Osmotically enhanced dewatering-reverse osmosis (OED-RO) hybrid system: Implications for shale gas produced water treatment

Journal of Membrane Science 554 (2018) 282–290

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Osmotically enhanced dewatering-reverse osmosis (OED-RO) hybrid system: Implications for shale gas produced water treatment Jungwon Kim, Jungbin Kim, Junghyun Kim, Seungkwan Hong

T

⁎

School of Civil, Environmental and Architectural Engineering, Korea University, 145 Anam-ro, Seongbuk-gu, Seoul 02841, Republic of Korea

A R T I C LE I N FO

A B S T R A C T

Keywords: OED-RO hybrid system Module-scale modeling Shale gas produced water (SGPW) treatment High water recovery Specific energy consumption (SEC)

Managing shale gas produced water (SGPW) is one of the greatest challenges for shale gas industry due to its high salinity and water volume. Osmotically enhanced dewatering (OED) has great potential for treating SGPW because of its higher water recovery and lower energy consumption. This study systematically investigated the effects of operating conditions on OED performance through numerical simulation of membrane modules. The simulation results first showed that OED achieved higher water recovery over forward osmosis (FO) due to less internal concentration polarization (ICP). Water recovery could be higher with decreasing feed flow fraction, increasing normalized membrane area, and increasing hydraulic driving force fraction. It was also demonstrated that OED-RO hybrid process was able to yield more water with similar energy efficiency as one-stage RO, for SGPW of 28.5 g/L total dissolved solids (TDS) under realistic conditions considering inefficiency associated with pump and energy recovery device (ERD). Lastly, to validate our findings, OED experiments were performed with pre-treated real SGPW as a feed solution, and exhibited good agreement with the simulation results. Specifically, water recovery was achieved up to 67% with a high rejection rate of over 97% for most ions at a hydraulic pressure of 30 bar. Our modeled and experimental observations suggest that the OED-RO process can be an energy-efficient process in concentrating high salinity wastewater.

1. Introduction Shale gas has recently emerged as an energy source that offers an opportunity to meet the increasing demand for energy around the world. Technical advances in drilling and fracturing have led to more profitable gas production and resulted in a 40% increase in the world technically recoverable gas resources [1]. However, during the extraction of shale gas from shale formations using new drilling and hydraulic fracturing, 11,500–19,000 m3 of water per well is commonly used for hydraulic fracturing [2], and after gas production begins, shale wells continue to generate wastewater, referred to as produced water. This wastewater is characterized by a high concentration of solids, metal, salts and organics originating from within the shale formation [3]. Due to the large volume of water, the high salinity and the concentration of dissolved substance, there is growing public concern about its potential harmful effects for human health and the environment [4]. Consequently, more stringent regulations have been enacted and have led to increased pressure on gas developers to manage the produced water in ways that minimize the volume to be treated or disposed of [5]. Therefore, the recent trend in wastewater management in the shale industry has focused on increasing its internal reuse by

⁎

means of on-site treatment technologies [6–8]. One proposed technology for reducing the volume of wastewater is forward osmosis (FO). In FO, a draw solution, which has higher osmotic pressure than the feed solution, is required to provide osmotic pressure gradient across the membrane. This osmotic pressure gradient acts as the driving force allowing water molecules to naturally transport from the feed solution into the draw solution. For the reuse of shale gas produced water characterized by high salinity, FO can offer attractive advantages including high water recovery and fouling reversibility [9–12]. However, after the FO process, a regeneration process of the diluted draw solution is required further to obtain fresh water, which is placed at an energetic disadvantage because of the high concentration of the draw solution [13,14]. To solve this problem caused by the regeneration of high salinity draw solution, a few studies have reported on the concept of using the balance of the two different driving forces (hydraulic pressure and osmotic pressure) with the aim of high water recovery and energy efficiency [15–18]. This particular operating mode has been referred to variously as osmotic assisted reverse osmosis (OARO) [15], draw solution assisted RO (DSARO) [16], and osmotically-enhanced dewatering (OED) [17]. We have used the term OED in this study, as it seems

Corresponding author. E-mail address: [email protected] (S. Hong).

https://doi.org/10.1016/j.memsci.2018.03.015 Received 23 December 2017; Received in revised form 2 March 2018; Accepted 8 March 2018 Available online 09 March 2018 0376-7388/ © 2018 Elsevier B.V. All rights reserved.

