Startup and Operation of a Supercritical Carbon Dioxide Brayton Cycle

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1 Eric M. Clementoni 1 Bechtel Marine Propulsion Corporation, West Mifflin, PA Eric.Clementoni.contractor@unnpp.gov Timothy L. Cox Bechtel Marine Propulsion Corporation, West Mifflin, PA Timothy.Cox.contractor@unnpp.gov Christopher P. Sprague Bechtel Marine Propulsion Corporation, West Mifflin, PA Christopher.Sprague.contractor@unnpp.gov Startup and Operation of a Supercritical Carbon Dioxide Brayton Cycle Bechtel Marine Propulsion Corporation (BMPC) is testing a supercritical carbon dioxide (S-CO 2 ) Brayton system at the Bettis Atomic Power Laboratory. The 100 kwe integrated system test (IST) is a two shaft recuperated closed Brayton cycle with a variable speed turbine driven compressor and a constant speed turbine driven generator using S-CO 2 as the working fluid. The IST was designed to demonstrate operational, control, and performance characteristics of an S-CO 2 Brayton power cycle over a wide range of conditions. Initial operation of the IST has proven a reliable method for startup of the Brayton loop and heatup to normal operating temperature (570 F). An overview of the startup process, including initial loop fill and charging, and heatup to normal operating temperature is presented. Additionally, aspects of the IST startup process which are related to the loop size and component design which may be different for larger systems are discussed. [DOI: / ] Introduction The S-CO 2 Brayton cycle is being actively developed for potential application in a wide range of energy conversion applications. S-CO 2 Brayton system development efforts at BMPC have primarily been focused on system thermodynamics, system control modeling, and the design, construction, and testing of the 100 kwe integrated system test [1 3]. The IST is a simple recuperated closed loop S-CO 2 Brayton system with a variable speed turbine-compressor and a constant speed turbine-generator (Fig. 1). The IST is designed to generate nominally 100 kwe at a relatively modest turbine inlet temperature of 570 F (299 C) as shown in the design full power heat balance (Fig. 2). Due to the small scale of the IST equipment, overall loop efficiency is much lower than is predicted for larger S-CO 2 Brayton cycles. The IST component layout and physical arrangement of the test loop are shown in Figs. 3 and 4. The heat source for the IST is a 1 MW electrically heated organic heat transfer fluid system, which transfers heat to the CO 2 through a standard shell-and-tube heat exchanger. The heat sink is a chilled water system, which rejects heat from the precooler and other heat loads to a refrigerated chiller. This chilled water system is broken into two loops so that cooling flow can always be provided to auxiliary heat loads throughout the system while the precooler can either be cooled from this chilled loop or heated during startup through a separate water loop to achieve supercritical conditions in the CO 2. System power level is controlled by changing the speed of the turbine-compressor while operating the turbine-generator at a fixed speed. A compressor recirculation valve (CCV4) is utilized to maintain compressor surge margin by bypassing a portion of the compressor flow through the precooler rather than to the turbines. The turbine-compressor and turbine-generator designs incorporate a motor-generator within the pressure boundary (Fig. 5). Two reciprocating compressors are used in series to lower the pressure in the motor-generator cavities to reduce windage and bearing 1 Corresponding author. Contributed by the Cycle Innovations Committee of ASME for publication in the JOURNAL OF ENGINEERING FOR GAS TURBINES AND POWER. Manuscript received January 9, 2014; final manuscript received January 14, 2014; published online February 18, Editor: David Wisler. losses as well as provide supplemental cooling flow in addition to the CO 2 flow that passes through the shaft seals. IST Operation BMPC has developed system models and control algorithms to evaluate plant control during IST operations including startup, steady-state and transient operation, and shutdown [4,5]. The model is used to generate pretest predictions for IST operations and provide input to the test procedures before testing is performed. Test results are periodically incorporated into the model to update predicted system response and modify system operating parameters. The IST is highly instrumented with 204 instruments, 53 controllable devices, and 318 channels of data logging. Proportionalintegral feedback controllers are available on most controllable devices to maintain a desired setpoint. The IST is operated by a single operator at the IST operating station with only two normal operations requiring a second person to perform a manual action at the component. While not yet utilized, the control system has the ability to input a scripted set of control actions, which would allow the system to be operated without any operator involvement once the script was initiated. Loop Startup. To fill the