Effect of Electrical Load and Operating Conditions on the Hydraulic Performance of a 10 kW Pelton Turbine Micro Hydropower Plant

Abstract

Micro-hydroelectric power plants play a fundamental role in microgrid systems and rural electrification projects based on non-conventional renewable energies, where the stability of the electricity supply and load variability are critical factors for efficient operation. This work focuses on analyzing the impact of electrical load variation on the performance of a 10 kW micro hydroelectric power plant equipped with a Pelton turbine coupled to an electric generator. The main objective is to characterize the behavior of the turbine–generator system under different operating conditions, evaluating the hydraulic performance of the turbine, the electrical performance of the generator, and the overall performance of the micro power plant. Key variables such as flow rate, pressure, shaft speed, mechanical torque, current, and electrical voltage are monitored, considering the effect of electrical consumption on each of them. The experimental methodology includes tests at different electrical loads connected to the generator, using the spear system, which allows the flow rate in the injector to be modulated. The results indicate that reducing the flow rate using the spear increases the torque on the shaft, as well as the electrical current and voltage, for the same energy demand. Likewise, it is observed that the electrical efficiency of the generator remains stable for shaft speeds above 400 rpm, while the overall efficiency of the turbine–generator improves by up to 25% at this same speed. However, a voltage drop of more than 8% is recorded when the electrical power consumption increases from 3 kW to 9 kW, which demonstrates the sensitivity of the system to load variations. This work provides a comprehensive view of the dynamic behavior of micro-hydraulic power plants under realistic operating conditions, proposing an experimental methodology that can be applied to the design, optimization, and control of small-scale hydroelectric systems. These results provide novel experimental evidence on how electrical load variations affect the global performance of P -based micro hydropower systems.

1. Introduction

Hydropower is one of the oldest renewable-energy technologies, with documented use since 1737 and the first Chilean installation dating back to 1897. The Pelton turbine—patented by Lester Pelton in 1889—remains the preferred option for high-head, low-flow applications because it converts hydraulic head directly into shaft power with minimal internal losses [1].

1.1. Micro-Hydro for Rural Electrification

Well-designed micro- and pico-hydro schemes can supply remote communities at a lower life-cycle cost than diesel, wind, or PV, provided that plant size matches demand and technical risks are mitigated. Williams and Simpson quantified these cost advantages in village-scale pico-hydro projects [2]. Long-term field studies in Bolivia further showed that success depends on continuous learning and strong community ownership [3]. In Peru, Domenech et al. combined wind, PV and micro-hydro microgrids to electrify 58 households, tailoring each technology to local resource availability and social constraints [4]. Kyriakarakos later demonstrated that integrating pumped-hydro storage into islanded microgrids can lower total costs and raise renewable penetration relative to battery-only configurations [5]. On a global scale, Uddin et al. reviewed more than 1800 microgrids and identified hybrid micro-hydro systems as a cornerstone for resilient rural electrification [6]. These examples underline the need for reliable performance data to guide plant sizing and operation; such information becomes critical when micro-hydro units operate as last-in-line sources within isolated microgrids [7].

