1

DEWATERING OF HIGH SALINITY BRINES BY OSMOTICALLY ASSISTED REVERSE OSMOSIS

Jason T. Arena, U.S. DOE National Energy Technology Laboratory, Pittsburgh, PA Timothy V. Bartholomew, Carnegie Mellon University, Pittsburgh, PA

Ashutosh Sharma, Carnegie Mellon University, Pittsburgh, PA Meagan S. Mauter, Carnegie Mellon University, Pittsburgh, PA

Nicholas S. Siefert, U.S. DOE National Energy Technology Laboratory, 626 Cochrans Mill Rd, Pittsburgh, PA 15236

nicholas.siefert@netl.doe.gov, Ph: 412-386-4404

Abstract

Geologic carbon dioxide storage (GCS) is the practice of capturing carbon dioxide (CO₂) from industrial plants and storing it in subsurface geologic formations for the mitigation of anthropogenic CO2 emissions. Saline formations have been identified as one type of subsurface formation suitable for CO2 storage and are typified by a porous interconnected rock strata, saturated with brine. In GCS, mitigation of subsurface pressure accumulation from CO₂ injection is essential. Formation pressure can be alleviated by extracting some of the brine saturating the formation prior to or during GCS. While GCS extracted brines can have large variations in TDS, a quantitative evaluation of select GCS brines show a total dissolved solids (TDS) of 100‒225 g·L-1. The high TDS of these brines limits the available methods to dewater them for volume minimization because reverse osmosis (RO) can only dewater a saline solution to a maximum concentration of about 90 g·L-1 TDS. This limitation has impeded the application of RO for dewatering high salinity brines and produced waters.

In reverse osmosis, the upper concentration limit to which a brine can be dewatered is intrinsically linked to its osmotic pressure alongside the membrane’s selectivity and the mechanical stability of associated equipment. A new approach to the reverse osmosis process, called osmotically assisted reverse osmosis (OARO), uses an engineered osmotic pressure difference between the feed solution and a less saline sweep solution that flows on the permeate side of a membrane. This engineered osmotic pressure difference allows a feed solution to be dewatered at hydrostatic pressures lower than its osmotic pressure. For sweep regeneration, a diluted sweep solution can be dewatered as a feed to subsequent OARO stages and returned to its initial concentration. A complete process will have a sufficient number of OARO stages to dilute the sweep solution until it can be desalinated by a final RO stage. This study will present experimental data on the water flux for a bench scale evaluation of OARO up to a salinity approximately 3.5 times greater than seawater using driving pressures up to 27.6 bar.

1. Introduction

Concern about globally increasing surface temperatures and its linked association to atmospheric concentrations of carbon dioxide (CO2) have become a driver for technologies to capture and store anthropogenic CO2. One challenge in geologic CO2 storage (GCS) is pressure management

2

within the selected storage formation. Saline formations suitable for GCS are typified by a porous interconnected rock strata, saturated with saline brine at depths greater than 800 m located far from exiting faults with a caprock seal [Holloway (2005), IPCC (2005), Varre et al. (2015)]. High pressure injection of CO2 into the subsurface can cause increased pressure within the storage formation and overpressure within the formation can result in seismic events [Cihan et al. (2015), Lee et al. (2016)]. Pressure management within a GCS formation can be achieved by the extraction of formation brine before and/or during injection of CO2 in the subsurface. One growing need of research linked to GCS is the management of brine extracted for pressure management within both the storage formation and overlying formations [IPCC (2005), Birkholzer and Zhou (2009), Gaus (2010), Buscheck et al. (2011), Cihan et al. (2015)].

When brine is extracted to manage formation pressure, a required next step is the disposition of the extracted brine. Typically, these brines are sufficiently saline such that they cannot be used for domestic, industrial, or agricultural purposes [Bourcier et al. (2011), Veil et al. (2011)]. Some domestic sandstone formations either serving as a part of GCS operations or under consideration have salinities ranging from approximately 100–225 g·L-1 total dissolved solids (TDS), which is around 3‒6 times greater than seawater [Sass et al. (1998), Knauss et al. (2005), Dilmore et al. (2008), Lu et al. (2012)]. These extracted brines are primarily comprised of dissolved sodium, chloride, and calcium ions. In the disposition of these brines, isolation from formations used for industrial, agricultural, and drinking water is paramount; therefore, GCS brine disposal into surface waters is not a viable option [Birkholzer and Zhou (2009), Lemieux (2011)].

One option for brine management is to dewater these brines and reduce their volume. The resulting brine concentrate, now having a higher concentration of dissolved solutes, can be reinjected with a net reduction in subsurface volume. The water generated by the dewatering process should be of sufficient quality that it could be used for industrial/agricultural purposes or discharged into surface waters [Aines et al. (2011), Bourcier et al. (2011)].

