What Will it Take to Get to 250,000 ppm Brine Concentration via Ultra-high Pressure Reverse Osmosis? And is it Worth it?

Arezou Anvari, Jishan Wu, Arian Edalat, Nikolay Voutchkov, Ahmed Al-Ahmoudi, Subir Bhattacharjee, Eric M.V. Hoek

Abstract

The possibility of ultra-high pressure reverse osmosis (UHPRO) to replace thermal desalination promises reductions up to 50% in both energy consumption and capital cost for brine concentration. However, commercially available reverse osmosis (RO) membranes suffer significant performance degradation under ultra-high pressure due to severe membrane compaction and embossing. High temperature operations such as in Middle East seawater desalination and brine mining further exacerbate both issues due to softening of the polysulfone support membrane. Assuming that pressure and temperature-related compaction issues can be resolved through innovations in membrane materials and module design, this study asks the questions, “What will  it take to get to 250,000 ppm via ultra-high pressure reverse osmosis? And is it worth it?” This work provides a comprehensive examination of the RO process at high pressure—specifically SWRO to HPRO to UHPRO. We elucidate the influence of thermophysical properties, hydraulic pressure, permeate flux, water permeability, feed spacer design, concentration polarization and salt rejection and stage design on recovery and specific energy consumption (SEC). Our analysis of high rejection (HR) and high flux (HF) RO membranes reveals a dual-edged nature of HF membranes: while they boast superior permeability, enabling reduced trans-membrane hydraulic pressure, they require higher system level water flux, which induces significant concentration polarization (CP). This in turn leads to CP-limited product water recovery, which is a novel finding. In addition, membrane compaction drives up the applied pressure and SEC, and also the capital cost of UHPRO systems. We further assess changes in feed spacer designs that could dramatically reduce CP and enhance product water recovery. This paper presents a holistic roadmap of how system design drives these parameters. The findings herein are poised to guide both technological refinements and best practices for RO membrane brine concentration, propelling the water treatment industry towards a more sustainable future.

1 Introduction

Over the last six decades, the emergence and evolution of reverse osmosis (RO) membrane technology has revolutionized water treatment, particularly desalination [1, 2]. Originally a conceptual framework confined to laboratories, RO technology has burgeoned into an incredibly versatile water treatment technique and its range of applications now extends to seawater and brackish groundwater desalination, various ultra-pure water and process separations in addition to industrial and municipal wastewater recycling [3]. Recent advancements in RO membrane materials [4] and process technologies [5] have further enhanced its capabilities. These improvements have driven the efficiency of seawater desalination to within 10 to 20% of the thermodynamic performance limits. Similarly, for brackish and low-salinity water treatment, these enhancements have achieved efficiency levels within 30 to 40% of the theoretical limits [6]. Despite these significant improvements in RO membrane materials and process technologies over several decades, the pursuit of innovations in brine concentration continues. Alternative separation concepts like osmotically-assisted reverse osmosis (OARO), membrane distillation (MD), forward osmosis (FO), capacitive de-ionization (CDI), microbial fuel cell-assisted electro-dialysis (MFCED), and humidification-dehumidification (HDH) are actively being pursued for industrial development [1]. These techniques retain considerable appeal due to their potential to deliver superior water quality and enhanced recovery rates for hyper-saline brines. They also demonstrate the possibility to desalinate waters that challenge the capabilities of RO, such as high-temperature waters and hyper-saline brines.

One major constraint of modern RO membrane technology is the limited availability of high-pressure RO membrane modules and pressure vessels. Typical seawater RO products are not designed to exceed 80 bar (1,200 psi) of operating pressure [7]. Notwithstanding potential recovery limitations imposed by mineral scaling, this pressure limitation prevents one from achieving economically viable product water recovery for feed waters with higher salinity than seawater. If an RO membrane product were developed to withstand continuous operating pressures needed to concentrate brine high enough that it is suitable for crystallization (>250,000 mg/L). Additionally, with a high-pressure tolerant RO membrane one could realize an FO/RO integrated process capable of seawater desalination at high recovery using simple salt-based draw solutions, i.e., no heat required to recover thermo-reversible polymers [8] or thermolytic salts [9]. Further, one could employ RO to further concentrate and valorize SWRO brine, which appears of great interest in the Middle East [10] to concentrate other hyper-saline brines such as oil and gas produced waters from hydraulic fracturing and brines targeted for direct lithium extraction [11].

Given the above review and simple analysis, one might ask the question, “Why cannot commercially available RO membrane technology withstand higher operating pressures and temperatures?” Most modern RO membranes are thin film composites (TFC). These membranes are constructed from polyamide (PA) coating film chemistries [12] polymerized in situ onto polysulfone (PSF) porous supports previously phase inverted onto polyester (PET) non-woven fabrics. Over 50% of performance (permeate production) loss of the composite RO membranes have been reported at high applied pressures [13]. This significant performance degradation is primarily ascribed to the compaction or densification of the mesoporous PSF support layer, a finding corroborated by several studies [13-15]. Pendergast et al.[14] examined hand-cast composite RO membrane compaction at pressures reaching 35 bars, documenting a reduction in the thickness of the PSF support layer between 20% to 30%, accompanied by up to 50% loss of water permeability. Davenport et al. [15] reported a 35% decrease in permeability and a 60% reduction in cross-sectional thickness of commercial seawater RO membranes when operated at 150 bars. In a more recent study, Wu et al. [13] explored commercial RO membrane compaction at pressures up to 207 bars, unveiling a thickness reduction in the PSF support layer between 38% 98 to 60%, resulting in over 50% water permeability loss[13].The densification and collapse of skin layer pores, situated directly beneath the PA coating film, extends the effective diffusion path length through the composite PA-PSF membrane structure, thereby diminishing the permeability of both salt and water [13].