1.2. Efficiency of Pelton Turbines

Extensive laboratory work confirms the high hydraulic efficiency of Pelton runners. Williams and Simpson reported values close to 80% for properly tuned pico-hydro units [2], whereas Cobb and Sharp showed that modest bucket-geometry refinements can add a further 10% [8]. Gupta et al. achieved up to 90% efficiency by optimising nozzle configurations [9], and Coleman and Steele provided a rigorous uncertainty framework to validate such measurements [10]. At prototype scale, Xiao and Wang combined CFD and field tests to demonstrate stable efficiencies above 75% [11]; Zeng et al. later improved those figures by refining the nozzle–pipeline layout [12]. Computational Fluid Dynamics (CFD) has become an essential tool in hydraulic turbine design and optimisation. By enabling detailed analysis of flow patterns, pressure distribution, and turbulence effects within the turbine and associated piping, CFD allows researchers to assess the influence of design and operating parameters on efficiency without the need for costly and time-consuming experiments. Karpenko [13] demonstrated the application of CFD to assess energy losses in hydraulic pipelines. Tiwari et al. [14] reviewed CFD methods in the context of hydraulic turbine design, highlighting their role in performance improvement. Zhu et al. [15] compared various turbulence models in the prediction of flow through Pelton turbines, showing how numerical predictions can guide experimental validations. Validated CFD approaches, such as those presented by Kim et al. [16], provide reliable estimations of hydraulic behavior under different operating conditions and geometries, and have become standard practice in the development of high-performance turbine systems. Brekke highlighted cavitation control as essential for maintaining both performance and safety [17], while Lubis et al. proved that low-cost spoon-based designs can still exceed 70% efficiency in rural settings [18].

1.3. Unresolved Aspects in Existing Turbine Testing

Most recent experimental studies focus on small laboratory turbines and use mechanical braking to emulate load [19,20,21]. Consequently, they neglect the interaction between the turbine and an electric generator operating under variable electrical demand—an aspect that becomes critical once the plant is integrated into a microgrid. Moreover, the power levels examined are often limited to 70–100 W, far below the 10 kW class typical of community micro-hydro schemes.

1.4. Authors’ Previous Work

Our earlier investigations analysed hydraulic efficiency, water head, and input power under varying flow conditions [22]. We subsequently developed an artificial neural network model capable of predicting shaft torque for the same turbine [23]. Those studies, however, did not explore how electrical variables influence global performance.

1.5. Aim of the Present Study

Therefore, the present work extends previous efforts by experimentally characterising a 10 kW Pelton turbine with its permanent-magnet generator and resistive loads in place. We quantify electrical efficiency, voltage, and current under controlled combinations of spear position, flow rate, and electrical consumption. The resulting data set enables (i) identification of the optimal operating window for both turbine and generator and (ii) guidance for dispatch strategies in microgrids that rely on micro-hydro as a flexible, high-head resource.

2. Materials and Methods

2.1. Experimental Setup

The experimental setup (see Figure 1) and methodology used in this study are based on previous laboratory work by the authors [22,23], where the performance of a Pelton turbine was evaluated under controlled conditions. The test bench includes a closed-loop hydraulic circuit driven by a KSB Etabloc 80–250 pump (KSB SE & Co. KGaA, Frankenthal, Germany), with flow and pressure regulated via a variable frequency drive. Flow rate and inlet pressure are measured using a Yokogawa Optiflux2000/IFC300 electromagnetic flowmeter (KROHNE Messtechnik GmbH, Duisburg, Germany), which has an accuracy of ±0.2%, and a Yokogawa EJX11A Pitot tube (Yokogawa Electric Corporation, Tokyo, Japan), with an accuracy of ±0.25%, respectively.

The turbine is coupled to a 10 kW permanent magnet synchronous generator. Torque and rotational speed are measured using a torque sensor (Model RWT 420, Sensor Technology Ltd, Banbury, UK), which has an accuracy of ±0.1%, and an optical tachometer, with an accuracy of ±0.02, respectively. Electrical power is determined through voltage and current measurements using a digital multimeter with accuracies of ±0.1% for both voltage and current. Different electrical loads (ranging from 1.5 kW to 9 kW) are applied using resistive fan heaters, as depicted in the diagram in Figure 2.

The flow rate varies between 25 and 115 ${\mathrm{m}}^3/\mathrm{h}$, and the maximum head and flow rate delivered by the pump are 40 m.c.a and 129.6 ${\mathrm{m}}^3/\mathrm{h}$, respectively. The circuit piping has a diameter of 4 in. For further details on the experimental configuration, the reader is referred to [23].