IST with the required mass to achieve the full-power heat balance operating conditions, CO 2 vapor is introduced just before the second, high-pressure reciprocating compressor, which boosts the CO 2 to loop pressure. The loop is initially filled to a fixed mass to get near normal operating mass and then adjusted during loop heatup to meet a state point at a given operating condition based on model predictions. After the initial fill, no further mass adjustments are needed to operate between startup conditions and full power. However, smaller amounts of CO 2 may be added or bled to trim system mass or make up for leakage. The IST turbomachinery is only operated when the CO 2 in the main loop is above the critical point of 88 F (31 C) and 1070 psi (absolute) (7.4 MPa) to avoid two-phase conditions at the compressor inlet. When shut down and cold, the CO 2 in the loop is two-phase with a pressure equal to the saturation pressure of the coldest location in the loop. To warm the loop to supercritical conditions, heat is added to the main loop through the precooler and intermediate heat exchanger (IHX). The secondary water loop Journal of Engineering for Gas Turbines and Power JULY 2014, Vol. 136 / Copyright VC 2014 by ASME

2 Fig. 1 Simple recuperated Brayton cycle is heated using an electric water heater to achieve a temperature of 100 F (38 C) throughout the precooler. Approximately 40% of the CO 2 loop volume is contained in the two shell-and-tube heat exchangers that form the precooler, so this heating greatly increases loop pressure as CO 2 density is reduced. Additionally, the oil heater temperature is increased to 165 F (74 C) to preheat the hot section of the main loop. A small centrifugal pump is used to circulate 0.2 lbm/s (0.1 kg/s) of CO 2 through each of the two turbines and preheat the piping from the IHX to the turbine inlets. Increasing the temperature at the turbine inlets ensures that the density is in the proper operating range so that forward turbine flow will be established when starting the turbine-compressor and turbine-generator. While the CO 2 loop is being heated to establish the required conditions for starting the turbine-compressor, the cavity pressure in both motor-generator cavities is reduced below loop pressure and maintained at 600 psi (absolute) (4.1 MPa). Prior attempts to start the turbine-compressor with a lower pressure in the motorgenerator cavity were unsuccessful due to the additional drag created by two-phase fluid conditions within the motor-generator cavity. Upon startup of the turbine-compressor and turbinegenerator, the pressure within the motor-generator cavities is further reduced to 400 psi (absolute). The final requirement for turbine-compressor startup is for the CO 2 pressure at the compressor inlet to be nominally 1230 psi (absolute) (8.5 MPa) at the nominal startup temperature distribution in the loop. This starting pressure is chosen so that supercritical conditions are maintained throughout the turbine-compressor startup, fluid conditions at the turbine inlets promote positive turbine flow upon startup, and the loop has sufficient mass to reach the target compressor inlet pressure at full power. CO 2 is added to the loop as needed to make up for system leakage using the same system used for the initial mass charge. Once all loop conditions are achieved for startup, the circulating pump is secured and its flow path is isolated from the main CO 2 loop. The turbine throttle valve (CCV2) on the turbinegenerator is closed to prevent the turbine-compressor startup flow from causing the turbine-generator to spin at low speeds, which could create excessive turbine-generator bearing wear. To startup the main loop, the turbine-compressor is motored to an idle speed of 37,500 rpm. Approximately 30 s to 1 min later, the turbine-generator is motored to an idle speed of 37,500 rpm following confirmation of stable turbine-compressor operation and minimum system flows. The turbine inlet throttle valve (CCV2) for the turbine-generator is set to open once the turbinegenerator speed exceeds 20,000 rpm, which indicates a successful startup. Motor generator cavity pressures are decreased to 400 psi (absolute) to reduce windage and bearing heating within the cavities. Figure 6 shows the speed and motor power for the turbine-compressor and turbine-generator during a normal startup. Figure 7 shows the CO 2 loop flows and Fig. 8 shows the loop pressure distribution during a normal startup. During startup, 70% of the compressor flow goes through the recirculation valve (CCV4) with the remaining flow split between the two turbines. Additionally, at these low flow rates the piping and heat exchanger pressure drops are small and the turbines are the main system resistance. Following turbomachinery startup, the secondary water loop flow is secured and precooler flow is transitioned over to the primary chilled water loop to provide a heat sink for the CO 2 loop. The water system flow rate is modulated to maintain the CO 2 temperature at the turbine-compressor inlet at 100 F (38 C). The compressor recirculation valve (CCV4) is then repositioned to send more compressor flow through the two turbines. This loop Fig. 2 IST design full power heat balance / Vol. 136, JULY 2014 Transactions of the ASME

