The development of the regions along the Northern Sea Route of the Russian Federation is difficult without an energy source possessing such qualities as autonomy, reliability, and maneuverability. A small nuclear power plant (SNPP) should become such an energy source (Belyaev et al. 2020; Petrunin 2021). Currently, the State Corporation Rosatom is developing an SNPP project based on the RITM marine reactor (Zverev et al. 2020b; Petrunin et al. 2023; Samoilov et al. 2023).

However, the use of marine reactor technologies for stationary power plants has required significant design changes to the reactor core, as well as increased accuracy of coolant temperature measurement at the core outlet (Kulakov et al. 2020 ; Zverev et al. 2020a; Zakharychev et al. 2021; Romanov et al. 2023). This is due to the more stringent requirements of regulatory documentation. To meet these requirements, it is necessary to increase the number of temperature monitoring points from three to twenty-four. The coolant temperature is measured by resistance thermometers (RTDs), with flow sampling accomplished through dedicated tubes installed at the level of the outlet openings of the FA upper nozzles.

The ingress of local cold or hot jet flows from the fuel rod bundle into the sampling tube can significantly affect the accuracy of coolant temperature determination using the RTD. To account for this effect, it is necessary to identify the regions of the fuel rod bundle from which coolant enters the sampling tube and to select the geometry of the FA upper nozzle that ensures intensive flow mixing.

Studying the process of coolant flow formation as it exits the FA will make it possible to determine the optimal upper nozzle geometry and to identify the regions of the fuel rod bundle from which flow is most likely to enter the sampling tube. The experimental data obtained are necessary to substantiate design solutions for core components of the RITM reactor for the SNPP.

The flow of water coolant was simulated using an air flow from the experimental test bench of the research laboratory for thermal hydraulics of nuclear power plants at Nizhny Novgorod State Technical University named after R.E. Alekseev, based on the theory of hydrodynamic modeling (Gukhman 1973, 1974).

The open loop of the aerodynamic test bench includes a high-pressure fan, an experimental model, and a measurement system (Dmitriev et al. 2019; Dmitriyev et al. 2021). The hexagonal channel of the model was assembled from two parts. The first part contained simulators of fuel rods, burnable absorber rods, a central displacer, six stiffening angles, and two spacer grids. The second part of the model included replaceable mock-ups of the upper nozzles, as well as a mock-up of the mounting plate for the FA with the sampling tube (Fig. 1). The dimensions of all model elements were scaled up relative to the actual design components according to a geometric similarity factor of 5.79.

Three types of FA upper nozzles were used in the model design. The first type is characterized by the presence of three wide drain windows for radial flow outlet and six circular holes for axial flow outlet. The second type of upper nozzle is characterized by the presence of six narrow drain windows and six circular axial holes. The flow passage area of the first and second type upper nozzles is the same. In the design of the third type of upper nozzle, the six circular holes for axial flow outlet are omitted (Fig. 2).

Measurement of axial flow velocity was performed using the pneumometric method. Pitot tubes were used to determine the static and dynamic pressure of the flow at a point. The pressure values, taking into account correction factors, were then recalculated into axial velocity values using a digital differential manometer. The uncertainty in determining the axial velocity did not exceed 7.5%.

To determine the effect of the upper nozzle design on coolant flow and to identify zones of the fuel rod bundle from which flow enters the sampling tube, the tracer injection method (propane gas) into the air flow was used. The procedure was as follows:

To identify the zones of the fuel rod bundle from which flow could enter the sampling tube, a corner cell and a central displacer cell were selected, as well as a pair of regular cells, one of which was located opposite the sampling tube cross-section, and the other on the opposite side of the fuel rod bundle (Fig. 3). The tracer was injected through these cells.

All measurements were performed before the outlet spacer grid, at the outlet of the fuel rod bundle, at the outlet of the upper nozzle, in the region between the upper nozzle and the FA mounting plate, in the mock-up of the sampling tube, and in the openings of the FA mounting plate. Based on the measurement results, maps (contour plots) of axial velocity and tracer distribution were constructed.