Having identified a significant performance gap in RO membrane technology, it is clear why interest in traditional thermal desalination processes persists, while interest in new alternative desalination processes and technologies continues to grow. This study evaluates SWRO, HPRO and UHPRO processes. The key areas we examine are:

1. The effect of the number of elements in series on the overall desalination performance.

2. The role of spacer designs, particularly focusing on porosity and thickness, in influencing the efficiency of desalination processes. membranes in desalination.

3. The comparative performance of high-rejection (HR) membranes versus high-flux (HF) long-term performance.

4. The potential advantages and drawbacks of compaction-resistant membranes in sustaining.

5. The energy consumption analysis on SWRO, HPRO and UHPRO processes at different recoveries and operating pressures.

2 Model development

The reverse osmosis (RO) membrane system is modeled by interpreting each spiral-wound element as a series of thin, parallel rectangular flow channels, with a feed spacer of a defined thickness and porosity. The curvature of the element is disregarded, as it was determined that assuming a rectangular geometry rather than a bent envelope doesn't notably affect the depiction of flow behavior within the feed channel [16]. Each flow channel is discretized into small incremental unit volumes. The local influent, effluent, residual pressures, flows, and concentrations are calculated from a forward difference solution to the differential mass and momentum balances presented. A local mass transfer coefficient, influenced by the local retentate flow rate, is calculated to determine the effects of solute concentration polarization on bulk, membrane, and permeate concentrations. Locally determined mass and momentum transport parameters are then integrated across the entire system to determine global system performance.

The schematic of the model is illustrated in Fig.1, where NF pretreated feed with 32 g/L NaCl is pumped into a series of SWRO elements with a recovery rate of 50%. The retentate from the first SWRO stage is then pumped into a second stage HPRO system, also at 50% recovery. Finally, the retentate is concentrated up to 250 g/L in the last UHPRO stage.

To construct the model, several key assumptions are made. The feed stream chemistry is considered to be pure NaCl brine, with the thermo-physical properties strictly applicable to the feed stream. The simulations do not consider any deleterious effects of fouling, scaling, membrane compaction, or chemical cleaning. Baseline water permeance (A) and solute permeability (B) values are assumed to be 1 LMH/bar and 0.06 LMH for seawater RO (SWRO), 0.8 LMH/bar and 0.08 LMH for high-pressure RO (HPRO), and 0.6 LMH/bar and 0.1 LMH for UHPRO, respectively. The lower A value for HPRO and UHPRO are attributed to membrane compaction due to the higher applied pressures. In other simulations, the A and B values of SWRO are simulated within HPRO and UHPRO systems. In scenarios where high-flux SWRO membranes are used, the A and B values are set at 2.4 LMH/bar and 0.0912 LMH, respectively.

Fig. 1. Schematic of the full-scale model setup. NF pretreated feed with 32 g/L NaCl pumped into a series of SWRO elements with 3057.56 gpm flow rate at 50% recovery. The retentate of the first stage of SWRO is pumped into the second stage HPRO system with 50% recovery. And then with final UHPRO stage, the retentate is concentrated up to 250 g/L.

2.1 Governing equations

In the membrane system, the primary driver for water permeation is the difference between the feed side hydraulic pressure and the trans-membrane osmotic pressure. For solute permeation, the concentration gradient across the membrane is the dominant force. Based on the solution diffusion model [17],water flux, Jw (=vp), the local permeate velocity) and solute flux ,Js (=vpcp ),are described by

where A is the intrinsic water permeability, B is the intrinsic solute permeability, ∆p is the available hydraulic pressure, ∆π is the trans-membrane osmotic pressure. In addition, ∆cm(=cm-cp) is the trans-membrane concentration difference, cm and cp are solute concentrations at the membrane-solution interface at the feed and permeate sides, and rs (=1-cp/cm) is the intrinsic local salt rejection.

According to Mulder [18], one can re-arrange eq (2) to obtain cp=cmB/(vp+B), which indicates that local permeate concentration varies with solute concentration, solute permeability, and permeate velocity. The expression for permeate concentration can be additionally re-arranged to produce a useful expression for the local (real) rejection by an RO membrane,

The rejection is seen to depend on permeate velocity. The system's local separation performance is shaped by variables like hydraulic pressure, retentate crossflow velocity, and concentration. These variables are further affected by transport within the spiral wound element and membrane fouling phenomena. In our model, while water and solute permeability serve as fitting coefficients and may change over time due to membrane compaction [13], we assume that water permeability, solute permeability, and solution osmotic coefficient are held constant throughout the system. Factors such as solute rejection, permeate velocity, osmotic and hydraulic pressures, solute concentration, and retentate velocity differ from inlet to outlet, as detailed below.

