Research on Wear Characteristics and Experiment on Internal Through-Passage Components for a New Type of Deep-Sea Mining Pump

Abstract

The conveyor electric pump for deep-sea mining is a key piece of equipment in deep-sea mineral transportation systems, as its flow capacity and wear resistance affect the reliability of the entire system. In this study, aimed at solving problems such as mining pumps being prone to clogging and wear, which could lead to performance degradation, a mining pump with wide flow paths and high performance was designed, and hydraulic performance tests were conducted on the new pump. The test results obtained were in good agreement with the numerical simulation results. Based on the reliability of the performance test results, using computational fluid dynamics (CFD) methods, and taking the wear model into consideration, a wear analysis was conducted on the internal through-passage components of the pump under different solid-phase particle parameters and operating conditions, and the average wear rate on the surface of the predominant through-passage components was calculated. The results showed that the hydraulic performance of the newly designed pump met the design requirements, with different particle parameters and operating conditions causing different degrees of wear on the through-passage components. The wear test was carried out with a test pump, and the comparison between the test results and the numerical calculation results showed that the numerical calculation of the wear of the deep-sea mining pump was accurate.

1. Introduction

With the acceleration of globalization, the demand for raw materials has increased, and land mineral resources have continued to be depleted, prompting people to look for alternatives in the oceans. As the deep sea contains a large volume of mineral resources—such as polymetallic nodules, cobalt-rich crusts, and polymetallic sulfides—more and more countries have begun to pay attention to the deep-sea field [1]. As one of the core pieces of equipment in deep-sea mineral transport systems, the mining pump is used to transport coarse-grained minerals. It needs to have multiple features—including good flow capacity, wear resistance, corrosion resistance, high pump head, and high efficiency—to avoid pump failure and improve the reliability of transport systems [2,3].

In 1978, the Ocean Management Company (OMI) consortium used an electric lifting pump developed in Germany to conduct a deep-sea polymetallic nodule mining test, but it was terminated due to the severe wear of the pump [4]. In 2002, after multiple design and testing processes on mining pumps, China also realized that the problem of mining pump wear needed to be solved [5]. Clark’s research found that the erosion of sediment particles depended primarily on five key factors—that is, the slurry concentration, flow velocity, impact angle, particle size, and particle shape [6]. Dong et al. studied the surface erosion problem of the centrifugal slurry pump’s through-passage components caused by coarse particles by calculating the maximum Von Mises stress and proved that the particle diameter, particle shape, particle impact angle, and medium flow velocity could all cause wear of the internal pump components [7]. Tarodiya’s research also mentioned that the characteristics of solid-phase particles, flow characteristics, and slurry concentration could affect the pump wear [8].

The pump wear caused by particles is complicated, and the wear caused by different solid-phase particle parameters and operating conditions is different. Noon et al. studied the wear of the pump casing and the pump’s through-passage components caused by the impact velocity, slurry concentration, and the diameter of slurry particles under slurry temperatures of 70 °C and 90 °C through simulations. The results showed that when the particle concentration was higher than 20%, the wear of the pump wall surface would slowly increase [9]. Wang et al. studied the wear of double-suction centrifugal pumps caused by sludge concentration and particle size, the results showing that owing to the presence of sludge, the pump head and shaft power were lower compared with pumping pure water, decreasing as the sludge concentration and sludge particle size increased [10]. Huang et al. analyzed the impeller wear caused by various sediment types and particle trajectories at the inlet, and they found that the wear was concentrated primarily on the rear cover and working surface [11]. Pagalthivarthi et al. analyzed the wear of the slurry pump under conditions of solid–liquid two-phase flow based on the characteristics of a slurry pump conveying low solid-phase particle concentrations, providing a reference for pump wear under low-concentration conditions [12]. In addition to causing wear, the particles can also affect the pump’s head and efficiency performances [13,14].