The structure of the axial flow in characteristic cross-sections was determined using dimensionless values of axial flow velocity. The dimensionless values were obtained by normalizing the local velocity values at measurement points to the average flow velocity (bulk velocity) in each measurement cross-section. The difference in the average flow velocity in each measurement cross-section is due to the significant difference in the area of each such characteristic cross-section. The value of the average flow velocity in the characteristic cross-sections was determined by taking the ratio of the axial flow rate of the air working fluid at the inlet of the experimental model channel to the flow passage area of the model in the measurement regions.

The flow rate at the inlet to the experimental model channel was determined using a measuring header (flow measurement manifold). The pressure drop across the measuring header was measured by a SAPFIR differential pressure sensor with an uncertainty of no more than 0.25%. Then, based on the appropriate conversion, the value of the axial flow rate was determined. To confirm the flow rate at the model inlet measured with the flow measurement manifold, additional velocity measurements were performed downstream of the manifold in the settling section before the model. The velocity was measured using an Omniport40-T30 hot-wire anemometer sensor with an uncertainty of no more than 0.1% according to the procedure described in GOST 12.3.018-79 “Ventilation systems. Methods of aerodynamic testing”. The velocity profile was also measured in the settling section before the model channel. The measurements yielded a uniform velocity profile.

During the research, the temperature of the air working fluid at the inlet to the experimental model channel was recorded. The measured temperature range of the air flow was from 25 °C to 27 °C, which did not significantly affect its other physical parameters.

The selection of the optimal geometry of the experimental model and the possibility of subsequently recalculating the air flow characteristics obtained during the experiments to full-scale coolant flow conditions are possible when hydrodynamic similarity is achieved, which requires simultaneous compliance with geometric, kinematic, and dynamic similarity.

Geometric similarity is ensured by matching the geometric dimensions of the model and the full-scale object, taking into account a scale factor of 5.79. The use of a large scale factor and, consequently, the use of fuel rod simulators with a diameter of 40 mm is due to the following factors:

Dynamic similarity of the fluid motion processes in geometrically similar structural elements is ensured by maintaining close values of the Reynolds number. The maximum Reynolds number in the model was 90000, while in the full-scale object it was 100000.

Kinematic similarity of the fluid motion processes in geometrically similar structural elements is ensured by maintaining proportionality of the dimensionless velocities at corresponding points in the model and the full-scale structure. To achieve this, it is necessary to conduct experimental studies in the similarity zone of the drag coefficient of the FA components.

The investigation of the flow structure in the experimental model was carried out in the similarity zone of the drag coefficient of the FA components. The onset of the similarity zone of the drag coefficient of the FA components was determined experimentally:

Analysis of the above dependencies shows that, starting from a certain value of the Reynolds number, the friction drag of the smooth tube bundle and the HDC of the spacer grid and FA upper nozzles cease to change significantly. This phenomenon indicates the onset of the similarity zone of the drag coefficient of the FA components. During the experiments, the similarity zone of the drag coefficient of the FA components began in the Reynolds number range from 30000 to 40000.

Furthermore, the measurement results of the HDC of the spacer grid and FA upper nozzles showed their agreement with the HDC values obtained for full-scale FA structural components at identical Reynolds numbers. This also confirms the representativeness of the experimental studies.

The Reynolds number was determined according to the formula:

Re = (W·d_h)/ν,

where W is the flow velocity at the inlet of the experimental model channel (m/s), dh is the hydraulic diameter in the characteristic cross-section of the experimental model (m), ν is the kinematic viscosity of the fluid (m²/s). The hydraulic diameter (dh) was determined for each measurement cross-section.

All measurements were carried out with the average flow velocity of air in the characteristic cross-sections of the model lying in the range of 20–40 m/s and a Reynolds number of 60000–90000.