2.2. Theoretical Basis of Performance

The equations used to calculate the power and efficiency of the turbine and the electric generator of the micro hydroelectric power plant are presented below [8,9,19,21,24,25]:

• The hydraulic power $P_h$ for an incompressible flow is calculated from the product of fluid density, gravitational acceleration, volumetric flow rate, and pressure head, as expressed by:

  $$P_h=\rho ·g·Q·H$$

  (1)
• The total electrical active power is obtained by summing the active power of each phase:

  $$P_{\mathrm{e}}=\sum _{z=1}^3{V}_z{I}_z$$

  (2)

  where $V_z$ and $I_z$ are the RMS voltage and current of phase z. In this experiment, all loads are purely resistive and balanced, so the power factor $cos\varphi$ is assumed to be unity. Hence, voltage and current are in phase and the active power in each phase reduces simply to the product $V_z{I}_z$.
• The electrical efficiency can be estimated as:

  $${\eta }_e={\displaystyle \frac{P_e}{P_b}}$$

  (3)
• The hydraulic efficiency of the turbine is estimated as:

  $${\eta }_t={\displaystyle \frac{P_b}{P_h}}$$

  (4)
• Overall efficiency considers the efficiency of the turbine and the efficiency of the electric generator. The overall efficiency can be expressed as:

  $${\eta }_o={\eta }_t{\eta }_e={\displaystyle \frac{P_e}{P_b}}$$

  (5)

2.3. Repeatability of the Tests

Measurement repeatability was assessed with the standard error of the mean (SEM), which quantifies the expected dispersion of sample means around the true population mean. For a set of n independent observations $x_i$, the SEM is [26]:

$$\mathrm{SEM}={\displaystyle \frac{\sigma }{\sqrt{n}}},\sigma =\sqrt{{\displaystyle \frac{1}{n-1}}\sum _{i=1}^n{\left(x_i-\overline{X}\right)}^2},\overline{X}={\displaystyle \frac{1}{n}}\sum _{i=1}^n{x}_i,$$

(6)

where $\overline{X}$ is the sample mean and $\sigma$ the unbiased sample standard deviation [26]. Unless otherwise stated, error bars in the figures correspond to $\pm \mathrm{SEM}$; this choice provides a visual estimate of the repeatability of each data point without inflating the interval width as would occur with $\pm \sigma$.

2.4. Measurement Uncertainty $u_x$

In addition to the statistical analysis based on the standard error of the mean (SEM), a combined uncertainty analysis was performed, considering the instrumental precision of the sensors used in the experiments. The uncertainty $u_x$ was computed using the standard method of uncertainty propagation, applying the root sum square (RSS) of the relative uncertainties of the measured variables involved in each calculated quantity:

$$u_x=x·\sqrt{\sum _i{\left({\displaystyle \frac{u_{x_i}}{x_i}}\right)}^2}$$

(7)

where x is the calculated quantity (e.g., torque or wheel speed), $x_i$ are the measured variables involved in the computation (e.g., Q, H, T, N), and $u_{x_i}$ are their respective absolute uncertainties. This formulation propagates the measurement errors to estimate the overall uncertainty in the derived quantity x.

2.5. Protocol for Experimental Test

The experimental protocol used in this study is based on a methodology previously published by the authors [22,23], in which the performance of a Pelton turbine was evaluated under varying flow conditions, nozzle openings, and electrical loads. The same flow range (25 to 115 ${\mathrm{m}}^3/\mathrm{h}$, in increments of 5 ${\mathrm{m}}^3/\mathrm{h}$) and electrical loads (from 1.5 to 9 kW) are maintained. As a novel contribution, this work includes the use of a spear needle controlled by a servomotor, which enables more precise regulation of the nozzle opening. The main function of the servomotor is to move the spear needle longitudinally, as illustrated in Figure 3 (in the X direction). The protocol from the previous study defined three types of tests, which are briefly described below:

• Variation of spear positions of 25, 50, 75, and 100% considering a constant electrical power consumption load on the generator. The tests are carried out at four spear position points: X = 0% indicates that the injector is fully open, then 25%, 50%, and 75% spear positions are evaluated. It has been decided that the maximum advance of the spear should be 75%, (which would be the same as saying a minimum nozzle opening of 25%), since beyond this point the water jet loses its characteristic shape, which is not relevant for the purposes of this study. It should be noted that the 100% spear travel position corresponds to a totally closed injector with zero water flow. We start with the spear position risen by 75%, with a flow rate of 25 ${\displaystyle \frac{m^3}{h}}$, and then increase the flow rate every 5 ${\displaystyle \frac{m^3}{h}}$ up to the maximum flow rate allowed in each test. Once this test is complete, the system is turned off, the spear position reduced by 50%, and the same procedure is repeated for the 25% and 0% spear positions.
• Variation of the electrical power consumption (3, 6, and 9 kW) with a fixed spear position (distance moved from x = 0; 0%). The generator is subjected to different electrical power consumption while maintaining the injector fully open (0% spear displacement). The flow rate is increased from 25 to 115 ${\mathrm{m}}^3/\mathrm{h}$ in increments of 5 ${\mathrm{m}}^3/\mathrm{h}$ for each load level.
• Electrical load variation (1.5 to 9 kW) at constant turbine–generator shaft speeds. In this test, the generator is subjected to increasing electrical loads (1.5, 3, 4.5, 6, 7.5, and 9 kW) while maintaining a constant shaft speed of 200 rpm. For each load level, the flow rate is adjusted accordingly to maintain the set speed. Once completed, the same procedure is repeated at 400 rpm. The position of the spear nozzle is set at 75%, which corresponds to a nozzle opening of 25%, as in the previously described configuration.

After completing the tests at 75% spear position, the same procedure is repeated at shaft speeds of 200 and 400 rpm for the remaining nozzle openings (50%, 25%, and 0%).

It is important to clarify that the term “maximum flow rate allowed in each test” refers to the nominal current and voltage of the heaters (electrical power consumption) 10 A and 220 V, respectively. However, during testing, these values were increased to 12 A and 340 V without causing any visible damage to the heaters.

3. Results and Discussion

3.1. Influence of the Spear Positions

The variation in flow rate directly affects the inlet pressure and hydraulic power, as previously described in [22] and remains consistent under the experimental conditions of the present study.

The torque increases almost linearly with shaft speed for all injector openings, a behavior consistent with previous results [22]. If we compare the different spear positions in the torque–flow graph, we can see in Figure 4a that increasing the spear position also increases the flow rate supplied to the turbine. This is obvious, since enlarging the nozzle area lowers the jet velocity, which in turn decreases the torque and raises the turbine–generator shaft speed, as shown in Figure 4b, thereby increasing the shaft power. As the nozzle opening grows, the water jet expands and loses coherence; part of the flow strikes outside the bucket centre line, reducing the effective momentum transfer and thus the measured torque (Figure 4a). This behaviour is consistent with the “fill-factor” theory discussed by Brekke [17]. Therefore, to obtain the same torque we must increase the flow rate whenever the spear position is increased. On the other hand, a nearly linear trend is observed in each test: the slope is steeper at larger spear positions and gradually diminishes as the spear is withdrawn.

If we compare the different spear positions in the torque vs. flow graph, we can see in Figure 4a that increasing the spear position also increases the flow rate supplied to the turbine, which is obvious since increasing the nozzle area decreases its speed, which in turn decreases the torque and increases the speed of the turbine–generator shaft, as shown in Figure 4b and increases the shaft power. Therefore, to obtain the same torque, we must increase the flow rate while increasing the spear position. On the other hand, it can be observed that there is a linear behavior in each of the tests for each of the spear positions, observing a higher slope in the case of higher spear position, which decreases as the spear position decreases.

Figure 4c shows a torque/shaft rotation speed relationship on the ordinate axis. It is observed that the torque has a preponderance over the shaft speed when the flow is low. This is due to the torque required to overcome the initial inertia, and as the flow is low, the shaft rotation speed is also low. This relationship decreases as flow increases, and the increase in torque and speed becomes proportionally similar.