Based on the importance and complexity of the pump wear caused by particles, many researchers have conducted experimental studies. Walker et al. conducted comparative analysis between the actual wear of a mud pump and a laboratory wear test, proposing that it would be necessary to further analyze the internal wear characteristics of the mud pump under specific working conditions—such as medium flow velocity, concentration, particle size distribution, and particle shape—to better understand the pump wear problem [15]. Serrano et al. analyzed the impeller wear of a water pump caused by the concentration of sediments in the Acre River, Brazil, using the ball-crater micro-scale abrasive wear test to study the abrasiveness of the sediments. The results showed that the impeller wear exhibited a linear increase as the concentration of the sediments increased [16]. Sugiyama et al. predicted the wear depth of an impeller based on numerical simulations as well as experimental data [17]. Peng et al. studied the wear problem of heavy-duty slurry pumps. Their analysis showed that when the slurry flow was small, the flow of the pump was unstable and backflow was obvious; with an increase in particle concentration, the flow path resistance increased, the backflow increased, and the wear of the local wall surface inside the pump was intensified. These analytical results have been verified experimentally [18]. The centrifugal pump impeller wear test showed that particle erosion was the main cause of impeller damage [19,20]. Some researchers have also studied the motion characteristics of solid-phase particles in centrifugal pumps from different angles through experiments and analyzed the impact of slurry concentration and particle size on the head and efficiency of slurry pumps [21].

It is extremely important to choose a suitable wear model to analyze pump wear caused by particles. After analyzing and summarizing a large volume of literature, Meng et al. found that the particle erosion wear model involved a total of 33 variables, with each model comprising five different variables on average [22]. However, establishing a suitable and reliable wear prediction model can be complicated. Bross et al. used a new model to predict the impact of different impeller design parameters on pump wear, calculated the velocity field on the suction side of the impeller and the seal gap, predicted the local wear intensity of the pump based on the acting velocity and fluid material characteristics, and conducted a comparative analysis between the computational data and actual wear data [23]. Based on the two-phase particle flow model, Liao et al. analyzed the wear of various sand particle sizes on the surface of pump blades [24]. Wang et al. used a particle model to analyze the internal flow field characteristics of a centrifugal pump conveying solid–liquid two-phase flow [25]. Several researchers have also used particle and heterogeneous models to analyze the flow characteristics around the rear blades of various slurry pump shapes, thereby predicting the wear characteristics of the pumps [26]. In analyzing the wear characteristics of a pump conveying solid–liquid two-phase flow, the Eulerian–Eulerian multiphase flow model has often been used, achieving good results [27,28,29]. Tarodiya et al. analyzed the impact of changes in the particle size distribution on the performance of the centrifugal slurry pump based on the particle Euler–Euler model, making experimental comparisons. The results showed that changes in particle size distribution significantly affected the performance characteristics of particles in the impeller and flow path. With an increase in fine-grained particles, the particle pressure strength, maximum particle viscosity, and head loss all decrease [30].

Since there are many factors that affect the wear performance of mining pumps, this article examines the wear of the internal through-passage components of a mining pump under different solid-phase particle parameters and operating conditions based on the self-designed mining pump.

2. Methods: Hydraulic Performance Test and Verification Analysis

2.1. Mining Pump Structure

The structure of the deep-sea mining pump is composed of hydraulic components, such as the impeller, guide vane, suction connection, water outlet connection, water guide jacket, and water inlet; motor motive components; and mechanical components, such as couplings and seals [31]. The slurry flows in from the suction connection section of the conveyor electric pump, then passes through the water guide jacket, the water inlet section, and the impeller guide vanes of the pump at all levels, before flowing out from the water outlet guide casing. The main through-passage components of the mining pump are shown in Figure 1.

The mining pump impeller designed has the geometric parameters shown in

Table 1. Geometric parameters of model pump impeller.

	Parameter	Value
Impeller	Inlet diameter/mm	216.8
	Outlet diameter/mm	440
	No. of blades	3
	Blade wrap angle/°	140
	Outlet placement angle/°	28
	Outlet width/mm	61
	Outlet inclination angle/°	10

.