3 Fig. 3 IST component arrangement Fig. 4 IST physical layout Fig. 5 IST turbomachinery internals state is referred to as cold idle and represents a condition that can be maintained indefinitely without the risk of damaging the turbomachinery. Heatup to Normal Operating Temperature. Heater power is increased to raise the temperature of the heat transfer fluid from 165 F (74 C) to the temperature required to achieve a turbine inlet temperature of 570 F (299 C). Due to thermal stress limitations of the IHX, this heatup is limited to a maximum rate of 200 F/h (111 C/h). As the loop heats up, the power produced by the turbines increases accordingly due to the favorable thermal hydraulic conditions. When the turbine-generator thermal hydraulic power exceeds that required to motor the shaft, the turbinegenerator controller transitions from motoring to generating. Figure 9 shows the motoring and generating power of the turbine-compressor and turbine-generator as a function of turbine inlet temperature with both units operating at 37,500 rpm. The turbine-generator transitions from motoring to generating with a turbine inlet temperature of 290 F (144 C). A small increase in motor power is seen shortly before the transition to generating due to increased windage losses as turbine-generator rotor cavity conditions are changed to maintain rotor cavity temperatures within desired operating limits. Once the turbine inlet temperatures get above 300 F (149 C), the compressor CO 2 inlet temperature is reduced to 96 F (36 C) for the remainder of the heatup to normal operating temperature. This reduction in compressor inlet temperature causes an increase in compressor inlet density and flow rate, which increases the turbine power generation faster than when limited to turbine thermal conditions alone. Once the turbine inlet temperatures are at the design inlet temperature of 570 F (299 C) with the turbomachinery operating at the idle speed of 37,500 rpm, the loop condition referred to as hot idle is achieved. Figure 10 shows the system heat balance at hot idle operating conditions. As shown in Figs. 9 and 10, the turbomachinery has a net power of approximately 0 kwe at hot idle operating conditions with turbine-generator output nearly equal to the motoring power of the turbine-compressor. Turbine-generator power output of 5.3 kwe at hot idle operating conditions is approximately half of what was predicted at this operating condition due to higher than expected windage and power conversion losses. Normal Power Generation. For power generating operations, the speed of the turbine-generator is increased from idle to the normal operating speed (65,000 or 75,000 rpm depending on Journal of Engineering for Gas Turbines and Power JULY 2014, Vol. 136 /

4 Fig. 6 IST turbomachinery startup Fig. 8 IST loop pressure distribution during startup Fig. 7 IST loop mass flow rates during startup method of speed control), and the turbine-compressor speed is adjusted to achieve the desired power level of the system. The rate at which the system power level is changed is controlled by limiting the rate of change of the power level setpoint in combination with the rate of change of the turbine-compressor shaft speed. Initially, the turbine-compressor speed is set based on a lookup table that uses the desired power level to determine the turbinecompressor speed setpoint based on model predictions, and the resulting actual generator power is measured and recorded. Normal operation will employ an updated lookup table based on the actual system operating data obtained in this initial testing. Off-Nominal Power Generation. In addition to normal operation, the IST has been tested at off-nominal conditions through a range of turbine-compressor speeds in order to confirm proper speed control of the turbomachinery. This testing was performed at a heat source heat exchanger CO 2 outlet temperature setpoint of 540 F (282 C) with the turbine-generator at a fixed speed of 55,000 rpm. Turbine-compressor speed was increased incrementally from idle speed (37,500 rpm) to 52,000 rpm. As the turbinecompressor speed is increased, the system mass flow rate and pressure ratio increase, which increase turbine-generator power generation. Data from this test is shown in Table 1 with a peak Brayton power generation of 17.6 kw. Operation at higher turbine-compressor speeds has resulted in unstable turbinegenerator operation, which is being investigated at the time of writing this paper. The issue is associated with the ability of the Fig. 9 Turbomachinery powers during heatup with both shafts at idle speed of 37,500 rpm motor-generator controller to maintain stable voltage regulation and commanded speed setpoint and is being evaluated with the turbomachinery vendor to improve system performance. This issue is not related to the characteristics of the S-CO 2 Brayton cycle but rather software and tuning related