The applied hydraulic pressure drives water permeation through RO membranes. Yet, as retentate flows over the membrane and feed spacer surfaces, hydraulic pressure losses occur from the system's inlet to its outlet [19, 20]. In our model, the Poiseuille equation is combined with the empirical friction factor of Schock and Miquel [16] to describe the available hydraulic pressure at each axial location within the feed side crossflow channel as

Here ux is the crossflow velocity through the spiral wound elements, fsp(= 6.23Re-⁰.³) is an empirical friction factor accounting for the presence of a feed spacer, p is the solution density, dH is the hydraulic diameter, and Re is the local Reynolds number.

We assume the solution's density and viscosity in each stage to be independent of temperature and pressure assuming the process to be isothermal and largely isobaric. Concentration dependence of these properties is considered (see Section 2.6). However, retentate velocity, Reynolds number, and spacer friction factor vary throughout the system. We also allow for local variations in hydraulic diameter to account for changes in channel cross-section due to fouling. Factors such as total feed flow rate, the number of parallel pressure vessels (modules), leaf count per element, leaf and spacer dimensions (including thickness and porosity), play a crucial role in describing the module's hydraulic pressure drop [16].

The local crossflow velocity is determined from a differential mass balance on water flowing through the spiral wound elements as

where u0 is the inlet flow velocity (at x = 0). In eq (5), inlet velocity for each stage is determined by dividing the influent volumetric flow rate by NM NL εsp WH*,where NM is the number of modules (pressure vessels containing spiral wound elements) into which the feed water is pumped in parallel, NL is the number of membrane leafs (crossflow channels) per spiral wound element, εsp is the effective porosity of the flow channel created by the feed spacer, W is leaf width, and H* is the effective local channel height. The local permeate volumetric flow rate, Qp, is  2.vp NM NL W. dx where dx is the length of membrane considered. The factor “2” accounts for water permeating through both top and bottom sides of a membrane leaf.

where c0 is the inlet solute concentration. Equation (6) is used with eqs (1 – 5) to predict the local trans-membrane osmotic pressure and solute permeation. It is necessarily iterative because cp depends on cx.

When the local salt rejection is known, the extent of concentration polarization (i.e., the CP factor) can be estimated from [18, 21]

 In eq (7), ks is the solute mass transfer coefficient, which is determined from

where Re (= uxdH/) is the crossflow Reynolds number, Sc (= /D) is the Schmidt number, D is the solute diffusivity, and  v(=u/p) is the solution kinematic viscosity. The hydraulic diameter, dH, for a spacer-filled thin rectangular channel is 2εspH. The local retentate fluid velocity governs the mass transfer coefficient, and therefore, influences the local CP factor, solute rejection, and trans-membrane osmotic pressure.

2.2 Thermo-physical properties

Thermo-physical properties were determined from polynomial regressions of OLI systems software projections at different concentrations up to 4M NaCl:

where c denotes molality (mol NaCl/kgwater).

2.3 Global System Performance

Key system performance parameters for full-scale RO processes include product water quality, product water recovery, applied hydraulic pressure, and specific energy consumption. Global average permeate water concentration,cp, is calculated from the ratio of solute and water flow through the membranes integrated over the entire system,

Here, W is the flow channel or membrane leaf width. The feed (inlet) and global average permeate concentration describe the observed rejection,X (=1-cp/c0 ), which is analogous to conversion in a chemical reactor.

The total product water recovery is defined from the ratio of product and feed water flow rates,

and is equivalent to the yield in a chemical reactor. A certain amount of water is recovered from each element in series. As recovery increases (from inlet to outlet), the local feed-side solute concentration increases and so does the trans-membrane osmotic pressure, the solute passage, and the risk of mineral scale formation. Therefore, the maximum product water recovery is determined by balancing the benefits of more water production against the costs of acid and anti-scalant addition, increased salt passage, and increased energy consumption.

In full-scale RO plants, the product water recovery is maintained at a value safely below the limiting concentration factor, i.e., the brine concentration at which mineral scale formation occurs. The brine concentration factor, CF (= cL/c0), defines the ratio of solute concentrations in the brine and feed streams and is determined from [22]

However, scale formation (by our definition) occurs on the membrane surface; hence, the limiting concentration factor, CFlim(CP.CF) , accounts for the combined effects of concentration polarization and product water recovery on mineral scale formation potential.

Total energy consumption by RO processes is the product of the applied pressure and feed flow rate divided by the electrical efficiency of the pump used to generate the pressure and flow. The specific energy consumption (SEC) is the energy consumed per unit volume of water produced. Substituting the definition of product water recovery yields

where  represents the combined pump and motor efficiencies. We assume a total pump efficiency of 80% in specific energy consumption calculations.

In case where high energy recovery device (HER) is used, the net energy consumed by the system, Enet is the difference between Epump (energy consumption of the pump) and EHER (energy recovered 264 by HER), where:

Qbrine, flow rate of the brine or concentrate stream; ∆Pbrine, the pressure of the brine stream before the HER; ηHER , efficiency of the HER (assumed to be 95%).