2.2. Hydraulic Performance Test

The performance testing system of the mining pump is composed of components to regulate the voltage of the water supply and for measurement, control, feeding, and lifting. Figure 2 shows a schematic diagram of the mining pump testing system. Figure 3 shows a physical map of the mining pump test bench.

Before the hydraulic performance test of the mining pump, a no-load test was required to check whether the motor was running normally and to determine whether the noise and vibration generated by the motor met the design requirements. The I₀ (A), P₀ (kW), R₀ (Ω), and temperature t (°C) were measured when n = 1450 r/min. The hydraulic performance test was performed after the no-load test had been completed.

2.3. Comparative Analysis of Performance Test Results and Numerical Simulations

Based on the performance requirements of mining pumps, the changes in the head, efficiency, and shaft power under different flow conditions were numerically simulated and analyzed. The flow–head curve, flow–efficiency curve, and flow–shaft power curve of the mining pump are shown in Figure 4a–c, respectively.

It can be seen from Figure 4a that both the test head and the numerically simulated head decrease with increasing flow and that the overall trends are in good agreement. At the rated flow point Q_d = 420 m³/h, the test head is 94.86 m, and the simulated head is 96.51 m, a relative error of just 1.7%. The maximum relative error occurs when the flow is Q = 100 m³/h, the relative error being 4.3%, which is within a reasonable range. It can be seen from Figure 4b that the test efficiency and the numerically simulated efficiency are also in good agreement. At the rated flow point Q_d = 420 m³/h, the relative error is 4.1%, while the maximum relative error occurring when the flow is Q = 100 m³/h is 7.3%, which is again within a reasonable range. Similarly, the overall trends of the tested shaft power and simulated shaft power are in good agreement. At the rated flow point Q_d = 420 m³/h, the relative error is −2.3%. Over the entire flow range, the maximum relative error is −2.8%, the relative error being within a reasonable range.

Based on the above analysis, the hydraulic performance of the newly designed mining pump certainly meets the design requirements. The head, efficiency, and shaft power obtained from the tests were consistent with the overall trends of head, efficiency, and shaft power obtained from the simulations, the errors being relatively small. This provides a solid basis for the analysis of the wear of mining pumps.

3. Theory: 3D Model and Numerical Computation Strategy

The main design parameter values of deep-sea mining pumps adopted in this study were a rated flow rate Q_d = 420 m³/h, a single-stage head H_d = 45 m, a rated efficiency η_d = 52%, and a rated rotational speed n = 1450 r/min.

3.1. Three-Dimensional (3D) Model

The main through-passage components in the mining pump include impellers, guide vanes, etc. Our analysis focuses primarily on the wear of the impeller and guide vanes of the pump caused by coarse particles. The main through-passage component model of the mining pump is shown in Figure 5, the overall structure model being shown in Figure 6.

In the numerical simulations, the full flow field computational domain includes the inlet section, the first-stage impeller, the first-stage guide vane, the second-stage impeller, the second-stage guide vane, and the outlet section, as shown in Figure 7. The entire computational area is divided with structured mesh. The ICEM application is used to generate six sets of meshes with different numbers before computation with an irrelevance test being performed. When the number of meshes is greater than 2.2 million, the rate of change of the pump head can be controlled to within 1%. Consequently, the final number of meshes in the computational domain was determined to be this value for subsequent research.

3.2. Numerical Computational Method

Ansys Fluent (18.0, Ansys, Pittsburgh, PA, USA, 2018) was used to perform steady numerical simulations of the computational domain of the mining pump. In the computational process, it is assumed that there is no energy and mass exchange between the particle phase and the fluid phase and that gravity is taken into account. The computation of the fluid domain adopts the SST k-ω turbulence model, the motion of the particle phase in the computational domain adopts the DPM model based on Euler–Lagrange coordinates, the discretization of the Navier–Stokes equation adopts the second-order upwind style, and the computational convergence precision is set to 10⁻⁴ in the numerical simulation.

