Evaluation of disinfection and cavitation performance of a cylindrical rotational hydrodynamic cavitation reactor: Influence of key geometric parameters of the cavitation generation unit

Highlights

• •Disinfection performance of CRHCR via hydrodynamic cavitation was investigated.
• •Cavitation disinfection mechanism was elucidated through a micromorphological analysis.
• •Rectangular grooves exhibited superior cavitation performance over other geometries.
• •Effect of groove width, depth, oblique and groove angles, and groove numbers on cavitation was analysed.

Abstract

Hydrodynamic cavitation offers a promising technological platform for diverse industrial applications, including water treatment and chemical process intensification, and holds significant potential for widespread adoption in future advanced processing systems. This study investigates the disinfection efficacy of a novel Cylindrical Rotational Hydrodynamic Cavitation Reactor (CRHCR) and elucidates the underlying mechanism of Escherichia coli (E. coli) inactivation induced by hydrodynamic cavitation. Microscopic analysis of E. coli post-treatment revealed that the intense mechanical shear forces produced by collapsing bubbles are primarily responsible for bacterial inactivation. In addition, the influence of key geometric parameters of the cavitation generation unit on the hydrodynamic performance of the CRHCR was systematically examined. The results demonstrate that rectangular grooves exhibit superior cavitation performance compared to trapezoidal and triangular configurations. An increase in groove width and number correlates positively with enhanced cavitation intensity. In contrast, changes in groove depth, oblique tooth angle, and groove angle exhibit a non-linear trend, with cavitation performance initially increasing with as these parameters rise, followed by a decline once optimal thresholds are exceeded. Optimal cavitation performance was attained with a groove depth of 2 mm, an oblique tooth angle of 68°, and a groove angle of 5°. The observed variation in cavitation efficiency across different CRHCR configurations is attributed to the distinct geometries of the cavitation generation units, which modulate the distribution of low-pressure zones. These findings provide valuable insights into the structural design, theoretical understanding, and practical application of advanced hydrodynamic cavitation systems.

1. Introduction

Cavitation is a phase transition-driven physical phenomenon that occurs when the local pressure within a liquid falls below its saturated vapor pressure, leading to the formation of vapor-filled cavities or bubbles. These cavitation bubbles subsequently undergo rapid and violent collapse during pressure recovery, releasing intense localized energy [1]. This collapse generates extreme transient conditions—temperatures approaching 5000 K, pressures up to 1000 bar, and high-velocity microjets—collectively known as sonochemical effects [2]. The energy released during bubble collapse can induce the homolytic cleavage of water molecules, generating highly reactive hydroxyl radicals (·OH). These radicals enhance reaction kinetics, disrupt microbial cell walls, and facilitate the synthesis of nanomaterials, among other effects [[3], [4], [5], [6], [7], [8]]. As such, cavitation exhibits significant potential for diverse applications across environmental, biological, and nanotechnological domains.

Based on the mode of energy input, cavitation can be categorized into acoustic cavitation, hydrodynamic cavitation, optic cavitation, and particle-induced cavitation [9]. Acoustic cavitation (AC), whose sonochemical effects were first discovered in 1927 [10], has been extensively studied and applied over the past two decades. The transition to industrial-scale implementation remains limited due to inherent challenges such as high energy consumption and low throughput [11]. In contrast, hydrodynamic cavitation (HC) induces cavitation via localized pressure drops within a flow field [12]. It offers superior scalability and cost-effectiveness, making it more suitable for large-scale industrial applications [13]. Owing to its numerous advantages, hydrodynamic cavitation has garnered significant attention in recent years across various fields, including disinfection [14], wastewater treatment [15], papermaking [16], and biodiesel production [17]. In particular, hydrodynamic cavitation has been extensively studied for its application in water sterilization and disinfection [18,19]. Among the various reactor designs, orifice plate, Venturi, and rotary-type configurations are currently the most widely utilized in practical applications [[20], [21], [22], [23]]. In the early stages of hydrodynamic cavitation research, conventional jet-based reactors such as orifice plates and Venturi devices were primarily employed for microbial inactivation, targeting organisms such as bacteria and algae [24]. However, recent advancements have demonstrated that novel rotational hydrodynamic cavitation reactors exhibit enhanced performance and efficiency in microbial disinfection applications [25]. Hou et al. [26] investigated the disinfection performance of a rotational hydrodynamic cavitation reactor, examining the influence of key operational parameters such as rotational speed, flow rate, and initial bacterial concentration on the inactivation efficiency. Xu et al. [27] utilized a jet-type hydrodynamic cavitation reactor to treat Microcystis aeruginosa, demonstrating its potential for harmful algal bloom control. The results indicated that the photosynthetic activity of Microcystis aeruginosa was significantly and immediately suppressed following hydrodynamic cavitation treatment. Filipic et al. [28] confirmed the effectiveness of hydrodynamic cavitation in inactivating eukaryotic viruses, attributing the primary inactivation mechanism to the intense mechanical forces generated during cavitation collapse. A comprehensive review of cavitation-induced disinfection is provided by Sun et al. [29]. While numerous studies have concentrated on evaluating bactericidal efficiency, the underlying disinfection mechanism remains insufficiently explored and warrants further investigation. Therefore, further investigation into the fundamental disinfection mechanisms of hydrodynamic cavitation remains essential. In addition, existing literature highlights that the structural parameters of rotational hydrodynamic cavitation reactors significantly impact disinfection efficacy [9,30]. Consequently, geometrical optimization of reactor components to enhance cavitation performance is a crucial area of ongoing research.

Extensive research on jet-type configurations has underscored the importance of structural parameters in cavitation performance [[31], [32], [33], [34]]. For example, Li et al. [35] studied the influence of Venturi geometry on cavitation behavior. They reported that a smaller outlet angle facilitates earlier cavitation inception and promotes the formation of a higher density of microbubbles. Simpson and Ranade [36] employed numerical simulations to examine the influence of key geometric parameters, including blowhole thickness, inlet edge sharpness, and wall angle, on cavitation dynamics. Similarly, Jia et al. [37] investigated the effect of throat-to-nozzle area ratio on cavitation flow characteristics within a jet-type hydrodynamic cavitation reactor. These findings collectively confirm that optimizing structural parameters is critical for enhancing cavitation performance. Notably, the rotary-type hydrodynamic cavitation reactor, an advanced design that has gained attention in recent years, has demonstrated superior cavitation intensity compared to conventional jet-type configurations [38]. While most studies on rotational HC reactors focus on performance evaluation [[39], [40], [41]], fewer address structural optimization of these systems [42,43]. However, advancing hydrodynamic cavitation technology requires deeper investigation into the role of reactor geometry. Therefore, systematic research on the influence of geometric design parameters in rotational hydrodynamic cavitation reactors on cavitation performance remains essential.

