A Novel High Vacuum MSF/MED Hybrid Desalination System for Simultaneous Production of Water, Cooling and Electrical Power, Using Two Barometric Ejector Condensers

Abstract

This work presents a novel trigeneration system for the simultaneous production of desalinated water, electrical energy, and cooling, addressing the challenges of water scarcity and climate change through an integrated and efficient approach. The proposed system combines an 8-stage Multi Stage Flash Distillation (MSF) process with a 6-effect Multiple Effect Distillation (MED) process, complemented by an expander-generator to optimize steam utilization. Cooling production is achieved through a dual ejectocondensation mechanism, which enhances energy recovery and expands operational flexibility. The system’s performance was analyzed using Aspen Plus simulations, demonstrating technical feasibility across a broad operating range: 28.3 to 0.8 kPa and 68 to 4 °C. In cogeneration mode, the system achieves a Performance Ratio (PR) of 12.06 and a Recovery Ratio (RR) of 54%, producing 67,219.2 L/day of desalinated water and reducing electrical consumption by 12.03%. In trigeneration mode, it achieves a PR of 17.81 and an RR of 80%, with a cooling capacity of 1225 kW, generating 99,273.6 L/day of desalinated water while reducing electrical consumption by 3.69%. These results underscore the system’s capability to significantly enhance the efficiency and capacity of thermal desalination technologies, offering a sustainable and high-performing solution for coastal communities worldwide.

1. Introduction

Access to clean water, electricity, and air conditioning is essential for ensuring an adequate quality of life and promoting sustainable development, particularly in marginalized communities. However, these resources are increasingly under threat from global challenges, including the projected population growth to 9.3 billion by 2050 [1,2], drinking water scarcity, and climate change. These challenges drive a growing demand for electrical energy, which is required to power air conditioning systems as well as water purification and desalination technologies.

The outlook for water resource availability is increasingly unfavorable. Although 97% of the Earth’s surface is covered by water [3], inadequate management and uncontrolled exploitation are expected to result in a 40% global water deficit by 2030 [4]. Simultaneously, global energy demand and CO₂ emissions have risen significantly [5], primarily due to the burning of fossil fuels for electricity generation, which powers conventional desalination and air conditioning systems. This exacerbates climate change, with projections indicating a global temperature increase of up to 1.5 °C by 2100 [6]. Furthermore, the demand for air conditioning and refrigeration systems is predicted to rise sharply, from 4.4% in 2010 to 35% by 2050, and potentially reaching 61% by 2100 [7].

The development of systems capable of simultaneously providing multiple services—such as desalinated water, electricity, and air conditioning—offers a comprehensive solution to these challenges. By operating with high energy efficiency, these systems can significantly reduce energy consumption, lower operating costs, and promote a more environmentally sustainable approach [8,9].

Desalination is now considered a key solution to global water scarcity. The two dominant technologies in the market are Reverse Osmosis (RO) and thermal desalination. RO systems account for approximately 70% of desalination plants worldwide, while Multiple Stage Flash (MSF) and Multiple Effect Distillation (MED) systems represent 21% and 7%, respectively [10]. RO systems report an electrical energy consumption of 3.5 to 5.5 kWh/m³ of desalinated water produced. In contrast, MSF and MED technologies consume 1.5 to 3.5 kWh/m³ of electrical energy, in addition to 80 kWh/m³ of thermal energy [11]. MSF and MED systems are more prevalent in regions where water contains high concentrations of dissolved solids, as they require less pretreatment of the feedwater and lower maintenance. Although RO systems consume up to 25% less energy than MSF and MED systems, the higher cost of pretreatment increases operational expenses, making RO less viable in these regions [12].

Continuous advancements in MSF and MED technologies have shown that lowering the top brine temperature (TBT) mitigates corrosion issues and enhances energy efficiency [13]. These systems are commonly referred to as LT-MED. Aly [14] introduced a hybrid system that extended the temperature range from a TBT of 63 °C to a bottom brine temperature (BBT) of 6 °C, achieving a record-low BBT, cost reductions, and high performance, with a Performance Ratio (PR) of 14.2 and a Gain Output Ratio (GOR) of 14.8. Subsequent studies by Shahzad et al. [15], and Saren et al. [16] investigated the hybridization of MED systems with adsorption desalination (AD) technologies, further reducing the BBT to 5 °C. This improvement increased production capacity while simultaneously decreasing the specific energy consumption (SEC).

Research on the hybridization of MED desalination systems has also focused on their integration with conventional or organic Rankine cycles (ORCs), gas turbines, combined cycles, and the utilization of industrial waste heat. These approaches aim to simultaneously produce potable water and electricity, enhancing both the efficiency and overall performance of the systems [17,18]. ORCs are particularly notable for their ability to operate with low-capacity, low-temperature thermal energy sources by utilizing low-boiling organic fluids. Baccioli et al. [19] proposed integrating waste heat recovery from industrial processes into ORC systems to power MED desalination. Two configurations were analyzed, showing improved performance based on the second law of thermodynamics, with the first configuration proving more efficient by utilizing waste heat from the ORC to activate the MED system. Furthermore, Aguilar-Jiménez et al. [20] demonstrated the technical feasibility of integrating the ORC evaporator with the first effect of the MED system, achieving simultaneous operation of both subsystems. This integration resulted in increased water production capacity and PR for the MED/ORC system, with PR improvements of 8.8%, 13%, 17.65%, and 22% for electrical generation capacities of 20, 30, 40, and 50 kW_power, respectively.

The simultaneous production of desalinated water and cooling has been extensively studied. Hybridization proposals combining absorption heat pump cycles or cooling systems with MED desalination technologies are among the most commonly discussed in the literature [1,14,21,22].

Open absorption cycles integrated with desalination systems were investigated by López-Zavala et al. [23], who introduced a LiBr-H₂O absorption cooling system coupled with a single-effect flash desalination system powered by solar thermal energy. Under design conditions, this system achieved a production of 838 L/day of desalinated water and a cooling capacity of 23 kW. Furthermore, they proposed a system integrating two subprocesses: a 14-effect FLASH/MED desalination unit and a LiBr-H₂O absorption cooling system. This integrated energy and mass configuration resulted in a cooling capacity of 2012 kW and a daily fresh water production of 73,569 L/day, with a coefficient of performance (COP) of 6.15 and a PR of 6.63 [24].

The development and conceptualization of trigeneration systems for the simultaneous production of water, cooling/heating, and electric power have been widely reported in the literature, integrating various technologies to achieve this. For example, Maraver et al. [25] presented a trigeneration system combining an organic Rankine cycle (ORC) powered by biomass combustion, a MED desalination system, and an absorption chiller cycle (ACC) for the production of electricity, potable water, and cooling. The analysis focused on the optimal distribution of heat generated by the ORC, with 40% allocated to the MED and ACC systems. The economic evaluation indicated a payback period of 4 to 20 years, depending on biomass resource costs. The study provides an optimistic outlook for the development of trigeneration systems based on renewable energy sources.

In the research by Ghaebi and Abbaspour [26], a novel trigeneration system is proposed, integrating a MED desalination system with thermal vapor compression (TVC), an open-cycle gas turbine as the thermal source, a heat recovery steam generator (HRSG) operating at three pressure levels, and an absorption chiller cooling system (ACC) to produce desalinated water, cooling, and electric power. Heat from the combustion chamber is distributed through the HRSG, adjusting to the three pressure levels. The results show a thermal efficiency of 84.6% and an exergy efficiency of 57.57%.

To provide a clear comparison of existing systems,

Table 1. Comparative Overview of State-of-the-Art Hybrid and Trigeneration Desalination Systems.

Reference	Key Findings	Desalination Technologies	Performance Metrics	Economic Metrics
Mehrpooya et al. [27]	CSP-powered system integrates steam power, multi-effect desalination, and absorption refrigeration, emphasizing exergy efficiency and economic viability.	MED	${\mathsf{\eta }}_{\mathsf{\psi }}$$=66.05\%.$ ${\mathsf{\eta }}_{\mathrm{Th}.}=80.70\%.$	Return Period = 5.738 years. Profit = $6.828 million/year.
Lawal et al. [28]	Heat pump-powered HDH system achieves simultaneous water production and cooling, with feasibility assessed through parametric and economic analysis.	HDH	PR = 4.07. RR = 4.86%. COP = 4.85. SEEC = 160.16 kWh/m3.	Water Cost: $10.68–$20.39/m3.
Abdelhay et al. [29]	Solar-driven polygeneration integrates MED desalination and absorption refrigeration with a steam Rankine cycle, focusing on thermodynamic and economic optimization.	MED	${\mathsf{\eta }}_{\mathsf{\psi }}$$=23.95\%.$ ${\mathsf{\eta }}_{\mathsf{\psi }, \mathrm{Losses}}$= 82.42% in PTC.	Water Cost: $1.247/m3. Cooling Cost = $0.003/kWh.
Ali et al. [30]	Combines adsorption desalination (AD) with humidification-dehumidification (HDH) and ejectors for energy reuse, validated through energy and cost modeling.	HDH AD	SDWP = 83.1 m3/ton. PR = 2.7.	Water Cost = $1.49/m3 (solar), $0.54/m3 (waste heat).
López-Zavala et al. [31]	Novel MED system uses an ejector for vapor condensation and cooling, validated through simulation for coastal applications.	MED	PR = 11.96. COP = 4.7. COPglobal = 9.325. RR = 53%	-
La Torre et al. [12]	High-vacuum MED system with barometric ejector improves efficiency and production via pressure and temperature cascading.	MED	PR = 34.1% higher than conventional MED. RR = 83%	-

summarizes key findings, performance metrics, and economic metrics for recent state-of-the-art single, hybrid and trigeneration desalination technologies.

Finally, this work presents a novel trigeneration system for the simultaneous production of water, cooling, and electric power, integrating a hybrid MSF/MED desalination and cooling subsystem. This system proposes the energetic and mass integration of two thermal desalination technologies: Multi Stage Flash (MSF) with 8 stages and Multi Effect Distillation (MED) with 6 effects. The technological contribution of this hybrid MSF/MED system lies in the extension of the operating range, with a top brine temperature (TBT) of 68 °C and a bottom brine temperature (BBT) of 4 °C in the sixth effect. The elimination of preheaters in the MED section allows all generated steam to be used for brine evaporation in the subsequent effects. Additionally, an electrical expansion-generation subsystem is integrated, which extracts a fraction of the mass flow from the first flash stage to generate electricity and meet the system’s electrical demand. The expanded steam is reinserted into the first effect of the MED section, increasing capacity and efficiency in both desalination and cooling. Furthermore, a condensation process using two barometric ejectocondensers is employed instead of a conventional condenser, facilitating the extraction and adaptation of cold steam to produce desalinated water. The 4 °C product water stream from the final MED effect is then used to generate a cooling effect, further enhancing system efficiency.