Using the x-axis direction as an example, the equation for particle motion in the computational domain can be expressed as follows:

$$\frac{d{u}_p}{dt}=F_D\left(u-u_p\right)+\frac{g_x\left({\rho }_p-\rho \right)}{{\rho }_p}+F_V+F_P+F_x$$

(1)

where $F_D\left(u-u_p\right)$ is the drag force on a unit mass of solid-phase particles, as follows:

$$F_D=\frac{18\mu }{{\rho }_p{d}_p^2}\frac{C_D{R}_e}{24}$$

(2)

where u is the fluid velocity (m/s), $u_p$ is the particle velocity (m/s), μ is the hydrodynamic viscosity (N⋅s/m²), $\rho$ is the fluid density (kg/m³), ${\rho }_p$ is the particle density (kg/m³), $d_p$ is the particle diameter (mm), and $g_x$ is the acceleration of gravity in the x-axis direction (m/s²).

To calculate the motion trajectory of the solid-phase particles after colliding with the surface of the pump’s through-passage components, it is necessary to obtain the velocity direction and the magnitude of the velocity after colliding with the wall. As there is a certain degree of energy consumption after the particles collide with the surface of the through-passage components and rebound and velocity attenuation occurs, it is necessary to introduce an energy recovery coefficient related to the surface material properties of the through-passage components to express this consumption. In this paper, the random particle–wall collision model proposed by Tabakoff is used to calculate the energy recovery coefficient and can be expressed as follows [32]:

$$e_N=1.0-0.0211\theta +0.00228{\theta }^2-0.000000876{\theta }^3$$

(3)

$$e_T=0.953-0.000446\theta +0.00000648{\theta }^3$$

(4)

where $e_N$ is the normal recovery coefficient, $e_T$ is the tangential recovery coefficient, and $\theta$ is the collision angle between the particle and the surface of the through-passage component.

Since the particle volume concentration used in this study was less than 12%, the fluid-to-particle unidirectional coupling method was used in the numerical simulation to calculate the motion trajectory of the solid-phase particle. The interval for computation of the particle trajectory was 20 steps—that is, every 20 steps of calculation for the fluid phase was used as the benchmark, and a particle trajectory computation was performed on this basis, and so on.

3.3. Boundary Condition Setting

The mining pump inlet adopts the form of a velocity-inlet. The inlet velocity is calculated based on the inlet flow and the inlet through-passage cross-sectional area. At the same time, the solid-phase particles at the inlet face enter the computational domain axially, their entry velocity being the same as the fluid velocity—that is, there is no relative velocity. The outlet of the mining pump adopts the boundary condition in the form of outflow and defines the flow velocity weighting to be 1. All through-passage surfaces in the computational domain adopt non-slip boundary conditions, the wall roughness being defined as 0.5 based on the real blades of the model pump. The data transmission between the inlet section and the impeller and between the impeller and the spatial guide vane adopts the interface method. In the setting of the boundary conditions for solid-phase particles, the escape boundary condition is used for the pump inlet and outlet, and the reflect boundary condition is used for the surfaces of other through-passage components.

3.4. Selection of Wear Model

The expression of the wear prediction model used in the computation of the surface wear of the mining pump through-passage components can be expressed as shown in Equation (5) [33]:

$$R_{erosion}={\displaystyle \sum _{n=1}^{N_{particles}}\frac{m_{p}C(d_p)f(\theta )v^{b(v)}}{A_{face}}}$$

(5)

where:

• $R_{erosion}$: erosion rate;
• $C\left(d_p\right)$: particle size function of the solid-phase particles;
• $\theta$: impact angle between the particle and the wall;
• $f\left(\theta \right)$: function of the impact angle;
• $v$: relative velocity between the particle and the wall, m/s;
• $b\left(v\right)$: function of the relative velocity between the particle and the wall;
• $A_{face}$: unit surface area of the wall, mm²;
• $N$: the number of particles impacting the cell surface area;
• $m_p$: mass of a single particle;
• $d_p$: particle size;
• $n$: number of particles.

