [
3]. The SAGD consists of two horizontal wells, about 800–1200 m in length, with a horizontal injection well typically five meters above the horizontal production well (
Figure 1) [
4,
5]. Steam is injected from the injection well into the formation to reduce the bitumen viscosity. After forming a steam chamber inside the reservoir, the melted bitumen flows along the steam chamber’s edge toward the production well located around 5 m below the injection well [
6,
7]. The operational temperature and pressure in the Alberta SAGD projects vary, depending on the depth. However, the typical operating temperature of the SAGD well is around 200–250 °C [
8,
9]. Li et al. [
10] reported different injection pressures for various SAGD projects in Alberta, ranging from 2000 to 5000 kPa, depending on the reservoir depth.
As the unconsolidated oil-bearing sands are inherently loose, SAGD wells require the implementation of stand-alone screens to support the wellbore and control sand production [
11,
12,
13,
14,
15,
16]. The slotted liner is currently widely used in Canada’s SAGD process due to its economy and adequate mechanical integrity [
17,
18,
19,
20].
Figure 2a illustrates the slotted liner configuration, and
Figure 2b shows different slot geometry designs, which can be manufactured into the straight cut, keystone cut, and seamed profiles.
Improperly designed stand-alone screens can compromise the sand control, damaging the production equipment and causing operational problems and high remediation costs [
21,
22,
23,
24]. Thus, proper design criteria are needed for the slotted liner to restrict sand production and maintain the flow performance at a desirable level for the SAGD well lifecycle [
25].
A literature review shows that various design criteria have been developed for the slotted liner’s optimal design. In 1937, Coberly [
27] proposed the first criterion for selecting the screen aperture size based on the formation particle size distribution (PSD). From the experimental tests using a single-slot coupon, Coberly found that the aperture size should be smaller than two times of the formation sand’s D10 (the sieve size that retains 10% of the material mass) to form a stable sand bridge. Subsequently, several researchers investigated and proposed the design criteria of the slotted liner for various applications. Based on field experience, Fermaniuk [
28] proposed an empirical design criterion for slotted liners in SAGD. The maximum slot size in this criterion is 3.5 times D50, and the minimum slot size is two times D70. Other researchers conducted prepacked SRT experiments to determine the proper slot size [
17,
18,
19,
29,
30].
During the preheat stage in SAGD, the formation sand collapses and accumulates on the stand-alone screen due to the loss of the bonding strength generated by the bitumen. Thus, researchers believe prepacked SRT reasonably represents the SAGD wellbore condition [
31,
32,
33,
34,
35]. Past experimental works recognized a few factors that should be incorporated in the slotted liner design criteria, including slot size, slot density, fluid velocity, fluid phase and PSD [
18,
19,
30,
36]. However, these factors were only partially incorporated in their experimental study. For example, Bennion et al. [
17] and Devere-Bennett [
29] neglected the impact of slot density and suggested slotted liner selection protocols based on a single-slot coupon test.
Mahmoudi [
37] and Wang et al. [
30] conducted single-phase prepacked SRT experiments with multislot coupons and generated design criteria for different formation PSD’s. However, the fluid phase change is not considered in their work. Mahmoudi et al. [
19] and Wang et al. [
30] presented their design criteria graphically by using the “Traffic Light System” (TLS). In the TLS, traffic light colors (green, yellow and red) are employed to a linear axis to indicate the optimal slot size window. The green color in the TLS means an acceptable slot size, in which the amount of sand production is lower than 0.12 lb/ft
2, and retained permeability (RP) is greater than 70% [
19,
30]. The yellow zone indicates a marginal performance, defined with sand production between 0.12 and 0.15 lb/ft
2, and RP between 50 and 70%. The red zone means an unacceptable slot size, in which sand production is over 0.15 lb/ft
2, and RP is below 50%. The procedure of creating the TLS is summarized as (1) conduct sand control tests, (2) obtain the sand production and retained permeability data, (3) plot the sand production and retained permeability data against the slot size, (4) curve to fit against the testing data with the optimal equations for the sand production and retained permeability, respectively, (5) use the equations to find the boundaries for the sanding and flow performance, (6) build individual TLS bar for the sand production and retained permeability, and (7) combine the two TLS bars to obtain the overall TLS.
Table 1 summarizes the contributions and limitations of the current design criteria for the slotted liner.
In summary, limited design criteria are currently available for SAGD production wells, and they often neglect essential factors in the design. This paper develops new design criteria for the slotted liner in SAGD production wells. The criteria are obtained by employing a multi-phase SRT and incorporating several critical factors such as the slot size, slot density, fluid velocity, fluid phase and PSD. The new design criteria account for more realistic flow scenarios than the current criteria.
4. Design Criteria for Slotted Liners
The design criteria for slotted liners are developed based on sand production and retained permeability results. Proper design criteria should meet both sanding and flow performance requirements. Mahmoudi et al. [
19] and Wang et al. [
30] used a graphical method called the traffic light system (TLS) to present their design criteria.