The proposed trigeneration system introduces an innovative technology that simultaneously produces desalinated water, cooling, and electrical power. Key innovations include the elimination of seawater preheaters in the MED section, improving overall system efficiency, and the integration of components that optimize energy use. This configuration provides a sustainable approach to meeting water, cooling, and electricity demands simultaneously.

2. System Description

Figure 1 shows the proposed trigeneration system, which consists of: (1) the hybrid MSF/MED thermal desalination and cooling subsystem, (2) the expansion and power generation subsystem, (3) the product water extraction and adaptation subsystem, and (4) the motive fluid cooling and brine dilution subsystem.

The operation of the hybrid MSF/MED thermal desalination and cooling subsystem begins when the seawater stream (18b) is drawn from the Seawater Tank and pressurized to the transport pressure (3 to 5 bar) by the hydraulic pump (P_SW2), exiting as stream (22). This is introduced to the condenser (Cond₈), where it absorbs the latent heat of condensation of the water vapor stream (51), preheating in the process. Then, the seawater stream (23) passes through the condensers of the subsequent stages, gradually increasing its temperature by acquiring the latent heat of condensation. This process continues until it reaches the maximum preheating temperature (65.5 °C) in the condenser (Cond₁), exiting as the seawater stream (30). Subsequently, it is introduced to the expansion valve (EV_gen), where a sudden pressure drop is performed until it reaches the saturation pressure of the TBT of the MSF/MED hybrid desalination subsystem. The stream (31) then enters the generator, where it receives both sensible and latent heat from the heat transfer fluid, which can come either from the solar collector field (stream 91) or from another thermal source. At the generator (GEN) outlet, the seawater stream (32) is in a two-phase (liquid-vapor) state and is introduced into a liquid-vapor separation chamber (S_Br1), where it is divided into two streams: the concentrated brine (33) and the vapor (34). The concentrated brine stream (33) is directed to the second flash distillation stage, where it is introduced into the expansion valve (EV_Br2). This valve suddenly reduces the pressure to the operating condition of stage 2, causing the flashing effect and steam generation, which leads to the concentrated brine stream (40) being in two phases. In turn, the mass flow of the water vapor stream (34) is divided by the splitter (ST₂) which directs a fraction of the mass flow to the expander (stream 35) and another to the condenser (Cond₁) (stream 37). The water vapor stream (37) is completely condensed in the condenser (Cond₁) and then directed to the second flash distillation stage, where it is introduced into an expansion valve (EV_PW2) to suddenly reduce its pressure to the operating condition of stage 2, which causes the flashing effect and vapor generation, resulting in a two-phase product water stream (39). The flashing vapors generated by the sudden pressure drop in the brine and product water of stage 2, are directed to their respective liquid-vapor separators (S_PW2) and (S_Br2), and then mixed in the mixer (Mix₂) and subsequently condensed in the condenser (Cond₂). The liquid streams of brine and product water are directed to the expansion valves of the next stage (EV_Br3) and (EV_PW3) to repeat this process successively until the eighth flash distillation stage.

At the end of the eighth flash distillation stage, multiple distillation effects (MED) are initiated. MED effects involve the use of heat exchangers for the evaporation or release of water from brine streams, using the water vapor generated in a previous effect. In addition, two liquid-vapor separators are used in each effect; one for the product water and one for the concentrated brine, as well as two expansion valves. It is important to note that, in the multiple distillation effects, due to their respective operating temperatures, seawater preheating equipment cannot be used, which is characteristic of high vacuum systems. The water vapor that activates the MED section is obtained from the sudden expansion of the brine and product water from stage eight of the MSF section, where the two-phase product water stream (49) enters the liquid-vapor separator (S_PW8) and splits into two: a vapor stream (53) and a product water stream (54).

The two-phase concentrated brine stream (50) enters the liquid-vapor separator (S_Br8), where it is divided into a vapor stream (57) and a brine stream (56). The streams (53) and (57) are mixed in the mixer (Mix₈) and exit as stream (58), which will activate the first MED effect by releasing its latent heat and condensing in the heat exchanger (HX₁), leaving as stream (60).

The concentrated brine stream (56) enters the expansion valve (EV_Br9), where this valve suddenly reduces the pressure to the operating conditions of effect 1, in addition to causing a flashing effect and the generation of steam, this pressure reduction also leads to a decrease in the temperature of the stream (59). The two-phase stream (59) is directed to the heat exchanger (HX₁) where it receives the latent heat of condensation from the steam streams (58). This heat exchange is possible since the aforementioned vapor stream (stream 58) are at a higher temperature than the two-phase brine stream (59), maintaining a minimum approach temperature of (∆T_min = 2.5 °C) to carry out the heat transfer.

After exiting the condenser (Cond₈), the product water stream (52) is directed to the mixer (Mix₉) together with the product water streams (54) and (60), exiting as the product water stream (64). The product water stream (64) enters the expansion valve (EV_PW9), where the pressure is suddenly reduced to effect 1 operating conditions. Simultaneously, the two-phase concentrated brine stream (61) is directed into the liquid-vapor separator (S_Br9) and splits into two: a vapor stream (62) and a concentrated brine stream (63), where the latter enters the expansion valve (EV_Br10), where the pressure is suddenly reduced to the operating conditions of effect 2. This process is repeated sequentially until the sixth multiple distillation effect is reached. The expansion and power generation subsystem begins by receiving a fraction of the vapor mass flow (stream 35) from the brine vapor-liquid separator (S_Br1). The vapor stream (35) is then introduced into the volumetric expander, where it is drawn in and admitted into an expansion chamber. During this expansion process, the specific volume of the vapor increases (v₃₆ > v₃₅), resulting in a simultaneous decrease of its pressure and temperature. The study of the expansion subsystem details two different modes of operation, characterized by working with different discharge pressures. In mode 1, the specific volume of the stream (35) is expanded to reach the operating pressure of MED effect 1, while in mode 2 it is expanded to reach the operating pressure of MED effect 6. Finally, the subsystem concludes with the insertion of the two-phase stream (36) in the liquid-product water vapor separator corresponding to the effect in which the outlet pressure of the volumetric expander coincides. By reinserting the two-phase stream (36), the vapor mass flow of that stream is combined with the vapor mass flow fraction of the two-phase stream (65) within the vapor-liquid separator (S_PW9). This results in an increase in the mass flow in the vapor stream (67) and, consequently, in a higher heat transfer within the exchanger (HX₂), which causes a higher generation of water as vapor from the two-phase sea brine stream (66). The described phenomenon is repeated successively in the subsequent MED effects, which increases the product water generation and cooling capacity of the proposed trigeneration system.

It is worth mentioning that the kinetic energy of the fluid (stream 35) is extracted by the expansion process and converted into mechanical work. This mechanical work is transferred to the rotary motion by the rotor to the shaft, then the shaft work is channeled to an electric generator for final conversion into usable electrical energy. The electrical energy produced by the expander is used for the self-consumption of the proposed system.

After leaving the sixth effect, the product water stream (84) is mixed in the mixer (Mix₁₃) with the accumulated product water stream (82) from previous distillation stages and effects, resulting in the final product water stream (85). The latter passes to a 3-way valve (3WV), whose main function is to control the direction and/or division of the product water stream (85). It is important to note that the product water stream (88a) is used as the cooling fluid because it is in a saturated state, with a temperature of 4 °C and a pressure of 0.81 kPa. The supply of the cooling fluid to the (stream 88a) is conditioned by the 3-way valve (3WV), allowing its partial or total use, depending on the cooling needs required by the user. On the other hand, the product water stream (88b) is directed to a barometric column (BC2) where its pressure is increased and discharged into the product water tank. The vapor streams (1) and (2) at low temperature and vacuum pressure are conducted to the ejector condenser (EJ1) for condensation by pressure increase.

Figure 2 and Figure 3 show the product water extraction and adaptation subsystem, which includes two ejectocondensers; one that cools the motive fluid with brine and the other with seawater. The principle of operation of the ejectocondenser, like a conventional ejector, begins with the motive fluid entering the ejector (streams 9 and 14), where it passes through the convergent-divergent motive nozzle, increasing its velocity and reducing its pressure (streams 9sc and 14sc). This process draws and entrains the secondary fluid (6 and 93) into the ejector suction chamber. Subsequently, both streams enter the mixing chamber, where they combine, resulting in the transfer of momentum, heat, and mass, exiting as streams (9m and 14m).

Inside the mixing chamber of the ejector, a shock wave is produced due to the deceleration of the mixed stream, resulting in an increase in pressure and temperature of the mixture, all as a result of the transport phenomena mentioned above. Subsequently, the mixture of motive and suctioned fluid (also called primary and secondary) enters the diffuser, where, by increasing the cross-sectional area of the ejector, its velocity decreases, causing a pressure increase and thus achieving total condensation of the water vapor. Finally, the condensed mixture (10 and 15) is directed towards the barometric columns (BC1 and BC3) because it is under vacuum pressure, before being deposited in the product water storage tank at atmospheric pressure.

The ejectocondensation process causes a temperature increase in the mixed streams at the outlet of the ejector due to the transfer of condensation heat from the secondary fluid to the primary fluid. This increase in temperature causes energy accumulation in the product water storage tank, from which the motive fluid is extracted. To counteract the temperature increase, the heat exchangers (HX_Brine) and (HX_SW) are used in their respective condensation circuits. The heat exchanger (HX_Brine) uses the concentrated brine stream (3) as a cold stream (4 °C), which makes it possible to reduce the temperature of the motive fluid stream (9) to 9 °C. On the other hand, the heat exchanger (HX_SW) uses the seawater stream (19) as the cold stream (25 °C), allowing a temperature of the motive fluid stream (14) of 28.5 °C. The two-phase brine stream (4) is split into two in the vapor-liquid separator (S_Br15). The concentrated brine stream (17) is directed to the barometric column (BC4), to increase its pressure and be subsequently deposited in the brine dissolution tank, where it joins with the seawater stream (20). The steam stream (5) enters the flow splitter (ST1) to control the steam mass flow to be condensed in both ejectocondensers. The ejectocondenser (EJ1) condenses the vapor stream (6), which consists of the sum of the vapors coming from the evaporator (1), the sixth effect (2) and the vapor stream (5b). On the other hand, the ejectocondenser (EJ2) is in charge of condensing the remaining vapor (88) that was not introduced in the first ejectocondenser.

Main Characteristics of the System

• Integration of Advanced Thermal Desalination Technologies

The system combines an 8 stage Multi Stage Flash (MSF) process and a 6 effect Multiple Effect Distillation (MED) process, operating under high vacuum conditions. With MSF pressures ranging from 28.3 to 5.1 kPa and MED pressures from 3.8 to 0.8 kPa, the system enhances desalination efficiency and expands operational capabilities compared to conventional designs.