A review of existing literature reveals that neither the disinfection mechanism underlying cavitation nor the structural optimization of rotational hydrodynamic cavitation reactors have been sufficiently explored. To bridge these gaps and build on our previous work, this study first evaluates the disinfection efficiency of a Cylindrical Rotational Hydrodynamic Cavitation Reactor (CRHCR) and elucidates its underlying mechanism by examining micro-morphological changes in E. coli following cavitation treatment. Furthermore, the influence of key structural parameters of the cavitation generation unit within the CRHCR on cavitation performance was systematically analyzed. The underlying mechanisms by which geometric configurations induce cavitation were also investigated. These findings offer a solid theoretical foundation for the rational design and further development of next-generation CRHCR systems.

2. Experimental and numerical methods

2.1. Geometric parameters

Fig. 1 illustrates the schematic configuration of the cylindrical rotational hydrodynamic cavitation reactor (CRHCR), which primarily comprises a grooved rotor enclosed within a stator housing. The rotor measures 80 mm in diameter and 100 mm in length, serving as the central cavitation-inducing component of the system. The rotor surface is engineered with 20 uniformly distributed grooves, each measuring 6 mm in width and 3 mm in depth, and is rotationally driven by an electric motor. The stator housing incorporates a designated fluid inlet and outlet, enabling continuous flow through the system. The annular gap between the rotor and the stator housing functions as the active cavitation chamber of the CRHCR. Additional structural and operational details of the reactor are available in our previous study [44]. Given that the stator and rotor are the primary elements responsible for cavitation generation, a systematic investigation of their geometric parameters is essential for optimizing cavitation performance and enhancing the overall efficiency of the CRHCR. Building upon our previously developed CRHCR design, this study systematically investigates the influence of key geometric parameters on cavitation performance. The parameters examined include groove shapes (non-interacting, rectangular, trapezoidal, and triangular), groove widths (w = 4, 5, 6, 7, and 8 mm), groove depths (d = 1.5, 2, 3, 4, and 5 mm), oblique tooth angles (θ = 60°, 64°, 68°, 72°, and 76°), groove inclination angles (β = 0°, 5°, 10°, 15°, and 20°), and the number of grooves (n = 12, 16, 20, 24, and 28). A detailed schematic of these geometric configurations is illustrated in Fig. 2.

Fig. 1.

Diagrammatic representation of the CRHCR assembly [44].

Fig. 2.

Geometric parameters of the cavitation generation unit in the CRHCR.

2.2. Experiment setup

Escherichia coli (E. coli), widely recognized as a global indicator organism for microbial water quality assessment, was selected to evaluate the disinfection efficacy of the CRHCR. Disinfection performance was assessed by determining the inactivation rate of E. coli (strain CGMCC1.234) through controlled experiments. The bacterial strain was sourced from Shanghai Angyu Biotechnology Co., Ltd. The bacterial culture was grown in Luria-Bertani (LB) nutrient broth and incubated at 37 °C for 24 h in a constant temperature incubator. To minimize the impact of osmotic pressure differences between the intracellular and extracellular environments on E. coli, a 0.01 M phosphate-buffered saline (PBS) solution was prepared and used as the working medium for the disinfection experiments. The PBS solution was prepared using analytical reagent (AR) grade chemicals, including potassium chloride (KCl, AR, CAS: 7447–40-7), sodium chloride (NaCl, AR, CAS: 7647–14-5), potassium dihydrogen phosphate (KH₂PO₄, AR, CAS: 7778–77-0), and disodium hydrogen phosphate dodecahydrate (Na₂HPO₄·12H₂O, AR, CAS:10039–32-4). All reagents were obtained from Sinopharm Chemical Reagent Co., Ltd.

A rotational hydrodynamic cavitation experimental setup was built, comprising a water tank, magnetic drive pump, CRHCR unit, rotor flowmeter, and associated connecting pipelines. The detailed specifications of the experimental system are available in our previous publication [45]. For each experimental run, 5 L of PBS solution was used as the working fluid, into which the pre-cultured E. coli suspension was introduced. Before each experimental run, the entire hydrodynamic cavitation system was sterilized using a 75 % ethanol solution to maintain aseptic conditions. Additionally, before initiating treatment with the CRHCR, the pump was run for 5 min to circulate the solution throughout the system, ensuring a homogeneous distribution of E. coli in the PBS solution. The CRHCR was then operated for 15 min, during which 1.5 ml samples were collected at 3-minute intervals to monitor the concentration of E. coli over time. The E. coli concentration for each experimental group was determined using the standard plate colony counting method. The final values reported represent the average of three independent trials. The disinfection efficiency of the CRHCR was evaluated using the disinfection rate, as shown in Eq. (1).

$$\mathrm{D}\mathrm{i}\mathrm{s}\mathrm{i}\mathrm{n}\mathrm{f}\mathrm{e}\mathrm{c}\mathrm{t}\mathrm{i}\mathrm{o}\mathrm{n}\mathrm{r}\mathrm{a}\mathrm{t}\mathrm{e}=\frac{C_0-C_i}{C_0}\times 100\%$$ (1)

where C_o and C_i represent the initial E. coli concentration and the E. coli concentration after treatment time i, respectively.

2.3. Numerical simulation

2.3.1. Governing equations

Cavitation is a classical phenomenon within multiphase flow systems. In this study, the Mixture multiphase flow model was employed to simulate cavitation flow. By incorporating a simplified governing equation and an efficiently coupled phase transition mechanism, the model effectively balances computational efficiency with simulation accuracy, making it well-suited for predicting cavitation behavior. The Mixture model serves as an ideal approach for capturing transient cavitation phenomena due to its computational efficiency and robustness. To ensure accurate prediction of fluid flow characteristics, the selection of an appropriate turbulence model is crucial for reliable numerical simulation [46]. Among the available approaches, the Reynolds-Averaged Navier-Stokes (RANS) method remains one of the most widely adopted frameworks for modeling turbulent flow fields [47]. However, the RANS approach exhibits limitations in capturing transient features and resolving complex turbulence structures. For cavitation simulations demanding high fidelity and detailed flow resolution, Large eddy simulation (LES) offers significant advantages [48]. With the continuous advancement in computational capabilities in recent years, LES has increasingly attracted attention from researchers as a more robust alternative for simulating unsteady cavitating flows [49]. In the subsequent simulations, the turbulence model is based on the LES approach. The corresponding governing equations are presented as follows (Eqs. 2–4).