Journal of Membrane Science 554 (2018) 282–290

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solution to assess the practical feasibility of OED in terms of water flux, water recovery and salt rejection. 2. Materials and methods 2.1. Determination of membrane characteristics A flat-sheet cellulose triacetate (CTA) FO membrane from HTI (Albany, OR) was used for this study. Polyester woven mesh is embedded within the porous support layer of the membrane to improve mechanical strength. All membrane samples were stored in DI water at 4 °C before use. The water permeability coefficient (A), salt permeability coefficient (B), and structure parameter (S) of the membrane were measured in bench-scale RO and FO tests. This RO and FO method is widely used for determination of membrane properties and is described in detail in earlier studies [22–24]. Briefly, the water and salt permeability coefficients (A and B) were first determined from

Fig. 1. Schematic diagram of OED-RO hybrid process.

to distinguish it more from RO as a post-draw regeneration system. Fig. 1 shows a schematic diagram of OED-RO hybrid process. In OED, a draw solution (DS), which has a lower concentration than the feed solution (FS), is used to replace a portion of the hydraulic pressure as the driving force with the osmotic pressure. Thus, the OED process can operate at a lower hydraulic pressure than the conventional RO process, resulting in more energy efficiency and higher water extractability from salinity wastewater. According to the recent study, it is estimated that the minimum energy consumption is 2.11 kWh/m3 for seawater at 35 g/L TDS to 70% recovery [18]. Table 1 shows an overview of recent studies on the concept of osmotically-assisted or enhanced RO process. Most research on such processes evaluated the specific energy consumption (SEC) by optimizing the draw solution concentration [15], number of stage and/or configuration [18]. However, the effect of design and operating parameters (e.g. water recovery, feed flow, membrane area, and driving force fraction) at module-scale has not yet been quantitatively analyzed, which is the key factors for determination of system-level performance. In addition, few studies simplify the model by assuming no salt passage through the osmotic membrane [15,18], even though the diffused salts have been reported to accumulate along the membrane module and thus have an impact on overall performance such as water flux and recovery [19–21]. Therefore, it is of paramount importance to analyze and optimize the key factors affecting system-level design of the OED process. The objectives of this study are to analyze the effects of key design and operating parameters on achievable water recovery through a module-scale OED model, and to verify experimentally its applicability for treatment of shale gas produced water by evaluating basic OED performance in laboratory tests. First, we developed a module-scale OED model to examine the effect of flow rate, hydraulic driving force fraction, and membrane area and characteristics on total water recovery. We then delineated the minimum specific energy required for the OED process and discussed how process inefficiency, including pump, ERD and frictional pressure loss, affects the SEC. Finally, OED experiments were carried out using shale gas produced water as a feed

A=

Jw ΔP

(1)

and

B = Jw ⎛ ⎝

1−r J ⎞ exp ⎛− w ⎞. r ⎠ ⎝ k⎠

r = 1 − Cp/ Cb

(2) (3)

where k is the feed solute mass transfer coefficient, Cp and Cb are the bulk concentration of the permeate and the feed solution, respectively, and r is the rejection rate. The structure parameter (S) was determined using the experimental flux obtained during FO test and calculated as follows:

S=

D ⎛ B + AπD ⎞ ln . Jw ⎝ B + Jw + AπF ⎠ ⎜

⎟

(4)

where D is the diffusivity of the draw solute, and πD and πF are the bulk osmotic pressure of the draw and the feed solution, respectively. In this study, pre-compaction was applied to investigate the effect of hydraulic pressure on membrane characteristics (Section 3.1.2). Before the measurement, membrane coupons were compacted at hydraulic pressures ranging from 10 to 40 bar with DI water over 4 h. Then A and B were determined in RO mode at 10 bar, using DI water and 2000 ppm NaCl solution, respectively. After the RO mode test, structure parameter, S, is measured in FO mode, using 1 M NaCl draw solution and deionized feed water. Membrane characteristics after pre-compaction at different hydraulic pressures are presented in Table 2. Apparent membrane characteristics vary with the applied pressure, as previously observed [23]. When applying a hydraulic pressure of 40 bar, a significant increase in salt permeability, B, was observed due to membrane deformation. Thus, the OED experiments were performed at a hydraulic pressure of 30 bar for stable operation.

Table 1 Summary of recent studies concerning the osmotically assisted or enhanced RO process from the literature. Ref.

Method

Feed Conc. (g/ L)

Brine Conc. (g/ L)

Recovery (%)

Hydraulic pressure (bar)

Considerations

[15]

M*

125

150–215

17–42

65

[16]

M

35

58–64

40–45

30–40

35–140

–

–

0–50

35 70

116 140

70 50

68.3 85

Sweep concentration for achievable recovery Membrane area and hydraulic pressure for SEC estimation Structure parameter, draw solution concentration, and flow rate ratio for SEC and cost estimation Driving force and membrane properties (A, B, S) for performance improvement Number of stage and recovery rate for SEC estimation

[17]

M and E

[18]

M

**

M* – Modeling; E** – Experiment.

283

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Table 2 Membrane characteristics after pre-compaction. Pre-compaction (bar)

10 30

Water permeability, A (L m−2 h−1 bar−1)

Salt permeability, B (L m−2 h−1)

0.74 ± 0.03 0.70 ± 0.04

0.49 ± 0.07 0.57 ± 0.06

( ) ( )

( ) ( )

⎛ CF , b exp Jw − CD, b exp − Jw S ⎞ k D Js = B ⎜ ⎟, ⎜ 1 + B ⎡exp Jw − exp − Jw S ⎤ ⎟ Jw ⎣ k D ⎦ ⎝ ⎠

Structure parameter, S (μm)

(7)

where A and B are the water and salt permeability coefficients, respectively, S is the structure parameter of the membrane, πF , b and πD, b are the osmotic pressures of the bulk feed and the osmotic pressure equalizer, respectively, D is the bulk diffusion coefficient for the solutions, and CF,b and CD,b are the concentrations of the bulk feed and draw solutions, respectively.

