Thermally Regulated Electrochemical–Membrane Integration for Hydrazine Wastewater Treatment

Unsymmetrical dimethylhydrazine (UDMH), a high-specific-impulse liquid rocket fuel [1], [2], [3], has been extensively utilized in aerospace applications, giving rise to increasingly severe environmental concerns. Since the 1960s, considerable efforts have been devoted to the treatment of UDMH-containing wastewater [4], [5], [6]. Early approaches were primarily based on chemical oxidation and physical processes, such as aeration, chlorination, and incineration. However, these methods generally suffer from low treatment efficiency, high energy consumption, and the generation of secondary pollutants, including chlorinated by-products and exhaust gases, which limit their applicability under increasingly stringent environmental regulations. UDMH wastewater mainly originates from two sources [7]: leakage and cleaning processes associated with storage tanks and pipelines, and combustion by-products generated during rocket launch and testing, which are subsequently entrained in firefighting water and discharged into drainage systems [8], [9], [10].

Since the beginning of the 21st century, more efficient combined treatment processes have been proposed to address these limitations. The development of UDMH wastewater treatment technologies can be broadly categorized into three stages. The first stage involves conventional chemical oxidation coupled with physical adsorption. The study employed chlorine dioxide oxidation to treat UDMH-containing wastewater and preliminarily elucidated the reaction mechanism based on experimental data [11], [12]. A multistage oxidation system combined with activated carbon adsorption was adopted to improve oxidant utilization. Nevertheless, the process requires strict control of the molar ratio between UDMH and oxidant (typically 1:9–1:11), resulting in excessive chemical consumption, high operational cost, and potential secondary pollution due to residual oxidants.

The second stage focuses on hybrid processes combining advanced oxidation processes with biological treatment. To reduce reagent consumption and enhance mineralization efficiency, Li et al. [13] proposed a combined activated carbon–microwave–Fenton process, achieving a removal efficiency of 99.3% at an initial concentration of 400 mg·L$^{-1}$. However, this approach still relies on Fenton reagents and generates substantial amounts of iron-containing sludge, which requires additional handling and increases overall treatment cost. Liu et al. [14] further explored the integration of acidic electrolyzed water with membrane bioreactors. Although water quality was improved, the system exhibited limited adaptability to high-toxicity wastewater and suffered from membrane fouling and operational complexity, restricting its engineering applicability.

The third stage is characterized by electrochemical in situ advanced oxidation technologies. In contrast to conventional processes, Electrocatalytic Oxidation (ECO) utilizes electrons as clean reactants and avoids the addition of external oxidants, thereby reducing the risk of secondary pollution. Yu et al. [15] prepared SnO$_\mathrm{x}$-La-Fe@Ti electrodes via micro-arc oxidation and achieved 90% UDMH removal through in situ hydrogen peroxide generation via the two-electron water oxidation reaction (2e$^-$-WOR). Huang et al. [16] further developed Zn$_2$SnO$_4$@Ti ceramic membrane electrodes, reaching a hydrogen peroxide production rate of 78.4 $\mu$mol·h$^{-1}$·cm$^{-2}$ at 3.3 V (vs. reversible hydrogen electrode). Despite these advances, most studies remain limited to electrode material performance at the laboratory scale, with insufficient attention to integrated system design, process stability, and energy utilization under practical operating conditions. In addition, boron-doped diamond electrodes, known for their high oxidation potential and superior electrochemical stability, have not yet been extensively explored in system-level engineering applications for hydrazine wastewater treatment.

Zeng and Ren [17] constructed a (001) facet-dominated TiO$_2$/CeO$_2$ S-scheme heterojunction and achieved 98.4% degradation with 84.5% mineralization of unsymmetrical dimethylhydrazine (100 mg·L$^{-1}$ within 140 min under simulated solar light. They attributed the enhancement to the synergistic effect of the facet and S-scheme heterojunction in promoting charge separation, identified intermediates to elucidate degradation pathways, and proposed an efficient strategy for aerospace wastewater treatment.

Shen et al. [18] showed that UV/PMS efficiently and safely degrades N, N-dimethylhydrazine compounds while controlling NDMA. They distinguished the roles of SO$_{4}$•, HO•, PMS, and UV: SO$_{4}$•/PMS generated NDMA, while UV assisted radical-mediated elimination. Degradation involved hydroxylation, bond cleavage, and nitrosation; NDMA was eliminated via photocleavage, radical addition, and hydroxylation. Toxicity and EE/O analyses confirmed safety and feasibility.

Song et al. [19] studied NDMA control during ozonation of daminozide and 2-F-DMH, where O$_3$/PMS achieved 50–65% reduction versus 10–25% by O$_3$/H$_2$O$_2$ at an oxidant/O$_3$ ratio of 8:1. Direct ozonation dominated due to high rate constants, and SO$_{4}$• (Rct) correlated linearly with NDMA suppression. Multiple small ozone doses further minimized NDMA. Bromate formation was higher with O$_3$/PMS, so both NDMA and bromate must be monitored in practice.

Su et al. [20] fabricated magnetic ZIF-67@C-ACPs microspheres, achieving a UDMH adsorption capacity of 351.96 mg·g$^{-1}$ via chemisorption with robust resistance to pH and coexisting substances, excellent reusability over six cycles, and low Co leaching. Pore filling, hydrogen bonding, and surface complexation were identified as the dominant mechanisms. A triple-stage tandem column device maintained effluent UDMH below China’s discharge standard ($<$0.5 mg·L$^{-1}$), demonstrating scalable potential for treating UDMH wastewater.