Continuity equation:

$$\frac{\partial {\rho }_m}{\partial t}+\frac{\partial ({\rho }_m{\overline{u}}_j)}{\partial x_j}=0$$ (2)

Momentum equation:

$$\frac{\partial ({\rho }_m{\overline{u}}_i)}{\partial t}+\frac{\partial ({\rho }_m\overline{u_i{u}_j})}{\partial x_j}=-\frac{\partial \overline{p}}{\partial x_i}+\frac{\partial }{\partial x_j}(\mu \frac{\partial {\overline{u}}_i}{\partial x_j})-\frac{\partial {\tau }_{\mathit{ij}}}{\partial x_j}$$ (3)

$${\tau }_{\mathit{ij}}=\rho \overline{u_i{u}_j}-{\overline{u}}_i{\overline{u}}_i$$ (4)

Based on the Boussinesq hypothesis, the subgrid scale (${\tau }_{\mathit{ij}}$) stresses are modeled as follows (Eqs. (5), (6).

$${\tau }_{\mathit{ij}}-\frac{1}{3}{\tau }_{\mathit{kk}}{\delta }_{\mathit{ij}}=-2{\mu }_t{\overline{S}}_{\mathit{ij}}$$ (5)

$${\overline{S}}_{\mathit{ij}}=\frac{1}{2}(\frac{\partial {\overline{u}}_i}{\partial x_j}+\frac{\partial {\overline{u}}_i}{\partial x_i})$$ (6)

where ${\mu }_t$ represents the subgrid-scale (SGS) turbulent viscosity. According to the WALE-Adapting Local Eddy-Viscosity (WALE) model, ${\mu }_t$ is represented in Eq. (7) and its $L_s$, $S_{\mathit{ij}}^d$, ${\overline{g}}_{\mathit{ij}}$ are presented in Eqs. 8–10, respectively.

$${\mu }_t=\rho L_s^2\frac{{(S_{\mathit{ij}}^d{S}_{\mathit{ij}}^d)}^{3/2}}{{({\overline{S}}_{\mathit{ij}}{\overline{S}}_{\mathit{ij}})}^{5/2}+{(S_{\mathit{ij}}^d{S}_{\mathit{ij}}^d)}^{5/4}}$$ (7)

$$L_s=min(kd,C_w{V}^{1/3})$$ (8)

$$S_{\mathit{ij}}^d=\frac{1}{2}({\overline{g}}_{\mathit{ij}}^2+{\overline{g}}_{\mathit{ji}}^2)-\frac{1}{3}{\delta }_{\mathit{ij}}{\overline{g}}_{\mathit{kk}}^2$$ (9)

$${\overline{g}}_{\mathit{ij}}=\frac{\partial {\overline{u}}_i}{\partial x_j}$$ (10)

where $L_s$ denotes the mixing length for subgrid scales. The parameters $d$, $k$, $C_w$, and $V$, represent the distance to the nearest wall, von Karman’s constant, the WALE model constant, and the local grid scale (mesh volume), respectively.

The Schnerr Sauer cavitation model was employed to predict cavitation behavior [50,51]. The liquid–vapor phase change due to mass transfer was modeled as follows (Eqs. (11), (12).

$$R=\frac{{\rho }_v{\rho }_l}{\rho }\alpha (1-\alpha )\frac{3}{{\mathfrak{R}}_B}\sqrt{\frac{2}{3}\frac{(P_v-P)}{{\rho }_l}}$$ (11)

$${\mathfrak{R}}_B={(\frac{\alpha }{1-\alpha }\frac{3}{4\pi }\frac{1}{n})}^{\frac{1}{3}}$$ (12)

where R, ${\mathfrak{R}}_B$, and $P_v$ represent the mass transfer rate, bubble radius, and saturation vapor pressure, respectively.

2.3.2. Solver setup and boundary conditions

The numerical simulation was conducted using ANSYS Fluent 2020R2. For spatial discretization of the momentum equations, the Bounded Central Differencing scheme was applied. The volume fraction was computed using the QUICK (Quadratic Upstream Interpolation for Convective Kinematics) scheme. For the transient formulation, the Bounded Second Order Implicit was employed to ensure temporal accuracy and stability. Additionally, the pressure–velocity coupling was handled using the Coupled scheme to enhance convergence stability. The convergence criterion for all governing equations was set to a residual threshold of 10^-5, with a time step size of 0.0001 s. The Courant number is set at 0.5. Boundary conditions were specified as a velocity inlet and a pressure outlet at the respective fluid domain boundaries. The rotor rotation was simulated using Mesh motion to capture the dynamic interaction within the computational domain. A rotational speed of 3000 rpm was applied. The interface between the rotating and stationary region was defined as a sliding mesh interface to ensure accurate coupling of the flow across the moving and static zones. The saturated vapor pressure was set to 3540P a, and the bubble number density was defined as 10¹². Further details regarding the solver configuration and boundary conditions are available in our previous study [45] and are therefore not repeated here.

2.3.3. Mesh independence study and validation of the CRHCR simulation model

The fluid domain of the CRHCR was discretized using a polyhedral mesh. This study builds upon our previous work, employing the same grid resolution and mesh configuration as established in that study. Given the structural differences among the cavitation generation units, a grid independence study was conducted using maximum grid sizes of 0.8 mm, 0.7 mm, 0.6 mm, and 0.5 mm to ensure the accuracy and reliability of the simulation results. The results indicated that a maximum grid size of 0.7 mm provided satisfactory accuracy and stability in the simulation. Consequently, a grid size of 0.7 mm was selected for all subsequent calculations. The detailed results of the mesh independence study are available in our previous work [45].

To validate the accuracy of the numerical model, the simulated cavitation cloud morphology and cavity length within the grooves were compared with experimental observations, as shown in Fig. 3 [45]. The left image in Fig. 3(a) illustrates the experimental flow field captured using a high-speed camera. The right image of Fig. 3(a) displays the iso-surface of the vapor phase volume fraction, as obtained from the numerical simulation. A comparison across different time points reveals a strong agreement between the experimental and simulated cavitation cloud morphologies within the groove, demonstrating the model's capability to capture the transient cavitation behavior accurately. In addition, the cavity lengths within the grooves obtained from both experimental observations and numerical simulations were compared, as shown in Fig. 3(b). The results exhibit good agreement between the two approaches, further confirming the reliability and predictive accuracy of the numerical model employed in this study.

Fig. 3.

(a) Temporal evolution of cavitation within a single groove of the non-interaction CRHCR: Left: experimental flow field captured via high-speed imaging; Right: numerical simulation showing iso-surface of vapor phase volume fraction (α = 0.2). (b) Comparison of cavitation cloud length within rotor grooves obtained from experimental observations and numerical simulations [45].