issues within the controller. The turbine-compressor is motoring during this test, reducing the net Brayton cycle power relative to turbine-generator output. The compressor recirculation valve (CCV4) position could be adjusted to increase the turbine mass flows, which would balance the compressor turbine and compressor powers to get a zero net turbine-compressor power and also increasing turbine-generator power, thereby increasing net Brayton power and cycle efficiency. Operating at a zero net turbine-compressor power is the target for normal power generation. The measured turbine-generator electrical power is approximately 20 kwe less than the calculated generator turbine power using the turbine state points included in Table 1. At 55,000 rpm and the cavity pressure utilized for this testing, the turbinegenerator windage losses are predicted to be approximately 9 kw. The remaining differences are being investigated to determine if there are additional loss mechanisms or if the measured electrical power output is lower than the actual power output of the turbinegenerator based on the method of measuring power. The compressor inlet pressure data shown in Table 1 indicates that compressor inlet pressure generally increases with increasing / Vol. 136, JULY 2014 Transactions of the ASME

5 Fig. 10 System heat balance at hot idle Table 1 Off-nominal power generation Turbine-generator speed (rpm) 55,000 55,000 55,000 55,000 Turbine-generator power (kw) Generator turbine flow (lbm/s) Generator turbine inlet temperature ( F) Generator turbine outlet temperature ( F) Generator turbine inlet pressure (psi (absolute)) Generator turbine outlet pressure (psi (absolute)) Turbine-compressor speed (rpm) 37,500 45,000 50,000 52,000 Turbine-compressor power (kw) Compressor turbine flow (lbm/s) Compressor turbine inlet temperature ( F) Compressor turbine outlet temperature ( F) Compressor turbine inlet pressure (psi (absolute)) Compressor turbine outlet pressure (psi (absolute)) Compressor flow (lbm/s) Compressor inlet temperature ( F) Compressor outlet temperature ( F) Compressor inlet pressure (psi (absolute)) Compressor outlet pressure (psi (absolute)) Net Brayton power (kw) IHX heat transfer to CO 2 (kw) Brayton efficiency 0% 4% 4% power level. This is due to the IST main Brayton loop not being a fixed mass system. As the turbine-compressor speed is increased, the windage heating increases the temperature of the CO 2 within the turbine-compressor motor-generator cavity and downstream tubing, thereby, reducing the density of the CO 2. Additionally, the cavity pressure is reduced with increasing turbine-compressor speed to reduce windage losses, which further reduces the density in the secondary CO 2 systems. As the CO 2 density in this part of the system is decreased, the mass is shifted to the main Brayton loop and, thereby, causes the compressor inlet pressure to increase. Normal operation through the entire power operating range would be performed with a constant cavity pressure and varying cavity cooling flow rates to maintain a relatively constant turbine-compressor cavity density and a fixed mass in the main Brayton loop. Loop Shutdown. The IST is normally shut down by first cooling the system back down to cold idle operating conditions. While the IST could be shut down from high temperatures, the turbomachinery needs to be cooled below 200 F to protect the motorgenerator components. A forced cooldown with the system operating is a much faster method of removing heat from the loop than allowing it to cool naturally. Cooling the system to cold idle also allows the auxiliary systems to be shut down as soon as the turbomachinery is shut down. During the cooldown, the turbinegenerator transitions back to motoring, and the loop behavior follows the heatup in reverse. Once the system is cooled down to a turbine inlet temperature of 165 F (74 C), the turbine-compressor and turbine-generator are simultaneously shutdown from 37,500 rpm by removing motor power and applying braking to the shafts. Figure 11 shows the rapid deceleration of the turbine-compressor and turbinegenerator shafts during a normal turbomachinery shutdown. At this point, the heaters are de-energized and cooling systems secured. The system is now completely shut down and can be walked away from within 10 min of turbomachinery shutdown. IST Operations Relative to Larger Systems Many of the IST operating conditions and limitations described in this paper are specific to IST equipment and may not apply to larger scale applications of the S-CO 2 Brayton cycle. The use of gas foil bearings in the turbine-compressor and turbine-generator result in many startup and operational procedures that may not be needed for larger scale turbomachinery with more capable Journal of Engineering for Gas Turbines and Power JULY 2014, Vol. 136 /