The experimental results of the axial flow velocity made it possible to identify the main patterns of its formation as it exits the FA.

In the fuel rod bundle, before the outlet spacer grid, the axial flow velocity is non-uniform. In the central region of the fuel rod bundle, located before the flow passage of the upper nozzle confusor, the dimensionless axial flow velocity was 1.05–1.2. Near the FA shroud, in the region one row of fuel rods away from it, the dimensionless velocity is 20% lower, amounting to 0.8–1.0. This flow pattern is characteristic of all upper nozzle types. Therefore, in this region, the influence of the specific shape of the upper nozzle confusor flow passage does not significantly affect the flow (Fig. 7).

Upon flow exit from the fuel rod bundle, a redistribution of the axial flow velocity is observed, caused by the geometric features of the outlet part of the bundle, the shape of the confusor, and the flow passage area of the upper nozzle openings.

The influence of the geometry of the fuel rod bundle outlet part is characterized by the formation of an axial flow structure in which the velocity behind the outlet part of each fuel rod is 35% lower than in the regions of the cross-section free of fuel rods.

The influence of the confusor shape and the flow passage area of the upper nozzle openings on the flow as it exits the fuel rod bundle is characterized by the formation of extensive regions with low and high axial velocity values. Before FA upper nozzles whose design includes six circular holes for axial flow outlet, the dimensionless axial flow velocity in the region of the upper nozzle confusor was 1.1–1.7, and before the six circular holes it was 1–1.3. Before the upper nozzle without six axial holes, the dimensionless flow velocity before the confusor cross-section is on average 30% higher, amounting to 1.4–2.2, which is due to the reduction in the flow passage area of the upper nozzle. For all upper nozzle types, the formation of a low axial flow velocity zone at the periphery of the fuel rod bundle is characteristic; the dimensionless axial velocity in this region was 0.2–0.7. The formation of this region is largely due to the geometry of the upper nozzle inlet cross-section (Fig. 8).

The magnitude of the axial flow velocity passing through the FA upper nozzles is influenced by the configuration and shape of the outlet openings, as well as their flow passage area. The influence of the configuration and shape of the upper nozzle outlet openings is clearly evident when comparing the flow structure in Type 1 and Type 2 upper nozzles, which have the same flow passage area but differ in the shape and number of the large drain windows. In the upper nozzle with three wide drain windows, the dimensionless axial flow velocity was 2.7–3.2, while in the upper nozzle with six narrow windows it was 2.5–3.1. Thus, the shape of the confusor flow passage of the upper nozzle has an effect on the magnitude of the axial flow velocity passing through it. The greatest contribution to the change in axial flow velocity comes from the reduction in the flow passage area of the upper nozzle, as can be seen when comparing the axial velocity structure in Type 2 and Type 3 upper nozzles. In the Type 3 upper nozzle without six circular holes and, consequently, with the smallest flow passage area, the dimensionless axial flow velocity is 20% higher, amounting to 3.2–4.0. In the six circular holes of Type 1 and Type 2 upper nozzles, the dimensionless axial flow velocity has similar values, ranging from 2.3 to 3.2 (Fig. 9).

In the region between the upper nozzle drain windows and the FA mounting plate, the flow has a pronounced jet-like character, with the formation of local axial velocity maxima and an extensive low-velocity region near the upper nozzle and the sampling tube. The size of this region depends on the upper nozzle design and the area of the flow passage blocked by the sampling tube, while the dimensionless velocity magnitude in this region is independent of the upper nozzle type and amounts to 0.3–0.6. Behind the upper nozzle with three wide windows, the low-velocity region is the most extensive, while behind the upper nozzle without six circular axial holes it has the smallest size. This is because behind the upper nozzle with three wide windows, the sampling tube blocks one-third of the flow outlet cross-section, whereas behind the upper nozzle with six narrow windows it blocks only one-sixth (Fig. 10).