Overall, existing technologies are constrained by two fundamental limitations. On the one hand, conventional Fenton-based processes are inherently associated with a “chemical dosing–sludge generation–sludge disposal” cycle, leading to high operational cost and environmental burden. On the other hand, standalone electrochemical oxidation, although environmentally favorable, faces challenges in treating high-salinity wastewater and often requires additional separation steps to achieve acceptable effluent quality. Meanwhile, heat generation during electrochemical reactions and its impact on system stability and energy efficiency have received limited attention in previous studies.

In view of these challenges, an integrated treatment system combining boron-doped diamond ECO with Disc-Tube Reverse Osmosis (DTRO) is developed and systematically configured. The boron-doped diamond electrode enables efficient mineralization of refractory organic pollutants, while DTRO facilitates the removal of suspended solids and dissolved salts. A recirculation strategy is further introduced to achieve internal salt balance, thereby reducing the need for external electrolyte addition. In addition to structural reliability, the system design accounts for heat dissipation and operational stability under coupled electrochemical and fluid flow conditions. The aim is to establish a technically feasible and energy-conscious approach for hydrazine wastewater treatment, with improved structural integrity and minimized secondary pollution.

The remainder of this study is organized as follows. Section 2 presents the structural design of the integrated system, including framework configuration and key component selection, supported by finite element analysis of static strength and random vibration. Section 3 focuses on the detailed design of the electrocatalytic reactor, including the reactor vessel, electrode assembly, cover structure, cooling system, and piping layout. Section 4 summarizes the main findings and discusses the engineering implications of the proposed design. Figure 1 shows the research framework.

The structural configuration provides the necessary spatial conditions for fluid circulation and heat dissipation, which are essential for maintaining stable operation of the integrated system. As the primary load-bearing unit of the ECO–DTRO integrated treatment equipment, the structural framework not only serves as a stable support for all functional components but also ensures structural integrity and reliability under various conditions, including static loading, vibration, and lifting operations.

This section focuses on the system-level structural design of the equipment, including the layout of the supporting framework and the selection of key structural components. Finite element analysis is conducted to evaluate the static strength and random vibration response of critical regions such as the framework, welded joints, and lifting structures under complex loading conditions, including equipment weight, operational loads, and transportation-induced vibration. Through optimization of weld geometry and lifting structures, the mechanical performance of the system is improved to meet strength and safety requirements, providing a reliable foundation for subsequent process integration and experimental validation.

Figure 2 shows the treatment process for hydrazine-containing wastewater. The wastewater is first delivered into the electrochemical reactor via a self-priming pump. During the initial start-up phase, an appropriate amount of sodium sulfate is introduced to enhance the electrical conductivity of the wastewater. Subsequently, mechanical stirring is initiated and the electrode plates are energized to commence the electrochemical reaction. Under these conditions, UDMH is effectively degraded. Upon complete degradation, the treated effluent is conveyed by a booster pump to a precision filtration unit for solid–liquid separation. The filtrate is then directed to a DTRO system for further concentration. The permeate obtained from the reverse osmosis unit meets the requirements for reuse. The concentrate, which contains a high concentration of salts, is recycled back to the electrochemical reactor to regulate the conductivity of the influent wastewater, thereby replacing the need for continuous sodium sulfate dosing and enabling internal salt circulation. The final reaction products of this integrated treatment process consist primarily of water, carbon dioxide, and nitrogen. The generated off-gas is non-toxic and can be safely discharged at elevated points without causing secondary atmospheric pollution. In addition, the process operates without the need for external heating or continuous chemical addition, and no secondary pollutants are generated during wastewater treatment.

The efficient and stable operation of the process units is fundamentally dependent on the rational spatial configuration and mechanical reliability of the supporting structural framework. Based on the process design, a systematic structural design of the equipment body was further carried out. Emphasis was placed on the optimization of the frame layout and the selection of key component parameters. The finite element method was employed to perform static strength analysis and random vibration simulations of critical structural components, including the frame, weld joints, and lifting lugs. The mechanical response characteristics under multi-source loading conditions were thereby evaluated, providing a structural foundation for subsequent system integration and experimental validation of the process units.

The base of the frame is required to support all internal devices. To reduce the number of structural members, the base was designed in a grid-like configuration analogous to a 4 × 4 layout. Consideration was also given to reserving sufficient space for pipeline routing and internal operational accessibility, in accordance with the principle of structural compactness. The base structure consists of short frame members with a length of 362.5 mm and long frame members with a length of 1600 mm. The upper structure is connected by four vertical frame members, each with a length of 1700 mm.

For the finite element analysis of the structure, the displacement-based method was adopted. The governing mathematical equilibrium equation is generally expressed as:

where, $M$ denotes the mass matrix, $C$ represents the damping matrix, $K$ is the stiffness matrix, $x$ is the displacement vector, and $F$ is the force vector.

In static analysis, when time-dependent loads, inertia effects, and damping contributions are neglected, the governing equation can be simplified as follows:

Under the condition that the damping term is neglected and the inertial term is formally transferred to the right-hand side of the equilibrium equation, the motion equation can be expressed as:

When $K$ is treated as a constant matrix, the formulation corresponds to linear static structural analysis, in which the material behavior is characterized as linearly elastic and the deformation is assumed to be infinitesimal. Conversely, when $K$ is considered a variable matrix, the formulation represents nonlinear static structural analysis.

Based on the aforementioned theoretical framework, the structural frame was simplified for modeling and simulation purposes to reduce computational cost while maintaining adequate analytical accuracy and improving solution efficiency. Since the bottom section of the frame directly bears the primary equipment loads and represents the region most susceptible to deformation and potential failure, simulation of this critical region was considered sufficient to characterize the overall mechanical response of the structure. The complete frame model was imported into the simulation module, and geometric simplification was performed by suppressing features that do not contribute significantly to the primary load-bearing behavior under actual operating conditions. As the model consists of multiple interacting components, bonded contact conditions were defined in accordance with the actual connection configuration, ensuring accurate load transfer.