1.1. Advantages and limitations of membrane technology Nearly all of the technologies suitable for dewatering extracted GCS brines were originally developed for the desalination of saline waters (i.e. brackish water and seawater) and so have matured for the production of freshwater in regions lacking sufficient natural supply. The earliest techniques of desalination, such as multi-stage flash (MSF) and multi-effect distillation (MED), mimic the natural water cycle by evaporating water from a brine, separating the steam, and condensing the water vapor to recover pure water [Miller (2003), Reddy and Ghaffour (2007)]. Distillation is advantageous for dewatering high salinity brines because feed water salinity has a relatively negligible impact on product water quality; however, both MSF and MED are reliant on an external steam sources for heating a feed water to drive the distillation process. Since supplying steam to disparate locales from a central source may be logistically challenging, it would be advantageous to operate a dewatering process using only electricity as a source of useful work. Another distillation technology, called mechanical vapor compression (MVC), is not reliant on an external steam supply and instead uses a motor driven compressor to drive the distillation process [Miller (2003), Al-Sahali and Ettouney (2007)]. Although MVC does not require an external steam supply, heat transfer limitations in a MVC evaporator adversely affect efficiency. The temperature gradients needed for distillation cause a significant irreversible generation of entropy, and the losses associated with the large transfer of thermal energy across a

3

temperature gradient is the main reason why distillation processes have a low thermodynamic efficiency [Nafey et al. (2008), Mistry et al. (2011)].

Because no heat transfer or phase change occurs, reverse osmosis (RO) has developed as a more efficient successor to distillation process for desalination and, like MVC, uses electricity as an energy source. RO uses a large applied hydrostatic pressure to push water across a semi-permeable membrane and typically has a higher electrical efficiency than distillation processes for desalination [Greenlee et al. (2009), Elimelech and Phillip (2011)]. Compared to distillation, RO is more sensitive to the salinity of the feed solution because of the link between salinity and the potential of a solution to draw pure solvent into it if that solution were separated from pure solvent by perfectly selective semi-permeable membrane, called the osmotic pressure [Wilson and Stewart (2013)]. As shown by Eq. (1), for a RO process to operate with positive water flux, the applied hydrostatic pressure must be larger than the osmotic pressure of the feed water. RO membranes can only operate at pressures up to 82.8 bar (1,200 psi); therefore, there are limits on the maximum salinity to which a RO process can dewater a brine [Dow, Bourcier et al. (2011)]. A plot showing the osmotic pressures of solutions of sodium chloride and extracted GCS brines is shown in Figure 1.

{ }ΔπΔPAJw −⋅= (1)

Figure 1 Osmotic pressure of a sodium chloride solutions calculated using osmotic

coefficients compiled by Hamer and Wu (1972) at 25°C. Also shown are the osmotic pressures of four extracted brines related to GCS, using the activity of water calculated by Geochemist’s

Workbench v9 with the thermo_phrqpitz database. The gold ‘Seawater’ line represents the range of global seawater concentrations [Dow].

4

1.2. Osmotically assisted reverse osmosis The semi-permeable membranes used in RO processes are inherently imperfect, meaning they are slightly permeable to dissolved solids. In RO, this inherent imperfection makes the permeate quality dependent on the passage of water and salt. Residual salinity in the permeate from salt passage lowers the osmotic pressure difference across a membrane. As shown by Eq. (1), a lower osmotic pressure difference means less hydrostatic pressure is needed for water to cross the membrane [Lee et al. (1981)]. Tailoring membrane permselectivity is not the only manner by which the osmotic pressure difference across a membrane can be engineered. For example, water flux in forward osmosis (FO) processes is predominately driven by an osmotic pressure difference between a feed solution and a higher osmotic pressure solution that is commonly positioned on the permeate or backside of a membrane [Cath et al. (2006), McCutcheon and Elimelech (2006)]. A drawback to FO is that water crossing a membrane in FO is mixed with and dilutes this high osmotic pressure solution thus requiring additional steps for the production of a low salinity product water [Van der Bruggen and Luis (2015)].

Figure 2 One possible configuration of a multi-stage OARO process to dewater a high

salinity brine. Shown here are two OARO stages followed by the requisite final RO stages to produce a low salinity product.