After combining the above formula and the particle motion trajectory for analysis, the wear on the surface of the pump’s various through-passage components caused by the movement of solid-phase particles in the computational domain can be quantitatively predicted. The impact angle function, $f\left(\theta \right)$, can be expressed using a piecewise polynomial as follows:

When $\theta$ ≤ 15°, then:

$$f\left(\theta \right)=b{\theta }^2+c\theta$$

(6)

When $\theta$ > 15°, then:

$$f\left(\theta \right)=x{\cos }^2\theta \sin \left(w\theta \right)+y{\sin }^2(\theta )+z$$

(7)

The values of the constants in the above formula are shown in

Table 2. Values of parameters in the impact angle equation.

b	c	x	y	w	z
−13.3	7.85	1.09	0.125	1	1

b, c, x, y, w, and z are empirical constants in the impact angle function equation.

.

4. Results: Wear Characteristic Analysis of Mining Pump

4.1. Impact of Solid-Phase Particle Parameters on Mining Pump Wear

The wear of the main mining pump through-passage components caused by solid-phase particle parameters was analyzed, and the solid-phase particle size, particle concentration, and particle density were selected for numerical computation. Based on Durand’s formula [34], the velocity of solid-phase particles in the entire flow field—that is, in the computation of solid–liquid two-phase flow—should usually be considered to ensure that the velocity of solid-phase particles is greater than the sedimentation rate of particles.

In this study, it was established—through calculations—that the inlet velocity of the particles was equal to 3.65 m/s under the design conditions. Consequently, it could be ensured that the velocity was greater than the critical velocity calculated using the Durand formula so that it would not be affected by the sedimentation of the particles.

4.1.1. The Impact of Different Particle Sizes

Based on the particle size of deep-sea nodule minerals, the following settings were selected in the numerical computation—the volume concentration of the conveyed solid phase was 7.5%, the flow was 420 m³/h, the pump revolution speed was 1450 rpm, and the particle sizes were 3, 6, 10, 14, 17, and 20 mm. Figure 8 and Figure 9 show the wear conditions of the impeller and guide vanes, respectively.

It can be seen from Figure 8 and Figure 9 that the area of wear damage of the impeller caused by solid-phase particles is mainly concentrated in the area near the blade inlet on the impeller front and rear covers. This is because after the particles enter the impeller axially, they continuously collide with the blade’s leading edge due to inertia, resulting in the area near the blade inlet being worn and damaged to varying degrees, with the local wear rate significantly increasing. As the particle size gradually increases, the area of wear damage on the surface of the front cover and the blade of the impeller caused by the particles’ changes, moving closer to the impeller outlet.

It can be seen from the wear distribution map of the first-stage guide-vane-edge area under different particle sizes that wear caused by smaller particles on the surface of the spatial guide vane edge is more concentrated and distributed in strips. As the particle size increases from 3 to 20 mm, the wear area increases, and the degree of wear becomes more serious.

To quantitatively analyze the wear on the surface of the through-passage components, the area-weighted average method was used to obtain the average wear rates on the surface of the four main through-passage components—including the first-stage impeller, the first-stage guide vane, the second-stage impeller, and the second-stage guide vane. The data obtained are shown in Figure 10.

