Steam  Car  Control  Analysis
                      E.A. Mayer
                  G.W.  Hurlong. Jr.
                      The Bendix Corporation
                       Research Laboratories
                       Southfield, Michigan

                        Final Report
                         July 1972
                          Prepared for

                   Steam Engine Systems Corporation
                      Watertown, Massachusetts
                    Prime Contract No. 68-04-0004
           Division of Advanced Automotive Power Systems Development
                    Environmental Protection Agency
                       Ann Arbor, Michigan
                                                  Copy No.

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Steam  Car  Control Analysis
                  E.A. Mayer

               G.W. Hurlong. Jr.
                  The Bendix Corporation
                   Research Laboratories
                   Southfield, Michigan


                    Final Report

                     July 1972
                     Prepared for


                Steam Engine Systems Corporation
                  Watertown, Massachusetts

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                            ACKNOWLEDGMENTS
The authors gratefully acknowledge assistance from the following people:

     Dr. Lawrence C. Hoagland and Dr. Joseph Gerstman of Steam
     Engine Systems Corporation, Watertown, Massachusetts, for
     their specific and detailed involvement in the development
     of the steam-generator analytical model and their continued
     review of the entire technical effort;

     Mr. Lewis J. Plumley of the Electrodynamics Division of The
     Bendix Corporation, North Hollywood, California, for the
     analysis and design of the electric servo actuators used in
     the facsimile and control-mode test experiments;

     Mr. James M. Kirwin and Mr. Elmer A. Haase of the Energy
     Controls Division of The Bendix Corporation,  South Bend,
     Indiana, for the design of the RSA Fuel Controller;

     Mr. Ray J. Brown of the Bendix Research Laboratories, South-
     field, Michigan, for the programming of the hybrid computer.

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                            TABLE OF CONTENTS
                                                                    Page

SECTION 1 - SUMMARY                                                 1-1

1.1   Program Definition and Summary of Results                     1-1
1.2   Conclusions                                                   1-3
1.3   Recommendations                                               1-4

SECTION 2 - PROTOTYPE CONTROL DESIGN                                2-1

2.1   Historical Background                                         2-1
2.2   Predictive Flow Control Concept                               2-3
2.3   Predictive-Control Concept Implementation with
      Variable-Speed Auxiliary Drive                                2-7

SECTION 3 - FACSIMILE CONTROL DESIGN                                3-1

3.1   Facsimile Control Concept                                     3-1
3.2   Burner Control                                                3-4
3.3   Fuel Control                                                  3-20
3.4   Bypass Valve                                                  3-31
3.5   Facsimile Design Verification Test Results                    3-39

SECTION 4 - ANALYTICAL CONTROL EVALUATION                           4-1

4.1   Hybrid Computer Model                                         4-1
4.2   Experimental Verification for Model SES-4                     4-25
4.3   Prototype Results                                             4-34

NOMENCLATURE                                                        4-59

SECTION 5 - CONTROL MODE SIMULATION EXPERIMENTS                     5-1

5.1   Experimental System Description                               5-1
5.2   Simulation-Hardware Design Test Results                       5-3
5.3   Closed-Loop Steam Generator Test Results                      5-6

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                               SECTION 1

                                SUMMARY


1.1   PROGRAM DEFINITION AND SUMMARY OF RESULTS

      Final Objective
      The final objective of the entire program is the development of a
control system for a steam-car power plant.  The achievement of the over-
all operational simplicity, safety and convenience that is currently
available in present-day spark-ignition automobiles serves as the guide-
line for design of the control system.  Power plant controls are to be
fully automatic with the only input required of the driver to be the
operation of the accelerator pedal.  The operating control for the power
plant must then control not only the input of steam to the expander,
but also automatically control the feedwater and heat inputs to the
steam generator to satisfy the operator's power demand.  The objectives
of the program phase covered by this final report were the selection of
the best control mode and experimental verification of the soundness of
the concept through pre-prototype control hardware.

      Accomplishments
      The program plan for the 16-month program is shown in Figure 1-1.
The plan denotes a good balance between analytical and experimental pro-
grams aimed at the fulfillment of the program objectives stated.  An
early examination of control concepts adaptable to the needs of a vehic-
ular steam power plant indicated the need for new development.  To fill
this need, a new control concept, the predictive flow-control system,
was proposed for the steam car.  The salient features of this controller
include the use of the expander as a positive-displacement machine to
give an indication of instantaneous steam-flow requirements.  A strong
open-loop control system then adjusts both the burner power level and
the feedwater supply system to satisfy this demand.  Secondary closed-
loop controls correct any deviations in the output pressure and tempera-
ture of the steam generator from the desired set points.  The corrections
are accomplished by an appropriate trim-control function in both the
burner and feedwater supply systems.