The material was defined as AISI 304 stainless steel, with the following properties: Young’s modulus of 1.9 × 10$^{11}$ N·m$^{-2}$, mass density of 8000 kg·m$^{-3}$, ultimate tensile strength of 517 MPa, yield strength of 206 MPa, and Poisson’s ratio of 0.27. During mesh generation, a gradual transition scheme was adopted. Following preliminary convergence testing, an optimal element size of 105 mm was selected. Boundary conditions were established to replicate the actual installation state. Specifically, fixed supports were applied at the welded joints between the four bottom corners of the frame and the vertical frame members (1700 mm in length). External loads were applied based on the actual weights of the equipment: each of the three support legs of the ECO reactor was subjected to a load of 3 kN; the inlet pump base plate of the filtration unit was subjected to 1500 N; each filter base plate was subjected to 750 N; and the vertical booster pump base plate was subjected to 3 kN. After solving, post-processing was conducted to generate contour plots of total deformation and equivalent stress. The simulation results are presented in Figure 3.

According to the equivalent stress distribution, the maximum stress observed at the bottom of the frame is 134.8 MPa, which is significantly lower than the allowable design stress of the material (185 MPa). Therefore, from the perspective of static strength, the designed frame satisfies safety requirements, and the structural strength is deemed adequate. Following structural optimization, the maximum equivalent stress at the lower weld joints was reduced from 192.4 MPa to 155.0 MPa, corresponding to a reduction of approximately 19.5%. This decrease indicates a substantial improvement in the safety margin.

A schematic representation of the critical weld locations is shown in Figure 4. The frame structure was fabricated using the shielded metal arc welding process, with AISI 304 stainless steel selected as the base material. Welding consumables of type A102 or A107 electrodes were employed, corresponding to the E308-XX classification. The deposited weld metal of this electrode type is characterized by a minimum ultimate tensile strength $R_{\mathrm{m}} \geq 550 \mathrm{MPa}$ and an elongation after fracture $A \geq$ 30%. Since all welded joints in the frame are fillet welds, the design strength of the weld metal was conservatively taken as $f_{\mathrm{wd}}=200 \mathrm{MPa}$. The thickness of the frame members is 2.5 mm. An I-shaped groove configuration (i.e., without groove preparation) was adopted. To prevent cracking caused by excessively rapid cooling of the weld metal, the fillet weld leg size $h_{\mathrm{f}}$ was constrained to be more than $1.5 t^{0.5}$. $t=2.5 \mathrm{~mm}$ represents the thickness of the welded components. Simultaneously, to avoid excessive heat input leading to significant welding distortion and residual stress, the fillet weld leg size was limited to no more than 1.2 times the plate thickness. Accordingly, it satisfied $1.5 t^{0.5} \leq h_{\mathrm{f}} \leq 1.2 t$ and $2.38 \mathrm{~mm} \leq h_{\mathrm{f}} \leq 3 \mathrm{~mm}$, and $h_{\mathrm{f}}$ was selected as 2.5 mm. The weld length $(L)$ was required to be no less than $8 h_{\mathrm{f}}$, and was therefore taken as 40 mm.

Considering the welding configuration at the bottom of the frame, the most critical weld location is identified as the front fillet weld at the junction between the frame base and the 1700 mm vertical frame member, where the applied load acts perpendicular to the weld length. According to the strength calculation formula for right-angle fillet welds, the normal stress perpendicular to the weld can be expressed as:

where, $\sigma_{\mathrm{f}}$ denotes the normal stress calculated based on the effective weld cross-section; $h_{\mathrm{e}}$ represents the effective throat thickness of the fillet weld, taken as 0.7 times the weld leg size; $l_{\mathrm{w}}$ is the effective weld length, defined as the actual weld length minus twice the weld leg size; and $\beta_{\mathrm{f}}$ is the strength enhancement factor for front fillet welds under static and intermittent dynamic loading conditions, with a value of 1.22. Based on the above formulation, $\sigma_{\mathrm{f}}=15.3 \mathrm{MPa}$, which is significantly lower than $f_{\mathrm{wd}}$ (200 MPa).

Taking into account the actual load transfer characteristics of the structure, the simulation model of the welded joints was further simplified rationally. Since the stress states of the critical welds on the four side faces are similar, one representative side was selected for detailed simulation analysis. As shown in the welding simulation contour plots (Figure 5), bonded contact conditions were applied between all interacting geometries, and a default mesh size was adopted. The simulation results indicate that the maximum stresses are 100.8 MPa and 196.51 MPa, respectively.

Local mesh refinement was performed on the geometric regions corresponding to the upper and lower welds on the left side, with the element size reduced to 4 mm. The boundary conditions were defined as follows: the end surface of the 1700 mm vertical frame member near the welded joint was constrained as a fixed support; a concentrated load of 3 kN was applied vertically downward on the upper surface of the 1600 mm horizontal frame member. After completion of the solution, post-processing was conducted to obtain the total deformation contours. Given that the stress distributions on the left and right sides are highly similar, the upper and lower welds on one side were selected to further evaluate the stress state of the welded joints. The corresponding equivalent stress contours were extracted for analysis, as shown in the weld simulation contour plots.

According to the refined equivalent stress results (Figure 6), the maximum equivalent stress in the upper weld is 97.8 MPa, whereas that in the lower weld reaches 192.4 MPa. Both values remain below $f_{\mathrm{wd}}$ (200 MPa), indicating that all welds satisfy the strength requirements. However, the lower weld exhibits a comparatively higher stress level, with its maximum value approaching the allowable design limit $f_{\mathrm{wd}}$. This observation suggests that this region represents a relatively weak zone within the structure.