3 Results & Discussion

3.1 Effect of number of elements on RO performance

This work investigates the impact of the number of seawater elements (i.e., 8, 6, and 4 elements in series) on the performance of three reverse osmosis (RO) processes, (SWRO, HPRO, and UHPRO). Table 1 outlines the process parameters used in the full-scale model for each process. In every case, the target recovery is set at 50%, and the necessary applied pressures are calculated accordingly. The feed concentrations and feed flow rates for HPRO and UHPRO are derived based on the corresponding values for SWRO and HPRO, considering the target recovery. The base case for each process includes a 34-mil feed spacer with an 88.7% porosity. The local and global performance of all processes, including factors such as solute concentrations, thermophysical properties, feed/differential pressure, and specific energy demand, are presented and discussed in the subsequent sections. This assessment aims to provide a deeper understanding of how the number of seawater elements in a series influences the overall performance of SWRO, HPRO, and UHPRO. This could potentially lead to insights for optimizing these processes and achieving better desalination results with lower energy demands.

3.2 Solute concentration and thermo-physical properties

The feed solute concentrations and normalized thermophysical properties (including viscosity, density, and diffusivity) over the length of the module are illustrated in Fig.2. For all three processes (SWRO, HPRO, and UHPRO), solute concentrations increase almost linearly across the length of the module, doubling to achieve a consistent recovery rate of 50%. The feed concentrations for the SWRO processes start at 32 g/L and the concentration of the UHPRO brine reaches 250 g/L. Despite different initial and final concentrations, the rate of solute concentration growth remains higher for all three processes.

The modeling results demonstrate an increase in viscosity and density, and a decrease in diffusivity across the length of the module for all processes, thus confirming the impact of salinity on thermophysical properties. Concentrating brine up to 64 g/L using SWRO results in a 5.6% increase in viscosity and a 2.1% increase in density, while reducing diffusivity by up to 4%. Meanwhile, HPRO, which concentrates the salt solution up to 125 g/L, sees these changes amplify, reaching up to 13% for viscosity, 4% for density, and a reduction of 8.1% for diffusivity. The most considerable change in thermophysical properties is observed in UHPRO, owing to the highest salinity (up to 250 g/L), with a 33% increase in viscosity, a 7.2% increase in density, and a 20% 300 decrease in diffusivity. These findings highlight the influence of high pressure and salinity on the thermophysical properties of the desalination process and offer valuable insights for the design and optimization of RO systems.

Fig. 2. (a) Model predicted local solute concentrations of SWRO, HPRO, and UHPRO; Model predicted local thermophysical properties: (b) viscosity, (c) density, and (d) diffusivity of SWRO, HPRO, and UHPRO.

3.3 Local pressures

The local hydraulic, osmotic, and trans-membrane pressures of SWRO, HPRO, and UHPRO systems at three different numbers of membrane elements in series are shown in Fig. 3.The model-anticipated pressures underscore the high hydraulic pressures for UHPRO compared to HPRO and SWRO. Concentrating brine up to 250 g/L using the RO system necessitates hydraulic feed pressures of up to 420 bar (as depicted by the blue line in Fig. 3c).

As shown in the results, attainment of a preset recovery of 50% per stage (module) with fewer elements in series necessitates higher feed-side hydraulic pressures (represented by circle markers).Therefore, to achieve a 50% recovery, higher hydraulic pressures are necessary with 4 elements — up to 51%, 41%, and 30% more for SWRO, HPRO, and UHPRO, respectively, when compared to scenarios with 8 elements.

The high hydraulic pressures increase fluxes and recoveries per unit length from the leading elements, resulting in a more rapid axial increase in feed side membrane surface concentrations and osmotic pressures (square symbols). The net transmembrane pressure driving force (represented by crosses in Figure 3) decreases more rapidly between the leading and trailing elements, resulting in a sharper difference between the leading and trailing element recoveries. The osmotic pressures attain high values toward the exit of the stages with fewer elements. It is worth noting that although the feed side membrane surface concentrations were identical at the end of the stage (See Figure 2a), the corresponding osmotic pressures are different. This is due to the different values of permeate concentrations, arising from the difference in salt passage from the leading to the trailing end of the stage. The salt passage is lower (effective rejection is higher) in case of fewer (4) elements per stage than for greater number (8) of elements per stage.

Increasing the number of elements in series helps achieve a desired stage recovery through application of lower hydraulic pressures. With higher number of elements in series, the rates of increase of axial concentration polarization and the axial decline in driving force are more benign, both of which tend to improve the overall efficiency of the RO stage performance. This is most prominent in SWRO (Figure 3a), but becomes relatively less prominent during UHPRO (Figure 334 3c). This observation has implications for both system design and operating philosophy of RO systems. The use of larger number of elements in series increases the capital expenditure for an RO system targeting a given recovery. However, longer pressure vessels with more elements in series achieve a given recovery with lower hydraulic pressures, reducing the overall operating expense of the RO system.

 Fig. 3. Model predicted local hydraulic, osmotic, and trans-membrane pressures of (a) SWRO; (b) HPRO; (c) UHPRO.

