This technical paper presents results of an air-cooled supercritical CO_{2} (sCO_{2}) finned-tube sink heat exchanger (HX) performance test comprising wide range of variable parameters (26–166 °C, 7–10 MPa, 0.1–0.32 kg/s). The measurement covered both supercritical and subcritical pressures including transition of pseudocritical region in the last stages of the sink HX. The test was performed in a newly built sCO_{2} experimental loop which was constructed within Sustainable Energy (SUSEN) project at Research Centre Rez (CVR). The experimental setup along with the boundary conditions are described in detail; hence, the gained data set can be used for benchmarking of system thermal hydraulic codes. Such benchmarking was performed on the open source Modelica-based code ClaRa. Both steady-state and transient thermal hydraulic analyses were performed using the simulation environment DYMOLA 2018 on a state of the art PC. The results of calculated averaged overall heat transfer coefficients (using Gnielinski correlation for sCO_{2} and IPPE or VDI for the air) and experimentally determined values shows reasonably low error of + 25% and – 10%. Hence, using the correlations for the estimation of the heat transfer in the sink HX with a similar design and similar conditions gives a fair error and thus is recommended.

## Introduction

In the nuclear power plant design, the consideration of multiple component failure scenarios is a motivator for the development of failure safe backup systems. One approach for a failure safe backup system currently under development is called supercritical CO_{2} heat removal (sCO_{2}-HeRo) [1]. It is designed for boiling water reactors and pressurized water reactors (PWRs) to prevent Fukushima-like accidents, where a combined station blackout, loss of ultimate heat sink, and loss of emergency cooling occurred. The sCO_{2}-HeRo is such an emergency cooling system. It transports the decay heat from the reactor core through a self-propellant, self-sustaining Brayton cycle, including compressor, heat exchanger (HX) (steam-sCO_{2}), turbine, and sink heat exchanger to the ambient air.

The main objective of this work was to provide evidence for the concept of the air-cooled finned-tube sink HX at laboratory conditions (technical readiness levels 3–4), develop and validate a new numerical Modelica-based model for the code ClaRa suitable for modeling steady/transient scenarios in sCO_{2} environment, and finally deliver valuable operational experience from the unique sCO_{2} facility at Research Centre Rez (CVR).

The measurement covered both the supercritical and subcritical pressures (7–10) MPa including transition of pseudocritical region (27–36) °C in the last stages of the sink HX. The nominal parameters of the sink HX were reached: 95 kW, 7.8 MPa, 166 °C/33 °C, 0.325 kg/s for the sCO_{2} side cooled by 25 °C forced air flow with ambient pressure.

A number of investigators have carried out experimental tests and analyses of the heat transfer performance of finned-tube sCO_{2} gas coolers. Majority of this work was focused only on steady-state analyses [1–4]. All of these authors use ε-NTU or LMTD (i.e., lumped method and distributed method) which has limitations, especially when it comes to modeling of rapidly varying thermophysical properties in the critical region. Therefore, e.g., LMTD has to be modified using an integral approach for LMTD [5] or finite methods need to be deployed, i.e., finite volume method utilized in this paper or finite element approach found in the work by Yin et al. [6] who performed stationary calculations and optimization.

Apart from an experimental research, there are numerous studies dedicated purely to simulation tools development for sCO_{2} energy systems. In the dynamic simulation software, there can be found a few in-house system codes analyzing nuclear reactors and experimental loops behavior with sCO_{2} [7,8] or system codes primary developed for light water reactors safety analyses like ATHLET, RELAP, and TRACE which has been upgraded for handling sCO_{2} simulations [9–12]. However, the validation of these codes in sCO_{2} environment has been lacking. Therefore, this study was conducted to present a new validated Modelica code as well as to submit a new set of sCO_{2} data for future benchmark.

To the best of our knowledge there has been no previous investigations reported in the literature on the sCO_{2} gas coolers performing experimental work together with both, the steady-state and transient analyses.