3. Results and discussion

3.1. Disinfection performance of CRHCR and bacterial regrowth behavior over extended storage

To evaluate the disinfection performance of the CRHCR, experimental tests were conducted using hydrodynamic cavitation for the inactivation of E. coli. Fig. 4(a) presents the bacterial disinfection rates after 15 min of treatment at various rotational speeds of the cavitation device. As shown in the figure, the disinfection rate increases progressively with rising rotational speed. The maximum disinfection efficiency of 90.35 % was achieved at a rotational speed of 4400 rpm, demonstrating the strong disinfection capability of the CRHCR. Table 1 presents a comparative analysis of the disinfection performance between the CRHCR and various previously reported cavitation reactors. The results demonstrate that the rotational hydrodynamic cavitation reactor achieves significantly higher disinfection rates within shorter treatment times compared to conventional jet-type hydrodynamic cavitation reactors. Furthermore, when compared with other rotational hydrodynamic cavitation reactors, the CRHCR exhibits superior disinfection performance.

Fig. 4.

Disinfection performance of the CRHCR: (a) Effect of rotational speed on E. coli disinfection rate with a treatment time of 15 min; (b) Bacterial regrowth rates following hydrodynamic cavitation treatment and 30-day storage.

Table 1.

Comparison of disinfection performance of hydrodynamic cavitation reactors.

Devices	Bacterial species	Treatment volume (L)	Treatment time (min)	Disinfection rate (%)	Ref.
Orifice	E. coil	50	120	32.7	Arrojo et al. (2008)[14]
Orifice	E. coil	21	120	85.3	Burzio et al. (2020)[52]
Venturi	E. coil	4	120	75.4	Sarc et al. (2018)[4]
Rotor-stator	E. coil	15	15	74.29	Hou et al. (2023)[26]
Rotor-rotor	E. coil	5	15	89.32	Xue et al. (2025)[53]
Present work	E. coil	5	15	90	−

In addition to this, a further investigation was conducted to explore a related aspect of the cavitation disinfection process. To investigate the bacterial regrowth behavior following cavitation treatment, the treated solution samples were stored for 30 days and continuously monitored for changes in bacterial concentration. All samples were maintained on a sterile operating platform throughout the monitoring period to eliminate potential interference from external environmental factors. The temporal variation in bacterial concentration over the 30 days is presented in Fig. 4(b). As shown, the bacterial concentration remained relatively stable, with only minor fluctuations observed over time. The figure also illustrates the bacterial regrowth rates corresponding to different rotational speeds after 30 days of storage, providing insights into the sustained disinfection effectiveness of the CRHCR system. At rotational speeds of 3600 rpm and 4000 rpm, the bacterial regrowth rate after 30 days of storage remained relatively low, with the highest observed increase reaching only 5.88 % compared to the concentration immediately after treatment. In contrast, at 3800 rpm and 4400 rpm, a decreasing trend in bacterial concentration was observed over the 30 days, indicating a potential delayed inactivation effect or sustained antimicrobial impact induced by cavitation at higher rotational speeds. Therefore, the CRHCR treatment enables the aqueous solution to maintain a low bacterial concentration over an extended period, highlighting its strong potential for long-term storage applications following drinking water disinfection.

To further elucidate the mechanism underlying cavitation-induced disinfection, E. coli cells treated with the CRHCR were characterized using SEM and TEM to examine their micro-morphological features. SEM images of untreated E. coli (Fig. 5a and 5b) show intact, rod-shaped cells, both in clusters and as individual organisms, serving as a baseline. As observed in Fig. 5(a) and 5(b), the untreated E. coli cells exhibit an intact, rod-shaped morphology characteristic of healthy bacterial structures. After 15 min of cavitation at 4400 rpm, SEM images (Fig. 5c and 5d) reveal severe structural damage: cells appear twisted and deformed, with surface ruptures, collapse, and membrane disintegration. Many bacteria are fragmented and lack recognizable features, confirming that CRHCR-induced cavitation causes extensive mechanical disruption and effective inactivation.

Fig. 5.

Morphological characteristics of E. coli observed via Scanning Electron Microscopy (SEM): (a, b) Cells before hydrodynamic cavitation treatment; (c, d) Cells after hydrodynamic cavitation treatment at 4400 rpm for 15 min.

TEM micrographs (Fig. 6) further illustrate these effects. Untreated cells (Fig. 6a and 6b) display well-defined cell walls and a uniform cytoplasmic matrix, indicative of healthy morphology. The untreated E. coli maintained good morphological integrity. However, following cavitation treatment, the cells exhibited significant structural deformation, including ruptured cell walls and leakage of cytoplasmic contents, ultimately resulting in severe cellular damage and lysis. These observations suggest that the cavitation effect disrupts the cell envelope, leading to cytoplasmic efflux and subsequent cell death.

Fig. 6.

Morphological characteristics of E. coli observed via Transmission Electron Microscopy (TEM): (a, b) Cells before hydrodynamic cavitation treatment; (c, d) Cells after hydrodynamic cavitation treatment at 4400 rpm for 15 min.

In summary, the observed micro-morphological alterations confirm that cavitation treatment severely compromises the structural integrity of E. coli, leading to the loss of physiological activity and cellular viability. As reported in the literature [29], the collapse of cavitation bubbles generates intense microjets, strong mechanical shear forces, localized high-temperature hotspots, and highly reactive hydroxyl radicals. Our findings indicate that mechanical shear from microjets is the primary mechanism of bacterial inactivation, although transient hotspots and radicals may also contribute. Further studies are required to elucidate the extent of their involvement in the disinfection mechanism.

3.2. Effect of geometric parameters of the cavitation generation unit on cavitation performance in CRHCR

The cavitation-generation unit is the primary region where cavitation is initiated, and it plays a decisive role in the CRHCR's overall performance. To investigate the influence of various structural parameters of the cavitation generation unit, including shape, width, depth, oblique tooth angle, groove angle, and the number of grooves on cavitation performance, the vapor phase volume generated during cavitation, total chamber volume, and the resulting cavitation ratio of the CRHCR were systematically characterized, as illustrated in Fig. 7. The cavitation ratio is defined as follows (Eq. (13):

$$\mathrm{C}\mathrm{a}\mathrm{v}\mathrm{i}\mathrm{t}\mathrm{a}\mathrm{t}\mathrm{i}\mathrm{o}\mathrm{n}\mathrm{r}\mathrm{a}\mathrm{t}\mathrm{i}\mathrm{o}=\frac{V_{\mathit{vapor}}}{V_{\mathit{total}}}$$ (13)

where V_vapor and V_total represent the cavitation vapor phase volume and the total volume of the CRHCR chamber, respectively.

Fig. 7.

Vapor phase volume (V_vapor), total chamber volume (V_total), and resulting cavitation ratio of the CRHCR under varying geometric parameters of the cavitation generation unit.