The general trend of hydraulic performance as a function of shaft speed and injector opening has been extensively documented in a previous work [22], and therefore, these results are not discussed in detail again.

In Figure 5a, when the flow rate and the spear nozzle position (from 0 to 75%) are kept constant, the voltage and current increase, which coincides with the torque increase shown in Figure 4c. This is due to the increase in pressure head generated by the reduction in nozzle diameter. On the other hand, it is also observed that both voltage and current exhibit the same behavior, behaving linearly for all spear positions with respect to the shaft speed (see Figure 5b). Furthermore, as expected, the voltage has a linear behavior as the current increases (see Figure 5c).

It is also observed that there is a zone before 110 V and 1.8 A, corresponding to a flow of less than 50 m³/h (Figure 5a) and shaft speed of less than 215 rpm (see Figure 5b, where a greater dispersion of results is obtained (also showing a greater instability in the electric current). This is due to the low speed of the shaft and, therefore, the turbine, where the frequency at which the flow impacts the turbine blades is low.

At 130 rpm the measured voltage fluctuates by $\pm 6$ V, i.e., an order of magnitude larger than the instrumental uncertainty of $\pm 0.2$ V; this confirms that the effect is physical rather than a measurement instrument.

With regard to the electrical efficiency of the generator, it is notable, from Figure 5d, that there is a marked increase in variation at shaft speeds below 200 rpm. Specifically, a range of 8–76.8% variation is observed for the spear position initial (0%). The increase of spear position leads to a decrease in variation, but even at a spear position of 75%, the variation remains high, ranging from 73.8–90.8%. These large variations are a consequence of low shaft speeds and the resulting voltage variations and instability. For shaft speeds above 200 rpm (equivalent to a flow rate of 45 m³/h), the variations in electrical efficiency become less pronounced, and this trend holds true across all spear positions tests. The highest electrical efficiency achieved was 94.6%, observed with the spear position set to the initial level (0%) and a shaft speed of 650 revolutions per minute (rpm), the variations in electrical efficiency are similar, increasing proportionally with both shaft speed and flow rate.

Figure 6 displays the overall efficiency of the system. Notably, the minimum efficiency is observed when the spear position is set at 25%, and this occurs at turbine shaft speeds below 700 rpm. Conversely, the highest efficiency attained is 67% for the 25 and 50% spear positions at shaft speeds of 482 and 561 rpm, respectively. Similarly to the electrical efficiency, a range below 200 rpm is identified where the test results exhibit significant scattering, owing to the low turbine speed. At such low rotational speeds, the jet only hits between 15 and 18 blades per second, producing torque pulses at a lower frequency. These periodic load variations increase hysteresis and eddy current losses in the generator, which explains the large spread observed in the overall efficiency data (${\eta }_{\mathrm{e}}$) within this low-rpm region.

3.2. Influence of Electrical Power Consumption

Figure 7a shows that the flow rate also increases with increasing electrical load, since it must supply more flow to increase the shaft speed of the turbine generator, so the slopes of the torque and shaft speed graphs increase with increasing electrical load on the generator (see Figure 7b). Furthermore, as can be observed in Figure 7a, the flow rate increases as the electrical consumption applied to the generator increases. This is due to the need to supply more flow to increase the shaft speed of the turbine–generator. Consequently, the slopes of the torque and shaft speed graphs also increase with increasing electrical consumption on the generator (as depicted in Figure 7b). As with the variation of the spear position, Figure 7d illustrates a torque/shaft speed ratio on the ordinate, revealing that torque is dominant over shaft speed at low flow rates. This is presumably due to the torque required to overcome the initial inertia of the turbine. As the flow rate and shaft speed are both low, this relationship diminishes as the flow rate is increased, resulting in a proportionally similar increase in both torque and speed.