• Dual Barometric Ejectocondensation Process

Two two-phase ejectors operating at different temperatures extract and condense low-temperature steam from the final MED effects, improving energy recovery and efficiency. The concentrated brine stream acts as a heat dissipating fluid in the first ejectocondensation circuit, further enhancing water production.

• Elimination of Preheaters in the MED Section

The system eliminates seawater preheaters, as the absence of a temperature gradient allows direct use of generated steam for brine evaporation in subsequent effects. This simplifies the system, reduces heat losses, and increases water production capacity.

• Expansion-Electrical Generation Subsystem

An expander-generator subsystem uses stage 1 flash steam to produce electrical power, with the expanded steam reintroduced into the MED effects. This dual-use of steam optimizes energy integration and supports the system’s trigeneration capabilities.

• Efficient Cooling Mechanism

The final product water, in a saturated state at 4 °C, is used as a cooling fluid for the evaporator, minimizing the need for external cooling resources and increasing system efficiency.

• Control of Steam Generation and Brine Concentration

The system incorporates control measures to limit steam generation and prevent calcium sulfate (CaSO₄) precipitation, ensuring operational stability and maintaining long-term system reliability.

3. Methodology

The parametric study of the trigeneration system proposed in this work focused on analyzing the integration of components such as heat exchangers, vapor-liquid separators, expansion valves, and an expander. An extensive literature review was conducted to identify the most suitable simulation software for evaluating the system’s operating behavior and design conditions, thereby determining the production limits for water, cooling, and power.

3.1. Software Simulation Process

The technical feasibility of the proposed system was evaluated using Aspen Plus simulation software, which enables the development, operational analysis, and economic assessment of chemical and petrochemical processes. The software includes extensive databases of thermochemical properties for substances and compounds, facilitating the calculation of the thermodynamic states of the study streams. Its modular architecture allows seamless interaction between specialized blocks, such as heat exchangers and separators, making it ideal for representing and simulating the thermodynamic processes and phenomena relevant to the operational study of the subsystems in this work. Figure 4 illustrates the general methodology used for the simulation process in Aspen Plus.

The simulation process for the proposed system begins with the calculation of thermodynamic properties using the selected methods: IAPWS-95 for water and Electrolyte for streams and modules interacting with seawater. These calculations are applied to the process streams and key modules (HEAT X, PUMP, FLASH2, HEATER, FSPLIT, MIXER, VALVE, COMPR) to ensure precise and reliable results. Subsequently, the system configuration is defined, and the simulation is carried out to model the interactions and behavior of the trigeneration system under specific operating conditions.

To ensure numerical stability and robust outputs, convergence blocks with iterative variables are employed throughout the simulation process. The outcomes of these variables are recorded and thoroughly analyzed to assess system performance and validate its operation against the predefined design criteria.

Table 2. Components, blocks in aspen plus and general equation of the proposed system.

Component	Block in Aspen Plus	Symbology	General Equation
Generator (Gen) Evaporator (EVA) Exchanger 1 (HXBr) Exchanger 2 (HXSW)	HEAT X		${\dot{\mathrm{m}}}_{31}{\mathrm{h}}_{31}+{\dot{\mathrm{m}}}_{88}{\mathrm{h}}_{88}={\dot{\mathrm{m}}}_{32}{\mathrm{h}}_{32}+{\dot{\mathrm{m}}}_{89}{\mathrm{h}}_{89}$ ${\dot{\mathrm{m}}}_{85\mathrm{a}}{\mathrm{h}}_{85\mathrm{a}}+{\dot{\mathrm{m}}}_{86}{\mathrm{h}}_{86}={\dot{\mathrm{m}}}_1{\mathrm{h}}_1+{\dot{\mathrm{m}}}_{87}{\mathrm{h}}_{87}$ ${\dot{\mathrm{m}}}_3{\mathrm{h}}_3+{\dot{\mathrm{m}}}_8{\mathrm{h}}_8={\dot{\mathrm{m}}}_4{\mathrm{h}}_4+{\dot{\mathrm{m}}}_9{\mathrm{h}}_9$ ${\dot{\mathrm{m}}}_{13}{\mathrm{h}}_{13}+{\dot{\mathrm{m}}}_{19}{\mathrm{h}}_{19}={\dot{\mathrm{m}}}_{14}{\mathrm{h}}_{14}+{\dot{\mathrm{m}}}_{20}{\mathrm{h}}_{20}$
Pump 1 (PPW1) Pump 2 (PPW2) Pump 3 (PSW1) Pump 4 (PSW2)	PUMP		${\dot{\mathrm{W}}}_1={\dot{\mathrm{m}}}_7{\mathrm{v}}_7({\mathrm{P}}_8 -{\mathrm{P}}_7)/\mathrm{Eff}$ ${\dot{\mathrm{W}}}_2={\dot{\mathrm{m}}}_{12}{\mathrm{v}}_{12}({\mathrm{P}}_{13} -{\mathrm{P}}_{12})/\mathrm{Eff}$ ${\dot{\mathrm{W}}}_3={\dot{\mathrm{m}}}_{18\mathrm{a}}{\mathrm{v}}_{18\mathrm{a}}({\mathrm{P}}_{19} -{\mathrm{P}}_{18\mathrm{a}})/\mathrm{Eff}$ ${\dot{\mathrm{W}}}_4={\dot{\mathrm{m}}}_{18\mathrm{b}}{\mathrm{v}}_{18\mathrm{b}}({\mathrm{P}}_{22} -{\mathrm{P}}_{18\mathrm{b}})/\mathrm{Eff}$
Stage n	FLASH2		${\dot{\mathrm{m}}}_{\mathrm{in}}{\mathrm{h}}_{\mathrm{in}}+{\dot{\mathrm{Q}}}_{\mathrm{in}}={\dot{\mathrm{m}}}_{\mathrm{vap}}{\mathrm{h}}_{\mathrm{vap}}+{\dot{\mathrm{m}}}_{\mathrm{liq}}{\mathrm{h}}_{\mathrm{liq}}$
Effect n	FLASH2		${\dot{\mathrm{m}}}_{\mathrm{in}}{\mathrm{h}}_{\mathrm{in}}+{\dot{\mathrm{Q}}}_{\mathrm{in}}={\dot{\mathrm{m}}}_{\mathrm{vap}}{\mathrm{h}}_{\mathrm{vap}}+{\dot{\mathrm{m}}}_{\mathrm{liq}}{\mathrm{h}}_{\mathrm{liq}}$
Condenser 1 (PH1) Condenser 2 (PH2) Condenser 8 (PH8)	HEATER		${\dot{\mathrm{m}}}_{29}{\mathrm{h}}_{29}+{\dot{\mathrm{m}}}_{37}{\mathrm{h}}_{37}={\dot{\mathrm{m}}}_{30}{\mathrm{h}}_{30}+{\dot{\mathrm{m}}}_{38}{\mathrm{h}}_{38}$ ${\dot{\mathrm{m}}}_{28}{\mathrm{h}}_{28}+{\dot{\mathrm{m}}}_{47}{\mathrm{h}}_{47}={\dot{\mathrm{m}}}_{29}{\mathrm{h}}_{29}+{\dot{\mathrm{m}}}_{48}{\mathrm{h}}_{48}$ ${\dot{\mathrm{m}}}_{22}{\mathrm{h}}_{22}+{\dot{\mathrm{m}}}_{51}{\mathrm{h}}_{51}={\dot{\mathrm{m}}}_{23}{\mathrm{h}}_{23}+{\dot{\mathrm{m}}}_{52}{\mathrm{h}}_{52}$
Mixer n	MIXER		$\sum {\dot{\mathrm{m}}}_{\mathrm{in}}{\mathrm{h}}_{\mathrm{in}}={\dot{\mathrm{m}}}_{\mathrm{out}}{\mathrm{h}}_{\mathrm{out}}$
Expansion Valve n	VALVE		${\dot{\mathrm{m}}}_{\mathrm{in}}={\dot{\mathrm{m}}}_{\mathrm{out}}$ ${\mathrm{h}}_{\mathrm{in}}={\mathrm{h}}_{\mathrm{out}}$
Expander	COMPR		${\dot{\mathrm{W}}}_{\exp }={\dot{\mathrm{m}}}_{\mathrm{in}}{\mathrm{v}}_{\mathrm{out}}({\mathrm{P}}_{\mathrm{in}} -{\mathrm{P}}_{\mathrm{out}})/\mathrm{Eff}$
Spliter 1 (ST1) Spliter 2 (ST2)	FSPLIT		${\dot{\mathrm{m}}}_5{\mathrm{h}}_5={\dot{\mathrm{m}}}_{5\mathrm{a}}{\mathrm{h}}_{5\mathrm{a}}+{\dot{\mathrm{m}}}_{5\mathrm{b}}{\mathrm{h}}_{5\mathrm{b}}$ ${\dot{\mathrm{m}}}_1{\mathrm{h}}_1={\dot{\mathrm{m}}}_{1\mathrm{a}}{\mathrm{h}}_{1\mathrm{a}}+{\dot{\mathrm{m}}}_{1\mathrm{b}}{\mathrm{h}}_{1\mathrm{b}}$

presents the names of the system components, the corresponding Aspen Plus blocks, the symbology, and the general equations applied in the system’s construction.

Table 3. Processes of the proposed system in the Aspen Plus simulation environment.

Process	Blocks in Aspen Plus
STAGE 1 AND ACTIVATION SOURCE
STAGE 2–8
EFFECT 1–5
EFFECT 6
SUBSYSTEM C-ECB

details the interconnection configuration of the blocks used to represent the subsystems of the proposed system in this work, as well as their organization within the main flow diagram in Aspen Plus. The hybrid MSF/MED thermal desalination and cooling subsystems, along with the Expander-generator, are represented in the processes STAGE 1 and ACTIVATION SOURCE, STAGES 2 to 8, EFFECTS 1 to 5, and EFFECTS 6. Meanwhile, the subsystems for extraction and product water adaptation, as well as cooling of the motive fluid and brine dilution, are represented in the SUBSYSTEM C-ECB (Cooling-Ejector Condenser Barometric) process.

3.2. System Parameters

In the operational study of the proposed system presented in this paper, the parameters outlined in

Table 4. Operating parameters of the MSF/MED thermal desalination subsystem at high vacuum.

Parameters	Units	Value
No. of Flash stages	-	8
No. of MED effects	-	6
No. of preheaters	-	8
Top brine temperature (TBT)	°C	68
Bottom brine temperature (BBT)	°C	4
ΔT stages/effects	°C	4.92
Feed seawater temperature	°C	25
Salt concentration of feed seawater	ppm	35,000

were defined, varying the seawater feed flow and the activation heat supplied to the generator. The seawater feed flow enters the system at 25 °C with a concentration of 35,000 ppm of NaCl, limiting the Top Brine Temperature (TBT) to 68 °C to mitigate corrosion issues, as reported by Palenzuela et al. [32].