Some studies have noted improved water flux in FO by applying a moderate pressure to the feed solution; however, in these pressure assisted FO applications water flux still occurs along an osmotic pressure gradient [Blandin et al. (2013), Coday et al. (2013), Oh et al. (2014), Banchik et al. (2016)]. As an alternative to these FO configurations, one can apply a moderate hydrostatic pressure to the feed solution and drive water across the membrane. In retaining a less saline solution on the permeate side of a membrane, the osmotic pressure difference across the membrane is engineered to be much lower than the osmotic pressure of the feed. By having the sweep solution on the permeate side of a membrane, a lower bulk osmotic pressure difference exists across a semi-permeable membrane. This lower osmotic pressure difference will in kind require less hydrostatic pressure to drive water against the osmotic pressure gradient; however, like FO, this osmotically assisted reverse osmosis (OARO) is not a means of direct desalination but one that dilutes the less saline solution contacting the membrane. A fundamental difference

5

from FO exists because OARO is driven by a reverse osmosis behavior; therefore, OARO stages can be linked together in series to incrementally step down the concentrations of the feed and sweep solutions as a continuous process. The engineering of the osmotic pressure difference in OARO means a feed solution can be dewatered well past the point where it is physically and mechanically unfeasible to dewater using only an applied hydrostatic pressure [Karode et al. (2000), Lakerveld et al. (2010), Lucas and Sawyer (2012), Wohlert (2012)]. As shown in Figure 2, the final step of a multiple stage OARO process would be a conventional RO stage and, as a critical part of the process, would be the ultimate source of a product water.

1.3. Objectives of this study The goal of this study is to characterize a commercially available membrane for performance in OARO with the purpose of generating a baseline for water flux across the membrane to assess the feasibility of OARO at the bench scale. This study will be conducted using a range of feed salinities with the sweep concentration adjusted to maintain a constant concentration difference between the feed and sweep across the range of salinities studied.

2. Materials and Methods

2.1. Materials The membrane used in this study was a woven supported asymmetric cellulose triacetate (CTA) membrane designed for forward osmosis, which was purchased from Fluid Technology Solutions (Albany, Oregon, USA). Sodium chloride (USP/FCC grade >99.0%) was purchased from Fisher Scientific (Hanover Park, Illinois, USA). Water used in this study was obtained from an in-house purified water (PW) source having a conductivity of ~12 μS·cm-1.

2.2. Experimental test system All experiments conducted in this study were performed on a bench scale test system functionally similar to a pressure retarded osmosis (PRO) test system described and used in prior studies [Anastasio et al. (2015), Arena et al. (2015), Huang et al. (2016)]. The test system used in this study and shown in Figure 3 consists of two recirculating fluid loops whose contact is mediated by a centrally located membrane cell (model CF042D-FO with the stainless steel filter support plate) purchased from the Sterlitech Corporation (Kent, Washington, USA). The two recirculated loops used in the test system will be referred to as the feed and sweep solutions. In general, the feed solution has a higher pressure and salinity than the sweep solution. In all tests, the selective layer of the CTA membrane would face the higher pressure feed solution. So the membrane might better withstand the applied hydrostatic pressures, it was loaded into the test cell supported by five pieces of permeate carrier recovered from a SW30-4040 spiral wound element (Dow Water and Process Solutions, Midland, Michigan, USA). The inlet and outlet pressure of the membrane cell were monitored using pressure gauges located on the inlet and outlet of the test cell. The volumetric flowrates of the feed and sweep solution were monitored using flowmeters. The salinity of the sweep solution was monitored using an HI7635 inline conductivity probe connected to a HI8936A transmitter (Hanna Instruments, Woonsocket, Rhode Island, USA). The mass of the solution was measured using an AX5202 precision balance (Ohaus Corporation, Parsippany, New Jersey, USA). All experimental data was recorded at forty second intervals using an in-house developed LabVIEW program.

6

Figure 3 Test system used in this study. System temperature was regulated by use of a recirculating chiller that recirculated cooling water through the heat exchanger on the sweep

solution loop and a coil submerged in the feed tank.

2.3. Experimental procedure The CTA FO membrane from Fluid Technology solutions was studied to evaluate the membrane’s performance in RO and PRO across a range of feed concentration and applied hydrostatic pressures and use these data to guide interpretations of water flux observed in OARO. A CTA membrane originally designed for FO was selected because of its innate hydrophilicity and better designed structure for reduced internal concentration polarization [McCutcheon and Elimelech (2008)]. The progenitor of this membrane manufactured by Hydration Technology Innovations has been widely studied in FO and PRO where it has been subjected to modest amounts of applied hydrostatic pressure [She et al. (2012), Coday et al. (2013), Anastasio et al. (2015)]. The initial setup of the test system, shown in Figure 3, was started using 6 kg and 2 kg of water for the feed and sweep. The membrane cell was loaded with the selective layer facing the feed solution.