To enhance the load-bearing reliability and safety margin of the critical weld region, geometric optimization was performed on the lower weld. Specifically, the fillet weld leg size hf was increased from 2.5 mm to 3.0 mm. This modification leads to a significant increase in the effective load-bearing cross-sectional area of the weld, thereby improving its resistance to bending and shear stresses. As a result, stress concentration effects are effectively mitigated, and both fatigue life and overall structural integrity are improved. The optimized weld configuration is therefore expected to more reliably satisfy the operational requirements under combined conditions of long-term static loading and transportation-induced vibration.

During practical operation, the hydrazine-containing wastewater treatment equipment is required to be frequently relocated between different sites. Consequently, road-induced excitations during transportation generate continuous vibration loads. These vibration loads are transmitted to the welded regions of the structural frame. Due to geometric discontinuities, welds inherently exhibit stress concentration effects, and their fatigue strength is generally lower than that of the base material. As a result, fatigue cracks are prone to initiation and propagation under cyclic loading, ultimately leading to fatigue failure. In particular, the welds located at the junction between the frame base and the 1700 mm vertical frame members are subjected to relatively high residual stresses (pre-stresses) introduced during the welding process, further increasing the susceptibility to fatigue damage. To evaluate the vibration reliability of the structure, a random vibration analysis based on the modal superposition method was adopted.

Modal analysis is fundamentally based on a coordinate transformation, in which a linear time-invariant system is transformed from the time or frequency domain into the modal domain. In this transformation, modal coordinates are introduced to replace the physical coordinates associated with the system of vibration differential equations. This approach enables the decoupling of the governing equations into a set of independent equations expressed in terms of modal coordinates and modal parameters, thereby facilitating the determination of the system’s modal parameters. Accordingly, the dynamic equation of motion for an N-degree-of-freedom system can be expressed as:

where, $M$ denotes the mass matrix, $C$ represents the damping matrix, and $K$ is the stiffness matrix.

When a lower level of computational accuracy is acceptable, structural damping may be neglected. Furthermore, by assuming that the external excitation force is zero, the governing differential equation can be reduced to:

Under this condition, the structure undergoes free vibration, and the nodal displacements can be expressed in the form of simple harmonic motion:

where, $x_0$ denotes the amplitude vector of the nodes (i.e., the mode shape), $\omega$ represents the natural circular frequency corresponding to the mode, and $\varphi$ is the phase angle.

Substituting this expression into the governing equation of motion yields the following equation:

During free vibration, the nodal displacement amplitude vector $\left\{u_0\right\}$ is non-zero; therefore, the determinant of the above equation is equal to zero. By solving the above equation, the $n$-th-order natural frequencies $\left(\omega_1, \omega_2, \omega_3 \ldots \omega_n\right)$ can be obtained.

Once the natural frequencies are obtained, they can be substituted back into the governing equation to determine the corresponding nodal displacement expressions.

Considering that, during transportation, the bottom of the frame continuously sustains the static loads of the mounted equipment and that the critical welds are subjected to residual pre-stress, the analysis was conducted under prestressed conditions. Accordingly, both modal extraction and random vibration response analysis were performed with the inclusion of prestress effects. The simulation procedure was implemented as follows. A static structural analysis module was first established in ANSYS Workbench, and was subsequently coupled in sequence with modal analysis and random vibration analysis modules. The material properties were defined consistently with the previously established database. Preprocessing was performed on the simplified structural model, followed by numerical simulation.

As illustrated in the random vibration contour plots (Figure 7), the equivalent stress results were evaluated using a scaling factor of 3$\sigma$. Under this criterion, it was determined that the maximum equivalent stress experienced by the upper weld does not exceed 22.2 MPa with a probability of 99.73%, while that of the lower weld does not exceed 23.6 MPa at the same confidence level. These results indicate that, under the specified transportation-induced vibration load spectrum, the maximum equivalent stress in the critical weld regions remains significantly lower than $f_\mathrm{wd}$ (200 MPa). Therefore, the structural integrity of the welded frame is confirmed to satisfy safety requirements under random vibration conditions.

Given the substantial overall weight of the hydrazine-containing wastewater treatment equipment and the requirement for frequent relocation between operational sites, lifting operations are necessary. Therefore, a lifting structure satisfying the required strength criteria must be designed. In the preliminary design, lifting lugs were symmetrically arranged and welded onto the four top frame members. Based on this configuration, a simplified mechanical model of a single lifting lug was established for load analysis, as illustrated in Figure 8.

During lifting operations, the hydrazine-containing wastewater treatment equipment is assumed to be in a non-operational state. For conservative design considerations, the total weight was taken as $1.5 t$, corresponding to a gravitational load of $G=15$ kN. In practical lifting conditions, the angle between the lifting sling and the vertical direction typically ranges from $45^{\circ}$ to $60^{\circ}$. When this angle reaches $60^{\circ}$, the lifting lugs experience the most unfavorable loading condition. Under this configuration, the maximum load acting on each lifting lug is given by $F_{\max}=G / 2=7.5 \mathrm{kN}$.

For design simplicity, AISI 304 stainless steel was selected for the lifting lugs. The material is characterized by an ultimate tensile strength of $\sigma_\mathrm{b} \geq$ 520 MPa. A safety factor of $n=4$ was adopted, resulting in an allowable stress of $[\sigma]=130 \mathrm{MPa}$. Accordingly, the tensile stress in the lifting lug must satisfy the condition $\sigma \leq$ 130 MPa. Based on the corresponding design equation, $d \geq 6.07$ mm.