3.3. Basic Assumptions

Simulation in Aspen Plus, especially when dealing with recirculation streams, requires the initialization and definition of iterative variables, as described in the methodology diagram. These variables iteratively solve the simulation process to achieve convergence. The operational simulation study of the proposed system considers the following basic assumptions, applicable to both trigeneration and cogeneration modes:

• The technical feasibility and the behavior of the proposed MSF/MED system to the high vacuum in steady state is analyzed.
• The hybrid MSF/MED to high vacuum thermal desalination subsystem uses a TBT of 68 °C and a BBT of 4 °C.
• The heat exchange between streams considers a minimum approach temperature (∆T_min) of 2.5 °C. for stages and effects.
• A maximum seawater preheating temperature of T₃₀ = 65.5 °C is established.
• An isentropic efficiency of 85% of the hydraulic pumping equipment (P_SW1, P_SW2, P_WP1, P_WP2) is considered.
• An isentropic efficiency of 85% is considered in the expander (Expander).
• Expander discharge pressure is 5.129 kPa.
• All the condensed product water is used as refrigerant fluid, in trigeneration operation.
• The pressure drop in the equipment and piping is negligible.
• An exergy efficiency of 85% is considered in the ejectocondenser (EJ1).
• An exergy efficiency of 28% is considered for the ejectocondenser (EJ2).
• The properties of the equilibrium state: temperature of 25 °C and absolute pressure of 1 atm.

3.4. System Energy Performance Indicators

The selected performance indicators evaluate in single, cogenerative and trigeneration production the system’s products. The equations are defined based on the energy balance by the first law of thermodynamics.

The following indicators were used in the desalination performance evaluation: Recovery Ratio (RR), Performance Ratio (PR) and global Performance Ratio (PR_global). The RR indicator (Equation (1)) represents the fraction of seawater that is converted to desalinated water. To evaluate the thermal supply of the system, the PR indicator (Equation (2)) was used, which describes the ratio between the latent heat of the product water and the thermal energy supplied. In addition, another indicator, the PR_global (Equation (3)), is proposed, which considers the thermal and electrical energy requirement supplied to the system (La Torre et al. [12]).

$$\mathrm{RR}={\displaystyle \frac{{(\dot{\mathrm{m}}}_{11}+{\dot{\mathrm{m}}}_{16})}{{\dot{\mathrm{m}}}_{22}}}\times 100$$

(1)

$$\mathrm{PR}={\displaystyle \frac{{(\dot{\mathrm{m}}}_{11}+{\dot{\mathrm{m}}}_{16}) \times (2326\frac{\mathrm{kj}}{\mathrm{kg}})}{{\dot{\mathrm{Q}}}_{\mathrm{GEN}}}}$$

(2)

$$\mathrm{P}{\mathrm{R}}_{\mathrm{global} }={\displaystyle \frac{{(\dot{\mathrm{m}}}_{11}+{\dot{\mathrm{m}}}_{16}) \times (2326\frac{\mathrm{kj}}{\mathrm{kg}})}{{\dot{\mathrm{Q}}}_{\mathrm{gen}}+\sum {\dot{\mathrm{W}}}_{\mathrm{Tot}. \mathrm{power}} -{\dot{\mathrm{W}}}_{\exp }}}$$

(3)

The cooling production performance of the system was analyzed using the Coefficient of performance (COP) indicator (Equation (4)), which relates the cooling capacity of the system to the energy supply (thermal and/or electrical) required. The COP_global equation (Equation (5)) proposed by López-Zavala et al. [31] was used to evaluate the performance of simultaneous desalination and cooling production.

$$\mathrm{COP}={\displaystyle \frac{{\dot{\mathrm{Q}}}_{\mathrm{EVA}}}{{\dot{\mathrm{Q}}}_{\mathrm{GEN}}+\sum {\dot{\mathrm{W}}}_{\mathrm{el}\text{é}\mathrm{c},\mathrm{tot}}-{\dot{\mathrm{W}}}_{\exp }}}$$

(4)

$${\mathrm{COP}}_{\mathrm{Global}}={\displaystyle \frac{{\dot{\mathrm{Q}}}_{\mathrm{EVA}}+{(\dot{\mathrm{m}}}_{11}+{\dot{\mathrm{m}}}_{16}) \times (2326\frac{\mathrm{kj}}{\mathrm{kg}})}{{\dot{\mathrm{Q}}}_{\mathrm{GEN}}+\sum {\dot{\mathrm{W}}}_{\mathrm{Tot}. \mathrm{power}}-{ \dot{\mathrm{W}}}_{\exp }}}$$

(5)

The following indicators were used to study energy consumption: Electrical Coverage (EC), Specific Electrical Energy Consumption (SEEC), Specific Thermal Energy Consumption (STEC) and Specific Energy Consumption (SEC). Electrical Coverage describes the fraction of the system’s electrical demand that is satisfied by the electrical generation of the expander (Equation (6)).

$$\mathrm{EC}={\displaystyle \frac{{\dot{\mathrm{W}}}_{ \exp .}}{\sum {\dot{\mathrm{W}}}_{\mathrm{el}\text{é}\mathrm{c},\mathrm{tot}}}}$$

(6)

SEEC and STEC are defined as the ratio between the amount of energy consumed (electrical or thermal) and the water capacity produced (Equation (7)). The total specific energy consumption (SEC) is defined as the sum of the specific consumption of electrical and thermal energy (Equation (8)).

$$\mathrm{SEEC} \mathrm{or} \mathrm{STEC}={\displaystyle \frac{\mathrm{Total} \mathrm{electrical} \mathrm{and} \mathrm{thermal} \mathrm{energy} \mathrm{consumption} \left(\mathrm{kW}\right)}{\mathrm{Water} \mathrm{product} \mathrm{capacity} ({\mathrm{m}}^3/\mathrm{h})}}$$

(7)

$$\mathrm{SEC}=\mathrm{SEEC}+\mathrm{STEC}$$

(8)

The operating performance of the ejectocondenser was evaluated with respect to the second law of thermodynamics. The exergy content of the currents entering the ejector was quantified using (Equation (9)). On the other hand, for the thermodynamic feasibility study, (Equation (10)) was used, which determines the exergetic destruction, likewise, for the calculation of the exergetic efficiency of the ejectocondenser, (Equation (11)) was used.

$$\dot{\mathrm{Ex}}=\dot{\mathrm{m}}\left[\mathrm{h} -{\mathrm{h}}_0-{\mathrm{T}}_0\left(\mathrm{s} -{ \mathrm{s}}_0\right)\right]$$

(9)

$$\sum \dot{\mathrm{Q}}(1-{\displaystyle \frac{{\mathrm{T}}_0}{\mathrm{T}}})+\sum {\dot{\mathrm{Ex}}}_{\mathrm{in}}-\sum {\dot{\mathrm{Ex}}}_{\mathrm{out}}-\sum \dot{\mathrm{W}}-{ \dot{\mathrm{Ex}}}_{\mathrm{Destroy}}=0$$

(10)

$${\mathsf{\eta }}_{\mathsf{\psi }}={\displaystyle \frac{{\mathrm{Exergy}}_{\mathrm{out}}}{{\mathrm{Exergy}}_{\mathrm{in}}}}$$

(11)

4. Validation of Simulation

The validation of the proposed system was performed in two parts: (1) the dual barometric ejectocondensation subprocess of the product water extraction and adaptation subsystem, and (2) thermal desalination subprocess concerning the MSF/MED subsystem to high vacuum. For the first part, the experimental work reported by Sağ et al. [33] which corresponds to the study of exergy destruction of the components of an ejecto-compression cooling system was used.

The proposed system, like the work reported by Sağ et al. [33], uses a liquid-vapor two-phase ejector in an ejectocompression cooling system. However, unlike the system proposed in this work, the author considers partial condensation at the ejector outlet as a design condition. Therefore, a comparison was made between the operating results of the experimental system and those obtained by simulation in Aspen Plus, with the objective of analyzing the operating performance and studying the partial condensation process in the two-phase ejector simulated in Aspen Plus.

As shown in

Table 5. Comparison of operational results of the experimental study reported by (Sağ et al. [33]) and this work.

Parameters	Units	Sağ et al.	This Work	Error (%)
Condenser outlet temperature	°C	57.64	59.02	2.394
Evaporator outlet temperature	°C	11.63	11.03	5.159
Ejector outlet temperature	°C	13.12	12.88	1.829
Ejector outlet pressure	kPa	456.00	456.00	0.000
Entrainment ratio	-	0.65	0.647	0.767
COP	-	2.12	2.115	0.236
Compressor duty	kW	2.11	2.10	0.332
Cooling capacity	kW	4.47	4.45	0.470
Average duty				1.398

, the average error of the results is less than 1.5%. When evaluating the operating performance of the system using the first law of thermodynamics, it is observed that the cooling capacity, compressor work and COP do not present errors greater than 0.5%. In addition, the operating performance of the two-phase ejector shows a maximum error of 1.829%, which demonstrates the feasibility of performing the ejector study in Aspen Plus.

The experimental study conducted by Sağ et al. [33] on the exergy destruction of each component of the ejectocompression cooling system was used to compare the experimental results with those generated by Aspen Plus.

Table 6. Comparison of exergy destruction between experimental results from (Sağ et al. [33]) and simulation in Aspen Plus.

Component	Units	Sağ et al.	This Work	Error (%)
Condenser	W	71.88	78.64	9.4
Compressor	W	974.3	917.5	5.82
Evaporator	W	87.03	84.78	2.58
Expansion Valve	W	0.46	0.5	8.76
Ejector	W	188.2	188.71	0.27
Separator	W	35.76	35.61	0.42
Exergy destroyed total:	W	1357.63	1305.74	3.82
Average error:				4.44

shows the exergy destruction of each component of the ejectocompression cooling system. The results show a total average error of 4.44%, highlighting the condenser as the component with the highest percentage of error, reaching 9.4%. In contrast, the ejector, which is the component of greatest interest, presents an error of 0.27%.

In the second part of the validation, the work presented by López-Zavala et al. [31] was used. The desalination and cooling system with barometric ejectocondenser that they describe shares similarities in the operating principles with the proposed hybrid MSF/MED to HV system. Detailed information on the key parameters and considerations for the simulation of this system is presented in

Table 7. Main considerations and value ranges of the system parameters proposed by (López-Zavala et al. [31]) and the simulation in Aspen Plus.

Parameters	Units	López-Zavala et al.	This Work
No. Effects	-	14	14
TBT	°C	70	70
BBT	°C	37	37
ΔP effects	kPa	3.13	3.13
ΔT effects	°C	2.5	2.5
Operating temperatura range	°C	70–35	70–35
Salt concentration of feed Seawater	ppm	35,000	35,000
Seawater feed water temperature	°C	27	27
Hot source Flow rate	kg/s	1.611	1.611
Hot source temperature	°C	95	95

.