Furthermore, to elucidate the underlying mechanisms through which different structural parameters affect cavitation behavior, we then analyzed how these parameters influence cavitation behavior by examining vapor-phase distributions, velocity fields, pressure maps, and cavitation patterns (Fig. 8, Fig. 9, Fig. 10, Fig. 11, Fig. 12, Fig. 13). Finally, we introduced a Cavitation Performance Index (CPI) to quantify the impact of each parameter on overall cavitation efficiency (Fig. 14). The CPI is defined as follows (Eq. (14):

$$\mathit{CPI}=\frac{V_{\mathit{vapor}}/V_{\mathit{total}}}{\mathrm{\Delta }p/p_v}$$ (14)

where Δp denotes the pressure drop across the cavitation generation unit, and p_v is the saturated vapor pressure of the working fluid.

Fig. 8.

Effect of groove shape in the cavitation generation unit on cavitation characteristics. (a-1)–(d-1) Vapor phase volume fraction and velocity vector distribution; (a-2)–(d-2) Pressure distribution with cavitation patterns visualized as iso-surfaces (vapor phase volume fraction = 0.2) rendered by pressure on the rotor surface; (a-3)–(d-3) Vapor phase volume fraction and pressure distribution along the middle line of the groove.

Fig. 9.

Effect of groove width in the cavitation generation unit on cavitation characteristics. (a-1)–(e-1) Vapor phase volume fraction and velocity vector distribution; (a-2)–(e-2) Pressure distribution with cavitation patterns visualized as iso-surfaces (vapor phase volume fraction = 0.2) rendered by pressure on the rotor surface; (a-3)–(e-3) Vapor phase volume fraction and pressure distribution along the middle line of the groove.

Fig. 10.

Effect of groove depth in the cavitation generation unit on cavitation characteristics. (a-1) –(e-1) Vapor phase volume fraction and velocity vector distribution; (a-2)–(e-2) Pressure distribution with cavitation patterns visualized as iso-surfaces (vapor phase volume fraction = 0.2) rendered by pressure on the rotor surface; (a-3)–(e-3) Vapor phase volume fraction and pressure distribution along the middle line of the groove.

Fig. 11.

Effect of oblique tooth angle in the cavitation generation unit on cavitation characteristics. (a-1)–(e-1) Vapor phase volume fraction and velocity vector distribution; (a-2)–(e-2) Pressure distribution with cavitation patterns visualized as iso-surfaces (vapor phase volume fraction = 0.2) rendered by pressure on the rotor surface; (a-3)–(e-3) Vapor phase volume fraction and pressure distribution along the middle line of the groove.

Fig. 12.

Effect of groove inclination angle in the cavitation generation unit on cavitation characteristics. (a-1)–(e-1) Vapor phase volume fraction and velocity vector distribution; (a-2)–(e-2) Pressure distribution with cavitation patterns visualized as iso-surfaces (vapor phase volume fraction = 0.2) rendered by pressure on the rotor surface; (a-3)–(e-3) Vapor phase volume fraction and pressure distribution along the middle line of the groove.

Fig. 13.

Effect of the number of grooves in the cavitation generation unit on cavitation characteristics. (a-1)–(e-1) Vapor phase volume fraction and velocity vector distribution; (a-2)–(e-2) Pressure distribution with cavitation patterns visualized as iso-surfaces (vapor phase volume fraction = 0.2) rendered by pressure on the rotor surface; (a-3)–(e-3) Vapor phase volume fraction and pressure distribution along the middle line of the groove.

Fig. 14.

Cavitation performance index (CPI) and pressure drop in CRHCRs with varying geometric parameters.

3.2.1. Effect of groove shape on cavitation performance of the CRHCR

It is evident that the shape of the cavitation generation unit significantly influences the cavitation performance of the CRHCR. Previous studies have reported various geometric configurations of cavitation-inducing structures, including rectangular grooves in hydrodynamic cavitation reactors [9,54,55] and conical dimples [56], among others. These findings collectively demonstrate that the shape of the cavitation generation unit is a critical determinant of the device’s cavitation performance. Consequently, selecting an optimal geometry is essential for maximizing cavitation efficiency within the CRHCR system. In this work, we compared four stator-rotor groove configurations: non-interaction, rectangular, trapezoidal, and triangular. As shown in Fig. 7(a), under identical operating conditions, the non-interaction type failed to induce cavitation, resulting in the weakest cavitation intensity among the four configurations. In contrast, the interacting groove designs significantly enhanced cavitation performance. A detailed investigation into the cavitation mechanisms of both interaction and non-interaction CRHCR configurations is available in our previous work [45]. Among the CRHCR configurations with rectangular, trapezoidal, and triangular grooves, the rectangular groove exhibited the largest cavitation vapor phase volume (V_vapor), while the triangular groove produced the smallest. A similar trend was observed in the cavitation ratio, indicating that groove geometry significantly influences cavitation intensity. Therefore, these results indicate that the CRHCR equipped with rectangular grooves exhibits the highest cavitation intensity among the tested configurations.

To gain deeper insight into the influence of groove geometry on cavitation behavior, Fig. 8 illustrates the vapor phase volume fraction and corresponding pressure distribution. Fig. 8(a-1) through 8 (d-1) illustrate the vapor phase volume fraction, which characterizes the spatial distribution of cavitation within the groove cross-sections of the CRHCR. The arrows in the figure indicate the velocity vectors, providing insight into the local flow dynamics. Fig. 8(a-2) through 8(d-2) present the pressure distribution across the corresponding cross-sections. The cavitation morphology on the rotor surface is visualized using iso-surfaces defined by a vapor phase volume fraction of 0.2, with pressure values applied for surface rendering to enhance spatial interpretation of cavitation intensity. This visualization simultaneously reveals the cavitation morphology at the cavitation generation unit and the corresponding pressure field in the same region. Fig. 8(a-3) through 8(d-3) display the distribution of phase volume fraction and pressure along the centerline of each groove. In the non-interaction CRHCR, cavitation was not generated, and no low-pressure zones were observed, indicating the poorest cavitation performance, as shown in Fig. 8(d). In contrast, the interaction-type grooves all generate cavitation, with the rectangular profile (Fig. 8 (a-1) to (a-3)) producing the most extensive cavitation zones. This superior performance arises because rectangular grooves maximize internal volume, whereas triangular grooves—the smallest in volume—yield the least cavitation, and trapezoidal grooves lie in between. This trend is consistent with the total chamber volume (V_total) of the CRHCR configurations, as illustrated in Fig. 7(a). Attached cavitation also forms within the grooves, and its development depends strongly on groove shape.