This membrane was evaluated under a variety of conditions with each test divided into four primary phases. Because the properties of sodium chloride solutions are well tabulated at 25°C, all tests were performed at 25°C. A volumetric flowrate of 1.0 L·min-1 for the feed solution and 0.5 L·min-1 for the sweep solution were used in all tests. The first portion of each test was a one hour long compaction of the membrane with a feed pressure of 31.0 bar. Following the compaction of the membrane the pure water permeance of each membrane was measured at feed pressures ranging from 27.6–6.9 bar. The third portion of each test was used to characterize the RO and/or the PRO performance of the membrane to observe water and salt flux at different external conditions such as the feed salinity and applied hydrostatic pressure. The final portion of each test was to measure the water flux in OARO across the CTA FO membrane and observe trends relating water flux to the feed and sweep salinities and/or applied hydrostatic pressure. Water flux in RO and OARO was observed using bulk concentration differences across a membrane of either 0.3 mol·kgH₂O-1 or 0.6 mol·kgH₂O-1. Because the osmotic pressure of a

7

sodium chloride solution deviates from a linear relationship at higher concentrations and to keep a mostly constant bulk osmotic pressure between the feed a sweep solutions (Figure 1), a maximum feed concentration of 2.1 mol·kgH₂O-1 (2.0 mol·L-1) was selected. A complete summary of all the test conditions used in this study are shown in Table 1. All data presented is the average of at least triplicate tests with fresh membrane samples used in each test.

Table 1 Summary of test conditions used in this study.

Test Regime Test Pressures Feed Sweep

Compaction 31.0 bar Purified Water Purified Water

RO/PRO Water and Salt Flux

27.6–6.9 bar in 6.9 bar increments

Purified Water 0.15 mol·kgH₂O-1 0.3 mol·kgH₂O-1

0.45 mol·kgH₂O-1 0.6 mol·kgH₂O-1 0.9 mol·kgH₂O-1 1.2 mol·kgH₂O-1 1.5 mol·kgH₂O-1 1.8 mol·kgH₂O-1 2.1 mol·kgH₂O-1

Purified Water “ ” “ ” “ ” “ ” “ ” “ ” “ ” “ ” “ ” “ ” “ ”

OARO Water Flux 27.6–6.9 bar in 6.9 bar increments

0.9 mol·kgH₂O-1 1.2 mol·kgH₂O-1 1.5 mol·kgH₂O-1 1.8 mol·kgH₂O-1 2.1 mol·kgH₂O-1

0.3 & 0.6 mol·kgH₂O-1 0.6 & 0.9 mol·kgH₂O-1 0.9 & 1.2 mol·kgH₂O-1 1.2 & 1.5 mol·kgH₂O-1 1.5 & 1.8 mol·kgH₂O-1

2.4. Equations describing water and salt flux in OARO In all osmotic membrane processes, the concentration of dissolved solids at the membrane’s selective layer is critical to calculating the water and solute flux. Like in RO, the convection of solution and diffusion of solutes causes an external boundary layer to develop in the feed solution at a membrane’s selective layer. In addition, the presence of the sweep solution coupled with the asymmetry of osmotic membranes results in a competition between solution convection and solute diffusion on the permeate side of the membrane within its support layer. This internal concentration polarization is fundamentally identical to the phenomena encountered in FO processes [McCutcheon and Elimelech (2006)]. One key difference between OARO and FO is that hydrostatic pressure is the primary driving force for the flow of water across the membrane, and so, for the purposes of convention, the positive direction for water flux is the same as RO, which is against the osmotic pressure difference across the membrane. The water and solute flux across a membrane in OARO are defined by Eq (2) and Eq (3) and are similar to the equations previously developed for PRO. It should be noted that there is a change in the sign for water flux

8

and Pf−Ps is larger than π(cf,m)−π(cs,m), such that water flux against the osmotic pressure difference would be positive [Lee et al. (1981), Yip et al. (2011)].

[ ] ( ) ( )[ ]{ }ms,mf,sfw cπcπPPAJ −−−⋅= (2)

{ }ms,mf,s ccBJ −⋅= (3)

Concentration polarization on both sides of a membrane in OARO greatly influences the concentration of solutes at a membrane’s selective layer. Due to non-linearity in the osmotic pressure shown in Figure 1, the significance of concentration polarization in OARO is of even greater significance should extremely high concentrations (i.e. >3 mol·kgH₂O-1) of sodium chloride be considered for the feed and sweep solutions. To most accurately calculate the transmembrane osmotic pressure, concentrations of the feed and sweep solutions at the membrane selective layer, defined by Eq. (4) and Eq. (5), must be known.