Figure 7e shows that the ratio of torque/wheel-speed remains constant for each case of applied electrical power consumption, regardless of the flow rate. However, this ratio becomes higher as the electrical consumption imposed on the generator increases, with the torque having a preponderance over the flow rate. On the other hand, previous laboratory work [22] demonstrated that the hydraulic efficiency of the turbine increases as the electrical energy consumption decreases, a result that is confirmed by the laboratory experiments conducted in the present study. This relationship can be explained by Equations (1)–(5), as the inlet pressure increases when more torque is required (see Figure 7c).

Before reaching a shaft speed of 550 rpm, the maximum hydraulic efficiency is 68.9% at an electrical power consumption of 3 kW. After reaching 550 rpm, the hydraulic efficiency drops to 65%, but the hydraulic efficiency reaches a maximum of 69% for a shaft speed of 580 rpm at an electrical power consumption of 9 kW.

Figure 8a shows the behavior of voltage at the output of one phase of the generator for different electrical consumption levels. We observe a linear increase in voltage and current with increasing shaft speed of the generator. As the electrical power consumption increases, the current also increases, even though the voltage is kept constant, the shaft speed remains practically constant at 110 V. However, when we work with a higher voltage, for example 220 V, the shaft speed increases by 10.97% as the electrical load increases from 3 to 9 kW, as mentioned earlier. This implies an increase in torque and current, as can also be seen in Figure 8c.

In Figure 8a, we can see that to obtain a voltage of 220 V (the voltage level used by electronic equipment in Chile), we need to operate in a range of 490–545 rpm, and for 110 V, we need to work in a range of 240–260 rpm. Additionally, the figure shows that there is an initial zone up to approximately 130 rpm where the voltage is unstable. This is due to the low revolutions of the turbine generator shaft and the low frequency at which the water jet hits the turbine blades, resulting in low vibrations.

In Figure 8b, we can see how the voltage changes with respect to the flow rate, with a higher slope for a lower electrical power consumption. On the contrary, the behavior of the current is higher when we apply a higher electrical power consumption to the electric generator.

Regarding the electrical efficiency of the generator, we can observe in Figure 8d that before a shaft speed of 170 rpm, there is a large dispersion of results, particularly when using a spear position of 75%. These variations were expected due to voltage variation and instability in low turbine shaft speed conditions. The highest electrical efficiencies were observed for shaft speeds above 170 rpm. Specifically, when using an electrical power consumption of 9 kW, the minimum and maximum efficiencies were 92.8–97.9%. For the case of applying an electrical power consumption of 6 kW, the minimum and maximum electrical efficiencies were 90.8–93.7%. Finally, the least favorable case was observed when using an electrical power consumption of 3 kW, with minimum and maximum efficiencies of 87.7–93.9%, with these last two cases being quite similar.

Figure 9 shows the behavior of the overall efficiency of the system, where we see a zone below 150 rpm of shaft speed that we will not consider for the analysis because it is a zone with dispersed speed values, due to the low speed. We see a zone between 400–580 rpm, in which we find the maximum values of total efficiency of the system; 65.7, 64.2, and 63.5 % for loads of 9, 6, and 3 kW, respectively. The maximum shaft speed limits are 727, 580, and 438 rpm for the electrical consumption of 3, 6, and 9 kW, respectively; this is due to the fact that we cannot exceed beyond the shaft rotation speeds, since this would lead to higher voltage and current consumption, which can damage the electrical collectors.

3.3. Influence of Constant Shaft Speed and Variable Electrical Consumptions

Figure 10 shows the torque/flow ratio for each electrical consumption applied to the electrical generator. An increase in the torque/flow ratio is observed with the increase in the electrical load in all cases. This means that the increase of this ratio is very similar in all cases, except for a delay between them due to the increase in shaft speed and decrease in spear position, as mentioned above.