410 ± 29 486 ± 40

2.2. Measurement of water flux and solute rejection A laboratory-scale cross flow OED unit described elsewhere [17] was used to measure the membrane performance. A custom-built test cell with symmetric rectangular channels (77 mm L × 26 mm W × 3 mm H) was used. Spacers were inserted in the draw channel to support it by preventing deformation of the membrane under high hydraulic pressure [23,25,26]. A variable speed gear pump (Cole-Parmer, Vernon Hills, IL) and a high-pressure pump (Hydracell, Minneapolis, MN) were used to recirculate the draw and feed solutions, respectively. The cross flow velocities of both channels were kept constant at 12.3 cm/s. The applied pressure was monitored with digital pressure meters and controlled with a back pressure regulator at the feed channel outlet. The temperatures of the feed and draw were maintained at 25.0 ± 1.0 °C.

3. Results and discussions 3.1. Module-scale analysis: performance in counter-current flow mode The previous studies have demonstrated that the counter-current crossflow mode of operation can achieve higher module water flux and feed water recovery [28,29]. The OED process is also partially driven by the osmotic pressure difference between the FS and the DS. During the OED process, the permeated water dilutes the DS until a thermodynamic equilibrium is reached between the feed and draw side. Therefore, final feed water recovery will be determined by the operating conditions including the feed flow rate fraction, the normalized membrane area, and the applied hydraulic pressure. In this section, we examine the effects of various operating parameters on the feed water recovery rate under the counter-current flow mode of OED and FO.

2.3. Feed and draw solutions A shale gas produced water sample was provided from Sichuan, China. Before use, raw SGPW was pretreated with a UF membrane with molecular weight cut-off (MWCO) of 100 kDa. An NaCl solution was used as the draw solution. All the chemicals used were reagent grade. The concentration of the draw solution was adjusted to yield various osmotic pressures to match the osmotic pressure gradient between the feed and the draw solution in the simulation.

3.1.1. Theoretical upper limit of water recovery The OED and FO modules consist of a feed and permeate stream separated by a semi-permeable membrane. We define the feed water recovery, R, and the feed flow rate fraction, ϕF , as follows:

R=

ΔQD QF , in QF , in QF , in + QD, in

(8)

2.4. Measurement of water flux and solute rejection

ϕF =

OED performance with FO membranes in terms of water flux and solute rejection was determined with the laboratory scale OED unit. The water flux (Jw) passing through the membrane was measured based on the weight change of the draw solution. To measure the rejection of cations from the mixture solution in OED mode, samples of the feed and the draw were taken after a complete OED run and the concentration of each solute was measured using inductively coupled plasma mass spectrometry (ICP-MS) and ion chromatography (IC). The observed solute rejection (rion ) was determined based on the concentration of the feed ions in the permeate or the draw solution and was calculated as follows:

where ΔQD is the volumetric rate of the permeated water, QF,in is the initial feed flow rate, and QD,in is the initial draw flow rate. To determine the theoretical upper limit of water recovery, we consider the ideal scenario in which the feed and permeate streams are in thermodynamic equilibrium (i.e., equal net driving force and feed concentration) with the counter-current flow. The following assumptions were made: 1) a module with an infinite membrane area, 2) no solute leakage across the membrane, 3) no hydraulic pressure drop along the module, 4) NaCl solutes for the feed and the draw solution, and 5) osmotic pressures proportional to the concentration using the van’t Hoff equation, π = iCRgT, where i is the van’t Hoff dissociation factor (i = 2 for NaCl), C is the molar salt concentration, Rg is the ideal gas constant, and T is the absolute solution temperature. Using Eqs. (8)–(9), we examined the dependence of the maximum achievable water recovery on ϕ for the counter-current flow. The counter-current flow configuration allows for two possible equilibrium conditions: equal driving force at the feed or at the permeate inlet of the membrane module, as described in [29]. The critical feed flow rate fraction, ϕ*, can be defined by satisfying simultaneous equilibrium conditions, and can be derived from the water and solute mass balance as follows:

CD, ion × VD ⎞ rion = ⎜⎛1 − ⎟ × 100, CF , ion × VP ⎠ ⎝

(5)

where CD, ion is the solute concentration in the draw solution after a given period of time, CF , ion is the average solute concentration in the feed, VP is the volume of water permeating across the membrane, and VD is the volume of the draw solution. Considering the hydraulic pressure applied on the feed side, the water flux (Jw) and reverse solute flux (Js) in OED can be expressed as [15,17,27]:

ϕ* =

( ) ( ) Jw k

( ) ( ) J S − wD

⎞ ⎛ πF , b exp − πD, b exp ⎟, Jw = A ⎜ΔP + B J J S ⎜ 1 + J ⎡exp kw − exp − wD ⎤ ⎟ w⎣ ⎦⎠ ⎝

(9)

CD, in + (ΔP / iRg T ) CF , in + CD, in

(10)