⋅⋅

−⋅

⋅⋅

−−

⋅⋅

⋅+

⋅⋅

−−

⋅⋅

⋅

⋅+

⋅⋅

⋅=κD

δJexp1

κDSJ

expκD

δJexp

JB1

κDSJ

expcκD

δJexpc

JB

κDδJ

expccf

fw

s

w

f

fw

w

s

wbs,

f

fwbf,

wf

fwbf,mf, (4)

⋅⋅

−−⋅

⋅⋅

−−

⋅⋅

⋅+

⋅⋅

−⋅−

⋅⋅

⋅⋅+

⋅⋅

−⋅=κDSJexp1

κDSJexp

κDδJexp

JB1

κDSJexpc

κDδJexpc

JB

κDSJexpcc

s

w

s

w

s

fw

w

s

wbs,

f

fwbf,

ws

wbs,ms, (5)

3. Result and Discussions

3.1. Water and salt flux in RO and PRO The initial water flux for a feed and sweep of purified water fell within a fairly narrow range between 16–18 L·m-2·h-1 for all the membranes characterized in this study at a feed pressure of 27.6 bar. The narrow range of observed water flux is significant because it is indicative of a similar water permeance amongst all the membranes samples used in this study. The average water permeance for all the membranes used in this study, indicated by the pink line in Figure 4, was 0.646±0.047 L·m-2·h-1·bar-1. The water flux observed for all of the RO/PRO tests conducted in this study are shown in Figure 4. These data have uniform trends with water flux universally increasing at higher feed pressures. One interesting feature present is that differences in water flux at each of the four pressures grows smaller with increasing concentrations. For example at 27.6 bar with a purified water sweep, the difference in water flux between the 0.15 mol·kgH₂O-1 and 0.45 mol·kgH₂O-1 feed is approximately 7.1 L·m-2·h-1 while the difference in water flux between the 1.8 mol·kgH₂O-1 and 2.1 mol·kgH₂O-1 feed is 2.4 L·m-2·h-1. This difference is the result of increasing internal concentration polarization with the higher values of bulk salt concentration.

9

Figure 4 Water flux for sodium chloride feed solutions of varying concentration at 25°C.

Positive water flux denotes RO while negative water flux denotes PRO. Note: Error bars that are not visible fall completely within the space occupied by data points.

Figure 5 Salt flux for a sodium chloride feed solutions of varying concentration at 25°C.

The salt fluxes observed for the RO/PRO characterization of this membrane are shown in Figure 5. As observed for the water flux, salt flux also increases with increasing feed pressure. If water flux opposite the osmotic pressure difference is considered to be positive, higher water flux will result in a higher solute concentration of the feed at the membrane’s selective layer. For

10

example, a 2.1 mol·kgH₂O-1 feed with a purified water sweep has a bulk concentration difference of approximately 2.1 mol·kgH₂O-1; however, in PRO, external concentration polarization dilutes the feed reducing the concentration difference across the membrane to approximately 0.93 mol·kgH₂O-1 at a feed pressure of 27.6 bar. A decrease in the hydrostatic pressure of the feed also decreases the water flux. So at a feed pressure of 6.9 bar, the dilutive effect of external concentration polarization increases and reduces the concentration difference across the membrane to approximately 0.75 mol·kgH₂O-1. The increased solute flux observed is a result of the increased salt concentration at the selective layer. Also, as would be expected by Eq. (3), higher salt concentrations in the bulk feed solution translates to higher salt flux through the membrane.

3.3. Water flux in OARO Figure 6 and Figure 7 show the water flux observed for the CTA FO membrane under OARO conditions with the concentration difference between the feed and sweep solutions fixed at 0.6 mol·kgH₂O-1 and 0.3 mol·kgH₂O-1, respectively. A significant drop in the water flux was observed when comparing water flux data for a 0.6 mol·kgH₂O-1 feed and purified water sweep with the other water flux data shown in Figure 6. This drop in water flux is the result of internal concentration polarization where water permeating the membrane dilutes the sweep solution, increasing the concentration difference between the feed and sweep. The corresponding increase in the osmotic pressure difference decreases the overall driving force across the membrane. This means that more of the applied hydrostatic pressure is needed to simply overcome mass transport limitations. Despite the low water flux measured, for all feed and sweep pairings with a 0.6 mol·kgH₂O-1 concentration difference, water permeates through the membrane by reverse osmosis opposite the osmotic pressure gradient across it. As shown by Figure 1, the osmotic pressure of even a 0.9 mol·kgH₂O-1 brine is sufficiently large such that a hydrostatic pressure of 27.6 bar is insufficient for water to permeate a membrane against the osmotic pressure gradient.