The results in this paper will benefit to researchers, designers, software engineers, thermal hydraulic specialists, and operators of sCO_{2} energy systems through the shared measured data and described operational procedures in a unique sCO_{2} facility.

### Description of the sCO_{2}-Hero System.

Figure 1 depicts the scheme of the sCO_{2}-HeRo retrofitted into the PWR. In case of a station black-out and the loss of ultimate heat sink accident, the reactor automatically shuts down, the turbine fast-driven valves close, and the safety valves open. However, the residual heat is produced. By nature, without the utilization of main circulation pump (MCP), natural circulation is established in the primary circuit, which transfers the decay heat to the steam generators (SG) and evaporates its water content. The steam flows into a heat exchanger (CHX), which must be very compact to fit into the limited space available in existing reactor building. The steam condenses and the liquid water, driven by gravity, flows back into the SG. Thus, the water content in the SG is preserved. Inside the compact HX the sCO_{2} heats up. It flows through a turbine, which is located on the same shaft as the compressor and the generator. Downstream of the turbine, the sCO_{2} gets cooled by the air in the sink HX and is delivered to the compressor and back to the compact heat exchanger. Over a large operating range, the turbine of the Brayton cycle shall produce more power than the compressor needs to operate. The excess power is transferred into electricity, which is used to power additional fans of the sink HX for better heat removal.

The sCO_{2}-HeRo system can be attached to both existing pressurized water reactors and boiling water reactors, since the thermodynamic parameters of steam are similar. Without having the sCO_{2}-HeRo system deployed, the water content in the SG would steadily decrease (by releasing the steam through pressure safety valve or pressure relief valve) causing overheating of the primary circuit which could eventually lead to fuel damage [13,14].

Within the European project “sCO_{2}-HeRo,” six partners from three European countries are working on the assessment of this cycle. The goal is to numerically and experimentally show evidence for the concept on a small-scale demonstrator of the sCO_{2}-HeRo system which shall be incorporated in the PWR demonstrator (a reproduction of a two-loop pressurized water reactor Siemens/Kraftwerk Union design at a scale of 1:10) at the Simulator Centre of KGS and GfS in Essen, Germany. Before assembling the small sCO_{2}-HeRo system in the Simulator Centre, each major component was tested in different institutions. The performance of the compact HX (microchannel type) was verified in the sCO_{2} test loop (SCARLETT) in University of Stuttgart, while the air-cooled sink HX, compressor, and turbine were measured in the CVR sCO_{2} experimental facility.

## Description of Sink HX for the Demonstrator

The design of the sink HX strongly influences the behavior of the whole sCO_{2}-HeRo system, as it is operated near the critical point region of CO_{2} (7.8 MPa, 33 °C). Underestimated size of the HX can lead to a not self-propellant sCO_{2}-HeRo design. This is due to the high outlet temperature of the HX (inlet to the compressor) resulting in excessive compression work.

According to the optimized cycle calculations of the sCO_{2}-HeRo system, the sink HX model for the small scale sCO_{2}-HeRo has been specified [13].

Table 1 shows the main thermodynamic parameters for the selected two identical sinks HX's working in parallel. Each designed as finned tube HX type cooled by forced air (fan with EC motor with speed control). One of them was selected for testing and implemented into the sCO_{2} loop in CVR.

The conceptual drawing with overall dimensions is shown in Fig. 2.

The internals of sink HX includes stainless steel AISI 304 tubes in staggered arrangement with rectangular aluminum fins (metal sheet). The arrangement is such that the flow on the sCO_{2} side is purely horizontal (except the inclined bends placed outside the air flow), while on the air side the flow is completely vertical. An illustrative scheme is shown in Fig. 3.

The overall heat transfer area for one sink HX is 361 m^{2}. The detail geometry of sink HX is included in Table 3.