3.4 Concentration polarization and permeate flux

The water permeate fluxes at any location within the membrane module are depicted in Fig. 4a. As previously mentioned, the driving force for permeation is the transmembrane pressure, which is the difference between the available feed hydraulic pressure and osmotic pressure. This explains the decrease in local permeate flux in all studied processes, resulting from the decreased transmembrane pressure due to increased local solute concentration (Fig. 2) and osmotic pressures (Fig. 3).Higher permeate fluxes at a lower number of membrane elements (i.e., 4 and 6 versus 8 elements) confirm the increased applied pressures required to recover 50% of the water from the feed solution. As the water permeates the membrane and leaves the module, the local cross flow velocity and Reynolds number decrease, as shown in Fig. 4b. Consequently, the mass transfer coefficient, which is governed by cross flow velocity, is lower across the length of the membrane and for lower numbers of membrane elements. These results suggest that a smaller number of membrane elements in series exacerbates concentration polarization. Model predictions of concentration polarization modulus are shown in Fig. 4d. As the solute concentration and mass  transfer coefficient impact the local CP factor, the highest CP factors are obtained at the lowest membrane elements (i.e., 4 elements) in the UHPRO process. In each process, the CP modulus hierarchy is as follows: 4 elements > 6 elements > 8 elements, and for each membrane element, it's UHPRO > HPRO > SWRO. This verifies that higher applied pressures for the UHPRO process, especially for lower element numbers, are entirely attributed to higher concentration polarizations.

Fig. 4. Model predicted local (a) permeate fluxes; (b) the crossflow Reynolds number; (c) the solute mass transfer coefficient; (d) concentration polarization factor of SWRO, HPRO, and UHPRO processes.

3.5 Local and global rejection and permeate concentration.

The salt rejection results, depicted in Fig. 5a, show a diminishing trend of salt rejection as one progresses along the module's length. This pattern aligns with the decline in permeate fluxes presented in Fig. 4a. Given that salt rejection is influenced by permeate velocity, a reduced flux subsequently leads to decreased salt rejection. Furthermore, operations characterized by elevated salt concentrations and osmotic pressures necessitate a more substantial hydraulic pressure, as evidenced in Fig. 3. This increased pressure results in augmented permeate fluxes, especially noticeable at the module's onset, as illustrated in Fig. 4a. Incorporating fewer elements intensifies both the hydraulic pressure and permeate flux, leading to heightened rejections, a phenomenon evident in Fig. 5a. For a more comprehensive comparison between processes utilizing varying element counts, Fig. 5b presents the average rejection values and permeate concentrations. Generally, augmenting the element count in a process tends to degrade both the permeate quality and the membrane's rejection efficiency. With the escalation of feed concentration and pressures, solute permeability (B) increases from SWRO to HPRO and continues to UHPRO processes. This shift results in diminished rejection and increased permeate concentrations.

Fig. 5. Model predicted (a)Local salt rejection;(b)Global rejections and permeate concentrations (Cp) of SWRO, HPRO, and UHPRO processes with 8,6, and 4 elements.

3.6 Effect of spacer design on RO performance

In RO spiral wound elements, feed spacers are used mainly to support the membrane and create paths for the feed to flow in the channel. However, spacers have important role in enhancing mass transfer and reducing concentration polarization as well as they increase the turbulence of the flow along the surface of the membrane that leads to enhancement of permeate production [6].Therefore, a systematic study on the effect of feed spacer designs on mass transfer and concentration polarization of RO processes, especially UHPRO, is required to optimize their geometry. Here, we studied the effect of commercial spacers’ porosity and thickness on the performance of RO processes. For this purpose, four different porosities of 88.7%, 83.0%, 77.4%, and 66.1% and four different brine spacer sizes of 34mil, 32mil, 28 mil, and 26mil are studied on an RO system with 8 modules in series configuration.

3.6.1 Effect of spacer porosity

The studied cases of spacer porosity are shown in Table 2. For all three processes of SW,HP, and UHP-ROs at constant flow rate, number of elements in parallel and series, and spacer size(i.e., 34 mil), the effect of spacer porosity on RO performance is investigated while setting the overall recovery target to 50%.

The local mass transfer coefficient, which is affected by the local cross flow velocity, is calculated to determine the impacts of spacer porosity on concentration polarization. In the spacer feed channel, the local feed flow velocity is directly dependent on the local flow rate and cross section area which is related to the spacer porosity as 𝑢𝑥= Qf/ H.w.ε.Therefore, lower spacer porosity increases the axial (cross-flow) velocity, resulting in higher Reynolds (Re) number and mass transfer coefficient. The model predicted mass transfer coefficient (Fig. 6a) shows that the value of this coefficient increases at lower porosity - the highest mass transfer coefficient is associated to the lowest porosity - 66.1% in this study. Figure 6 demonstrates reduced CP factor as a result of improved mass transfer coefficient associated with reduced spacer porosity. The impact of spacer porosity on increasing mass transfer coefficient (Fig. 6a) and decreasing CP factor (Fig. 6b) is remarkable for concentrating high salinity solutions through HP and UHPRO processes.Therefore, optimization of spacer design in terms of decreasing porosity up to 25%, improves mass transfer coefficient up to 35%, resulting in decreased CP up to 11%.Model predicted permeate fluxes, Re numbers, solute and permeate concentrations, and rejections at different spacer porosities for all three processes are presented in supplementary information (SI).