Water fluxes for feed and sweep compositions have a concentration difference of 0.3 mol·kgH₂O-1 are shown in Figure 7. Here, a relationship, identical to that shown in Figure 6 , was also observed with respect to increasing hydrostatic pressure of the feed. For example, larger increases in water flux were observed for a 0.3 mol·kgH₂O-1 feed with a purified water sweep as the feed hydrostatic pressure increases than for higher salinity feed and sweep solutions. Once again, this decrease in water flux is attributable to internal concentration polarization effects whereby, as water flux is increased, the solute concentration in the membrane support layer is further diluted. This increases the osmotic pressure difference between the sweep and feed solution and negates the higher applied hydrostatic pressure to the feed solution, greatly reducing any increase in water flux across the membrane. The reduced effect of increasing hydrostatic pressure to improve water flux indicates that one future path for optimizing an OARO process would be balancing the pressure and flowrates for the sweep and feed solutions. This path for improvement is in addition to the natural need for improved membranes with reduced internal concentration polarization. This is a focus that OARO shares with the development of membranes for FO processes, including PRO.

11

Figure 6 Water flux for sodium chloride feed solutions of varying concentration at 27.6 bar and 25°C with a bulk concentration difference between the feed and sweep

solutions of 0.6 mol·kgH₂O-1.

Figure 7 Water flux for sodium chloride feed solutions of varying concentration at 25°C with

a bulk concentration difference between the feed and sweep solutions of 0.3 mol·kgH₂O-1.

4. Conclusions

The need for subsurface pressure management for GCS is going to be a critical aspect in the future deployment of carbon dioxide storage and technologies that minimize the energy costs of

12

brine management will be an important part in reducing the energy demands of GCS. The OARO process has, at the bench scale, shown that a high salinity feed can have water removed from it using a hydrostatic pressure much lower than its osmotic pressure. This would be done by circulating a less saline sweep solution on the permeate side of the membrane. This fundamental proof of OARO as a concept opens up the motivation to study other membranes for OARO. Also, essential to the development of OARO is the development of computational models to design and perform energy estimates on the OARO process.

5. Symbols

Term Description cf,b Feed molar concentration in the bulk, mol·L-1

cf,m Feed molar concentration at the membrane, mol·L-1

cs,b Sweep molar concentration in the bulk, mol·L-1

cs,m Seep molar concentration at the membrane, mol·L-1

A Water permeance of a membrane, L·m-2·h-1·bar-1

B Solute permeability of a membrane, L·m-2·h-1

Df Diffusion coefficient of sodium chloride in feed solution, m2·s-1

Ds Diffusion coefficient of sodium chloride in sweep solution, m2·s-1

Js Solute flux across membrane, mol·m-2·h-1

Jw Volumetric water flux across membrane, L·m-2·h-1

Pf Feed hydrostatic pressure, bar

Ps Sweep hydrostatic pressure, bar

S Structural Parameter, m

T Absolute Temperature, K

δf Feed external boundary layer thickness, m

κ Conversion factor, 3.6·106 L·s·m-3·h-1

π(c) Function relating osmotic pressure to concentration, bar

6. Disclaimer

This manuscript was prepared as an account of work sponsored by an agency of the United States Government. Neither the United States Government nor any agency thereof, nor any of their employees, makes any warranty, express or implied, or assumes any legal liability or responsibility for the accuracy, completeness, or usefulness of any information, apparatus, product, or process disclosed, or represents that its use would not infringe privately owned rights. Reference therein to any specific commercial product, process, or service by trade name, trademark, manufacturer, or otherwise does not necessarily constitute or imply its endorsement, recommendation, or favoring by the United States Government or any agency thereof. The views

13

and opinions of authors expressed therein do not necessarily state or reflect those of the United States Government or any agency thereof.

7. Acknowledgements

This work was completed as part of National Energy Technology Laboratory (NETL) research for the U.S. Department of Energy’s (DOE) Fossil Energy (FE) program. Support for Jason Arena and Tim Bartholomew was provided by appointment to the National Energy Technology Laboratory Research Participation Program, sponsored by the U.S. Department of Energy and administered by the Oak Ridge Institute for Science and Education.

8. References

Aines, R D; T J Wolery; W L Bourcier; T Wolfe and C Hausmann (2011), Fresh water generation from aquifer-pressured carbon storage: Feasibility of treating saline formation waters. Energy Procedia 4, pp 2269-2276.

Al-Sahali, M and H Ettouney (2007), Developments in thermal desalination processes: Design, energy, and costing aspects. Desalination 214 (1), pp 227-240.