When ϕ is smaller than ϕ*, thermodynamic equilibrium can be reached only at the permeate inlet of module, which is defined as the feed limiting regime (FLR) in the previous study [29]. The maximum achievable water recovery during the FLR can be given by:

(6)

and 284

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Table 3 Parameters used in the numerical simulation at module-scale. Parameters

Value

Unit

Membrane area Channel height Channel hydraulic diameter Pump efficiency Energy recovery device (ERD) efficiency

10 0.76 1.5 80 95

m2 mm mm % %

recovery were numerically determined by solving the mass conservation equations:

Fig. 2. Maximum achievable recovery as a function of feed flow rate fraction (ϕ ) in FO and OED processes for feed concentration of 28.5 and 70 g/L NaCl. Solid and dashed lines represent OED and FO mode, respectively. Simulation conditions include DSOED,in = 65 g/L NaCl at hydraulic pressure of 30 bar; DSFO,in = 101 g/L NaCl at 0 bar; countercurrent flow configurations; no salt leakage.

RFLR =

(11)

When ϕ is greater than ϕ*, thermodynamic equilibrium can be reached only at the feed inlet of module, which is defined as the permeate limiting regime (PLR). The maximum achievable water recovery during the FLR can be given by:

RPLR =

(1 − ϕ)(CP, in − CF , in + (ΔP / iRg T )) ϕ (CP, in − (ΔP / iRg T ))

(13)

d (QF cF ) d (QD cD ) = = Js dAm dAm

(14)

The variations of concentrations and feed water recovery have been simulated along a counter-current module using 28.5, 65 and 101 g/L NaCl as FS, DSOED and DSFO, respectively. Modeling was carried out with Matlab, and its procedures are described in detail in Appendix B. Key parameters used in the simulation are provided in Table 3. Since feed water recovery is directly influenced by the membrane area, the membrane area is an important parameter in system design. However, the membrane module and its replacement account for a substantial portion of capital and operating costs. To evaluate the impact of membrane area on water recovery, normalized membrane area is plotted against water recovery in Fig. 3. Normalized membrane area is the membrane area per inlet feed flow rate and it provides the membrane area required to achieve the specific water recovery depending on the inlet feed flow rate. The draw solution concentration in OED and FO are 65 and 101 g/L, respectively. A hydraulic pressure of 30 bar is applied on the feed side only for the OED mode. The initial feed flow rate fraction is 0.77, which is lower than the critical fraction of 0.78. The red line represents the maximum achievable water recovery obtained from the mass balance, as shown in Fig. 3. For both OED and FO modes, water recovery increased with increasing normalized membrane area. However, it must be noted that the OED mode consistently obtained higher water recovery than that obtained in FO mode. For example, when the normalized membrane area was set to 0.32 m2 h L−1 for CF,in of 28.5 g/L, the OED attained a maximum achievable recovery of 72% for a given set of operating

CD, in − CF , in + (ΔP / iRg T ) CD, in + (ΔP / iRg T )

dQF dQD = = Jw dAm dAm

(12)

In this study, the general equations presented by [29] have been modified to exclude the influence of the solute leakage rate for simplicity. All derivations are provided in Appendix A. Based on Eqs. (10)–(12), the maximum achievable water recovery has been estimated and presented as a function of the feed flow rate fraction in Fig. 2. The concentrations of FS, DSOED, and DSFO are 28.5 (24 bar), 65 (54 bar), and 101 g/L (84 bar), respectively. The hydraulic pressure applied in the OED process is 30 bar (i.e., the effective net driving force for water permeation is 60 bar). A feed flow rate fraction of 0 means that the draw solution continues to flow along the membrane module while the feed flow rate is zero. A feed flow rate fraction of 1 describes the opposite case. Fig. 2 clearly demonstrated that there is no difference between the maximum achievable water recovery in the OED and FO processes due to the same amount of net driving force. However, interestingly, the critical flow rate fraction for OED differs from that of the FO process. Over the critical point, the recovery rate drops sharply because the capacity of the draw stream is insufficient to continue the water extraction from the feed stream. Therefore, both processes will operate with less than the critical feed flow rate fraction. In the OED process, the critical flow rate fraction was 0.74 and 1.08 for feed inlet concentrations of 70 and 28.5 g/L, respectively. It was found that a higher feed inlet concentration results in a lower critical feed flow rate fraction. A value of 1.08 means that maximum recovery can be achieved regardless of flow feed fraction. Compared to the FO process, the maximum recovery in OED can be obtained over a larger range of feed flow rate fraction, which results from the hydraulic pressure not being affected by dilution.

Fig. 3. Water recovery as a function of normalized membrane area in OED (solid line) and FO mode (dashed line) under counter-current flow mode. Green and gray lines represent a feed concentration of 28.5 and 70 g/L NaCl, respectively. Red dotted lines represent the maximum achievable water recovery for a given feed concentration obtained from Eq. (10). The concentration of DSOED and DSFO is 65 and 101 g/L NaCl, respectively. Membrane properties are as presented in Table 2.