It can be seen that under the same particle size, the wear rate on the surface of the spatial guide vane is higher than that of the impeller surface, except that the average wear rate of the second-stage guide vane decreases slightly when the particle size is 14 mm. Moreover, as the particle size increases, the wear rate of each part of the surface of the through-passage components increases accordingly, the degree of wear being positively correlated with the solid-phase particle size. Due to the larger mass of the particles themselves, the impact damage to the surface materials of the through-passage components increases correspondingly, making frequent collisions with the blade surface more likely—further increasing erosion and wear damage. When the particle size is 20 mm, the average wear rate on the surface of the first-stage impeller reaches 1.53 × 10⁻⁷ kg/s⋅m², while the average wear rate on the surface of the first-stage guide vane reaches 2.77 × 10⁻⁷ kg/s⋅m². This is because particles with larger diameters obtain greater kinetic energy in the blade exit area, causing wear damage in the form of a cut shape at the blade exit. The wear rate of this part is caused mainly by the tangential velocity component of the large-size particles [35].

4.1.2. The Impact of Particle Volume Concentration

In addition to particle size, the particle volume concentration has a significant impact on the particle motion trajectory and the wear law of through-passage components. In the numerical computation, the volume concentration of the conveyed solid phase was set to be 4%, 7.5%, and 11.5%; the flow was set to 420 m³/h; the pump revolution speed was set to 1450 rpm; and the particle size was set to 6 mm.

Figure 11 and Figure 12 show the wear distribution of the impeller surface and the spatial-guide-vane surface under different particle volume concentrations. It can be seen from the figures that the wear rate of either the blade or the spatial guide vane surface is in direct proportion to the volume concentration of particles. When the volume concentration of solid-phase particles increases from 4–11.5%, the wear area on the surface of the through-passage component increases, the wear rate of the same parts increasing accordingly. This is due to the increase in the particle concentration increasing the number of particles entering the pump inlet, resulting in an increase in the number of collisions between them and the blade surface, thereby leading to increased wear.

Figure 13 shows the changes in the wear rate of the two-stage impeller-guide-vane through-passage components in quantitative form.

It can be seen from Figure 13 that whether it is an impeller or a spatial guide vane, the wear rate increases with increasing particle volume concentrations. Under the range of particle volume concentrations of 4–11.5%, the average wear rate of the impeller surface increases from 4.29 × 10⁻⁸ kg/s⋅m² to 1 × 10⁻⁷ kg/s⋅m², and the average wear rate of the guide vane surface increases from 1.05 × 10⁻⁷ kg/s⋅m² to 3.37 × 10⁻⁷ kg/s⋅m². Under the same solid-phase particle volume concentration, the average wear rate of the guide vane surface is much higher than that of the impeller surface. Conversely, the kinetic energy of the particles significantly improves under the rotation of the impeller, as they are thrown out of the impeller at a higher speed before entering the guide vane. It can be seen from the wear computation formula that the wear rate on the surface of the through-passage components is directly proportional to the 1.73rd power of the particle velocity, so its increase is bound to aggravate wear. Conversely, from the geometric structure of the spatial guide vane, it can be seen that the radial size decreases along the axial direction, and the overall shape is tapered. This also increases the frequency of high-energy particles entering the spatial guide vane and making contact with the surface of the guide vane edge, thereby accelerating wear.

4.1.3. The Impact of Particle Density

To study the effect of particle density, the pump revolution speed was set to 1450 rpm, the flow was set to 420 m³/h, the conveyed particle size was set to 6 mm, the particle volume concentration was set to 7.5%, and the particle density was set to 1400 kg/m³, 1900 kg/m³, and 2800 kg/m³, respectively.

Figure 14 and Figure 15 show the wear distribution on the surface of the impeller and the spatial guide vane at the three densities. It can be seen from the figures that the area of wear damage on the surface of the impeller and the spatial guide vane caused by the solid-phase particles is mainly distributed on the surface of the front cover plate and the area near the blade inlet on the rear cover plate.

Since contact between particles of lower density and the surface of the through-passage components is based primarily on friction, the wear distribution on the surface of the impeller and the spatial guide vane is more uniform. Conversely, when the density increases further, the collision angle between the particles and the surface of the through-passage components increases as the area of high wear is more concentrated, which is dominated by impact damage.

Figure 16 shows the change law of the impact of solid-phase particles on the wear rate on the surface of the impeller and the spatial guide vane under three density characteristics.