Anastasio, D D; J T Arena; E A Cole and J R McCutcheon (2015), Impact of temperature on power density in closed-loop pressure retarded osmosis for grid storage. J. Membr. Sci. 479, pp 240-245.

Arena, J T; S S Manickam; K K Reimund; P Brodskiy and J R McCutcheon (2015), Characterization and performance relationships for a commercial thin film composite membrane in forward osmosis desalination and pressure retarded osmosis. Ind. Eng. Chem. Res. 54 (45), pp 11393-11403.

Banchik, L D; A M Weiner; B Al-Anzi and J H Lienhard (2016), System scale analytical modeling of forward and assisted forward osmosis mass exchangers with a case study on fertigation. J. Membr. Sci. 510, pp 533-545.

Birkholzer, J T and Q Zhou (2009), Basin-scale hydrogeologic impacts of co 2 storage: Capacity and regulatory implications. Int. J. Greenh. Gas Control 3 (6), pp 745-756.

Blandin, G; A R Verliefde; C Y Tang; A E Childress and P Le-Clech (2013), Validation of assisted forward osmosis (afo) process: Impact of hydraulic pressure. J. Membr. Sci. 447, pp 1-11.

Bourcier, W; T Wolery; T Wolfe; C Haussmann; T Buscheck and R Aines (2011), A preliminary cost and engineering estimate for desalinating produced formation water associated with carbon dioxide capture and storage. Int. J. Greenh. Gas Control 5 (5), pp 1319-1328.

Buscheck, T A; Y Sun; Y Hao; T J Wolery; W Bourcier; A F Tompson; E D Jones; S J Friedmann and R D Aines (2011), Combining brine extraction, desalination, and residual-brine reinjection with co 2 storage in saline formations: Implications for pressure management, capacity, and risk mitigation. Energy Procedia 4, pp 4283-4290.

Cath, T Y; A E Childress and M Elimelech (2006), Forward osmosis: Principles, applications, and recent developments. J. Membr. Sci. 281 (1), pp 70-87.

Cihan, A; J T Birkholzer and M Bianchi (2015), Optimal well placement and brine extraction for pressure management during co 2 sequestration. Int. J. Greenh. Gas Control 42, pp 175-187.

14

Coday, B D; D M Heil; P Xu and T Y Cath (2013), Effects of transmembrane hydraulic pressure on performance of forward osmosis membranes. Environ. Sci. Technol. 47 (5), pp 2386-2393.

Dilmore, R M; D E Allen; J R M Jones; S W Hedges and Y Soong (2008), Sequestration of dissolved co2 in the oriskany formation. Environ. Sci. Technol. 42 (8), pp 2760-2766.

Dow, Filmtec™ reverse osmosis membranes technical manual. Dow Chemical Company, http://msdssearch.dow.com/PublishedLiteratureDOWCOM/dh_095b/0901b8038095b91d.pdf?filepath=liquidseps/pdfs/noreg/609-00071.pdf (accessed June 23, 2016).

Elimelech, M and W A Phillip (2011), The future of seawater desalination: Energy, technology, and the environment. Science 333 (6043), pp 712-717.

Gaus, I (2010), Role and impact of co 2–rock interactions during co 2 storage in sedimentary rocks. Int. J. Greenh. Gas Control 4 (1), pp 73-89.

Greenlee, L F; D F Lawler; B D Freeman; B Marrot and P Moulin (2009), Reverse osmosis desalination: Water sources, technology, and today's challenges. Water Res. 43 (9), pp 2317-2348.

Hamer, W J and Y C Wu (1972), Osmotic coefficients and mean activity coefficients of uni‐univalent electrolytes in water at 25° c. J. Phys. Chem. Ref. Data 1 (4), pp 1047-1100.

Holloway, S (2005), Underground sequestration of carbon dioxide—a viable greenhouse gas mitigation option. Energy 30 (11), pp 2318-2333.

Huang, L; J T Arena; M T Meyering; T J Hamlin and J R McCutcheon (2016), Tailored multi-zoned nylon 6, 6 supported thin film composite membranes for pressure retarded osmosis. Desalination 399, pp 96-104.

IPCC (2005), Carbon dioxide capture and storage. Cambridge University Press: New York, New York, USA.

Karode, S; S Kulkarni and M Ghorpade (2000), Osmotic dehydration coupled reverse osmosis concentration: Steady-state model and assessment. J. Membr. Sci. 164 (1), pp 277-288.

Knauss, K G; J W Johnson and C I Steefel (2005), Evaluation of the impact of co 2, co-contaminant gas, aqueous fluid and reservoir rock interactions on the geologic sequestration of co 2. Chem. Geol. 217 (3), pp 339-350.