In this paper, the new criteria are also presented in TLS. The procedure for the TLS generation follows the same in Wang et al. [
30]. Proposed criteria consider two SAGD scenarios: normal SAGD condition and aggressive SAGD condition. The three-phase flow was regarded as the aggressive SAGD condition to emulate the steam breakthrough in real SAGD.
Figure 11,
Figure 12,
Figure 13,
Figure 14,
Figure 15 and
Figure 16 show the new criteria for slotted liners for different PSD’s and different slot densities.
TLS results in
Figure 11 show an absence of the green slot window and only narrow yellow windows for DC-I. This is attributed to the undesirable flow performance of the slotted liner due to the low open-flow-area. DC-I contains the highest fines concentration compared to DC-II and III, making it more vulnerable to the pore plugging.
Due to the limited amount of fines content in DC-II and III and coarser PSD characteristics, the slotted liner can provide a desirable sanding and flow performance, resulting in the green windows for DC-II and III. Particularly in DC-III (
Figure 15), the coarsest sand containing the least amount of fines, the green window is wider than the DC-II (
Figure 13).
Another finding is with the increase of slot density, the yellow windows in DC-I becomes wider and shifts to the left (
Figure 11). A similar observation is made for DC-II and III (
Figure 13 and
Figure 15), attributed to the increase of the open-flow-area leading to a higher sanding and retained permeability. Thus, both the upper and lower bound of the safe window shift to the left. It is also found that the upper bound of the safe aperture windows for all three PSD types become smaller to keep sanding within the acceptable level during the SAGD steam breakthrough. The finding is justified with the additional sand production caused by the steam breakthrough, which needs to be mitigated by using a narrower slot size. Thus, the design window shrinks for the aggressive SAGD operational conditions compared to the normal SAGD condition.
5. Comparison between Improved and Current Design Criteria
This section compares the new set of design criteria with the criteria in the literature for these three PSD’s.
Table 5 shows the comparison results for each PSD. As per Coberly [
27], the slot size of 2 × D10 could form a stable sand arch and prevent sanding for unconsolidated sands. Based on field data, Fermanuik [
28] suggested 2 × D70 to 3.5 × D50 as the safe slot window. However, both criteria are based on one or two points on the PSD curve. Mahmoudi et al. [
19] and Wang et al. [
30] proposed design criteria for SAGD production wells by conducting single-phase SRT tests. Their criteria considered the PSD curve, slot density, and operational conditions.
Table 5 compares the design criteria results for the normal condition. It can be seen that the upper bounds of the proposed new design criteria are close to the same obtained by Coberly and Fermaniuk’s criteria. When comparing the proposed criteria with criteria obtained from single-phase SRT experiments by Mahmoudi et al. [
19] and Wang et al. [
30], it is found that the upper limits of the previous criteria are slightly larger than the new ones. The difference is due to the different testing procedures, where the new testing procedure includes the multi-phase flow at varying water- and gas-cut.
However, as shown in
Table 6, the new design criteria’ upper bounds for the aggressive condition are much smaller than those in all current criteria. This deviation is attributed to the steam breakthrough impact on sand production considered in the new design criteria. The multiphase liquid–gas flow condition (brine, oil, and gas) resulted in much more sand production than the single-phase brine flow testing condition, hence a narrower slot in the new design criteria for the aggressive condition.
The slot window’s lower bounds that are governed by the plugging, obtained from single-phase and multi-phase SRT, show consistent results. However, the lower bound Fermanuik [
28] does not compare well, which can be attributed to not incorporating the fines content in the design criterion. The DC-I formation contains a large amount of fines, which makes it susceptible to plugging. Based on the testing results, if the slot size of 0.006′′ was used in DC-I, the retained permeability would be lower than 50%, leading to undesirable flow performance.
The outcomes from the proposed design criteria and those in the literature seem to agree on the upper bound for the normal condition. The proposed criteria provide further guidance regarding the plugging and steam breakthrough scenario.
6. Conclusions
Comparing the design criteria for slotted liners in the literature with those in this work shows limitations and uncertainties for the literature design criteria. The slot sizes designed based on one point of the whole PSD curve are not appropriate and could result in biased results. Therefore, it is strongly recommended to consider the impact of PSD in the design analysis. The laboratory SRT testing on field samples could provide a superior design than when using the criteria. However, it is obvious that the testing procedure strongly affects the slot design criteria. Therefore, the testing procedure should be designed consistently with field conditions. The proposed design criteria incorporate a multiphase flow testing procedure in the slotted liner design criteria. The new criteria are presented graphically for normal SAGD conditions and steam breakthrough conditions.
This paper improved current slotted liner design criteria by incorporating several influential factors into the testing design. With more key parameters involved, the testing results were more reliable, robust, and representative of the SAGD conditions. However, there were still some other key factors, including temperature, corrosion, erosion, asphaltene precipitation, and clay mineralogy, which were not fully represented in the current testing design. Therefore, field data are needed to validate the proposed design criteria.
SI Metric Conversion Factors
1 cp = 10−3 Pa∙s; 1 inch = 2.54 cm; 1 ft = 0.3048 m; 1 pound = 453.592 g; 1 psi = 6894.76 pa