## Test Facility at Research Centre Rez

The heat transfer investigations in the sink HX test configuration took place at CVR, using sCO_{2} experimental loop which was constructed within Sustainable Energy (SUSEN) project. This unique facility enables to study key aspects of the cycle (heat transfer, erosion, corrosion, etc.) with wide range of parameters: temperature up to 550 °C, pressure up to 30 MPa, and mass flow rate up to 0.35 kg/s.

Figure 4 shows the piping and instrument diagram (P&ID) of the loop. A part of the primary circuit used for the sink HX measurement is represented by thick line, and it consists of a low temperature regenerative heat exchanger (LTR) and high temperature regenerative heat exchanger (HTR), a main piston pump, and four electric heaters of the total maximum power of 110 kW. Heat exchangers HTR and LTR are designed as a counter-flow shell and tube-type from stainless steel (SS).

The electrical heater H3 with nominal power 20 kW is positioned at the bypass of the LTR in order to simulate the behavior of a recompression cycle.

For cooling purposes, two shell and tube type coolers CH_{1} and CH_{2} are connected to the loop. The cooler CH_{1} from SS is cooled by water (temperature 20 °C, 1.4 kg/s flow rate of water), and the cooler CH_{2} also from SS is used as the main cooling medium oil (Malotherm SH, Sasol, Sandton, South Africa), because of the high temperatures of the exhaust heat. Next part of the primary loop consists of two parallel electric heaters H_{1/1} and H_{1/2} from SS with 30 kW each, followed by one Inconel electrical heater H_{2} with 30 kW. Behind the heaters, a test section TS (pressure tube which enables to insert samples) and reduction valve RV is positioned. It is used to represent a turbine expansion. The main operating parameters of the primary circuit are shown in Table 2.

For testing of the sink HX, the low pressure side (behind the reduction valve) of the LTR and the HTR as well as the oil cooler CH_{2} were by-passed in order to achieve desired inlet temperatures (max. 170 °C) to the sink HX. The by-pass is marked in thick red line with squares. The omitted piping is marked in thin gray line. Pressure in the system is controlled either by the electric heaters, i.e., by the temperature in the circuit, or by the filling compressor/release valves (to the outside atmosphere) by which it is possible to control the amount of CO_{2} in the loop, thus the pressure.

Figure 5 shows the sCO_{2} loop and the installed sink HX configuration, which is outside of the experimental hall.

Component geometry of the sCO_{2} loop is summarized in Table 3.

## Measurements

This section contains the measurement procedure of the performed tests on sink HX within sCO_{2} experimental facility in CVR.

### Limits of the Test Facility at Research Centre Rez.

Operational limits of the test facility (Table 4) must be taken into account and they should not be exceeded during the performance test.

For carrying out the experiments, the primary circuit was first evacuated and then filled by CO_{2} (99.995%).

Figure 6 shows the sink HX outside of the experimental hall with in-coming and out-going pipelines together with all measurement devices.

### Measurement Parameters and Procedure.

The measurement campaigns covered both supercritical and subcritical regions including transition through the pseudocritical region in the last stages of the sink HX. The critical point of the CO_{2} is 7.39 MPa and 31.1 °C. The controlled (independent) and resulted (dependent) parameters are summarized in Table 5.

Measurement campaigns were carried out with different inlet conditions on both sides of the sink HX. The measurement time took about 15 min at each measurement point in order to reach stable conditions. The operational procedure was as follows:

- (1)
hold

*p*_{_sCO2_in}= 7.8 MPa at nominal - (2)
hold

*ṁ*_{_sCO2}and*T*_{_sCO2_in}at certain value (0.1, 0.2, or 0.32) kg/s and (50, 100, 166) °C, respectively - (3)
vary

*V̇*_{_air_out}, i.e., frequency of the fan (50, 75, 100) % of nominal 50 Hz, while for each frequency a measurement was recorded - (4)
increase/decrease

*m*_{_sCO2}while keeping the*T*_{_sCO2_in}and repeat step 3 and repeat this procedure for all variants of*ṁ*_{_sCO2}(0.1, 0.2 or 0.32) kg/s - (5)
increase/decrease