Table 8. Simulation validation between the theoretical results of López-Zavala et al. [31] and the simulation in Aspen Plus.

Parameters	Units	López-Zavala et al.	This Work	Error (%)
Seawater temperature at condenser outlet	°C	32	31.96	0.125
Seawater temperature at the PH@12 outlet	°C	64	64.1	0.156
Product water capacity	kg/s	0.865	0.859	0.694
PR	-	11.6	11.53	0.603
RR	-	0.53	0.5274	0.491
COP	-	4.7	4.67	0.638
COPglobal	-	9.325	9.291	0.365
Cooling capacity	kW	2042.87	2028.59	0.699
Heat exchanged (condenser)	kW	152.82	151.76	0.694
Entrainment ratio	-	0.0025	0.00243	2.800
Heat exchanged (heatsink)	kW	2297.54	2310.49	0.564
Average error:				0.47

shows the comparison of both simulations with respect to the desalination and cooling performance indicators, as well as the internal energy integration and heat transfer in the barometric condensation ejection subsystem. A maximum error of 0.638% is observed in the desalination and cooling performance indicators, a maximum error of 0.156% in the internal energy integration (preheating) and a maximum error of 2.8% in the barometric ejectocondensation subsystem section.

5. Analysis and Discussion of Results

The parametric study presented in this work on the proposed trigeneration system is divided into two parts: the analysis of the system’s behavior in cogeneration mode and its behavior in trigeneration mode.

5.1. Operational Study of the System in Cogeneration Mode

To operate in cogeneration mode, cooling production is eliminated in the MSF/MED hybrid thermal desalination subsystem under high vacuum conditions. This is achieved through the control of the 3-way valve (3WV), which diverts the condensed product water stream (88), preventing its entry into the evaporator (EVA) and thereby stopping the generation of the cooling effect in the system. The bypassed condensed product water stream (88b) is directed to the barometric column (BC2). The function of this column is to maintain the hydraulic vacuum seal of the system and increase the pressure to atmospheric levels, preparing the water for storage in the product water tank.

5.2. Operational Performance Results of the MSF/MED Thermal Desalination Subsystem Under High Vacuum in Cogeneration/Trigeneration Modes

The amount of heat transferred in the stages of the MSF section and the effects of the MED section, along with the heat gained by the seawater to reach the maximum preheating temperature of 65.5 °C, is presented in Figure 5, in relation to the variation in activation heat.

In the MSF section, heat transfer occurs only in the first stage due to the presence of the generator, which produces the steam necessary to reach the preheating temperature. The excess steam generated when operating above Q_gen = 50 kW is used for the expander to produce electrical energy. As activation heat increases, the heat transferred for seawater preheating decreases, since the required heat is supplied by the mixture of flashed vapors from the product water and brine within the MSF. More steam generated in the generator reduces the brine flow to the second stage, decreasing the amount of flashed steam and, consequently, the heat transfer for preheating. This results in an increased steam flow to the condenser (Cond₁) to achieve the maximum preheat temperature.

After the MSF section, the flashed vapors are no longer used to preheat the seawater due to the lack of a minimum temperature (ΔT_min = 2.5 °C) difference Instead, they are used to boil off the brine in the first effect of the MED section, gradually increasing heat transfer. Additionally, the reinsertion of steam expanded by the expander into the eighth stage of the MSF increases the mass of available steam, enhancing heat transfer in the MED section and creating a cascading effect throughout the other effects in the system.

The vapors generated by boiling and flashing in each stage or effect of the thermal desalination subsystem are shown in Figure 6. In the first stage of the MSF section, vapor is primarily produced through boiling to supply the condenser (Cond₁) and reach the maximum preheating temperature (65.5 °C). Subsequently, the desalination capacity decreases due to the flashing of water vapor caused by the pressure differential between stages.

By increasing the activation heat, more steam is generated in the generator. The excess steam produced, which is not required to feed the Condenser (Cond₁) in stage 1, is used as a working fluid for the expander and is reintroduced in the eighth stage. The MED section is activated by utilizing the expanded steam from the expander, along with the vapors generated by the flashing of product water and the brine flow from the eighth stage. This creates a cascading effect, where the vapors generated by flashing are no longer used for preheating, but instead for boiling, allowing more water to be extracted from the brine stream as it passes through the effects. The increase in heat added to the generator leads to an increase in product water production capacity and, consequently, to a higher cooling capacity of the system in trigeneration operation.

Figure 7 shows the production capacity of condensed water, water vapor directed to the expander and to Cond1, brine flow, and water vapor produced by flashing and boiling in effect 6. The internal variables in Figure 7 refer to the vapors generated in the first stage of the desalination subsystem, specifically the vapors from the condenser (Cond₁) and the expander, as well as the brine flow at the outlet of the thermal desalination subsystem. The vapors (Cond₁ and Expander) tend to increase as more heat is added to the generator. This occurs because the water vapor directed to Cond1 supplies the remaining heat required to reach the maximum preheating temperature of the seawater, which had previously been preheated by the vapors generated through flashing in stages 2 to 8, which released their condensation heat for this purpose. As a result, a higher release of vapor from the brine leads to a reduced brine flow directed to the MSF section, causing an increase in condenser (Cond₁). On the other hand, the steam flow to the expander increases because the water vapor originating from the generator, which is not used by condenser (Cond1), is directed to the expander.

The water vapor and condensed product water from effect 6 tend to increase, as shown in Figure 7. This occurs because the expanded steam at the expander outlet mixes with the steam from the eighth stage, and this mixture is used to boil the brine flow in the first heat exchanger of the MED section. This process gradually increases the desalination capacity as the steam flow from the expander rises, leading to enhanced heat exchange. Finally, the water vapor extracted from the brine in cogeneration or trigeneration operation modes tends to decrease as heat is added to the generator. This is due to the reduction in heat dissipated in the brine exchanger (HX_Brine), as the heat transfer depends on the brine flow (stream 3) and the amount of steam to be condensed in the ejectocondenser (EJ1).

5.3. Operational Performance Results of Subsystems: Product Water Extraction, Adequacy, and Cooling of the Motive Fluid in Cogeneration Mode

Figure 8 shows the brine concentrations at the outlet of the thermal desalination subsystem and the heat exchanger (HX_Brine). The brine concentration (X_ppm@effect6) in the brine stream (3) at the MED section outlet increases due to the intensified boiling of water vapor in the generator and the higher steam flow to the expander, used in effect 1 (HX₁). Similarly, the brine concentration (X_ppm@HXBrine) also rises. This is attributed to the increased brine concentration at the MED section outlet and the dissipation of condensation heat from the motive fluid by the saturated brine stream (3). This process causes water boiling and a subsequent rise in concentration within the heat exchanger (HX_Brine), with an average increase of 32.1% compared to the concentration at the outlet of effect 6, reaching a maximum of 225,087 ppm when operating at Q_gen = 350 kW. Furthermore, heat transfer in the heat exchanger (HX_Brine) and secondary flow (Fs_EJ1) decrease due to the reduced brine flow, diminishing the capacity to remove condensation heat. This lowers the condensing capacity of the ejectocondenser (EJ1) from 0.1758 to 0.0869 kg/s, redirecting the remaining cold vapor to the second ejectocondenser (EJ2).

Cold vapors from the brine separator (S_Br15) and effect 6 are used as secondary flows, while product water from a reservoir tank serves as the motive flow in the ejectocondenser. Figure 9 shows a decrease in the motive flow (Mf_PW1) from 155 to 56.56 kg/s in ejectocondenser (EJ1), and an increase in the motive flow (Mf_PW2) from 75 to 125 kg/s in (EJ2), along with an increase in the seawater flow (F_Sea) from 50 to 91.58 kg/s. This is due to the reduction in secondary flow (Sf_EJ1), driven by a decrease in brine flow, which reduces the demand for motive flow in (EJ1) and directs more steam to be condensed in (EJ2). The increase in secondary flow (Sf_EJ2) requires a higher motive flow, increasing the electrical demand of the hydraulic equipment (P_PW2) by 10.77%, while the reduction in (Sf_EJ1) decreases the demand on (P_PW1) by 15.33%. Heat exchange follows the trends of the motive flow: the heat exchanger (HX_Brine) shows a reduction from 526.61 to 248.12 kW, while the heat exchanger (HX_Sea) increases from 277.57 to 509.47 kW.

5.4. Overall Operating Performance Results in Cogenerative Mode

As shown in Figure 10, increasing the activation heat boosts the amount of steam produced in the generator, which is used as the working fluid by the expander-generator. This leads to a rise in the expander’s mechanical work, increasing power generation from 4.19 to 25.24 kW_power, thereby improving the system’s electrical self-consumption coverage from 5.94% to 35.16%. The reinsertion of expanded steam into the eighth stage of the MSF section in the MSF/MED hybrid thermal desalination subsystem, operating under high vacuum, enhances steam boiling, which improves both desalination capacity and steam condensation in the product water extraction and treatment subsystem, significantly raising the system’s electrical power demand.

Additionally, as depicted in Figure 10, the SEC indicator increases with the heat supplied to the generator, starting at 60.71 kWh/m³ and reaching 90.86 kWh/m³. This positions the proposed system above conventional MED and MSF desalination technologies in terms of SEC, as noted by Al-Karaghouli and Kazmerski [34].

As shown in Figure 11, the behavior of the PR and PR_global indicators reveals a decrease in the system’s desalination performance. This effect is observed because the system maintains constant product water production through the operation of the high-vacuum hybrid MSF/MED desalination subsystem and the motive fluid cooling and brine dilution subsystem, allowing the concentration process to reach a product water output of 1.212 kg/s. However, the increase in activation heat and the power allocated to the expander-generator’s self-consumption do not enhance desalination performance, leading to a decline in efficiency as more heat is added to the system.

According to Figure 7 an increase in activation heat correlates with a rise in water production capacity, escalating from 0.548 to 1.212 kg/s due to heightened seawater boiling (31). This phenomenon results in an increase in product water generation and RR from 38.2% to 84.4%.

As discussed in Figure 11 this increase also contributes to higher electrical demand. However, the power input from the expander has a notable impact on desalination performance. The PR indicator diminishes from 25.51 to 8.05, representing a reduction of 68.4%, while the thermal power output (Q_gen) increases to 350 kW. The PR_global decreases from 10.64 to 6.68, indicating a reduction of 37.21%. This decrease is less pronounced than that of the PR, attributed to the electrical self-consumption of the expander.

The proposed system attains a PR of 25.51, which is 1.83 times higher than the value reported by Aguilar-Jiménez et al. [20] for a MED/ORC cogenerative system. Additionally, the recovery rate of 38.2% to 84.4% is comparable to that observed in cooling desalination systems, which exhibit the highest recovery rates among phase-change desalination technologies [35,36].