Among the configurations, the triangular groove exhibited the smallest cavitation generation area, likely due to the absence of a horizontal surface, which limits the formation and growth of attached cavitation. More importantly, the pressure distribution in Fig. 8(a-2) to 8(c-2) indicates that, compared to the trapezoidal and triangular grooves, the rectangular groove's perpendicular walls create a larger flow obstruction, causing a sharper velocity drop and a broader low-pressure region that promotes cavitation. This further explains why the rectangular groove exhibits the largest cavitation region. In contrast, the trapezoidal and triangular grooves feature inclined side walls at their front faces, which alter the flow impact dynamics. Compared to the perpendicular walls of the rectangular groove, these inclined surfaces reduce the abruptness of pressure drop upon fluid impingement, thereby reducing the resulting cavitation intensity. Moreover, centerline plots of vapor fraction and pressure (Fig. 8(a-3) to 8(c-3)) further confirm that rectangular grooves maintain higher vapor-phase fractions and lower pressures over longer distances than the other shapes. Together, these qualitative and quantitative observations establish that rectangular grooves deliver the highest cavitation intensity among the configurations tested. Given the higher cavitation intensity associated with rectangular grooves, all subsequent investigations into the structural parameters of various cavitation generation units are conducted using the rectangular groove configuration as the baseline.

3.2.2. Effect of groove width on cavitation intensity in CRHCR

Based on the above analysis, it is evident that the groove within the cavitation generation unit plays a critical role in cavitation formation. Therefore, the following section investigates the effect of groove width on the cavitation performance of the CRHCR. As shown in Fig. 7(b), both the vapor phase volume and cavitation ratio increase with increasing groove width, indicating that CRHCRs with wider grooves exhibit higher cavitation intensity. Notably, the cavitation ratio increases sharply when the groove width reaches 6 mm, with the cavitation intensity more than doubling compared to that at a width of 5 mm. This suggests that a groove width of 6 mm represents a critical threshold influencing the cavitation performance of the CRHCR.

To evaluate the influence of groove width on the cavitation performance of the CRHCR, Fig. 9 presents the vapor phase volume fraction distribution and pressure distribution across the cross-section of cavitation generation units with varying groove widths. As the groove width increases from 4 mm to 8 mm, the cavitation cloud within the rotor groove expands progressively, indicating enhanced cavitation activity. In the stator groove region, cavitation is nearly absent at smaller groove widths (Fig. 9(a-1)). However, when the groove width is sufficiently large, a cavitation distribution similar to that observed in the rotor groove emerges in the stator region, as illustrated in Fig. 9(e-1). Wider rotor grooves create a larger disturbance zone and offer more surface area for vapor accumulation, facilitating the formation and growth of attached cavitation clouds. Simultaneously, the corresponding pressure distribution in Fig. 9(a-2) to 9(e-2) shows that increasing the groove width significantly enlarges the low-pressure region within the rotor groove. This effect is likely due to the wider groove allowing a greater volume of fluid to impinge on the front wall, thereby enhancing the localized pressure drop and promoting cavitation formation.

Moreover, larger groove widths lead to more pronounced cavitation in the stator groove region. Wider rotor grooves generate stronger vortices during rotation, which induce intensified vortex cavitation in the stator chamber (Fig. 9(b-2) to 9(e-2)). The vapor-phase iso-surfaces in Fig. 9(a-2) to 9(e-2) further illustrate that, within the CRHCR chamber, cavitation clouds at smaller widths are markedly less developed than those at larger widths. These observations confirm that increasing groove width enhances overall cavitation intensity. This conclusion aligns with previous studies: Zhang et al. [9] reported significant cavitation performance gains with larger radial grooves in a rotational HCR, and Sun et al. [57] showed that increasing circular hole diameter in the cavitation unit improves cavitation. Based on the above analyses, it can be concluded that increasing the lateral area of the cavitation generation unit effectively enhances cavitation intensity. This finding also highlights the superior potential of the rotational hydrodynamic cavitation reactor for scale-up and industrial applications.

3.2.3. Effect of groove depth on cavitation performance in CRHCR: Identification of an optimal depth range

Fig. 10 illustrates the influence of groove depth on cavitation distribution and the associated pressure field. As the groove depth increases from 1.5 mm to 3.0 mm, a progressive enlargement in the cavitation cloud area within the groove of the cavitation generation unit is observed, as evidenced in Fig. 10(a-1) to 10(c-1). Beyond 3.0 mm, however, the cloud no longer expands proportionally with groove volume (Fig. 10(d-1) and 10(e-1)), indicating a non-monotonic trend. Fig. 7(c) confirms this behavior: both the vapor-phase volume and the cavitation ratio rise with increasing depth up to an optimal point, then decline despite the total chamber volume continuing to increase. To explain this, Fig. 10(a-2) through 10(e-2) present pressure maps and vapor-phase iso-surfaces (volume fraction = 0.2). The iso-surfaces of vapor phase volume fraction serve to characterize the spatial morphology of cavitation on the rotor surface. The results reveal that both insufficient and excessive groove depths negatively impact cavitation intensity, indicating the existence of an optimal depth range. These observations align with Sun et al. [57], who reported that increasing hole height in circular cavitation units exhibits the same non-linear effect. Shallow grooves limit fluid expansion, producing thinner cavitation clouds, while overly deep grooves confine cavitation near the groove bottom and reduce fluid impingement on the walls. This reduced impingement weakens mechanical shear and suppresses low-pressure zone formation, impairing bubble inception. Additionally, the cavitation morphology observed through the iso-surfaces of vapor phase volume fraction in Fig. 10(a-2) to 10(e-2) reveal that at smaller groove depths, the limited chamber space restricts fluid expansion, thereby hindering the formation and development of extensive cavitation clouds. For deeper grooves, although the cavitation generation unit offers a large volume for potential cavitation development, the actual cavitation cloud generated is relatively limited in size. This indicates that an optimal groove depth is essential to maximize cavitation intensity in the CRHCR. Both insufficient and excessive groove depths adversely affect cavitation performance, highlighting the importance of precise geometric optimization for achieving high-efficiency operation.