*T*_{_sCO2_in}to new value and repeat steps 3 and 4 to record all variants of*T*_{_sCO2_in}(50, 100, 166) °CWith this procedure the influence of

*m*_{_sCO2},*T*_{_sCO2_in}, and*V̇*_{_air_out}was studied. In order to see impact of*p*_{_sCO2_in}following steps were taken: - (6)
hold

*T*_{_sCO2_in}at certain value (100 °C) - (7)
hold

*m*_{_sCO2}and*p*_{_sCO2_in}at certain value (0.1, 0.2, or 0.3) kg/s and (7, 7.4, 8.5, 9.4, 10) MPa, respectively, and vary*V̇*_{_air_out} - (8)
increase/decrease

*m*_{_sCO2}while keeping the*p*_{_sCO2_in}and repeat step 3 repeat this procedure for all variants of*ṁ*_{_sCO2}(0.1, 0.2 or 0.32) kg/s - (9)
increase/decrease

*p*_{_sCO2_in}to new value and repeat step 3 and 7 to record all variants of*p*_{_sCO2_in}(7*, 7.4, 8.5, 9.4*, 10*) MPa.

*Not all ṁ_{_sCO2} (0.1, 0.2 or 0.32) kg/s were possible to implement due to the limited power of filling pump.

### Measurement Devices and Experimental Errors.

Figure 4 shows the piping and installation diagram (P&ID diagram) of the modified sCO_{2} loop with the main components together with all installed measurement devices, such as a mass flow meter, volume flow meter, Pt-100 sensors, thermocouples, and pressure sensors. The nomenclature of the measurement devices respects the KKS identification system for power plants.

The uncertainties provided by the measurement devices, transducer, input card, and control system are summarized in Table 6. The errors correspond to calibration certificates and manufacturer's instructions.

The error propagations are described in Annex A.

The results for the design (nominal) conditions of the sink HX have shown 15% error propagation of the heat transfer on the sCO_{2} side *Q*_{sCO2} and 8% for the air side *Q*_{_air.}

### Experimental Results and Discussion.

This section contains experimental results for steady-state and transient operation.

#### Steady State Operation Results.

Figure 7 shows the experimental results of $Q_air=m\u02d9_air\xb7cp_air\xb7(T_air\u2212out\u2212T_air_in)$ and $Q_sCO2=m\u02d9_sCO2\xb7(h_sCO2\u2212in\u2212h_sCO2\u2212out)$. For all of the 34 measurements, the heat transfer ratio *R = Q_*_{air}/*Q*__{sCO2} stayed within the limits (115%/85%). The base source of the errors propagation for the *Q*__{sCO2} is the uncertainty of the thermocouple measurement of the outlet sCO_{2} (far less than at the inlet). This is due to the fact that the pseudocritical region (around 34 °C) is crossed here and each small error of the temperature determination leads to high errors in evaluation of enthalpies (up to 60 kJ/kg), i.e., heat power (15 kW). Figure 6 shows the sink HX standing outside of the experimental hall with pipelines and measurement devices.

The honey combs are utilized to stabilize the flow at the outlet of the air pipe and more importantly, in front of the Wilson grid which is used to measure volumetric flowrate throughout the pitot arrays. These consist of a row of vertical tubes, with alternate rows of holes facing up and down stream, measuring the total and substatic pressures from which dynamic pressures are calculated. As shown in Fig. 7, the air side heat flow rate *Q*__{air} exceeds the CO_{2} heat flow rate *Q*__{sCO2}. by max. 15%.

#### Comparison of Measurements With Correlations From the Literature.

The potential of the sCO_{2}-HeRo system to deal with a range of different accident scenarios and beyond-design accidents will need to be proven with the help of thermal hydraulic codes. Therefore, heat transfer models were compared with the experimental data.