Fig. 6. Model predicted local (a) Solute mass transfer coefficient; (b) Concentration polarization factor of SWRO, HPRO, and UHPRO processes at four different spacer porosities of 88.7%, 83%, 77.4%, and 66.1%

These enhanced mass transfer and reduced CP factor can efficiently decrease the required hydraulic pressure of RO processes to produce 50% recovery, especially for UHPRO processes which required extremely high pressures to concentrate brine up to 250 g/L. However, in the spacer 423 filled channels, the enhanced mass transfer due to the impact of advanced spacer designs on the created turbulence of flow is reported to be mostly accompanied by an increase in pressure drop. Therefore, it is required to study the influence of spacer designs on the axial pressure drop as well.

In Fig. 7, we present model predicted local hydraulic and osmotic pressures while decreasing spacer porosity in all RO processes to evaluate the effect of spacer porosity on the axial pressure drops and osmotic pressure losses. The lowest and highest axial hydraulic pressure drops (i.e., 2.01 bars and 4.85 bars) were observed for spacers with highest and lowest studied porosities in this work as 88.7% and 66.1%, respectively (Fig. 7a). Therefore, the pressure drop increased up to 2.41 times by decreasing porosity up to 25%. However, lowering spacer porosity mitigated osmotic losses up to (Fig. 7b) due to the improved mass transfer and reduced CP factor (as seen in 433 Fig. 6).

Model results of required hydraulic pressures at different spacer porosities (shown in Table 3) display minimum pressure at 66.1% for HP and UHP-RO processes while for SWRO, the minimum pressure was observed at spacer porosity of 77.4%. This confirms the tradeoff between increased pressure drop and improved mass transfer along with decreased concentration polarization as a result of spacer design changes. For SWRO process, the enhanced mass transfer is dominant when decreasing the porosity up to 77.4% while pressure drop is dominant by further decrease of porosity to 66.1%. For HPRO and UHPRO processes, the influence of spacer porosity on enhancing mass transfer is more predominant than on increasing pressure drops, leading to lowering applied pressure.

Fig. 7. Model predicted local (a) hydraulic and (b) osmotic pressures of SWRO, HPRO, and UHPRO processes at four different spacer porosities of 88.7%, 83%, 77.4%, and 66.1%

3.6.2 Effect of spacer size

The other important parameter of spacer design to improve RO performance is the brine 451 spacer size. The studied cases of spacer size are shown in Table 4 where for all three processes of ROs at constant flow rate, number of elements in parallel and series, and recovery, the effect of spacer size is examined. For this part, the spacer porosity was chosen as 66.1% for all process as the main goal of this work is optimizing UHPRO process in terms of lowering required hydraulic pressure.

Similar to the spacer porosity, the spacer size directly impacts the local flow velocity as smaller spacer size increases the cross-section velocity, resulting in higher Re numbers and subsequently improved mass transfer coefficient. The model predicted mass transfer coefficients and CP factors are shown in Fig 8. Results show enhanced mass transfer coefficients and decreased CP factors at lower thickness in which the highest mass transfer coefficient and lowest CP are attributed to the lowest thickness, which is 26 mil in this study. However, changes in the CP factor of UHPRO are significantly higher than changes in CP factors of SWRO and HPRO processes at different spacer thicknesses. In other words, concentration polarization decreased up to 5% in UHPRO system by decreasing spacer thickness to 26 mil from the base case with 34 mil (i.e., 24% decrease in spacer thickness) whereas the CP in SWRO and HPRO decreased only 2.5-3%. This observation underscores the importance of mass transfer on CP during UHPRO.

Fig. 8. Model predicted local (a) Solute mass transfer coefficient; (b) Concentration polarization factor of SWRO, HPRO, and UHPRO processes at four different spacer thickness of 34 mil, 32 mil, 28 mil, and 26 mil.

As shown in Table 5, replacing the 34-mil spacer with 26 mil spacers reduced the required hydraulic pressure of UHPRO process due to the enhanced mass transfer and reduced CP factor of channel with thinner spacer. However, in SWRO and HPRO processes, the required hydraulic pressure increased using thinner spacers. This is similar to the effect of spacer porosity, which is attributed to the effect of higher cross flow velocity on pressure drop. For SWRO and HPRO processes, the influence of smaller spacer filaments on increased pressure drop is dominant in comparison to its effect on enhanced mass transfer while it is opposite in UHPRO process. Model predicted local pressures, permeate fluxes, Re numbers, solute and permeate concentrations, and rejections at different spacer thickness for all three processes are presented in SI.

Table 5. Applied hydraulic pressure (bar) of SWRO, HPRO, and UHPRO at different spacer thickness.