3.1.2. Membrane area requirement in OED and FO modules This section demonstrates the superior performance of the OED process in terms of the final water recovery and the membrane area required. We considered both water and solute mass transfer across a limited membrane area. Module-scale concentration profiles and water 285

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Fig. 4. Normalized feed recovery as a function of feed flow rate fraction and hydraulic driving force fraction (horizontal and vertical axis, respectively) for (a) an ideal membrane, (b) a perfect rejection membrane with a support layer, and (c) a realistic membrane. The feed water recovery is normalized by the corresponding ideal maximum water recovery of a given net driving force, Rmax. A water permeability, A, of 2 L m−2 h−1 bar−1 is used for all membranes, and other parameters (salt permeability, B and structure parameter, S) for the calculations are provided in each subfigure. The feed solution is a 28.5 g/L NaCl solution. The normalized membrane area, Am/Qm is 0.2 m2 h L−1 and the net driving force is 50 bar.

above. However, the analysis shows that using the realistic membrane with B= 0.2 L m−2 h−1 and S = 100 µm would allow for more 100% of the normalized water recovery to be obtained in the OED mode. Because the feed concentration is higher than the draw concentration in OED mode, feed solutes continue to be diffused into the DSOED along the membrane module. As a result of salt leakage, the lower concentration of brine at the module outlet allows the OED mode to exceed the maximum achievable recovery obtained from the calculation, assuming a perfect salt rejection membrane in Section 3.1.1. Consequently, compared to the FO process using only an osmotic gradient as a driving force, the OED will always result in a higher water recovery by reducing the detrimental effect of ICP and forward salt diffusion.

conditions. In FO mode, only 65% of feed water could be extracted even when the normalized membrane area was 1.2 m2 h L−1. A similar trend was also found when the feed concentration increased. This simulation result clearly demonstrates that the OED process will be more effective for dewatering than the FO process. Such observation from the simulation is attributed to the dilutive internal concentration polarization (ICP) effect that occurs in the membrane support layer, and substantially reduces the DS concentration in the active layer compared to that in the bulk [30]. Based on Eq. (6), the adverse effect of the dilutive ICP is more severe at higher DS concentrations, resulting in greater loss of osmotic driving force in FO mode. The reduction of osmotic driving force by the ICP is an inevitable consequence of using the asymmetric membrane with the support layer. Therefore, from a practical perspective, the data in Fig. 3 emphasizes the need to utilize hydraulic pressure to reduce total membrane area.

3.2. Energy use in comparison with reverse osmosis 3.2.1. Minimum specific energy of OED-RO and conventional RO To realize the ideal SEC of the OED-RO hybrid process, we evaluated the minimum specific energy and compared it with the energy requirement of the single RO process. The minimum specific energy represents the energy required for a unit volume of permeate when the applied hydraulic pressure is equal to the osmotic pressure difference between the brine and the permeate. Assuming no salt permeation and no inefficiencies of process and ERD, the minimum specific energy is simply given by the osmotic pressure gradient at the RO or OED module exits [31,32]:

3.1.3. Effect of operating conditions on water recovery When the normalized membrane area is fixed, the driving force ratio and feed flow rate fraction can be modified for obtaining maximum water recovery. Thus, we investigate the optimal operating conditions for different membrane modules. Fig. 4 shows contour plots of water recovery as a function of hydraulic driving force fraction, defined as the ratio of hydraulic pressure to net driving force, ΔP /(ΔP + Δπ ) , and feed flow rate fraction for three counter-current membrane modules. The modules shown in Figs. 4a and 4b represent ideal membranes with perfect rejection (B = 0 L m−2 h−1) and/or no internal concentration polarization (S = 0 µm). Fig. 4c depicts a realistic membrane with salt leakage and a support layer (B = 0.2 L m−2 h−1 and S = 100 µm). The water recovery is normalized by the corresponding maximum achievable recovery for a given net driving force, Rmax. A water permeability of 2 L m−2 h−1 bar−1 is used for all membranes. For all cases, the effect of ECP was taken into account. Fig. 4 shows that the feed water recovery slightly increases with a decrease in the feed flow fraction and an increase in the hydraulic driving force fraction. A higher feed flow fraction enables the draw solution to maintain a relatively high osmotic pressure even when the solution is diluted by the permeated water from FS. For a similar reason, hydraulic pressure is not affected by the dilution of the draw solution, thereby allowing a higher hydraulic driving force fraction to obtain higher water recovery [17]. The membrane properties also have a significant impact on the attainable water recovery, as shown in Fig. 4. For an ideal membrane, a normalized water recovery of 98% can be achieved. With a structure parameter of 100 µm, a lower water recovery was observed compared to the ideal membrane due to the detrimental effect of ICP, as discussed

SEC =

Δπin (1 − R)

(15)