It can be seen from Figure 16 that when the particle density is 1400 kg/m³, the surface of the impeller and the spatial guide vane have a higher degree of wear, among which the average wear rate of the first-stage guide vane reaches 2.46 × 10⁻⁷ kg/s⋅m². As the particle density increases, the wear of both the first- and second-stage guide vanes is reduced, while the wear of the first-stage and second-stage impellers reaches its lowest at 1900 kg/m³, increasing slightly when the density further increases to 2800 kg⋅m³—thus indicating that the model pump is more suitable for conveying solid-phase particles of density 1900 kg/m³.

4.2. Impact of Operating Conditions on Mining Pump Wear

4.2.1. The Impact of Different Flow Conditions

The three flow conditions were set to 0.68 Q_d, 1.0 Q_d, and 1.33 Q_d; the volume concentration of conveyed solid-phase particles was set to 7.5%; and the pump revolution speed was set to 1450 rpm.

From the wear distribution on the surface of the impeller and the spatial guide vane shown in Figure 17 and Figure 18, it can be seen that the surfaces of the impeller and the spatial guide vane have a larger wear damage area and a higher degree of wear under the small flow condition of 0.68 Q_d. Under the action of a turbulent flow field structure, the frequency of contact and collision between solid-phase particles and the surface of the through-passage components increases correspondingly, resulting in intensified wear damage on the surface of the through-passage components.

Figure 19 shows the change law of the average wear rate on the two-stage impeller-guide-vane surface under the three flow conditions.

It can be seen from Figure 19 that, as the flow increases, the degree of wear on the surface of the impeller decreases to varying degrees. The wear rate of the impeller surface decreases from a high of 1.22 × 10⁻⁷ kg/s⋅m² to 6.38 × 10⁻⁸ kg/s⋅m, the wear rate of the guide vane surface reaching the lowest value at the designed flow, and the lowest wear rate being located on the second-stage guide vane under 1.0 Q_d with an average wear rate of 1.77 × 10⁻⁷ kg/s⋅m².

4.2.2. Impact of Different Revolution Speed Conditions

The pump revolution speed was set to 900, 1450, and 1800 rpm; the flow was set to 420 m³/h; the volume concentration of conveyed solid-phase particles was set to 7.5%; and the particle size was set to 6 mm.

Figure 20 and Figure 21 show the wear distribution on the surface of the first-stage impeller and the spatial guide vane under the three revolution speed conditions, respectively. It can be seen that as the speed of the pump impeller increases, the wear rate on the surface of the impeller and the spatial guide vane is proportional to the pump speed where increases in speed aggravate the wear damage on the surface of the pump through-passage components.

Figure 22 shows the change law of average wear rate of the two-stage impeller-guide-vane surface under the three revolution speed conditions.

It can be seen from Figure 22 that, compared with the second-stage impeller and the second-stage guide vane, the increase in the pump revolution speed has a greater impact on the changes in the average wear rate on the surface of the first-stage impeller and the first-stage guide vane, with this being more prominent when the speed is changed from 1450 r/min to 1800 r/min.

5. Test: Wear Model Verification

Due to the long manufacturing process for this deep-sea mining pump, the conditions for the wear test are not yet available. In order to verify the effectiveness of the wear model used in this article, we made a test pump with a specific speed and a similar structure for wear tests.

5.1. Test Principle

The main design parameter values of the test pump adopted in this study were a rated flow rate Q_d = 20 m³/h, a single-stage head H_d = 15 m, a rated efficiency η_d = 64%, and a rated rotational speed n =2860 r/min.

The main work is the qualitative wear test. The test pump is coated with water-based paint and the wear of the solid particles on the water-based paint in the pump is observed before and after the test. The test was carried out with 3mm particles with a concentration of 7.5%. After running for two hours under rated conditions, the pump was disassembled to observe the wear position and compare the test results with the numerical analysis results.