Process	Membrane Spacer size	34 mil	32 mil	28 mil	26 mil
SWRO	Applied Hydraulic Pressure (bars)	81.2	81.6	82.9	84.0.0
HPRO		144.2	144.2	145.0	145.8
UHPRO		316.6	315.8	314.9	315.0

 3.7 HF membranes vs. HR membranes

The previous sections were focused on the optimization of membrane elements and spacer design of full-scale RO process with HR membranes. The baseline A and B values of HR membranes, respectively, are considered as 1 LMH/bar and 0.0380 LMH (SWRO), 0.8 LMH/bar and 0.0304 LMH (HPRO), and 0.6 LMH/bar and 0.0228 LMH (UHPRO). The logic of reducing the A value as pressure/salinity go up is to account for compaction up to 20 % and 40 % for HPRO and UHPRO processes, respectively [13]. The goal of this section is to compare the performance of HR and HF RO membranes, especially to concentrate brine up to 250 g/L. For this purpose, A and B values of HF membranes, respectively, are considered as 2.4 LMH/bar and 0.0912 LMH (SWRO), 1.92 LMH/bar and 0.07296 LMH (HPRO) and 1.44 LMH/bar and 0.05472 LMH (UHPRO). Table 8 presents the studied cases for comparison of HF and HR membranes at three processes of SW, HP and UHP RO. In our model, the hydraulic pressure was calculated in each case to produce 50% recovery. The base case for SWRO process is the optimized module condition as 34 mil spacer with 77.4% porosity and 8 membrane elements in series while the base case of HPRO and UHPRO, respectively, are 34 mil spacer with 66.1% porosity along with 8 membrane elements and 26 mil spacer with 50% porosity with 4 membrane elements. All these conditions are same for both HF and HR membranes to allow comparison of their performance. It should be noted that the spacer porosity for HPRO and UHPRO along with the number of elements for UHPRO are lower than the optimized condition values which was obtained in the previous sections for HR membranes. This is because of high permeability of HF membranes and their limitation of producing 50% recovery at high spacer porosity and small modules.

Local feed solute concentrations of both membranes are shown in Fig. 9. As discussed before, solute concentrations increase over the length in all three processes to produce a constant recovery of 50% and ultimate brine concentrations of 250 g/L. Here, in each process, local solute concertation of HF membranes is higher than HR membranes, while reaching to the same brine concentrations. This is attributed to the difference of HF and HR membranes in terms of higher permeability of HF membranes.

Fig. 9. Model predicted local feed-brine solute concentrations of SWRO, HPRO, and UHPRO for HF and HR membranes.

Our model confirms the different performance of HF and HR membranes in all three RO processes (i.e., SWRO, HPRO, and UHPRO), as shown in Fig. 10. Higher applied pressures of HR membranes compared to the HF membranes (Fig 10a-c) are related to the higher permeability of HF membranes. Using HF membranes, lower trans-membrane pressures are required to produce same amount of water permeate as HR membranes (i.e., producing 50% recovery). Even though, local osmotic pressures of HF membranes are higher than HR membranes, which is the result of higher solute concertation as seen in Fig. 9, lower hydraulic pressures are needed compared to HR membranes. Therefore, to concentrate brine solution to 250 g/L while having a module with 4 membrane elements, minimum hydraulic pressure will be required using HF membranes without any changes in module design. However, for modules with higher numbers of membrane elements, HF membranes are not efficient to produce 50% recovery.

As discussed before, HF membranes have higher permeability compared to the HR membranes. Our model predicted permeate fluxes (Fig 10 d) proves higher permeate fluxes for HF membranes compared to HR membranes, resulting in lower mass transfer coefficient and higher CP factors of HF membranes. However, the reduction rate of local permeate fluxes, mass transfer coefficient, and CP factors for HF membranes are higher than HR membranes, which is the result of higher solute concertation (Fig. 9).

Fig. 10. Model predicted local Pressures of (a) SWRO; (b) HPRO; (c) UHPRO processes and local (d) Permeate fluxes; (e) Solute mass transfer coefficient; (f) Concentration polarization factor of SWRO, HPRO, and UHPRO processes with HF and HR membranes.

3.8 Compaction-resistant UHPRO membranes

Recent studies on membrane compaction at high applied pressures in RO processes have shown significant decrements to membrane water permeability [13, 15]. Compaction is likely more severe in HPRO and UHPRO due to high applied hydraulic pressures, resulting in higher energy consumption due to the required higher applied pressures to produce a constant water flux. Therefore, high pressures associated with HPRO and UHPRO processes will cause design challenges including novel membranes and pressure vessel designs to resist high applied pressures. In this study, baseline A values of HR membranes in HPRO and UHPRO, respectively, are considered 0.8 LMH/bar and 0.6 LMH/bar, accounting for 20% and 40% reduction in membranes’ permeability due to compaction at high applied operating pressures. In this section, we discuss the merits of hypothetical compaction resistant HPRO and UHPRO membranes by using the same A value as was used for SWRO (i.e., 1 LMH/bar). In addition, we compare and highlight the advantages of a compaction resistant membrane in HPRO and UHPRO processes relative to the advantages of improved spacer design in terms of porosity and thickness. Table 7 shows the studied cases of compaction resistant membranes with base spacers along with improved spacers’ design.

Fig. 11 shows the predicted local pressures in HPRO and UHPRO processes when using compaction-resistant membranes. Different spacer designs (34 mil and 26 mil with porosities of 88.7% and 66.1%) are compared. With same membrane permeabilities, the processes using low porosity spacers require less hydraulic pressure, although the spacer design is less significant than the effect of membrane permeability. Fig. 12 presents the predicted performance metrics of HPRO and UHPRO systems with compaction-resistant membranes, such as permeate fluxes, solute mass transfer coefficients, and concentration polarization factors. These metrics are crucial for assessing the efficiency of the membrane processes. The decreasing flux is due to increasing osmotic pressure along the channel. The decreasing flux and mass transfer coefficients cause decreasing concentration polarization. It is shown that process with more permeable membrane requires less hydraulic pressure to achieve target recovery. With thinner and less porous spacer, higher mass transfer is observed, in turn, lower CP is induced.