The outer diameter of the lifting lug was selected as $D_1=45$ mm, and the inner diameter was defined as $D_2=20$ mm. Accordingly, $d=D_1-D_2=23$ mm, which significantly exceeds the minimum required value calculated from the strength criterion.

Given that the prototype operates under variable site conditions and possesses considerable weight, manual handling is impractical. Consequently, relocation must be performed through lifting operations, which necessitates further evaluation of the weld strength at the lifting lugs through simulation analysis. Since the load distribution among the four lifting lugs is approximately uniform during lifting, the model was simplified by selecting a single representative lifting lug for simulation. Bonded contact conditions were applied between all interacting geometries, and a default mesh size was adopted for discretization.

Fixed supports were applied at both ends of the 1600 mm frame member. External loads of 3750 N were applied at two cross-sectional locations of the lifting lug, with the force direction specified as vertically upward. After completing the solution, post-processing was conducted to obtain contour plots of total deformation and equivalent stress in the weld region (Figure 9).

In the simulation analysis, the default mesh size for the weld region was initially set to 4.2 mm. To further improve the accuracy of the stress evaluation, the mesh was refined by reducing the element size to 3.5 mm. A localized mesh control was applied by selecting the weld region and independently assigning the refined mesh size of 3.5 mm. Figure 10 shows the equivalent stress contours obtained after mesh refinement.

According to the refined equivalent stress contour results, the maximum stress in the weld region was found to be 214.5 MPa, which exceeds $f_\mathrm{wd}$ (200 MPa). This indicates that the initial weld configuration does not satisfy the strength requirements, and thus modification of the weld geometry or optimization of the lifting structure is necessary. Following iterative numerical simulations, the final design configuration was established. Specifically, the bottom arc radius of the lifting lug was increased to 13 mm, and the fillet weld leg size was increased to 3 mm. These modifications effectively increased both the weld length and the effective load-bearing cross-sectional area. The results indicate that the maximum equivalent stress in the weld region was reduced to 190.1 MPa, which is lower than $f_\mathrm{wd}$ (200 MPa). Therefore, the optimized design satisfies the required strength criteria. Figure 11 shows the equivalent stress contour of the optimized weld.

Building upon the structural design of the equipment framework, further emphasis was placed on the structural design and parameter optimization of key components. Detailed design efforts were concentrated on the electrocatalytic reactor and its associated subsystems. These include the reactor vessel, electrode assembly, cover structure, cooling system, and piping network. The structural configuration and parameter selection of these core components were systematically refined to ensure reliable hardware support for subsequent process integration and experimental validation. In addition to mechanical reliability, the design of key components also considers heat transfer and fluid flow behavior, which are essential for maintaining stable operating conditions in electrochemical treatment systems.

To satisfy the overall dimensional constraint of the hydrazine-containing wastewater treatment equipment (1.5 m × 1.5 m × 1.5 m), a vertically stacked configuration of electrode assemblies was deemed infeasible due to height limitations. Therefore, three electrode assemblies were arranged within the same horizontal plane. Given that the dimensions of a single electrode plate are 150 mm × 285 mm and that a central agitator must be installed within the reactor, the internal diameter of the vessel was required to exceed 653 mm. Accordingly, the reactor vessel diameter was set to 700 mm. To facilitate complete discharge of treated wastewater after the reaction process and to prevent liquid retention, the bottom of the vessel was designed with a composite curved geometry. Specifically, the longitudinal section of the bottom consists of two arc segments smoothly tangent to the vessel wall. The radius of the lower arc was set to 593 mm, while that of the upper arc was defined as 152 mm, ensuring a smooth transition. The vessel wall thickness was determined based on the following expression:

where, $s$ is the wall thickness, $P_1$ is the internal water pressure (calculated as 0.0061 MPa), $D$ is the internal diameter of the vessel, $\sigma$ is the allowable stress of the material (taken as 130 MP ), and $\Phi$ is the weld joint efficiency factor (taken as 0.8). The calculated wall thickness was approximately 0.02 mm. Therefore, to ensure manufacturability, computation, and an adequate design margin, a wall thickness of 5 mm was adopted. The longitudinal cross-sectional configuration of the reactor vessel is illustrated in Figure 12.

Each electrode assembly consists of 15 layers, with a thickness of 1.5 mm per layer. Accordingly, the volume of each electrode assembly was estimated to be approximately 0.003 m$^3$, while the volume of the cooling pipes was also estimated to be 0.003 m$^3$. Based on a total liquid volume of 200 L, the liquid level of the wastewater was calculated to be 610 mm above the vessel bottom. Therefore, both inlet ports were positioned above the liquid level, while the upper edge of the electrode assemblies was required to remain below the liquid surface. Additionally, a clearance of at least 90 mm was reserved at the vessel bottom for the installation of the outlet pipe. Supporting legs were also required to elevate the reactor structure, and the agitator motor has a length of 263 mm. Consequently, the overall vessel height was required to exceed 180 + 610 = 790 mm. Considering sufficient design margin, the total vessel depth was selected as 773 mm, resulting in a cylindrical wall height of 580 mm.

The supporting legs were arranged at an inclination angle $\theta$ relative to the central axis. To avoid extending beyond the vessel diameter and occupying additional external space, the angle $\theta$ was determined based on the following relationship:

where, $h_1$ represents the height of the center of gravity of the entire electrocatalytic reactor, calculated to be approximately 1300 mm , and $r$ denotes the radial distance from the central axis to the support leg, taken as 200 mm. The calculated angle is $\theta=81.25^{\circ}$, and a design value of $\theta=80^{\circ}$ was adopted. Each support leg has a length of 247 mm, elevating the vessel bottom to a height of 180 mm above the ground to accommodate the outlet piping. The base of each support leg was designed with dimensions of 230 mm × 230 mm × 15 mm. Three support legs were arranged symmetrically at $120^{\circ}$ intervals, consistent with the subsequent electrode configuration. The distribution of the support legs is illustrated in Figure 13.