The minimum specific energy of a single RO and OED-RO hybrid process can be graphically represented in Fig. 5. The blue or green area under the osmotic pressure curves indicate the theoretical minimum energy for water separation (equal in magnitude but opposite in sign to the free energy of mixing). According to Eq. (15), the rectangular area below the dashed pressure line is proportional to the minimum specific energy. The OED-RO hybrid process requires the same minimum specific energy as the single RO process because of mass conservation of water in the whole system. For the condition with 28.5 g/L NaCl feed and 75% recovery, the minimum specific energy in both processes was 2.69 kWh/m3. In a counter-current OED-RO hybrid system, the minimum hydraulic pressure in OED is the osmotic gradient between the final brine and the initial DSOED (i.e., POED = πFS, out − πDS, in ). After the OED process, the diluted DSOED is sent off to the post RO process as the feed solution, and re-concentrated by the RO process to maintain the same initial DSOED concentration at the module inlet. The minimum hydraulic pressure in post RO is the osmotic pressure of the initial 286

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Fig. 5. The osmotic pressure as a function of water recovery in (a) a one-stage RO and (b) an OED-RO hybrid process. Theoretical minimum energy for dewatering and draw regeneration as a function of water recovery is represented by the blue and green area under the osmotic pressure curve, respectively.

the feed channel. Fig. 6a shows how the process inefficiencies affect the specific energy of each system. As anticipated, the SEC of both configurations increases with increasing total water recovery. For the whole range of water recovery, the one-stage RO process has a lower SEC compared to the OED-RO system, in which an additional pump and ERD are incorporated. However, as total water recovery increases, the energy loss associated with pressure pumps and ERDs decreases, and its contribution to SEC also decreases. For a high total water recovery of 75%, the SEC of the three configurations, an ideal one-stage RO, a one-stage RO, and an OED-RO, are 2.69, 3.4, and 3.6 kWh/m3, respectively. In these conditions, the thermodynamic energy efficiency, defined as the ratio of theoretical minimum energy consumption to the actual energy, was calculated as 32% for one-stage RO and 30% for the OED-RO process. Since the efficiency of thermal-driven processes is estimated to range from 6% to 8% [13] due to the mechanism of water transfer through the phase change, the OED-RO process would be energy-efficient in treating high salinity wastewater with a higher osmotic pressure than the burst pressure of the RO. Fig. 6b compares the profiles for hydraulic pressure required for the one-stage RO and OED-RO. If typical RO operating pressure is 60 bar, the one-stage RO achieves 55% water recovery for a 28.5 g/L NaCl feed solution. Typical RO membranes are designed to withstand high hydraulic pressure. In contrast, osmotic membranes are designed with a thin support layer to reduce ICP, but recent studies have reported that osmotic membranes could be operated at a hydraulic pressure of up to 70 bar [36]. Thus, if the maximum allowable hydraulic pressure of the

DSOED (i.e., PpostRO = πDS, out ), and consequently the total hydraulic pressure in the OED-RO process is equal to the minimum hydraulic pressure in the single RO process (i.e., POED + PpostRO = PRO ). 3.2.2. Effect of inefficiency on specific energy consumption SEC modeling of the processes outlined above used some ideal assumptions (e.g., no salt leakage and no process inefficiency) for simplicity and by allowing a more intuitive understanding of the results. The energy consumption was further evaluated by considering process inefficiency, including concentration polarization, frictional pressure losses, pump efficiency and ERD efficiency. For the conventional RO process, a typical one-stage RO configuration was considered, in which ERD is installed. The OED process also requires a high-pressure pump and ERD. Therefore, it is expected that the energy loss from such process inefficiency is almost doubled in the OED-RO hybrid system, compared to the one-stage RO. The concentration of DSOED at the module inlet was set to half the concentration of FS at the module outlet. The efficiency of the highpressure pump and ERD was set to 80% and 95%, respectively [33]. The frictional pressure loss in the feed channel is calculated using a semiempirical equation [34,35]:

Re 2 ρν 2 ⎞ dP = 0.8Re−0.19 ⎛ 3 dL ⎝ dh ⎠ ⎜

⎟

(16)

where Re is the Reynolds number, ρ is the solution density, ν is the kinematic viscosity of the solution, and dh is the hydraulic diameter of

3.5

Hydrauilc pressure (bar)

Specific energy consumption (kWh/m 3)

100

Ideal one-stage RO One-stage RO OED-RO

2.5

1.5

Δ P one-stageRO Δ P OED + Δ P post RO

80

Δ P post RO

60

Allowable hydarulic pressure

40

20

(a)

(b)

0.5

0

5

15

25

35

45

55

65

75

Total recovery, R (%)

5

15

25

35

45

55

65

75

Total recovery, R (%)

Fig. 6. (a) The SEC as a function of water recovery in ideal one-stage RO, one-stage RO and OED-RO process, and (b) the hydraulic pressure required for each process. The feed concentration of CF,in is 28.5 g/L, simulating SGPW. The feed flow fraction rate is 0.5. The properties of the osmotic membrane were set to water permeability, A of 2 L m−2 h−1 bar−1, salt permeability, B of 0.2 L m−2 h−1 and structure parameter, S of 100 µm.