The heat transfer at the tube side where sCO_{2} flows is geometrically characterized by the inner diameter and shape of the tubes and has been thoroughly studied. Numbers of correlations are discussed in the literature [16–18].

_{2}) of the heat exchanger, it is suitable to use well-known Gnielinski correlation for the forced convection [18]. Although, some investigators [19–21] modified this correlation, as indicated by Zilio et al. [22], these correlations often predict similar results for CO

_{2}gas coolers

The air, which is pulled through the cooler by a fan mounted at the top of the unit, flows around the tube bundle with fins. This is geometrically much more complex. It includes definition of transverse and longitudinal tube spacing, tube outer diameter, number of tube rows, fin spacing, fin thickness, and fin type. Besides this complexity, the air local heat transfer coefficient is by one order of magnitude smaller than of the sCO_{2} side. Thus, the air side determines the size of the whole HX.

*α*

_{ideal}is then calculated from the Nusselt number using equivalent diameter

*d*

_{outer}. Since the design of the HX contains fins for increasing the heat transfer area, the real local heat transfer coefficient efficiency of the fin needs to be taken into account. The real local heat transfer coefficient is calculated according to the following equation:

For the calculation of efficiency of the rectangular fins *η*_{fin}, a formula stated in Ref. [24] was used. For the given geometry it resulted in *η*_{fin} = 0.95.

The graph in Fig. 8 shows a comparison of resulted averaged overall heat transfer coefficients *k*_{_calc_avg} calculated (using Gnielinski [18] for sCO_{2} and IPPE [23] for the air) and experimentally determined *k*_{_exp_avg} for all the 34 measurement points. The overall *k*_{_exp_avg} was calculated from the measured temperatures, pressures, mass flow rates on both the sCO_{2} and air sides using the following formula $Q=kexp_avg\xb7Aouter\xb7\Delta T\u2032(W)$ describing the heat transferred in each control volume of the sink HX. The positive errors suggest that the calculated values, using correlations, overestimate the experimental values for the negative errors and vice versa. It can be seen that the discrepancy is reasonable low + 25% and −10%.

From the graph Fig. 9, it can be concluded that both correlations according to IPPE and VDI are in perfect match.

The effect of the mass flux on the local heat transfer coefficient of sCO_{2} is illustrated in Fig. 10. At the same pressure, the local heat transfer coefficient of sCO_{2} increases with mass flux due to higher Reynolds number.

Figure 11 presents the local heat transfer coefficient of sCO_{2} for different cooling pressures ranging from 7.1 MPa to 9.4 MPa at a given mass flux. For the supercritical pressures (higher than 7.4 MPa), the peak values in the local heat transfer coefficient are shown at the same pseudo-critical temperatures. Higher pressure has lower local heat transfer coefficient because the specific heat is lower. At the subcritical pressure (7.1 MPa), the local heat transfer coefficient increases toward colder temperatures and even exceeds the values of supercritical pressure due to the higher specific heat at this region. There has been considerable prior research done in the area of sCO_{2} coolers with similar findings [20,21].

#### Transient Operation.

During the performance measurement of the sink HX a transient test was performed. The volumetric flow rate of the air was stepwise changed from the value 12,250 m^{3}/h through 9400 m^{3}/h (75% fan speed) to 6400 m^{3}/h (50% fan speed) while keeping the nominal sCO_{2} mass flow rate at 0.32 kg/s. Before each change a steady-state was reached such that *p*_{_sCO2_in} = 7.8 MPa *T*_{_sCO2_in} = 166 °C. Each drop of *V̇*_{_air_out} resulted in a rise of pressure (2–4 bars) in the primary circuit due to a higher mean temperature in the system, particularly in the sink HX. This was compensated with the pressure control system feeding additional sCO_{2} by a booster compressor. At time 1450 s (6400 m^{3}/h, 0.32 kg/s), frequency of the main circulation pump started to stepwise decrease the *ṁ*_{_sCO2}. As consequence of the *ṁ*_{_sCO2} reduction, the inlet temperature to the sink HX *T*_{_sCO2_in} abruptly increased, until it reached its maximum limit 170 °C at 1820s, even though the air fan was switched back to its nominal 100%. The automatic control system switched off all heaters which were at this time almost at their maximum, i.e., H_{1/1}—28 kW, H_{1/2}—30 kW, H_{2}—26 kW, and H_{2}—20 kW. Switching off the electric heaters resulted in sudden drops of the temperatures and pressures in the system. However, there was some reaction time of the control system, and the inlet temperature to the sink HX was slightly exceeded. The controlled parameters are summarized in Table 7.