The width of the vessel rim was designed as 50 mm. Eight uniformly distributed holes were drilled along the rim to enable bolted fastening between the vessel body and the cover. Among these, three holes were designed as outward-opening slots to allow the cover to be lifted and opened from one side. When combined with wing nuts, this configuration facilitates convenient partial opening of the cover for inspection of the internal condition of the vessel. The design is illustrated in Figure 14.

Each electrode assembly was inserted into the reactor vessel through a square opening of 240 mm × 275 mm. To secure the assembly, a corresponding frame of dimensions 370 mm × 330 mm was employed. The frame thickness was selected as 5 mm. A total of 20 uniformly distributed holes were arranged along the perimeter of the frame to facilitate bolted connections between the frame and the electrode assembly. The three square openings and their corresponding frames were arranged symmetrically at 120° intervals. The radial distance R from the frame to the central axis of the vessel was determined based on the following relationship:

where, $d$ represents the reserved diameter for the agitator. Since an agitator must be installed at the center of the vessel, sufficient clearance must be provided to avoid interference with the electrode assemblies. Given that the impeller diameter is 88 mm , a design value of $d=98$ mm was adopted. The parameter $l$ denotes the length of the internal electrode support structure, taken as 320 mm . Based on these parameters, $R=418$ mm (Figure 15).

The wastewater outlet pipe at the bottom of the vessel was designed with an inner diameter of 40 mm and a wall thickness of 5 mm. The pipe extends vertically downward by 90 mm and then bends at $90^{\circ}$ toward the side of the vessel opening corresponding to the electrode assembly. It is positioned along the central line between two adjacent electrode assemblies. The horizontal section has a length of 78 mm, and a flange with a diameter of 144 mm and a thickness of 15 mm was installed at the terminal end to ensure reliable pipe connection.

On the same side of the vessel, two additional ports were arranged in the upper region: one for the return of concentrated reverse osmosis brine after filtration and the other for the cooling pipe connection. The return pipe for concentrated brine has a diameter of 28 mm and a wall thickness of 2 mm. The vertical position ($H$) of the pipe opening relative to the vessel rim was determined according to:

where, $H_1$ is the total vessel depth (773 mm), $h_2$ is the liquid level height (610 mm), $l$ is the extension pipe length (70 mm), and $x$ is the clearance between the pipe opening and the liquid surface, taken as 23 mm . Based on these parameters, the calculated value is $H=70$ mm. Due to the relatively high pressure and flow velocity within this pipe, a $90^{\circ}$ elbow was incorporated inside the vessel, followed by a vertical extension pipe of 70 mm length directed downward. This configuration serves to guide the flow and prevent direct impingement on the opposite electrode assembly. On the external side of the vessel, the horizontal pipe section has a length of 80 mm , and a flange with a diameter of 144 mm and a thickness of 15 mm was installed at the terminal end. To enhance the structural stiffness of the connection between the pipe and the flange, four reinforcing ribs with a thickness of 5 mm were arranged between the flange and the vessel wall. These ribs were designed to withstand elevated internal pressure and to suppress vibration effects.

Both cooling pipe nozzles were designed with an inner diameter of 22 mm and a wall thickness of 4 mm. The two nozzles were symmetrically arranged on the upper section of the reactor vessel, each located at a distance of 120 mm from the vessel rim. The center-to-center spacing between the two nozzles was set to 177 mm, and they were arranged symmetrically about the axis of the concentrated reverse osmosis brine return port. The total length of each nozzle was 44 mm, with approximately 5 mm extending into the interior of the vessel, allowing the nozzle ends to slightly protrude from the inner wall surface. This design facilitates precise alignment and sealing between the cooling pipes and the nozzles while also promoting a smooth internal flow transition. A schematic diagram of the concentrated reverse osmosis brine return port and the cooling pipe nozzles is shown in Figure 16.

The pH sensor port and the wastewater sampling port were both located on the side corresponding to the electrode assembly adjacent to the vessel cover opening, specifically on its right-hand side. These two ports were symmetrically arranged along the central line between two adjacent electrode assemblies, as illustrated in Figure 17. The pH sensor port was designed with a diameter of 20 mm and a wall thickness of 6 mm. A pH sensor was integrated into this port to enable real-time monitoring of the pH value of the wastewater within the reactor. The measured data were transmitted to the automated control system, thereby providing critical feedback for maintaining stable operation of the electrochemical process. The wastewater sampling port was designed with a diameter of 14 mm and a wall thickness of 3.5 mm. A manual sampling valve was installed at the port to facilitate periodic sample collection for subsequent offline analysis and process validation. A schematic diagram of the pH sensor port and the wastewater sampling port is presented in Figure 17.

The wastewater inlet was designed with a diameter of 25 mm and a wall thickness of 3.5 mm, with a total nozzle length of 110 mm. A flange with a diameter of 144 mm and a thickness of 15 mm was installed at the external end of the pipe to ensure reliable connection with external piping systems. On the internal side, the inlet pipe extends approximately 5 mm into the reactor, allowing the nozzle tip to slightly protrude above the inner wall surface. This configuration facilitates flow guidance and reduces inlet turbulence. The inlet was positioned at a location $90^{\circ}$ clockwise from the wastewater sampling port, with a vertical distance of 80 mm from the vessel rim.

Three handles were mounted on the external surface of the vessel. These handles were symmetrically arranged at $120^{\circ}$ intervals around the vessel circumference, with one handle positioned directly opposite the wastewater inlet. Each handle has a length of 100 mm and extends approximately 60 mm outward from the vessel surface. The cross-section of each handle is circular, with a diameter of 20 mm. The vertical distance between the handles and the vessel rim was set to 150 mm. The arrangement of the handles is illustrated in Figure 18.