5.5. Operational Study of the System in Trigenerative Mode

To operate the proposed system in trigeneration mode, the condensed product water stream (84) exits the sixth effect in a saturated state at 4 °C and 0.81 kPa. This stream is directed to a three-way valve (3WV) that regulates its flow based on cooling demand. When cooling is required, the valve redirects stream (88a) to the evaporator (EVA), where its low temperature efficiently absorbs heat to produce the cooling effect. If cooling is not needed, the valve diverts stream (88b) to a barometric column (BC2), where it undergoes a controlled pressure increase before being discharged into a product water tank. Additionally, low-temperature vapor streams (1) and (2) from earlier stages are processed in the ejectocondenser (EJ1), where they mix with a high-pressure driving fluid. This process condenses the vapor, releasing latent heat and maintaining vacuum conditions, while the condensed fluid is directed to BC2. By combining chilled product water and vapor condensation via the ejector, the system achieves efficient and flexible cooling production, optimizing its performance in trigeneration mode.

This trigeneration operation, involving the simultaneous production of product water, cooling, and electrical energy, has a direct impact on all subsystems, particularly in the cooling of the motive fluid and brine dilution. Power generation requires the production of additional steam in the generator, which is used as the working fluid to feed the expander. Subsequently, the expanded steam at the expander outlet is reinserted into the vapor-liquid separator (S_PW9), increasing product water production and cooling capacity, as a greater mass flow of steam is available for boiling the concentrated brine stream (66) entering the heat exchanger (HX₂). This leads to a higher concentration of brine at the outlet of the sixth effect. These factors result in a reduction of the maximum limit of the reinjected steam mass flow percentage and a decrease in the final concentration of the brine entering the brine dilution tank.

5.6. Operational Performance Results of Subsystems: Product Water Extraction, Adequacy, and Cooling of the Motive Fluid in Trigeneration Mode

The brine concentrations, measured in parts per million (ppm), at the outlet of the thermal desalination subsystem and the concentrated brine stream at the outlet of the heat exchanger (HX_Brine) are presented in Figure 12. The brine concentration (X_ppm@effect6) represents the concentration in stream (3). The increase in this brine concentration is caused by two factors: the increased boiling of water vapor in the generator and the increased steam flow directed to the expander, which is then used for brine vapor boiling in the heat exchanger of effect 1 (HX₁). Similarly, the brine concentration (X_ppm@HXBrine) also rises, primarily due to the increase in brine concentration at the outlet of the MED section.

Additionally, the dissipation of condensation heat from the motive fluid by the saturated brine stream (stream 3) causes this stream to enter the saturation region, releasing water vapor and resulting in a higher concentration. The brine concentration at the outlet of the heat exchanger (HX_Brine) shows an average increase of 203.01% compared to the concentration at the outlet of effect 6, exceeding the limiting concentration of 350.00 ppm when operating at QQ_gen = 350 kW, with a reported concentration of 386.959 ppm.

Furthermore, heat transfer in the heat exchanger (HX_Brine) and the secondary flow (Fs_EJ1) tend to decrease due to the reduction in brine flow, resulting in a diminished capacity to remove condensation heat. This reduces the condensing capacity of the ejectocondenser (EJ1) from 0.1758 to 0.0869 kg/s, directing the remaining cold vapor to the second ejectocondenser (EJ2).

The cold vapors from the brine separator (S_Br15), evaporator, and effect 6 are used as secondary flows, while the product water, coming from a reservoir tank, is used as the motive flow in the corresponding ejectocondenser. In Figure 9, a decrease in the motive flow (Mf_PW1) of the ejectocondenser (EJ1) from 145.9 to 52.794 kg/s is observed, alongside an increase in the motive flow (Mf_PW2) of ejectocondenser (EJ1) from 272.8 to 817.05 kg/s, and an increase in the seawater flow (F_Sea) from 176.8 to 529.4 kg/s. This occurs because the motive flow is closely linked to the secondary flow; a decrease in the secondary flow indicates a lower amount of motive flow needed to ensure condensation.

The reduction in secondary flow (Sf_EJ1), as shown in Figure 13, is caused by the decrease in brine flow, which limits the amount of heat transferred in the heat exchanger (HX_Brine). With an increase in activation heat, the steam flow to be condensed in the ejectocondenser (EJ1) is reduced, transferring the remaining steam to the ejectocondenser (EJ2) for condensation.

The increase in secondary flow (Sf_EJ2) requires a higher motive flow to achieve condensation, leading to an average increase of 17.86% in the electrical demand of the hydraulic equipment (P_PW2). Conversely, the reduction in secondary flow (Sf_EJ1) results in an average decrease of 24.9% in the electrical demand of the hydraulic equipment (P_PW1). Additionally, heat exchange follows the trends of both secondary and motive flows. As mentioned, condensing a larger amount of steam requires a higher motive flow, resulting in increased heat transfer in the heat exchanger to remove the condensation heat generated in the ejection-condensation process. The heat transfer in the heat exchanger (HX_Brine) decreases from 1974.25 kW to 459.35 kW, while the heat transfer in the heat exchanger (HX_Sea) increases from 979.02 kW to 2954 kW.

5.7. Overall Operating Performance Results in Trigenerational Mode

The power demand of the proposed system is primarily driven by the hydraulic pumping equipment in the product water extraction and adaptation subsystem. In trigeneration mode, the electrical demand increases due to the higher flow of motive fluid required for the condensation of the product water stream (88a) evaporated in the ejectocondensers (EJ1 and EJ2).

The expander increases its electrical generation capacity as more heat is supplied, utilizing the additional steam as the working fluid. The reintroduction of the expanded vapor into the vapor-liquid separator (S_PW9) enhances both water production and cooling, which subsequently raises the system’s electrical demand. This behavior is shown in Figure 14, where both electrical demand and electrical coverage tend to increase.

This is because the electrical generation provided by the expander is less than the increase in the electrical demand of the system’s hydraulic equipment, which is caused by the impact of the expanded steam on the MSF/MED thermal desalination subsystem operating at high vacuum, as well as the increase in cold steam mass flow from effect 6 and the evaporator. The system’s electrical demand increases to a maximum of 351.14 kW, but the power generated by the expander reduces the system’s overall electrical consumption by an average of 5.32%. Additionally, the self-consumption electrical coverage varies between 2.11% and 7.18%.

Although the product water capacity increases from 1128 to 1304 m³/day, this increase is not significant. The additional steam production in the generator leads to a rise in both water production and cooling capacity, increasing the power demand, which the expander fails to fully compensate. As a result, SEC indicator increases from 53.16 to 143.95 kWh/m³, reflecting higher energy consumption relative to the activation heat.

The system’s cooling capacity also increases with the additional heat supplied to the generator, rising from 665.07 to 2320.3 kW, as shown in Figure 15. This is due to the reinsertion of expanded steam into the thermal desalination subsystem, which enhances product water capacity. The liquid product water at 4 °C is used as the cooling fluid. After producing the cooling effect, the product water transitions to the vapor phase and is divided into two streams (ST₂), fractionating the vapor mass flow for the proper operation of the ejectocondensers (EJ1 and EJ2).

The maximum COP achieved is 3.43, which is 4.70 times higher than the COP of 0.73 reported by Gomri [37] for single-effect absorption systems. Additionally, the COP_global indicator proposed by López-Zavala et al. [31] was 9.325, which is 1.643 times lower than the COP achieved by the proposed system, which reached 15.231.

The PR decreases from 52.51 to 8.66 as the heat input to the generator increases, while water production rises from 1.128 to 1.304 kg/s. The maximum PR of 52.51 is 3.76 times higher than that reported by Aguilar-Jiménez et al. [20]. The PR_global, which accounts for both the thermal and electrical energy of the system, drops from 12.15 to 4.32, as the reduction in electrical demand from expander generation does not offset the increased heat input to the generator. The RR reaches a maximum of 86.7%, which is 2.22 times higher than the MED-HDH system reported by Tahir and Al-Ghamdi [38]. This suggests that future studies should investigate the technical feasibility of incorporating a crystallizer in the motive fluid cooling subsystem and brine dilution process.

5.8. Operational Performance Results of the Ejectocondensation Process

A study based on the second law of thermodynamics was conducted on the barometric ejectocondensation process, with the objective of determining the operating parameters of the motive pressure for the hydraulic pumping equipment (P_PW1 and P_PW2). The study assumed constant thermodynamic properties for the secondary fluid, with a pressure of 0.804 kPa, temperature of 4 °C, and mass flow of 0.2167 kg/s. For the motive fluid, an initial mass flow rate of 130 kg/s was considered. However, as mentioned, the hydraulic pumping equipment features a flow variator that adjusts to ensure condensation.

The exergy destruction results, shown in Figure 16, indicate that regardless of the temperature condition of the motive fluid, an increase in motive pressure at each temperature condition leads to an increase in exergy destruction. Additionally, it is observed that higher driving temperatures result in greater exergy destruction, suggesting a decrease in exergy efficiency. This occurs because increasing the driving pressure requires a higher motive flow to ensure complete condensation of the steam inside the ejectocondenser. A higher motive flow in the ejectocondenser leads to greater energy losses due to friction.

5.9. General Comparison

5.9.1. Operating Mode Comparison

The technological system developed in this work exhibits a distinctive operating behavior in terms of water production and electrical demand, depending on the operation mode.

In trigeneration mode, the water production capacity ranges from 1.128 to 1.304 kg/s, while in cogeneration mode it ranges from 0.548 to 1.212 kg/s, reflecting an average increase of 38.18% in trigeneration mode (Figure 17).

In trigeneration mode, the condensed water in the MED section is sent to the evaporator (EVA), where it changes phase to steam. This additional steam is directed to the ejectocondenser (EJ1) for condensation, increasing the motive flow (Fm_PW1) and heat dissipation in the heat exchanger (HX_Brine). The saturated brine (stream 3) is further heated and separated in the liquid-vapor separator (S_Br15). Water vapor extraction in trigeneration mode varies from 0.791 to 0.1915 kg/s, while in cogeneration mode it ranges from 0.211 to 0.099 kg/s. This process reduces the final brine flow (stream 17), enhancing the potential for valuable mineral extraction and achieving zero liquid discharge.

5.9.2. Literature-Based Operating Comparison

To emphasize the main distinctions and advantages of the proposed MSF/MED-HV system, a comparative analysis is conducted against the MED-HV system reported by La Torre et al. [12] and the MED system described by López-Zavala et al. [31]. The comparison evaluates operational performance and design capacities, considering the specific design conditions established by the respective authors and those defined for the MSF/MED-HV system in this study. Key metrics such as desalination capacity, cooling output, and energy efficiency are presented in

Table 9. Results of comparative table of the MSF/MED-HV and MED systems reported in literature.