## Benchmark With Clara Numerical Code

### ClaRa Source Code Overview.

The pipe model includes equations derived from the general form of the conservation equations by the finite volume approach. The finite volume approach was used to derive a set of ordinary differential equations from partial differential equations, such that they can be implemented in a computer and numerically solved. In many situations (e.g., pipe model which is our case), it is reasonable to simplify models by restricting to one-dimensional mass flows which can be then spatially discretized and modeled by number of control volumes. For each control volume, we can write mass, momentum, and energy balance equations which are implemented in ClaRa.

### ClaRa Source Code Extension.

### Description of the Test Facility Implementation With ClaRa.

The dynamic sCO_{2} loop model includes all major components of the CVR test facility according to the P&ID. The main circulation pump MP is speed-controlled with preset input parameters. Heaters with PID controllers provide desired temperatures at the sink HX. The outlet temperature of cooler CH_{1} is handled with PID-operated water flow rate. The pressure in the system is controlled by feeding additional sCO_{2} (by a booster compressor) or releasing sCO_{2} through orifices, modeled in the computational model in a simple manner by the sCO_{2} source, and the PID controller. The air flow rate through the sink HX is handled with defined input.

### Results.

The main resulted parameters from both, the measurement and transient simulation, are shown in Fig. 14. They show fair agreement, demonstrating reasonable accuracy of the simulation tool. There is an evident deviation at the peak inlet temperature of sCO_{2} to the sink HX (by 13 K) leading to 3 bar pressure difference and 2 K discrepancy at the sink HX outlet. Apparently, this results from a smaller heat capacity of the numerical model than in reality. A faster temperature change (sCO_{2}) at the sink HX inlet justifies that. The model neglects all pipe supports, flanges, and bolts.

## Conclusions

This paper reports the performance tests of the supercritical air-cooled finned-type sink HX (tube Ø 12 mm x 0.7 mm) and presents a high quality numerical model. Altogether 34 measurement points were collected which were used for system code validation. Additionally, transients were logged, aiming to understand the energy and mass storage effects in the component.

The following conclusions can be drawn from the experimental results:

The pressure, mass flux, and temperature of sCO

_{2}have significant effects on the local heat transfer coefficient, especially near pseudo-critical region. The local heat transfer coefficient is decreased when cooling pressure is increased (for p_{sCO2}> 7.4 MPa) otherwise increased when mass flux is increased. The local heat transfer coefficient along the sink HX changes rapidly with the temperature of the fluid. It reaches a peak near the pseudo-critical temperature due to the highest heat capacity.The experimentally determined heat balances from the measured parameters on both sides (sCO

_{2}and air)*Q*__{air}and*Q*__{sCO2}are in good agreement (±15%) with each other.The results of calculated averaged overall heat transfer coefficients

*k*_{_calc_avg}using correlations (Gnielinski [18] for sCO_{2}and IPPE [23] or VDI [24] for the air) and experimentally determined values*k*_{_exp_avg}show for the performed tests reasonably low error of + 25% and −10%. Therefore, using the correlations for the estimation of the heat transfer in the sink HX with a similar design and similar conditions gives a fair error and thus is recommended. It is straightforward. Utilizing the measured data for look up tables for the HT of the sink HX is rather complicated to program.The analyzed correlations for heat transfer on the air side according to IPPE and VDI are in perfect match with each other.