287

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Fig. 7. (a) Modeled and experimental concentration profiles as a function of time, and (b) modeled and experimental water flux as a function of water recovery for the SGPW dewatering experiment. By injecting highly-concentrated stock solution, the DS concentration was adjusted according to the concentration profile for DS from the numerical simulation. Red star symbol represents the concentration of the feed solution at the end of the experiment. Experiment conditions include: membrane = CTA FO; feed solution = shale gas produced water (SGPW) pretreated by UF (100 kDa); cross flow velocity = 8.5 cm/s; hydraulic pressure = 30 bar; temperature = 25.0 ± 1.0 °C. Simulation conditions are as presented in Tables 2, 3.

which accounts for 67% of feed water recovery. The results obtained from the experiments were close to the data from the numerical simulation. If sufficient membrane area or operation time were available, more water could be extracted from the FS until the maximum achievable recovery was reached. Fig. 7b shows the modeled and experimental water fluxes as a function of time (or relative position along the membrane module). The solid line indicates the simulated water fluxes at varying FS and DS concentrations used in the experiment. Declining permeate water fluxes over time reflect the effect of the elevated osmotic gradient between the FS and DS. This is because the effect of the dilutive ICP becomes more severe as solute concentrations increase, resulting in further decrease in the concentration of DS on the membrane surface. Furthermore, the experimental observations for the water flux also agreed well with the simulation results. To ensure the quality of water obtained from SGPW by the OED process, the rejection rate of the most prevalent ions in the SGPW was measured. Table 4 shows the ionic composition of the feed and draw solutions before and after the OED experiment operated at a recovery rate of 67%. Interestingly, the rejection rates for Na+ and Cl− were higher than for divalent ions. This observation is attributed to a low concentration gradient caused by the use of NaCl as a draw solution. For this reason, the rejection rate of potassium, a monovalent ion like Na and Cl, was much lower than for the divalent ions. The rejection rates for divalent ions were above 97%, which were lower than those reported in previous studies. The relatively low rejection rate may be explained by the elevated feed concentration under high recovery operation. Such an elevated feed concentration leads to a decrease in water flux without a subsequent decrease in salt flux, and hence the rejection rate decreased. In addition, membrane deformation and stretching under high hydraulic pressure may affect the salt permeability, causing a decrease in rejection rates. While a subsequent RO process for regeneration of diluted DS solution rejects these ions more effectively, different rejection rate for these solutes in the OED and RO processes would result in their accumulation in the draw solution. This accumulation of salts in the DS not only reduces water flux in the RO, but also degrades the final product water quality [37]. Thus, the draw streams of the integrated process may be treated by an additional process such as ion exchange to selectively remove accumulated solutes before being re-concentrated in the RO. In addition, using a highly selective membrane in the first separation step (OED) can minimize the accumulation of undesired feed solutes in the closed-loop draw solution.

Table 4 Inorganic composition of the feed, initially, and of the draw, terminally. Items

Initial SGPW (mg/L)

SGPW brine (mg/L)

Final DS (mg/ L)

Rejection (%)

Na+ Ca2+ Ba2+ Sr2+ Mg2+ K+ ClTDS

6901 431 220 91 52 246 18,464 28,596

22,427 1285 664 317 172 1431 57,480 85,215

17,114 14.2 6.5 3.1 0.5 88.2 39,599 56,833

100 ± 0.0 96.9 ± 0.4 97.3 ± 0.2 98.2 ± 1.2 96.1 ± 4.9 87.8 ± 8.4 99.9 ± 0.0

osmotic membrane is set to 60 bar, the OED-RO is capable of over 75% water recovery for these given conditions. Therefore, to be competitive with the conventional RO process, the OED-RO must be operated within a recovery range in which the RO cannot operate.

3.3. Application of OED for dewatering shale gas produced water To evaluate the feasibility of OED-RO as a dewatering process, we carried out further laboratory experiments with shale gas produced water as a feed solution and compared it to the model output. A membrane cell used in the laboratory test unit may be not suitable for demonstrating module-scale simulation results under counter-current flow mode, because the operating conditions of the module simulation, such as normalized membrane area, pressure drop, and mass transfer, are quite different from those of the laboratory test cell. Thus, we assumed that the membrane module consists of a row of numerous small membrane coupons used in the tests. In addition, since the draw solution was diluted as the experiments proceeded, we manually injected the concentrated stock solution to adjust the DS concentration based on the concentration profiles from the simulation results. Fig. 7a shows the modeled and experimental concentration profiles of the FS (blue) and DS (green) as a function of time (or relative position along the membrane module) in counter-current flow mode during the OED runs. The initial feed flow rate fraction in the simulation was set to 0.77, which is lower than the critical fraction of 0.78. Other parameters are listed in Table 3. The solid and dashed green lines indicate the FS and DS concentration profile obtained from numerical modeling under counter-current flow mode. The feed concentration gradually increased as the DS concentration increased. The final water recovery obtained from the simulation was 72%. In comparison, the experimentally determined FS concentration reaches about 85 g/L, 288

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4. Conclusions In this study, we performed numerical modeling to fundamentally demonstrate the advantages of the OED process for high salinity wastewater and to practically assess the impact of operating conditions, such as flow fraction, hydraulic driving force fraction, normalized membrane area, and feed concentration on its performance at the module-scale. The OED experiments using real shale gas produced water were carried out to validate the simulation results. The results of this study advocated that the OED-RO hybrid process can be a promising technology for dewatering high salinity wastewater because of its higher water recovery and energy efficiency. The major findings are summarized as follows:

•

• Hydraulic

•

driving force fraction plays a significant role in determining performance for dewatering of high salinity wastewater. Our modeling showed that the OED process can offer advantages over typical FO operation, including higher water recovery, lower concentration of the diluted draw, and significant reduction in membrane area requirements. This is attributable to the enhanced water transport through the membrane without adverse effects of ICP in OED process. Through numerical modeling, minimizing energy losses from the inefficiency of pump and ERD is relatively important in the OED-RO process since additional pump and ERD are required in OED.