Each electrode assembly was composed of eight anode plates and seven cathode plates, which were separated by insulating spacers to form an integrated electrode stack with overall dimensions of 308 mm in length, 190 mm in width, and 75 mm in height. Two conductive strips were inserted between each layer of electrode spacers to electrically connect the cathode plates in series. The front and rear electrode spacers were secured using specially designed clamping plates, and the entire assembly was integrated using four anode conductive rods. These conductive rods were further enclosed within protective sleeves to prevent direct exposure to the wastewater. The length of the anode conductive rod $H_2$ was determined using the following relationship:

where, $x_1$ represents the length of the electrode stack (308 mm), d denotes the thickness of the mounting plate (10 mm, as detailed in a subsequent subsection), and $l_1$ is the external extension length of the conductive rod outside the reactor vessel. To facilitate electrical connections and interconnection among conductive rods, $l_1$ was selected within the range of 80–100 mm, and a value of 92 mm was adopted. Based on these parameters, $H_2$ = 400 mm was obtained.

Each anode plate was designed with seven through-holes. The holes located on the main surface of the plate were used to fasten the plates together into a complete electrode assembly using bolts. The holes positioned on the protruding sections of the plates were used to connect adjacent anode plates and to establish electrical conduction between the anode plate group and the anode conductive rods. To prevent deformation of the protruding sections during bolt tightening, conductive rings were introduced between adjacent layers of anode plates. This configuration ensures uniform distribution of clamping forces while maintaining reliable electrical connection. The detailed arrangement is illustrated in Figure 19.

A mounting plate with dimensions of 370 mm × 330 mm × 10 mm was installed at the rear of the electrode assembly to provide structural support and fixation. The mounting plate was secured to the reactor vessel using bolted connections, ensuring reliable mechanical integration. The electrode rods were designed to pass through the mounting plate and extend to the exterior of the vessel, enabling connection to external power cables. Effective sealing measures were implemented at the penetration points to prevent leakage. At the interface between the electrode rods and the electrical cables, a protective enclosure was installed to prevent accidental contact with live conductors. The protective enclosure was designed with a quick-release mechanism, allowing convenient installation and removal to facilitate routine inspection and maintenance. The detailed structural configuration is illustrated in Figure 20.

The reactor cover was designed with the same thickness as the vessel wall (5 mm) and a diameter of 800 mm, allowing complete coverage of the vessel rim. A split-type configuration was adopted, consisting of a fixed cover and a hinged movable cover plate. These two components were connected through hinges, enabling convenient opening for inspection and maintenance of the internal reaction conditions. Three openings were arranged on the main body of the cover. A central opening with a diameter of 230 mm was provided, above which a flange was installed. The distance $X$ between the flange and the upper surface of the cover was determined according to:

where, $H_3$ is the length of the agitator shaft (548 mm), $L$ is the vertical distance from the bottom of the electrode frame to the vessel opening (488 mm), and $h_3$ is the thickness of the cover (5 mm). Since the lower end of the agitator shaft was designed to be aligned with the bottom of the electrode assembly to ensure effective mixing of the wastewater around the electrodes, the resulting distance (from the flange plate to the outer surface of the tank cover was determined by subtracting the height from the electrode frame bottom to the cylinder opening and the cover thickness) was calculated as $X$ = 55 mm. The flange was designed with a thickness of 5 mm and a diameter of 300 mm. Two exhaust ports were symmetrically arranged on either side of the central opening. These ports were intended for the release of gases generated during the treatment process. The right-side exhaust port was connected to a fan via a flexible hose to continuously remove gases produced during operation. The left-side exhaust port remained sealed under normal conditions but could be connected directly to an overhead pipeline in emergency situations where rapid gas discharge is required. Each exhaust port has a diameter of 100 mm. A flange was installed above each port, located 55 mm above the cover surface, with a thickness of 3 mm and a diameter of 220 mm. A temperature measurement port was positioned at an angular offset of $25^{\circ}$ from the central axis on the right side, at a radial distance of 280 mm. The diameter of this port was 55 mm. Two handles were symmetrically arranged along the vertical axis at positions 250 mm above and below the center, respectively. These handles were designed to facilitate lifting of the entire cover during disassembly or partial lifting of the lower cover section for inspection. The cross-sectional geometry of each handle consisted of two semicircles with a diameter of 10 mm combined with a rectangular section of 15 mm × 10 mm. Each handle has a length of 100 mm and is positioned 40 mm above the cover surface. A 10 mm wide inner flange was incorporated on the underside of the cover to ensure proper alignment during closure. Additionally, a grid-like stiffening structure was introduced to enhance the load-bearing capacity of the cover. The reactor cover model is illustrated in Figure 21.

Due to the geometric constraints imposed by the vessel diameter, as well as the spatial arrangement of the agitator and electrode assemblies, the cooling pipe was configured to be coiled beneath the electrode assemblies to avoid mechanical interference. The inlet temperature of the cooling water was assumed to be 20 ℃. The heat dissipation capacity of the cooling pipe per unit time can be expressed as:

where, $Q_1$ denotes the heat transfer rate of the cooling pipe (W), $h_3$ represents the convective heat transfer coefficient (W·m$^{-2}$·℃$^{-1}$), $A$ is the surface area of the cooling pipe (m$^2$), and $\Delta T$ is the temperature difference between the pipe surface and the surrounding fluid (℃).