The relationship between hydraulic efficiency, flow rate, and electrical load has been previously reported by the authors in controlled laboratory experiments [22], showing that efficiency drops significantly as the electrical load increases beyond 3 kW for constant shaft speeds of 200 and 400 rpm. These observations are consistent with the current experimental findings and are not repeated here.

In Figure 11a, we can see that the voltage does not vary significantly under low shaft speed conditions (200 rpm). However, in Figure 11b, when we use shaft speeds of 400 rpm, we observe that the voltage remains practically constant for the same electrical power consumption and different spear positions, but we see a decrease in voltage as the electrical load applied to the generator increases. On the contrary, we observe that the current significantly increases with increasing electrical power consumption, which could also be partially observed in the previous results (Section 3.2).

Regarding the electrical efficiency shown in Figure 11d, we can see that the highest electrical efficiency of the generator, using a shaft speed of 400 rpm, is obtained by applying an electrical consumption of 4.5 kW for a spear position of 25 and 50%, achieving an efficiency of 93.4%. However, with a spear position of 0%, an electrical efficiency of 97.7% can be achieved. On the other hand, with a spear position of 75%, a slightly lower efficiency ranging from 96–97% is obtained, but it is more stable as the applied electrical power varies from 3 to 9 kW.

For cases in which low speeds of rotation of the turbine–generator shaft (200 rpm) are used, it is observed that the efficiency is considerably lower, reaching values of less than 80% for an applied electrical consumption of less than 4.5 kW. The best electrical efficiency is obtained in all cases when an electrical consumption of 7.5 kW is applied, although this yields an efficiency between 95–97%. However, the low speeds are far from the speed at which this type of generator operates and are subject to instabilities caused by the low shaft speeds, as mentioned in the previous cases.

Figure 12a,b shows the overall efficiency of the turbine–generator system. The overall efficiency increases as the position of the spear is reduced. The best efficiency is obtained with an electrical consumption of 4.5 kW at a shaft speed of 400 rpm. Using a spear position of 0%, the efficiencies are 63.7% and 64.5%, respectively. On the other hand, if we observe the efficiency for a speed of 200 rpm, we see that the best performance in this scenario is obtained with the spear position of 0% and 25%, with efficiencies of 55.7% and 54.0%, respectively. In addition, the best performances for a spear position of 50% and 75% are obtained when applying an electrical consumption of 3.0 kW. Thus, the best overall performance of the system can be found in the ranges of 3 to 4.5 kW of electrical consumption applied to the generator. It is also important to mention that the lowest performance can be observed when applying extreme electrical consumption of 1.5 kW and 9 kW. As we increase the flow rate, the efficiency increases up to the maximum value, and then begins to decrease as we increase both the electrical consumption and the flow rate.

4. Conclusions

This experimental study assessed the performance of a 10 kW Pelton micro-hydropower plant under different spear positions, shaft speeds, and electrical loads. The main findings are:

• Spear position. For a given load, increasing the percentage of the spear opening raises the torque, and consequently the generator’s voltage and current—at the same flow rate.
• Torque–speed balance. In all tests, the torque/speed ratio exceeds unity at low rpm, reflecting the inertia required to start the runner; the ratio stabilises as speed rises.
• Electrical efficiency. Large scatter occurs below 200 rpm. Above 450 rpm the generator maintains ${\eta }_e\ge 94\%$, reaching a maximum of ≈98% at 400 rpm with spear positions of 25–50% and loads of 4.5–7.5 kW.
• Overall efficiency. The best global performance (${\eta }_{\mathrm{overall}}\approx 65\%$) is obtained with electrical consumptions of 3–4.5 kW, spear openings $\le 25\%$, and shaft speeds > 400 rpm.
• Low-speed operation. Speeds below 150 rpm cause noticeable voltage/current fluctuations and shaft vibrations, attributed to the low jet-bucket impact frequency; such conditions should be avoided in regular operation.
• Practical implication. The quantified operating window provides guidelines for dispatching micro-hydro units within islanded or grid-connected microgrids.