The sink HX heat exchanger configuration is able to remove planned 95 kW under design conditions, 7.8 MPa, 166 °C/33 °C, 0.325 kg/s (for the sCO

_{2}side) and 24 °C (design is 25 °C), 3.65 kg/s for the forced air flow with ambient pressure.Air-cooled finned-tube sink HX is suitable for the sCO

_{2}-HeRo system.For a transient scenario—step-wise drop of ṁ_sCO

_{2}followed by loss of electric heating power, a Modelica code with newly implemented sink HX model was used. Simulation matches the measurement results well with mean deviations (*ṁ*_{_sCO2}5%,*V*̇_{_air_out}5%,*T*_{_sCO2_in}2%,*T*_{_sCO2_out}3%,*p*_{_sCO2_in}3%,*T*_{_air_out}3%).

## Acknowledgment

Authors thank Johannes Brunnemann and Timm Hoppe from XRG Simulation who provided insight and expertise of Modelica/ClaRa and wish to acknowledge the help of Martina Fruhbauerova with the final editing and proof read.

## Funding Data

European Union's Horizon 2020 research and training/research and innovation programme (662116/No 764690).

Ministry of Education, Youth and Sport Czech Republic – project LQ1603 Research for SUSEN.

## Nomenclature

*A*=area, m²

*c*=_{p}specific heat capacity, J·kg

^{−1}·K^{−1}*d*=diameter, m

*h*=enthalpy, J·kg

^{−1}*h′*=height of fin, m

*H*_{flow}=enthalpy flow, W

*K*=overall heat transfer coefficient, W·m

^{−2}·K^{−1}*L*=length, m

*ṁ*=mass flow rate, kg/s

*n*=number of fins of 1 tube

- Nu =
Nusselt number

*p*=pressure, Pa

*P*=electric power, W

- Pr =
Prandtl number

*Q*=heat power, W

- Re =
Reynolds number

*s*_{1}=pitch of tubes perpendicular to the air flow direction, m

*s*_{2}=pitch of tubes of HX above each other from the air flow sense, m

*s*_{3}=pitch of tubes behind each other (diagonal) from the air flow sense, m

*T*=temperature, K

*u*=gap between fins of 1 tube, m

*V̇*=volumetric flow rate, m

^{3}·s^{−1}*w*=velocity, m/s

- Δ
*p*=pressure drop, Pa

- Δ
*T′*=difference in temperatures of the mediums (air/sCO

_{2}) within one segment of a heat exchanger, K

### Greek Symbols

*α*=coefficient of heat transfer, W·m

^{−2}·K^{–1}*β*=auxiliary variable to calculate an efficiency of a fin

*δ*=thickness, m

- ζ =
pressure drop coefficient

*η*=dynamic viscosity, Pa·s

*η*_{fin}=efficiency of a fin

*λ*=thermal conductivity of a medium, W·m

^{−1}·K^{−1}*ρ*=density, kg·m

^{−3}*σ*_{cp}=error propagation of specific heat capacity, J·kg

^{−1}·K^{−1}*σ*_{h}=error propagation of enthalpy, J/kg

*σ*=_{m}error propagation of mass flow rate, kg·s

^{−1}*σ*_{Q}=error propagation of heat power transferred, W

*σ*_{ρ}=error propagation of density, kg·m

^{−3}*σ*_{V̇}=error propagation of volumetric flow rate, m

^{3}·s^{−1}

### Subscipts

- air =
air

- adv =
advection

- calc_avg =
calculated + averaged

- cross =
cross section

- e =
equivalent

- exp_avg =
experimentally determined + averaged

- fin =
fin of the heat exchanger

- fric =
frictional

- grav =
gravitational

- h =
hydraulic

- H
_{1/1}, H_{1/2, }H_{2}, and H_{3}=heaters H

_{1/1}, H_{1/2}, H_{2}, and H_{3} - Ideal =
ideal (e.g.,

*α*_{ideal}is coefficient heat transfer for*η*_{fin}= 1) - in =
inlet

- inner =
inner side (of tube/HX)