In practical engineering calculations, the convective heat transfer coefficient $h_1$ is typically determined through empirical correlations or experimental measurements. For conventional cooling pipes, $h_1$ can be estimated using the following relationship:

where, $Nu$ is the Nusselt number (dimensionless), representing the convective heat transfer capability of the fluid within the pipe; $k$ is the thermal conductivity of the fluid (W·m$^{-1}$·℃$^{-1}$); and $D_1$ is the equivalent diameter of the cooling pipe (m). The Nusselt number Nu can be further determined using empirical correlations, as expressed in Eq. (19).

where, $Re$ denotes the Reynolds number, a dimensionless parameter characterizing the flow regime of the fluid, which was taken as 1500. The Prandtl number $Pr$, also dimensionless, represents the thermal diffusivity characteristics of the fluid and was taken as 7. Based on these values, the Nusselt number $Nu$ was calculated as $Nu$ = 77.28. By adopting a cooling pipe diameter of 20 mm, the convective heat transfer coefficient was determined as 9.2736.

The surface area of the cooling pipe $A$ can be calculated using the following expression:

where, $\pi$ is the mathematical constant (approximately 3.14), $D$ is the diameter of the cooling pipe (m), and $L_1$ is the total length of the pipe (m). The cooling coil was designed with four helical turns and a pitch of 40 mm , resulting in a total pipe length of $L_1=4.145 \mathrm{~m}$. The resulting surface area was $A$ = 0.3 m$^2$. The corresponding heat dissipation rate of the cooling pipe per unit time was therefore obtained as $Q_1$ = 5564.16 W . The heat generation of the liquid within the reactor during operation can be expressed as:

In this expression, $Q_2$ denotes the heat generated by the water (W·s), $m$ is the mass of water (g), $c$ is the specific heat capacity of water (J·g$^{-1}$·℃$^{-1}$), and $\Delta T_1$ is the temperature change (℃). The values were taken as m = 200,000 g, c = 4.2 J·g$^{-1}$·℃$^{-1}$, and $\Delta T_1$ = 1 ℃, yielding $Q_2$ = 840,000 W per second. Therefore, by continuously monitoring the internal temperature of the reactor, the cooling performance can be regulated through adjustment of the coolant flow rate or composition, enabling effective control of the reaction temperature within the vessel. The cooling pipe model is illustrated in Figure 22.

From a system-level energy perspective, the balance between heat generation and heat dissipation plays a decisive role in maintaining stable operation. The calculated cooling capacity of 5564.16 W indicates that the designed cooling system is capable of effectively removing the heat generated during electrochemical reactions under typical working conditions. This thermal balance prevents excessive temperature rise within the reactor, thereby avoiding deterioration of reaction efficiency and ensuring stable performance over prolonged operation. It should be noted that the heat generation estimation is based on a simplified assumption, while actual operating conditions involve heat dissipation through multiple pathways, including fluid circulation, reactor walls, and heat exchange with surrounding components, resulting in a significantly lower effective heat accumulation within the system.

A hydrostatic pressure corresponding to a liquid level of 620 mm from the vessel bottom was applied to the inner wall of the reactor and internal components. Based on the numerical solution, the stress distribution of the electrocatalytic reactor was obtained, as illustrated in Figure 23.

When the reactor was filled with 200 L of wastewater, the maximum stress was observed at the cover region supporting the agitator and at the interface between the cooling pipe and the vessel. The maximum stress value was calculated as 8.2302 MPa, which is significantly lower than the yield strength of stainless steel (310 MPa).

Subsequently, a rotational speed of 88 rpm was applied to the agitator. Under operating conditions in which the agitator was rotating and the reactor contained 200 L of wastewater, the resulting stress distribution of the electrocatalytic reactor is shown in Figure 24. The maximum stress remained concentrated at the cover region supporting the agitator and at the interface between the cooling pipe and the vessel, with a value of 8.2303 MPa. This result indicates that the vibration induced by the motor-driven rotation of the agitator blades does not produce a significant influence on the overall structural strength and can therefore be considered negligible.

This study presents the design and optimization of an integrated treatment system combining ECO with DTRO for hydrazine-containing wastewater. The work addresses both structural reliability and system-level operational stability under practical engineering conditions.

(a) An integrated treatment system based on ECO coupled with DTRO was developed. The structural performance of key load-bearing components, including the framework, welded joints, and lifting structures, was evaluated through finite element analysis under static loading and transport-induced vibration conditions. The optimized configuration reduced the maximum equivalent stress of the framework to 134.8 MPa, while the peak stress in the lower weld decreased from 192.4 MPa to 155.0 MPa, and that in the lifting-ring weld decreased from 214.5 MPa to 190.1 MPa, all meeting the required strength criteria. Random vibration analysis further confirmed the reliability of the welded regions under transportation conditions, ensuring stable mechanical performance during operation and handling.

(b) A comprehensive design and parameter selection of the electrocatalytic reactor and its key components were carried out, including the reactor vessel, electrode assembly, cover structure, cooling system, and piping layout. The reactor was designed with a diameter of 700 mm and a wall thickness of 5 mm, and a double-arc bottom geometry was adopted to facilitate complete drainage. The electrode assembly, consisting of eight anode plates and seven cathode plates, was integrated with conductive rings, mounting plates, and protective structures to ensure reliable electrical connection and operational safety. The split-type cover structure improves accessibility for maintenance. In particular, the coiled cooling system, with a heat dissipation capacity of 5564.16 W, enables effective removal of heat generated during electrochemical reactions, thereby maintaining thermal stability within the reactor. The coordinated design of structural layout and flow passages also supports stable fluid circulation, which is beneficial for both mass transfer and temperature regulation during operation.

Overall, the proposed system provides a feasible and reliable engineering solution for the treatment of hydrazine-containing wastewater. The results highlight the importance of integrating structural design with thermal management considerations, offering practical guidance for the development of stable and energy-conscious electrochemical treatment systems.