Parameters	Unitz	MSF/MED-HV (1)	MED-HV La Torre (2)	MED Lopéz-Zavala (3)	(Δ1–2%)	(Δ1–3%)
Desalination capacity	kg/s	1.14	1.11	0.86	2.70	32.55
Cooling	kW	1224.32	-	2042.87	-	40.06
Qgen	kW	150	191.61	173.43	−21.71	−13.50
Wpumps total	kW	231.98	39.3	261.37	490.27	11.24
PR	-	17.81	13.49	11.60	32.02	53.53
RR	-	0.80	0.83	0.53	3.61	50.94
PRglobal	-	6.99	11.20	4.65	37.58	50.32
COP	-	3.27	-	4.7	-	30.42
COPglobal	-	10.43	-	9.32	-	11.84

to highlight the relative performance of each system.

Among the key advantages of the proposed MSF/MED-HV system is its PR of 17.81, surpassing the systems by La Torre et al. [12] and López-Zavala et al. [31] by 32.02% and 53.53%, respectively. The system by La Torre et al. also operates under high-vacuum (HV) conditions (28.3 to 0.8 kPa; 68 to 4 °C), while López-Zavala et al. operates under conventional conditions (30.87 to 5.63 kPa; 70 to 35.37 °C). Despite similar or less demanding operational ranges, the proposed system achieves superior desalination performance, with a RR of 0.8, exceeding López-Zavala et al. RR of 0.53 by 50.94%.

In terms of energy demand, the MSF/MED-HV system requires 150 kW of generator heat input (Q_gen), 21.71% lower than the Q_gen reported by La Torre et al. and 13.5% lower than that of López-Zavala et al., highlighting its improved thermal efficiency. However, the total power requirement for hydraulic pumps (W_{pumps total}) is higher (231.98 kW), primarily due to the dual barometric ejectocondensation process. Despite these demands, the system achieves superior global performance indicators, with a PR_global of 6.99 and a COP_global of 10.43, representing an improvement of 50.32% and 11.84%, respectively, compared to López-Zavala et al. system. These advances highlight the significant thermodynamic and operational efficiency of the proposed design.

After the parametric study, the following system design conditions were determined for both cogeneration and trigeneration modes.

Table 10. Thermodynamic and operational characteristics of the MSF/MED at high vacuum trigenerative system.

No. Stage/Effects	Pressure (kPa)	Temperature (°C)	Accumulated Product Water (kg/s)	Brine Concentration (ppm)	Qcond (kW)	Qeffects (kW)	Seawater Temperature (°C)
1	28.3	68.00	-	36,467	38.65	150	65.5
2	22.73	63.08	0.0164	36,785.2	28.3	-	56.55
3	18.1	58.15	0.0281	37,104.2	28.34	-	52.05
4	14.32	53.23	0.0398	37,423.4	28.27	-	47.53
5	11.23	48.31	0.0512	37,743.6	28.24	-	43.03
6	8.72	43.38	0.0625	38,065.5	28.32	-	38.52
7	6.72	38.46	0.0737	38,387.9	28.24	-	34.01
8	5.13	33.54	0.0847	38,711.4	28.23	-	29.5
1	3.87	28.61	0.1364	40,245	-	94.13	25.0
2	2.89	23.69	0.1854	42,298.2	-	123.15	25.0
3	2.14	18.77	0.2455	45,014.5	-	152.01	25.0
4	1.56	13.85	0.3166	48,617	-	181.17	25.0
5	1.12	8.92	0.3986	53,451.1	-	210.72	25.0
6	0.79	4.00	0.4913	60,092	-	240.09	25.0
HXBrine,cog	0.79	4.00	0.1788	76,443	-	445.61	-
HXBrine,trig	0.79	4.00	0.5498	175,605	-	1371.1	-
Totalcog					236.63	1596.8
Totaltrig					236.63	2422.5

,

Table 11. Energy evaluation of the proposed MSF/MED at high vacuum system under design conditions in cogenerative operation.

Components:	Energy Balance	Gains (kW)	Losses (kW)	Internal Energy Integration (kW)
Generator	${\dot{\mathrm{m}}}_{31}{\mathrm{h}}_{31}+{\dot{\mathrm{m}}}_{88}{\mathrm{h}}_{88}={\dot{\mathrm{m}}}_{32}{\mathrm{h}}_{32}+{\dot{\mathrm{m}}}_{89}{\mathrm{h}}_{89}$	150
HXBrine	${\dot{\mathrm{m}}}_3{\mathrm{h}}_3+{\dot{\mathrm{m}}}_8{\mathrm{h}}_8={\dot{\mathrm{m}}}_4{\mathrm{h}}_4+{\dot{\mathrm{m}}}_9{\mathrm{h}}_9$		445.618
HXSea	${\dot{\mathrm{m}}}_{13}{\mathrm{h}}_{13}+{\dot{\mathrm{m}}}_{19}{\mathrm{h}}_{19}={\dot{\mathrm{m}}}_{14}{\mathrm{h}}_{14}+{\dot{\mathrm{m}}}_{20}{\mathrm{h}}_{20}$		361.69
Condensers	${\dot{\mathrm{m}}}_{\mathrm{in}}{\mathrm{h}}_{\mathrm{in}}+{\dot{\mathrm{m}}}_{\mathrm{vap}}{\mathrm{h}}_{\mathrm{vap}}={\dot{\mathrm{Q}}}_{\mathrm{out}}+{\dot{\mathrm{m}}}_{\mathrm{liq}}{\mathrm{h}}_{\mathrm{liq}}$			236.63
Effects	${\dot{\mathrm{m}}}_{\mathrm{in}}{\mathrm{h}}_{\mathrm{in}}+{\dot{\mathrm{Q}}}_{\mathrm{in}}={\dot{\mathrm{m}}}_{\mathrm{vap}}{\mathrm{h}}_{\mathrm{vap}}+{\dot{\mathrm{m}}}_{\mathrm{liq}}{\mathrm{h}}_{\mathrm{liq}}$			1051.39
Expander	${\dot{\mathrm{W}}}_{\exp }={\dot{\mathrm{m}}}_{\mathrm{in}}{\mathrm{v}}_{\mathrm{out}}({\mathrm{P}}_{\mathrm{in}}-{\mathrm{P}}_{\mathrm{out}})/\mathrm{Eff}$		8.568
PPW1	${\dot{\mathrm{W}}}_1={\dot{\mathrm{m}}}_7{\mathrm{v}}_7({\mathrm{P}}_8-{\mathrm{P}}_7)/\mathrm{Eff}$	32.43
PPW2	${\dot{\mathrm{W}}}_2={\dot{\mathrm{m}}}_{12}{\mathrm{v}}_{12}({\mathrm{P}}_{13}-{\mathrm{P}}_{12})/\mathrm{Eff}$	33.26
PSW1	${\dot{\mathrm{W}}}_3={\dot{\mathrm{m}}}_{18\mathrm{a}}{\mathrm{v}}_{18\mathrm{a}}({\mathrm{P}}_{19}-{\mathrm{P}}_{18\mathrm{a}})/\mathrm{Eff}$	7.505
PSW2	${\dot{\mathrm{W}}}_4={\dot{\mathrm{m}}}_{18\mathrm{b}}{\mathrm{v}}_{18\mathrm{b}}({\mathrm{P}}_{22} -{\mathrm{P}}_{18\mathrm{b}})/\mathrm{Eff}$	0.176
Streams:
${\dot{\mathrm{m}}}_{22}{\mathrm{h}}_{22}$	Seawater intake	−22,298.3
${\dot{\mathrm{m}}}_{17}{\mathrm{h}}_{17}$	Brine outlet		−10,002.7
${\dot{\mathrm{m}}}_{10}{\mathrm{h}}_{10}+{\dot{\mathrm{m}}}_{15}{\mathrm{h}}_{15}+{\dot{\mathrm{m}}}_{82\mathrm{b}}{\mathrm{h}}_{82\mathrm{b}}$	Product water outlet		−12,401.4
${\dot{\mathrm{m}}}_{35}{\mathrm{h}}_{35}$	Expander inlet steam		−550.98
${\dot{\mathrm{m}}}_{36}{\mathrm{h}}_{36}$	Expander outlet steam		−559.55
Total:		−22,074.92	−22,698.75	1287.99
Product Water	${\dot{\mathrm{m}}}_{ \mathrm{prod}. \mathrm{water}}={\dot{\mathrm{m}}}_{15}+{\dot{\mathrm{m}}}_{10}+{\dot{\mathrm{m}}}_{85\mathrm{a}}$			0.778 kg/s
PR	$\mathrm{PR}=({\dot{\mathrm{m}}}_{\mathrm{prod}. \mathrm{water}}\times 2326 {\displaystyle \frac{\mathrm{kJ}}{\mathrm{kg}}})/({\dot{\mathrm{Q}}}_{\mathrm{gen}})$			12.062
PRglobal	${\mathrm{PR}}_{\mathrm{global}}=({\dot{\mathrm{m}}}_{\mathrm{prod}. \mathrm{water}}\times 2326 {\displaystyle \frac{\mathrm{kJ}}{\mathrm{kg}}})/({\dot{\mathrm{Q}}}_{\mathrm{gen}}+\sum {\dot{\mathrm{W}}}_{\mathrm{tot}.})$			8.179
Recovery Ratio	$\mathrm{RR}=\left({\displaystyle \frac{{\dot{\mathrm{m}}}_{\mathrm{prod}. \mathrm{water}}}{{\dot{\mathrm{m}}}_{\mathrm{Feed}}}}\right)\times 100$			54.2%
Total power demand	$\sum {\dot{\mathrm{W}}}_{\mathrm{tot}. \mathrm{cog}.}={\dot{\mathrm{W}}}_1+{\dot{\mathrm{W}}}_2+{\dot{\mathrm{W}}}_3+{\dot{\mathrm{W}}}_4$			71.20 kW
E.C.Cog	$\mathrm{E}.\mathrm{C}{.}_{\mathrm{cog}.}={\dot{\mathrm{W}}}_{\exp }$$/\sum {\dot{\mathrm{W}}}_{\mathrm{tot}. \mathrm{cog}.}$			0.1203
BC1	${\mathrm{H}}_1=({\mathrm{P}}_{\mathrm{atm}}-{\mathrm{P}}_{\mathrm{dis}})/({\mathsf{\rho }}_{10}\times \mathrm{g})$			10.179 m
BC2	${\mathrm{H}}_2=({\mathrm{P}}_{\mathrm{atm}}-{\mathrm{P}}_{\mathrm{dis}})/({\mathsf{\rho }}_{88\mathrm{b}}\times \mathrm{g})$			10.247 m
BC3	${\mathrm{H}}_3=({\mathrm{P}}_{\mathrm{atm}}-{\mathrm{P}}_{\mathrm{dis}})/({\mathsf{\rho }}_{15}\times \mathrm{g})$			9.952 m
BC4	${\mathrm{H}}_4=({\mathrm{P}}_{\mathrm{atm}}-{\mathrm{P}}_{\mathrm{dis}})/({\mathsf{\rho }}_{17}\times \mathrm{g})$			11.475 m

,

Table 12. Energy evaluation of the proposed MSF/MED at high vacuum system under design conditions in trigenerative operation.