The non-uniformity of the jet flow structure from the upper nozzle windows and within the internal volume of the upper nozzle confusor is determined by the presence of the sampling tube. By blocking part of the flow passage area of the upper nozzles, the sampling tube redistributes the flow such that behind the upper nozzles with three and six windows, two and five high-velocity jets are formed, respectively. Behind the upper nozzle with three windows, the dimensionless velocity in the high-velocity jet region is 1.7–2.1; behind the upper nozzle with six windows and six axial holes, it is 1.4–1.7; and behind the upper nozzle without six axial holes, it is 1.5–2.0. Upon flow exit from the upper nozzle with three windows and six circular holes, the greatest flow non-uniformity is observed, manifested as an extensive low-velocity region and high velocity values in the local jet regions.

In the openings of the FA mounting plate fragments, the jet flow from the upper nozzles persists, with local velocity maxima recorded in the openings opposite the windows of the upper nozzles without the sampling tube. Behind the upper nozzle with three windows, a pair of local dimensionless velocity maxima is formed in the openings of the FA mounting plate, with values of 1.2–1.3. Behind the upper nozzles with six windows, five local dimensionless velocity maxima are formed in the openings of the FA mounting plate, amounting to 1.0–1.2. In the openings of the FA mounting plate behind all types of upper nozzles, a low axial velocity region is formed near the sampling tube, caused by flow deceleration at the tube walls. The dimensionless velocity in this region was 0.4–0.8 (Fig. 11).

The uniformity of the axial flow in the sampling tube largely depends on the merging of the radial flow from the upper nozzle windows and the axial flow from the six circular holes. Behind Type 1 and Type 2 upper nozzles, this merging process occurs, and the dimensionless axial flow velocity in the sampling tube was 0.6–0.9. Behind the Type 3 upper nozzle, there is no merging of flows of different directions, and the dimensionless axial flow velocity was 0.95–1.0.

The results of the experimental study of flow mixing made it possible to select the optimal upper nozzle design that minimizes the influence of jet flows with different temperatures on the RTD reading, and to draw the conclusions presented below.

The design of the FA upper nozzle did not have a significant effect on the formation of cross-flows upon exit from the fuel rod bundle. This is confirmed by the absence of tracer smearing in the cross-section and its predominant localization in the injection cell. Such tracer distribution is characteristic of all upper nozzle design types. The tracer concentration values in all injection cells, with all types of upper nozzles installed, amounted to 2100–2500 ppm.

During flow through the upper nozzle confusors, a mixing process is observed, expressed as a displacement of the tracer toward the sampling tube simulator. The extent of tracer displacement largely depends on the flow asymmetry caused by the blockage of the upper nozzle confusor flow passage by the sampling tube. For all tracer injection zone variants, the greatest transverse tracer displacement toward the sampling tube, and consequently the greatest flow mixing, occurs in the confusor of the three-window upper nozzle. For this upper nozzle design type, tracer is detected in the sampling tube region for all tracer injection zones in the fuel rod bundle.

In the other upper nozzles, tracer is detected in the windows located opposite or near the injection cells. Thus, the flow entering the sampling tube comes predominantly from the region of regular cells and displacer cells located opposite or in close proximity to the flow passage of the sampling tube. The tracer concentration values for all injection cells, with all types of upper nozzles installed, amounted to 200–500 ppm. Examples of tracer distribution at the upper nozzle outlet when tracer is injected into a regular cell remote from the sampling tube are shown in Fig. 12.

The presence of tracer from the injection zones in the fuel rod bundle within the sampling tube is visualized by tracer distribution maps before the inlet cross-section of the sampling tube itself (Fig. 13). From each characteristic injection zone in the fuel rod bundle, an average of 10–15% of the initial tracer flow rate enters the sampling tube.

Thus, to minimize the influence of flows with different temperatures on the RTD readings, it is necessary to use upper nozzles with three windows asymmetrically arranged relative to the six circular holes.