3.2.4. Effect of oblique tooth angle on shear cavitation performance in CRHCR: Identification of an optimal angle

Within the cavitation generation unit of the CRHCR, cavitation primarily occurs in the groove region and the oblique tooth region. Based on the preceding analysis, it is evident that attached cavitation forms within the rotor grooves, while vortex cavitation predominantly arises in the stator grooves. In the oblique teeth region, high-velocity fluid shearing gives rise to shear cavitation. Since shear cavitation predominantly occurs in this region, investigating the influence of the oblique tooth's structural parameters on its formation is essential. Accordingly, Fig. 11 illustrates the vapor phase volume fraction and pressure distributions within the cavitation generation units of CRHCRs configured with different oblique tooth angles. Consistent with the previous figures, the morphology of the cavitation cloud was visualized using iso-surfaces of vapor phase volume fraction, with pressure mapping applied to the surface, as shown in Fig. 11(a-2) to 11(e-2). The oblique tooth angle is defined as the angle between the inclined surface and the vertical wall of the adjacent front groove. As shown in Fig. 7(d), an increase in the oblique tooth angle initially leads to a rise in both the cavitation vapor phase volume and cavitation ratio, followed by a subsequent decline. The maximum cavitation intensity is achieved at an oblique tooth angle of 68°, indicating the presence of an optimal angle for enhanced shear cavitation performance. At smaller angles, the impact area on the groove's front wall is reduced and the rotor–stator gap widens, both of which weaken shear forces and limit cavitation (Fig. 11(a-2) and 11(b-2)). A high-pressure zone was observed at the front end of the stator groove, which suppresses cavitation formation and consequently reduces cavitation intensity. When the angle exceeds 68°, the venturi-like channel between the oblique tooth and the stator narrows, shrinking the low-pressure zone and diminishing cavitation spread (Fig. 11(c-2) to 11(e-2)). The cavitation distribution region progressively reduces, resulting in a decline in cavitation intensity, as shown in Fig. 11(c-1) to 11(e-1). Notably, in the oblique tooth region, the cavitation generation zone decreases markedly, as shown in Fig. 11(d-1) and 11(e-1). A comprehensive analysis of the above findings confirms that the oblique tooth angle is a critical parameter influencing the cavitation strength of the CRHCR. Thus, both excessively small and large oblique-tooth angles impair cavitation intensity, confirming that 68° is the optimal angle for maximizing cavitation performance in the CRHCR.

3.2.5. Effect of groove inclination angle on cavitation intensity in CRHCR: Identification of an optimal tilt

Subsequently, the influence of the groove inclination angle within the cavitation generation unit on the cavitation intensity of the CRHCR was investigated. As shown in Fig. 7(e), the total internal chamber volume of the CRHCR decreases progressively with increasing groove inclination angle. The cavitation ratio exhibits a non-monotonic trend, initially increasing and then decreasing with the groove inclination angle. The maximum cavitation ratio is observed at an inclination angle of 5°, indicating that this configuration yields the most intense cavitation within the CRHCR.

To elucidate the underlying mechanism behind the enhanced cavitation intensity observed at a groove inclination angle of 5°, Fig. 12 presents the vapor phase volume fractions, velocity vector fields, and pressure distributions across the corresponding cross-sections of the cavitation generation unit. As illustrated in Fig. 12(a-1) to 12(e-1), the cavitation cloud distribution region expands as the groove inclination angle increases from 0° to 5°. However, with further increases beyond 5°, the cavitation cloud distribution region gradually reduces, indicating a peak in cavitation performance at this intermediate angle. The observed increase in cavitation intensity can be attributed to the enhanced vortex flow induced by smaller groove inclination angles within the stator groove region. This intensified swirling motion promotes the formation of a larger low-pressure zone, thereby facilitating cavitation generation, as shown in Fig. 12(a-2) and 12(b-2). However, when the groove inclination angle is increased to 15°, the excessive tilt alters the flow trajectory such that cavitation is primarily confined to the lower left corner of the groove. This localized cavitation occurs after the fluid impinges on the front end of the rotor groove wall during rotor rotation, thereby reducing the overall cavitation intensity. Moreover, the vortex cavitation region within the stator groove predominantly forms at the upper right corner of the groove. As shown in Fig. 12(d-2) and 12(e-2), the extent of the low-pressure zone in this region is significantly reduced, leading to a weakening of cavitation intensity. In the oblique tooth region, excessively large groove inclination angles reduce the spacing between the rotor’s oblique tooth and the adjacent non-grooved stator region. This restricted gap hinders the formation and development of cavitation, as evidenced in Fig. 12(c-2) to 12(e-2), and is a contributing factor to the observed decline in cavitation intensity. Therefore, the slight inclination of the groove is beneficial for enhancing the cavitation intensity of the CRHCR. This finding aligns with the study by Sun et al. [57], who investigated the effect of various tilt angles of dimples on the stator and rotor and similarly concluded that moderate tilting promotes stronger cavitation performance. In addition, Fu et al. [54] investigated the effect of the inclination angle of rectangular bumps on the inner wall surface of the CRHCR chamber. Their results indicated that a high inclination angle promotes the generation of cavitation bubbles, which contrasts with the findings of the present study. This discrepancy may stem from structural differences between the devices studied. The groove inclination angle within the cavitation generation unit plays a critical role in determining the cavitation intensity of the CRHCR. However, to date, limited research has been conducted on the specific influence of groove inclination in CRHCRs, indicating the need for further in-depth investigations in this area.

3.2.6. Effect of number of grooves on cavitation performance in CRHCR: Balancing shear and vortex cavitation

Since cavitation primarily occurs within the cavitation generation units of the CRHCR, the number of these units plays a critical role in determining the overall cavitation intensity and operational performance of the device. An increase in the number of cavitation-generating units is expected to enhance the cumulative cavitation effect, thereby improving the reactor's efficiency and effectiveness. For interaction-type rotational hydrodynamic cavitation reactors, the influence of the number of cavitation generation units on cavitation performance has been investigated in previous studies. Notably, Gostišai et al. [58], Sun et al. [[59], [60]], and Chipurici et al. [17] have reported that increasing the number of cavitation generation units can significantly enhance cavitation intensity and improve reactor efficiency, highlighting its importance as a key design parameter. In this study, the cavitation vapor phase volume and cavitation ratio of CRHCRs with varying numbers of grooves were analyzed, as presented in Fig. 7(f). The results indicate a positive correlation between the number of grooves. A greater number of stator and rotor grooves is beneficial for enhancing cavitation intensity. This enhancement is likely due to the increased frequency of fluid impingement on the front walls of the rotor grooves, which promotes the formation of more attached cavitation.