However, such energy losses could be reduced with an increased water recovery in the OED-RO hybrid process. For 28.5 g/L feed at a recovery of 75%, thermodynamic energy efficiency of one-stage RO and OED-RO hybrid process were projected to be 32% and 30%, suggesting that the OED-RO process can enhance the recovery of high salinity wastewater with the similar energy efficiency as onestage RO. For real shale gas produced water, the OED process was found experimentally to achieve 67% water recovery at a hydraulic pressure of 30 bar, which produced a brine concentration of 85 g/L. The results obtained from the laboratory scale experiments closely matched the data from the numerical simulation. The OED process demonstrated a high rejection of most ions even when the feed water recovery reached 67%.

Future studies should aim to optimize the system design based on economic assessments, effectively mitigate fouling and scaling, and develop an osmotic membrane customized for the OED process.

Acknowledgments This work is supported by the Korea Agency for Infrastructure Technology Advancement (KAIA) grant funded by the Ministry of Land, Infrastructure and Transport (Grant 18IFIP-B116952-03).

Appendix A. Derivations of maximum achievable recovery The conservation equation for the water and solute in the feed and draw streams can be written as:

QF , in − QF , out = QD, out − QD, in = QP = RQF , in

(A.1)

QD, in CD, in = QD, out CD, out

(A.2)

QF , in CF , in = QF , out CF , out

(A.3)

When thermodynamic equilibrium is reached only at the permeate inlet of the module, referred to as the feed limiting regime (FLR) in the main manuscript, this concentration can be described as:

CF , out = CD, in +

ΔP iRg T

(A.4)

When thermodynamic equilibrium can be reached only at the feed inlet of module, referred to as the permeate limiting regime (PLR) in the main manuscript, the concentration can be described as:

CF , in = CD, out +

ΔP iRg T

(A.5)

By combining Eqs. (A.1), (A.3) and (A.4), we can express QF,in in terms of the initial concentration and flow rate conditions, which can be described as:

(

QP CD, in + QF , in =

ΔP iRg T

CD, in − CF , in +

)

ΔP iRg T

(A.6)

Using Eqs. (A.1), (A.2) and (A.5), QD,in can also be expressed in terms of the initial concentration and flow rate conditions.

(

QP CF , in − QD, in =

ΔP iRg T

CD, in − CF , in +

)

ΔP iRg T

(A.7)

At the critical feed flow rate fraction, ϕ*, both the boundary conditions, Eqs. (A.4) and (A.5), should be satisfied, and therefore ϕ* is expressed as Eq. (10) in the main manuscript. By combining Eq. (A.3) with Eq. (A.4), the maximum achievable water recovery during the FLR can be rewritten in terms of the initial concentration and hydraulic pressure, expressed as Eq. (11) in the main manuscript. Similarly, using Eqs. (A.2) and (A.5), the maximum achievable water recovery during the FLR can be determined in terms of the initial concentration, feed flow fraction and hydraulic pressure, expressed as Eq. (12) in the main manuscript. Appendix B. Modeling strategy at counter-current flow mode A numerical solver via computer modeling was employed for performance prediction in the counter-current flow mode. With the target recovery 289

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of the system, the objective function is to minimize the difference between the target and the actual water volume to be drawn. This is determined by a pressure applied in OED associated with energy consumption. A given membrane area is discretized with sufficient integers, which determines the resolution of the optimal solution. The membrane properties such as the water permeability coefficient (A), salt permeability coefficient (B), and structural parameter (S) are set based on the membrane applied to the OED. For the initial iteration, each side of the OED is filled with each solution supplied. The calculation is performed from the feed inlet section/component. Based on the solution-diffusion model and hydraulic conditions, the water is extracted from the feed to the draw, and the solute transports across a membrane. The concentrated feed is then fed to the next discretized system, and the diluted draw solution moves in the direction of the RO feed system. A single iteration is conducted and run until the calculation of last discretized element ends. The next iteration is carried out with the saved data from the previous iteration, excluding the inlet solution concentration. After the iteration time exceeds the limit, the calculated result can be obtained from the program. Thus, the number of iterations should be sufficient for OED to reach a steady-state condition. In the steady-state condition, the total water volume obtained is calculated and then compared with the target water volume. Until the objection function reaches the minimum within a tolerance, the applied pressure is adjusted. From the optimal applied pressure (feasible solution), the SEC of OED is then calculated considering both the energy recovered by ERD and lost by pump efficiencies.

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