- out =
outlet

- outer =
outer side (of tube/HX)

- outer_tube_fin =
outer side among fins

- sCO
_{2}=supercritical CO

_{2} - tube =
tube of the heat exchanger

### Acronyms

- CAD =
computer-aided design

- CH
_{1}=water cooler

- CH
_{2}=oil cooler

- CVR =
Research Centre Rez

- EC =
electronically communicated

- GfS =
The Simulator Centre in Essen, Germany

- H
_{1/1}, H_{1/2}, H_{2}and H_{3}=electric heaters

- HT =
heat transfer

- HTR =
high temperature regenerative heat exchanger

- HX =
heat exchanger

- IPPE =
Institute of Physics and Power Engineering

- KKS =
identification system for power plants

- LMTD =
logarithmic mean temperature difference

- LTR =
low temperature regenerative heat exchanger

- LWR =
light water reactor

- MP =
main pump

- MCP =
main circulation pump

- NTU =
number of transfer unit

- P&ID =
piping and installation diagram

- PID =
proportional–integral–derivative

- PWR =
pressurized water reactor

- sCO
_{2}=supercritical carbon dioxide

- sCO
_{2}-HeRo =supercritical carbon dioxide heat removal system

- SG =
steam generator

- SS =
stainless steel

- SUSEN =
Sustainable Energy project

- TG =
turbine generator

- VDI =
VDI - Heat Atlas

### Appendix

When a function (e.g., enthalpy) is a set of nonlinear combination of the variables, an interval propagation could be performed in order to compute intervals which contain all consistent values for the variables. In a probabilistic approach, the function (e.g., enthalpy) must usually be linearized by approximation to a first-order Taylor series expansion.

_{2}enthalpies at the inlet and outlet of the sink HX were calculated with RefProp [28] as a function of two independent parameters, the measured temperatures and pressures. Therefore, the sCO

_{2}inlet temperature

*T*

_{_sCO2_in}, the outlet temperature

*T*

_{_sCO2_out}, the inlet pressure

*p*

_{_sCO2_in}and the outlet pressure

*p*_

_{sCO2_out}were used. Due to the reason, that the enthalpy equation from RefProp is not available, the above-mentioned standard deviation equation was simplified to following:

For the calculation of the sCO_{2} enthalpy uncertainty at the inlet of the sink HX $\sigma hsCO2\u2212in$ four enthalpies were used. The first one $h_sCO2\u2212inT_sCO2_in/p_sCO2_in_max$ was calculated with the measured sCO_{2} inlet temperature *T*_{_sCO2_in} and the maximum possible inlet pressure *p*_{_sCO2_in_max} = *p*_{_sCO2_in} + 0.11 MPa, the second one $h_sCO2\u2212inT_sCO2_in/p_sCO2_in_min$ with the measured sCO_{2} inlet temperature *T*_{_sCO2_in} and the minimum possible inlet pressure *p*_{_sCO2_in_min} = *p*_{_sCO2_in} – 0.11 MPa, the third one $h_sCO2\u2212inT_sCO2_in_max/p_sCO2_in$ with the measured sCO_{2} inlet pressure *p*_{_sCO2_in} and the maximum possible inlet temperature *T*_{_sCO2_in_max} = *T*_{_sCO2_in} + 1.75 K and the fourth one $h_sCO2\u2212inT_sCO2_in_min/p_sCO2_in$ with the measured sCO_{2} inlet pressure *p*_{_sCO2_in} and the minimum possible inlet temperature *T*_{_sCO2_min} = *T*_{_sCO2_in} − 1.75 K. The propagated sCO_{2} enthalpy uncertainty at the outlet of the sink HX $\sigma hsCO2\u2212out$ was calculated in the similar manner as for $\sigma hsCO2\u2212in$.