Fig. 11. Model predicted local pressures of (a) HPRO; (b) UHPRO processes with compaction resistant HR membranes in module using 34 mil and 26 mil spacers with porosity of 88.7% and 66.1 %, respectively.

Fig. 12. Model predicted of permeate fluxes, solute mass transfer coefficient, and concentration polarization factor of (a) HPRO; (b) UHPRO processes with compaction resistant HR membranes in module using 34 mil and 26 mil spacers with porosity of 88.7% and 66.1 %,respectively.

 3.9 Specific Energy Consumption

Based on the projected data for energy consumption (Fig. 13) analysis, it becomes evident that two pivotal parameters influencing RO processes are the operating pressure and the process recovery. Across all processes—SWRO, HPRO, and UHPRO—there's a tangible opportunity to significantly reduce the Specific Energy Consumption (SEC) by enhancing process recovery. In contrast, an elevated operating pressure combined with a low recovery rate results in inefficient energy utilization, leading to a marked increase in SEC. The integration of an energy recovery device has been observed to reduce the SEC by as much as a factor of five. Notably, as we transition from SWRO to HPRO, and subsequently from HPRO to UHPRO, the operating pressure approximately doubles, which in turn doubles the SEC. Given this trend, the implementation of an energy recovery device becomes indispensable, especially for high-pressure and ultra-high pressure operations.

At high recovery operation, the decrease in SEC leads to reducing the operating expenditure (OPEX). However, this may inflate the capital expenditure (CAPEX) due to the need for more membrane area and the associated costs. Similarly, at lower pressures, while the OPEX is reduced due to decreased SEC, more membranes (CAPEX) are needed to achieve a particular recovery rate. Conversely, in high-pressure scenarios, while the OPEX increases due to higher SEC, less membranes (CAPEX) are needed to treat the brine and reach the desired recovery rate. This delicate balance between OPEX and CAPEX is a crucial consideration for process design.

Fig. 13. Specific energy consumption (SEC) and specific energy consumption with high energy recovery device (SEC w/ HER) in different processes at different operating pressures and recoveries.

Conclusions

In our comprehensive exploration, we delved into the intricacies of SWRO, HPRO, and UHPRO systems, revealing the complex interplay between membrane efficiency, design nuances, and operational pressures. A detailed examination of high rejection (HR) and high flux (HF) RO membranes unveiled the dichotomous character of HF membranes. Despite their enhanced permeability and reduced trans-membrane pressures, HF membranes provoke increased system level water flux, inadvertently elevating concentration polarization (CP) and curbing product water recovery. Short module configurations, particularly with HF membranes, show promise in UHPRO systems by achieving high rejection rates and recoveries, while also proving advantageous from a lifecycle operational expenditures (OPEX) standpoint.

Membrane compaction remains a formidable barrier in high-pressure systems, significantly affecting water permeability. Addressing this issue calls for the development of compaction resistant membranes and robust pressure vessels, especially as hydraulic pressures rise in HPRO and UHPRO applications. When paired with advanced spacer designs that enhance mass transfer and decrease axial pressure drops, these membranes show potential for overcoming compaction challenges and improving overall system performance.

The design of elements to endure high-pressure compaction is paramount, with a focus on lead elements subjected to the greatest pressures. In UHPRO systems, strategic staging of low and high flux membranes can mitigate flux gradients and optimize hydraulic efficiency (low flux membrane at lead element and high flux membrane at tail element). Additionally, it is important to restrict UHPRO to non-interacting ions (preferably only Na, K, Cl), so pretreatment processes that remove divalent cations and other fouling factors is mandatory.

Energy consumption is a primary metric in RO system evaluation, intricately linked to pressures and recovery rates. The shift from SWRO to HPRO and to UHPRO often doubles specific energy consumption (SEC), underscoring the need for energy recovery technologies in high-pressure settings. This increase in SEC necessitates a careful balance of OPEX and capital expenditures (CAPEX) to ensure economic viability. The interplay between membrane performance, compaction resistance, and energy metrics demands a holistic approach to optimization, essential for advancing sustainable and efficient desalination processes.

 Declaration of competing interest

All authors declare that there are no financial and personal relationships with other individuals or organizations that can inappropriately influence this work; there is no professional or other personal interest of any nature or kind in any product, service or company that could be construed as influencing the position presented in, or the review of, this manuscript.

Acknowledgements

The authors are grateful for financial support for this study provided by Pacifica Water Solutions, UCLA Samueli Engineering School, the UCLA Department of Civil & Environmental Engineering, and the UCLA Sustainable LA Grand Challenge. This material is based upon work supported by the National Alliance for Water Innovation (NAWI), funded by the U.S. Department of Energy, Advanced Manufacturing Office under Funding Opportunity Announcement DE-FOA 0001905. J.W. thanks National Water Research Institute (NWRI), Southern California Salinity Coalition (SCSC), American Membrane Technology Association (AMTA), and North American Membrane Society (NAMS) for the fellowships support.

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