To gain deeper insight into how the number of grooves affects cavitation intensity in the CRHCR, Fig. 13 illustrates the vapor phase volume fraction and the corresponding pressure distribution across the cross-section of the cavitation generation unit. As shown in Fig. 13(a-1) and (a-2), a smaller number of grooves leads to a substantial increase in the inclined surface area within the oblique teeth region. Previous analysis indicates that shear cavitation originating in this area contributes significantly to overall cavitation generation. However, a reduction in the number of grooves enlarges the inclined surface area in the oblique tooth region, which in turn weakens the shear cavitation intensity. As a result, only limited cavitation cloud regions are observed, as illustrated in Fig. 13(a-1) and (b-1). This phenomenon can be attributed to the fluid shear induced by the enlarged inclined surface of the oblique tooth, which generates a localized low-pressure zone confined to the area adjacent to the inclined surface. In contrast, no such low-pressure region develops near the non-grooved section of the stator, as evident in Fig. 13(a-2). In regions where the inclined surface of the oblique teeth is smaller, the low-pressure zone extends across the entire oblique tooth region, as shown in Fig. 13(c-2). Furthermore, increasing the number of grooves markedly enhances the cavitation cloud area within both the rotor and stator groove regions, as illustrated in Fig. 13(a-1) to 13(e-1). More rotor grooves provide greater surface area for attached cavitation, while more stator grooves intensify vortex flows, supporting vortex cavitation onset. While increasing the number of grooves enhances the cavitation intensity of the CRHCR, it is constrained by the limited circumferential area of the rotor. Moreover, the oblique tooth region remains a critical component of the cavitation generation unit, playing a vital role in fluid shearing and contributing to overall cavitation performance. During rotor rotation, vortex flow is generated within the stator grooves. Therefore, when increasing the number of grooves to enhance the cavitation intensity of the CRHCR, it is also crucial to preserve a portion of the oblique tooth region. This ensures the coexistence of multiple cavitation mechanisms, such as shear and vortex cavitation, within the cavitation generation unit, thereby maximizing overall cavitation performance.

In summary, the geometrical configuration of the cavitation generation unit plays a key role in determining the cavitation intensity of the CRHCR. Among the groove geometries examined, rectangular grooves exhibit the highest cavitation intensity compared to their trapezoidal and triangular counterparts. Both groove width and the number of grooves show a positive correlation with the cavitation intensity of the CRHCR. In contrast, groove depth, oblique tooth angle, and groove inclination angle do not exhibit a direct or consistent positive correlation with cavitation intensity. Cavitation intensity was found to exhibit a non-linear relationship with geometric parameters, initially increasing and then decreasing as these parameters increased. The maximum cavitation intensity was achieved at a groove depth of 2 mm, an oblique tooth angle of 68°, and a groove inclination angle of 5°.

However, evaluating cavitation performance solely based on cavitation vapor phase volume and cavitation ratio may present certain limitations. To address this, a new dimensionless parameter has been proposed, the Cavitation Performance Index (CPI), which provides a more comprehensive assessment by incorporating both the cavitation ratio and the corresponding pressure drop.

The CPI and corresponding pressure drop for CRHCRs with varying geometric parameters of the cavitation generation unit are presented in Fig. 14. As shown in Fig. 14(a), among the four different groove configurations, the non-interaction CRHCR characterized by the absence of grooves in the stator exhibits the poorest cavitation performance. Although the non-interaction CRHCR exhibits a relatively low-pressure drop, it generates the weakest cavitation intensity, leading to the poorest overall cavitation performance. In contrast, the introduction of groove structures in the stator facilitates stator-rotor interaction, which significantly enhances cavitation intensity. This enhancement is also reflected in the increased cavitation ratio, as shown in Fig. 7(a). Among the rectangular, trapezoidal, and triangular groove designs, the rectangular configuration delivers the highest CPI with the lowest pressure drop, making it the most effective. For the CRHCR with rectangular grooves, the CPI increases progressively with the width of the groove, as shown in Fig. 14(b). This trend can be attributed to the reduction in flow resistance associated with wider grooves, which facilitates fluid entry and exit, resulting in a lower pressure drop and consequently reducing energy consumption. Therefore, the CRHCR with larger groove widths exhibits superior cavitation performance. The influence of groove depth, oblique tooth angle, and groove inclination angle on CPI is shown in Fig. 14(c)-14(e). In each case, CPI rises to an optimal value before declining, indicating ideal geometric conditions for maximum cavitation efficiency. The results show that the CRHCR achieves optimal cavitation performance at a groove depth of 3 mm, an oblique tooth angle of 68°, and a groove inclination angle of 5°. Finally, Fig. 14(f) illustrates the influence of the number of grooves on the CPI. It was found that, although the number of grooves does not exhibit a direct positive correlation with pressure drop, the CRHCR demonstrates a higher CPI with an increased number of grooves. This indicates that a greater number of grooves enhances the overall cavitation performance of the device, thereby improving the cavitation efficiency of the CRHCR.

4. Conclusion

In this study, the influence of key geometric parameters of the cavitation generation unit on both the disinfection efficiency and cavitation performance of the Cylindrical Rotational Hydrodynamic Cavitation Reactor (CRHCR) was systematically investigated through a combination of experimental and numerical approaches. The primary findings are summarized as follows. Experimental results using Escherichia coli (E. coli) confirmed the CRHCR's effective disinfection capability, achieving a disinfection rate of 90 % after 15 min of treatment at 4 400 rpm. Furthermore, the CRHCR-treated sample exhibited a low bacterial regrowth rate even after 30 days, indicating sustained antimicrobial effects. The disinfection mechanism attributed to cavitation was elucidated through micro-morphology analysis of E. coli, revealing significant structural damage induced by the cavitation effect. Scanning and transmission electron microscopy (SEM and TEM) showed that cavitation-induced microjets and mechanical shear forces led to cell wall rupture and cytoplasmic leakage—identified as the primary drivers of bacterial inactivation. The effects of key geometric parameters—including groove shape, groove width and depth, oblique tooth angle, groove inclination angle, and the number of grooves—on cavitation performance were investigated in detail. The results demonstrated that CRHCRs equipped with rectangular grooves outperform other groove geometries in terms of cavitation efficiency. Both groove width and groove count were found to be positively correlated with cavitation intensity; wider grooves and a greater number of grooves enhanced cavitation significantly. In contrast, for parameters such as groove depth, oblique tooth angle, and groove inclination angle, cavitation performance exhibited a non-linear trend, initially increasing and then decreasing with further increases in these geometric values. The CRHCR achieved optimal cavitation performance at a groove depth of 3 mm, an oblique tooth angle of 68°, and a groove inclination angle of 5°. These findings not only highlight the strong potential of CRHCR technology for drinking water disinfection but also offer valuable theoretical guidance for the design and optimization of next-generation rotational hydrodynamic cavitation reactors.

CRediT authorship contribution statement

Licheng Xue: Writing – original draft, Software, Investigation, Data curation. Zongrui Hao: Resources, Conceptualization. Sivakumar Manickam: Writing – review & editing, Validation. Gang Liu: Visualization. Haizeng Wang: Validation. Xun Sun: Supervision, Project administration. Haiyan Bie: Supervision, Project administration.

Declaration of competing interest

The authors declare the following financial interests/personal relationships which may be considered as potential competing interests: One of the authors is an Executive Editor of this journal, Sivakumar Manickam. If there are other authors, they declare that they have no known competing financial interests or personal relationships that could have appeared to influence the work reported in this paper.