Lakerveld, R; J Kuhn; H J Kramer; P J Jansens and J Grievink (2010), Membrane assisted crystallization using reverse osmosis: Influence of solubility characteristics on experimental application and energy saving potential. Chem. Eng. Sci. 65 (9), pp 2689-2699.

Lee, J-Y; M Weingarten and S Ge (2016), Induced seismicity: The potential hazard from shale gas development and co2 geologic storage. Geosciences J. 20 (1), pp 137-148.

Lee, K; R Baker and H Lonsdale (1981), Membranes for power generation by pressure-retarded osmosis. J. Membr. Sci. 8 (2), pp 141-171.

Lemieux, J-M (2011), Review: The potential impact of underground geological storage of carbon dioxide in deep saline aquifers on shallow groundwater resources. Hydrogeol. J. 19 (4), pp 757-778.

Lu, J; Y K Kharaka; J J Thordsen; J Horita; A Karamalidis; C Griffith; J A Hakala; G Ambats; D R Cole and T J Phelps (2012), Co 2–rock–brine interactions in lower tuscaloosa formation at cranfield co 2 sequestration site, mississippi, USA. Chem. Geol. 291, pp 269-277.

Lucas, A L and J E Sawyer (2012), Treatment of waters with multiple contaminants. United States Patent Application 2012/0145635 A1, June 14, 2012.

15

McCutcheon, J R and M Elimelech (2006), Influence of concentrative and dilutive internal concentration polarization on flux behavior in forward osmosis. J. Membr. Sci. 284 (1), pp 237-247.

McCutcheon, J R and M Elimelech (2008), Influence of membrane support layer hydrophobicity on water flux in osmotically driven membrane processes. J. Membr. Sci. 318 (1), pp 458-466.

Miller, J E (2003), Review of water resources and desalination technologies. Sandia National Laboratories, http://prod.sandia.gov/techlib/access-control.cgi/2003/030800.pdf (accessed June 8, 2016).

Mistry, K H; R K McGovern; G P Thiel; E K Summers; S M Zubair and J H Lienhard (2011), Entropy generation analysis of desalination technologies. Entropy 13 (10), pp 1829-1864.

Nafey, A; H Fath and A Mabrouk (2008), Thermoeconomic design of a multi-effect evaporation mechanical vapor compression (mee–mvc) desalination process. Desalination 230 (1), pp 1-15.

Oh, Y; S Lee; M Elimelech; S Lee and S Hong (2014), Effect of hydraulic pressure and membrane orientation on water flux and reverse solute flux in pressure assisted osmosis. J. Membr. Sci. 465, pp 159-166.

Reddy, K and N Ghaffour (2007), Overview of the cost of desalinated water and costing methodologies. Desalination 205 (1), pp 340-353.

Sass, B; N Gupta; J Sminchak and P Bergman (1998), In Geochemical modeling to assess the capacity of a midwestern united states geologic formation for co2 sequestration, Proceedings of the Fourth International Conference on Greenhouse Gas Control Technologies, Interlaken, Switzerland, pp 1079-1086.

She, Q; X Jin and C Y Tang (2012), Osmotic power production from salinity gradient resource by pressure retarded osmosis: Effects of operating conditions and reverse solute diffusion. J. Membr. Sci. 401, pp 262-273.

Van der Bruggen, B and P Luis (2015), Forward osmosis: Understanding the hype. Rev. Chem. Eng. 31 (1), pp 1-12.

Varre, S B; H J Siriwardane; R K Gondle; G S Bromhal; V Chandrasekar and N Sams (2015), Influence of geochemical processes on the geomechanical response of the overburden due to co 2 storage in saline aquifers. Int. J. Greenh. Gas Control 42, pp 138-156.

Veil, J A; C B Harto and A T McNemar (2011), In Management of water extracted from carbon sequestration projects: Parallels to produced water management, SPE Americas E&P Health, Safety, Security, and Environmental Conference, Society of Petroleum Engineers.

Wilson, A D and F F Stewart (2013), Deriving osmotic pressures of draw solutes used in osmotically driven membrane processes. J. Membr. Sci. 431, pp 205-211.

Wohlert, C W (2012), Apparatus and methods for solution processing using reverse osmosis. United States Patent 8,216,473 B2, July 10, 2012.

Yip, N Y; A Tiraferri; W A Phillip; J D Schiffman; L A Hoover; Y C Kim and M Elimelech (2011), Thin-film composite pressure retarded osmosis membranes for sustainable power generation from salinity gradients. Environ. Sci. Technol. 45 (10), pp 4360-4369.