The test system of this test pump is composed of a test pump, a water tank, a regulating valve, a pipeline, a data measurement system, etc. The principle of the test is shown in Figure 23.

The test pump is shown in Figure 24. The comparison of the main flow-passing parts in the pump before and after the test is shown in Figure 25.

The test site during the test is shown in Figure 26.

5.2. Comparative Analysis of Results

During the numerical calculation of the test pump, the wear model used is consistent with that of the deep-sea mining pump. The wear results of the test pump under rated conditions are compared with the numerical analysis results.

Figure 27 is a comparison of the wear position of the pump inlet position.

Figure 28 is a comparison of the wear results of the first stage guide vane. Figure 29 is a comparison of the wear results of the secondary guide vanes.

It can be seen from Figure 27, Figure 28 and Figure 29 that, under the wear and destruction of solid particles, the surface wear of the impeller blades is mainly concentrated near the blade inlet and the blade outlet. At the blade inlet, the solid particles enter the impeller inlet along the axial direction, and the speed direction changes from axial to radial under the rotation of the impeller. During this process, it will collide with the inlet edge of the blade at a large impact angle, which will cause serious wear and damage in this area. The wear near the blade outlet is mainly due to the increase in the kinetic energy of the particles under the impeller’s work and the impact at the impeller outlet. The tangential component of the velocity reaches the maximum at this point, which intensifies the wear of the blade exit edge. The wear form at this point is mainly abrasive wear. The result is consistent with the research results of Shen et al. on pump wear [36,37,38,39,40]. It should be noted that there are few research results on the abrasion of deep-sea mining pumps, especially the experimental results. This paper focuses on the abrasion characteristics and experiments of deep-sea mining pumps, while other researchers mainly focus on the abrasion characteristics and experiments of centrifugal pumps and slurry pumps.

The maximum wear of the guide vane blades appears near the inlet edge of the blade. At the same position, the surface wear rate of the first stage guide vane is higher than that of the secondary guide vane. After the rectification of the first stage impeller and the guide vane, it enters the second. The particle flow is more stable and the number of collisions with the surface of the guide vane is reduced, thereby effectively reducing the wear damage to the surface of the guide vane.

In general, the same wear model as the deep-sea mining pump is used to perform numerical calculations on the test pump. The numerical calculation results are consistent with the test results, indicating that the effectiveness of the wear model has been verified. This wear model can be used to perform numerical calculations on deep-sea mining pumps.

6. Conclusions

The deep-sea mining pump is among the key pieces of equipment in mining systems. Different particle sizes, particle volume concentrations, particle densities, and different flow and revolution speed conditions all cause different degrees of wear on the internal through-passage components of the mining pump.

Through a comparative analysis between the hydraulic performance test and numerical simulation, the head, efficiency, and shaft power obtained by the test are in good agreement with the overall trends of the head, efficiency, and shaft power obtained via numerical simulation. Accordingly, the newly designed wide flow path and high-performance mining pump met the design requirements.

Under the conditions of different solid-phase particle parameters, with the increase in particle size and particle volume concentration, the wear area of the blades and guide vanes continues to change, the wear area gradually increasing and the degree of wear continuing to be aggravated. The higher the particle density, the more concentrated the area of severe wear, which is dominated by impact damage.

Under different operating conditions, the wear damage of the through-passage components of the pump is aggravated with increasing pump speed, the wear damage area on the surface of the through-passage components being larger under lower flow conditions.

The average wear rate on the surface of the main through-passage components can be obtained by using the area-weighted average method. The larger the particle size, particle volume density, and flow, the greater the average wear rate. The new mining pump is more suitable for conveying coarse particles with a particle density of approximately 1900 kg/m³.

Numerical calculations and wear tests on test pumps have verified the effectiveness of the wear model. In future research, this wear model can be used to analyze the wear of deep-sea mining pumps, which can expand the research in this field. In the next step, we will also focus on research on the internal wear characteristics test and performance optimization of deep-sea mining pumps.