6 the time it takes to go from cold shutdown to normal operating temperature and full power. The IST turbomachinery motor-generators are located within the pressure boundaries and, thus, operate in a high pressure CO 2 environment. Abradable shaft seals to limit leakage into the cavities and auxiliary compressors are used to reduce the pressure in the motor-generator cavities relative to the CO 2 loop pressure; yet, the density of the CO 2 within the cavity is still times that of air at atmospheric pressure, which greatly increases windage losses. As a result, approximately 10% of the power produced by the turbine at full power is predicted to be consumed by windage and bearing losses. Better seals would allow the motorgenerator to be located outside the pressure boundary, providing better system efficiency and less operational complexity than is associated with motor-generator cavity pressure and temperature control. Fig. 11 Turbomachinery shutdown bearings. Startup of the IST turbomachinery requires that the turbine inlets be preheated to achieve positive turbine flow and maintain thrust loads within thrust bearing capacity. A system with a more capable thrust bearing and/or a different turbine and nozzle design may not have this same need. Additionally, the gas foil radial bearings have a minimum liftoff speed of approximately 6000 rpm, and the gas foil thrust bearings have limited thrust load capability below 35,000 rpm. As a result, the turbomachinery is rapidly motored up to a minimum speed of 37,500 rpm during startup. More capable bearings may allow startup of the turbinegenerator via loop thermal hydraulic conditions with an operating turbine-compressor or could support other startup methods, which do not require motoring of the turbomachinery. The IST is only operated when the CO 2 at the compressor inlet is above the critical point, with normal operating conditions of 96 F (36 C) and 1340 psi (absolute) (9.2 MPa). This operating point was chosen to keep the fluid at the compressor inlet in the supercritical regime for all anticipated loop transients and to minimize the density change associated with small offsets from the normal operating conditions. Operating the system closer to the critical point or as a condensing cycle with two-phase fluid in the loop could improve system efficiency but make system control during transients more difficult. An S-CO 2 Brayton cycle could also be designed to be started with two-phase conditions and then heated to supercritical throughout the system which could enable a faster loop startup. The maximum heatup rate of the IST is currently limited to 200 F/h (111 C/h) due to thermal stress limitations of the shelland-tube IHX. Due to this heatup rate limit, the IST takes over 4 h to startup and heatup from a cold shutdown state to the normal operating temperature of 570 F (299 C). Once normal operating conditions are achieved, the power level can be quickly increased to full power, limited only by turbomachinery shaft acceleration limits and loop thermal hydraulics. A larger scale system could be designed with higher heatup rate capability, drastically reducing Conclusions Operation of the integrated system test has proven that a supercritical CO 2 Brayton cycle can successfully be controlled and operated through startup, heatup, power generation, and shutdown. While many of the turbomachinery technologies utilized in this 100 kwe test loop may not be prototypic of what would be used in larger scale systems, the basic system control and operation are independent of size. Testing to date has not found any inherent issues associated with the supercritical CO 2 Brayton cycle. Further testing is planned which would lead to achieving full power operation independent of motoring the turbomachinery and demonstrating the potential of the S-CO 2 Brayton cycle as a viable power conversion technology. Acknowledgment This paper summarizes work that has been done by a number of devoted engineers, scientists, and support personnel at the Bettis Atomic Power Laboratory, Knolls Atomic Power Laboratory, and our subcontractors. This paper would not be possible without the efforts of this team. The submitted manuscript has been authored by a contractor of the U.S. Government under contract No. DE-NR References [1] Ashcroft, J. A., Kimball, K. J., and Corcoran, M. R., 2009, Overview of Naval Reactors Program Development of the Supercritical Carbon Dioxide Brayton System, Supercritical CO 2 Power Cycle Symposium, Troy, NY, April [2] Kimball, K. J., 2011, Overview of Naval Reactors Program Development of the Supercritical Carbon Dioxide Brayton System, Supercritical CO 2 Power Cycle Symposium, Boulder, CO, May [3] Kimball, K. J., and Clementoni, E. M., 2012, Supercritical Carbon Dioxide Brayton Power Cycle Development Overview, ASME Turbo Expo 2012, Copenhagen, Denmark, June 11 15, ASME Paper No. GT [4] Hexemer, M. J., Hoang, H. T., Rahner, K. D., Siebert, B., and Wahl, G. D., 2009, Integrated Systems Test (IST) S-CO 2 Brayton Loop Transient Model Description and Initial Results, Supercritical CO 2 Power Cycle Symposium, Troy, NY, April [5] Hexemer, M. J., 2011, Supercritical CO 2 Brayton Cycle Integrated System Test (IST) TRACE Model and Control System Design, Supercritical CO 2 Power Cycle Symposium, Boulder, CO, May / Vol. 136, JULY 2014 Transactions of the ASME