Components:	Energy Balance	Gains (kW)	Losses (kW)	Internal Energy Integration (kW)
Generator	${\dot{\mathrm{m}}}_{31}{\mathrm{h}}_{31}+{\dot{\mathrm{m}}}_{88}{\mathrm{h}}_{88}={\dot{\mathrm{m}}}_{32}{\mathrm{h}}_{32}+{\dot{\mathrm{m}}}_{89}{\mathrm{h}}_{89}$	150
Evaporator	${\dot{\mathrm{m}}}_{85\mathrm{a}}{\mathrm{h}}_{85\mathrm{a}}+{\dot{\mathrm{m}}}_{86}{\mathrm{h}}_{86}={\dot{\mathrm{m}}}_1{\mathrm{h}}_1+{\dot{\mathrm{m}}}_{87}{\mathrm{h}}_{87}$	1224.32
HXBrine	${\dot{\mathrm{m}}}_3{\mathrm{h}}_3+{\dot{\mathrm{m}}}_8{\mathrm{h}}_8={\dot{\mathrm{m}}}_4{\mathrm{h}}_4+{\dot{\mathrm{m}}}_9{\mathrm{h}}_9$		1371.12
HXSea	${\dot{\mathrm{m}}}_{13}{\mathrm{h}}_{13}+{\dot{\mathrm{m}}}_{19}{\mathrm{h}}_{19}={\dot{\mathrm{m}}}_{14}{\mathrm{h}}_{14}+{\dot{\mathrm{m}}}_{20}{\mathrm{h}}_{20}$		1640.41
Condensers	${\dot{\mathrm{m}}}_{\mathrm{in}}{\mathrm{h}}_{\mathrm{in}}+{\dot{\mathrm{m}}}_{\mathrm{vap}}{\mathrm{h}}_{\mathrm{vap}}={\dot{\mathrm{Q}}}_{\mathrm{out}}+{\dot{\mathrm{m}}}_{\mathrm{liq}}{\mathrm{h}}_{\mathrm{liq}}$			236.63
Effects	${\dot{\mathrm{m}}}_{\mathrm{in}}{\mathrm{h}}_{\mathrm{in}}+{\dot{\mathrm{Q}}}_{\mathrm{in}}={\dot{\mathrm{m}}}_{\mathrm{vap}}{\mathrm{h}}_{\mathrm{vap}}+{\dot{\mathrm{m}}}_{\mathrm{liq}}{\mathrm{h}}_{\mathrm{liq}}$			1051.39
Expander	${\dot{\mathrm{W}}}_{\exp }={\dot{\mathrm{m}}}_{\mathrm{in}}{\mathrm{v}}_{\mathrm{out}}({\mathrm{P}}_{\mathrm{in}}-{\mathrm{P}}_{\mathrm{out}})/\mathrm{Eff}$		8.568
PPW1	${\dot{\mathrm{W}}}_1={\dot{\mathrm{m}}}_7{\mathrm{v}}_7({\mathrm{P}}_8-{\mathrm{P}}_7)/\mathrm{Eff}$	45.29
PPW2	${\dot{\mathrm{W}}}_2={\dot{\mathrm{m}}}_{12}{\mathrm{v}}_{12}({\mathrm{P}}_{13}-{\mathrm{P}}_{12})/\mathrm{Eff}$	152.32
PSW1	${\dot{\mathrm{W}}}_3={\dot{\mathrm{m}}}_{18\mathrm{a}}{\mathrm{v}}_{18\mathrm{a}}({\mathrm{P}}_{19}-{\mathrm{P}}_{18\mathrm{a}})/\mathrm{Eff}$	34.24
PSW2	${\dot{\mathrm{W}}}_4={\dot{\mathrm{m}}}_{18\mathrm{b}}{\mathrm{v}}_{18\mathrm{b}}({\mathrm{P}}_{22}-{\mathrm{P}}_{18\mathrm{b}})/\mathrm{Eff}$	0.176
Streams:
${\dot{\mathrm{m}}}_{18\mathrm{b}}{\mathrm{h}}_{18\mathrm{b}}$	Seawater intake	−22,298.3
${\dot{\mathrm{m}}}_{17}{\mathrm{h}}_{17}$	Brine outlet		−4079.88
${\dot{\mathrm{m}}}_{10}{\mathrm{h}}_{10}+{\dot{\mathrm{m}}}_{15}{\mathrm{h}}_{15}$	Product water outlet		−18,258
${\dot{\mathrm{m}}}_{35}{\mathrm{h}}_{35}$	Expander inlet steam		−550.98
${\dot{\mathrm{m}}}_{36}{\mathrm{h}}_{36}$	Expander outlet steam		−559.55
${\dot{\mathrm{m}}}_{88\mathrm{a}}{\mathrm{h}}_{88\mathrm{a}}$	Evaporator coolant inlet		−7843.97
${\dot{\mathrm{m}}}_1{\mathrm{h}}_1$	Evaporator refrigerant fluid outlet	−6619.65
Total:		−27,311.604	−28,272.28	1288.02
Product water	${\dot{\mathrm{m}}}_{\mathrm{prod}. \mathrm{water}}={\dot{\mathrm{m}}}_{15}+{\dot{\mathrm{m}}}_{10}+{\dot{\mathrm{m}}}_{85\mathrm{a}}$			1.149 kg/s
PR	$\mathrm{PR}=({\dot{\mathrm{m}}}_{\mathrm{prod}. \mathrm{water}}\times 2326 {\displaystyle \frac{\mathrm{kJ}}{\mathrm{kg}}})/({\dot{\mathrm{Q}}}_{\mathrm{gen}})$			17.81
PRglobal	${\mathrm{PR} }_{\mathrm{global}}=({\dot{\mathrm{m}}}_{\mathrm{prod}. \mathrm{water}}\times 2326 {\displaystyle \frac{\mathrm{kJ}}{\mathrm{kg}}})/({\dot{\mathrm{Q}}}_{\mathrm{gen}}+\sum {\dot{\mathrm{W}}}_{\mathrm{tot}.}-{\dot{\mathrm{W}}}_{\exp })$			6.995
Recovery ratio	$\mathrm{RR}=\left({\displaystyle \frac{{\dot{\mathrm{m}}}_{\mathrm{prod}. \mathrm{water}}}{{\dot{\mathrm{m}}}_{\mathrm{Feed}}}}\right)\times 100$			80.06%
COP	$\mathrm{COP}={\dot{\mathrm{Q}}}_{\mathrm{eva}}/({\dot{\mathrm{Q}}}_{\mathrm{gen}}+\sum {\mathrm{W}}_{\mathrm{tot}.})$			3.278
COPglobal	${\mathrm{COP}}_{\mathrm{global}}=({\dot{\mathrm{Q}}}_{\mathrm{eva}}+{\dot{\mathrm{m}}}_{\mathrm{prod}. \mathrm{water}}\times 2326 {\displaystyle \frac{\mathrm{kJ}}{\mathrm{kg}}})/({\dot{\mathrm{Q}}}_{\mathrm{gen}}+\sum {\dot{\mathrm{W}}}_{\mathrm{tot}.}-{\dot{\mathrm{W}}}_{\exp })$			10.435
Total power demand	$\sum {\dot{\mathrm{W}}}_{\mathrm{tot}. \mathrm{trig}.}={\dot{\mathrm{W}}}_1+{\dot{\mathrm{W}}}_2+{\dot{\mathrm{W}}}_3+{\dot{\mathrm{W}}}_4$			231.983 kW
E.C.Trig	$\mathrm{E}.\mathrm{C}{.}_{\mathrm{trig}.}={\dot{\mathrm{W}}}_{\exp }$$/\sum {\dot{\mathrm{W}}}_{\mathrm{tot}. \mathrm{trig}.}$			0.0369
BC1	${\mathrm{H}}_1=({\mathrm{P}}_{\mathrm{atm}}-{\mathrm{P}}_{\mathrm{dis}})/({\mathsf{\rho }}_{10}\times \mathrm{g})$			10.185 m
BC2	${\mathrm{H}}_2=({\mathrm{P}}_{\mathrm{atm}}-{\mathrm{P}}_{\mathrm{dis}.})/({\mathsf{\rho }}_{88\mathrm{b}}\times \mathrm{g})$			-
BC3	${\mathrm{H}}_3=({\mathrm{P}}_{\mathrm{atm}}-{\mathrm{P}}_{\mathrm{dis}.})/({\mathsf{\rho }}_{15}\times \mathrm{g})$			9.952 m
BC4	${\mathrm{H}}_4=({\mathrm{P}}_{\mathrm{atm}}-{\mathrm{P}}_{\mathrm{dis}.})/({\mathsf{\rho }}_{17}\times \mathrm{g})$			13.069 m

and

Table 13. Ejectocondenser design conditions.

Ejectocondenser	Parameters	Units	Motive Fluid	Secondary Fluid	Ejectocondenser Output
EJ1	Temperature	°C	9	4	12.19
	Pressure	kPa	350	0.79	1.5
	Mass flow	kg/s	102.236	0.525	102.761
EJ2	Temperature	°C	28.5	4	29.38
	Pressure	kPa	400	0.79	4.10
	Mass flow	kg/s	457.192	0.623	457.815

provide the thermodynamic properties, operating characteristics, and energy evaluation of the system under the selected design conditions.

6. Conclusions

This study presents a novel trigeneration system that simultaneously produces desalinated water, electrical energy, and cooling, addressing critical challenges related to water scarcity and energy demand in coastal regions. The system integrates hybrid MSF and MED technologies under high vacuum, coupled with ejector-condensation mechanisms and an expander-generator subsystem. These innovations enhance energy recovery, operational flexibility, and resource efficiency.

Simulation results using Aspen Plus demonstrated the system’s technical feasibility and superior performance across different operating modes. In trigeneration mode, the system achieved a RR of 80%, a PR of 17.81, and a cooling capacity of 1225 kW while producing 99,273.6 L/day of desalinated water. In cogeneration mode, the system achieved a PR of 12.06 and reduced electrical consumption by 12.03%, emphasizing its adaptability to seasonal and demand-based operational shifts.

Key innovations include the elimination of seawater preheaters in the MED section, which reduces heat losses and increases desalination capacity, and the use of barometric ejector condensers, which optimize steam utilization and cooling production. These features not only improve energy efficiency but also simplify system design, making it more cost-effective and scalable for real-world applications.

The proposed system demonstrates significant potential for mitigating water and energy challenges while promoting sustainability. Future research should focus on validating the system under experimental and field conditions, assessing its long-term reliability, and exploring opportunities for scaling up to meet larger community needs. Additionally, economic analysis and life-cycle assessments would provide deeper insights into the system’s commercial viability and environmental impact.