      A wide-range, analytical model of the vapor generator was combined
with an analytical description of the expander and all auxiliaries to
form a hybrid-computer simulation,.  This computer simulation then served
as the design tool for the evaluation of the various control concepts.
In order to increase the credibility of the analytical results, initial
results were compared with available experimental data on the Steam
Engine Systems Model-4 vapor-generating system.  The experimental
                                                                     1-1

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I
N3
STARTING DATE: MARCH 15. 1971
TASK1
COMPUTER MODEL DEVELOPMENT
DEVELOP COMPUTER MODEL OF STEAM SYSTEM
REPROGRAMMING AND OPEN LOOP CHECK
MODIFY MODEL TO COMPLY WITH SES DATA
TASK 2.
CONTROL STUDIES
STUDY CONTROL MODES
PROGRAM CLOSED LOOP
SELECT CONTROL MODE
TASK 3.
CONTROLS HARDWARE PRELIMINARY DESIGN
DEVELOP PRELIMINARY HARDWARE SPECIFICATIONS
DESIGN PRELIMINARY HARDWARE
FINALIZE CONTROL HARDWARE DESIGN
COMPLETE DESIGN
TASK 4.
CONTROL MODE SIMULATION HARDWARE
SELECT CONTROL SIMULATION HARDWARE
ANALYZE AND INTERPRET CLOSED LOOP TEST RESULTS
TASK 5.
FACSIMILE HARDWARE
SYSTEM DESIGN
FEEDWATER BYPASS CONTROL VALVE DESIGN
MANUFACTURE
TESTING
FUEL CONTROLLER
EVALUATION OF RSA UNIT
DESIGN INTEGRATION
TESTING OF INTEGRATED UNIT
AIR THROTTLE VALVE
DESIGN
FABRICATE
TEST
DELIVER FACSIMILE HARDWARE TO SES
REPORTING
MONTHLY REPORTS
QUARTERLY REPORTS
FINAL REPORT
SUBMIT DRAFT TO SES
1971
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                                                 Figure 1-1 - Program Plan

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verification included both steady-state and dynamic comparison between
analytical predictions and known hardware performance.  In addition to
these open-loop verifications, a closed-loop control-mode simulation
test program also was conducted to verify the soundness of the proposed
control concept.

      The facsimile control-system hardware was specially designed to
allow a more extensive evaluation of the proposed control mode without
the development costs essential for an in-car installation.  The facsimile
controllers made maximum use of commercially available components to test
the concept.  An electronic analytical signal processor was used to pro-
vide the necessary computations for the predictive signals.  In addition,
a laboratory-type feedwater meter was included in the vapor-generator
feedwater system to predict the power level without relying on the expan-
der for this function.  A commercially available fuel controller was
integrated with the system to provide the primary fuel-metering function
to the burner.  In the facsimile hardware program a feedwater bypass
valve was also developed.  This unit will serve to provide experimental
design information for future, more advanced feedwater control hardware.

      The program, thus concluded, will provide baseline information and
the design tools necessary for the determination of design details needed
for the development of the in-car control system.

1.2   CONCLUSIONS

      On the basis of the completed program, the following conclusions
are justified:

      (1)  Conventional closed-loop control of temperature and pres-
           sure is not satisfactory for the steam generator system.
           The complex interrelationship between pressure and tem-
           perature in the boiling section are combined with extremely
           long and variable response time in the vapor generator
           system.  This combination becomes a source of control
           instability, particularly during operation at the low
           power levels and gain settings essential for safe and
           efficient operation of the power plant.

      (2)  The predictive control system with a closed-loop trim con-
           trol in the temperature and pressure circuits can provide
           a satisfactory power plant operation.

      (3)  The analytical model of the power plant has been verified
           by comparison with experimental data obtained by Steam
           Engine Systems Corporation.  The model has sufficient
           steady state and dynamic accuracy for detailed design
           of the control system.
                                                                     1-3

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       (4)  The simplified hardware implementation of the  predictive-
           control  system does not provide satisfactory control over
           a  20:1 range of steam flow.  To date, the control studies
           have not identified all the details of a system capable
           of operation over a flow range of at least 20:1.
       (5)  The program completed to date has developed an analytical
           tool that  offers a realistic evaluation of details of the
           control  design.  The tradeoff studies conducted during
           the program also indicate  that the optimization of the
           control  concepts depend very strongly on component perfor-
           mance characteristics.  In other words, the complexity
           of control can be traded against component complexity.
           For example, the possible  choice of a variable-displacement
           feedpump can reduce the control complexity through the
           elimination of the variable-ratio auxiliary drive and its
           control  system.

1.3   RECOMMENDATIONS

      On the  basis  of the experimental and analytical control studies
conducted during the  program covered by this report, a significant first
step was made toward  the successful proportional control of Rankine-
cycle power plants.   The most significant achievement, the establishment
of a useful,  analytical model of the Rankine-cycle power plant, is a
basis for recommendations for continued work.  The use of this tool
offers a considerable savings both in time and development costs for
the design of the Rankine-cycle power plant controls.  The following
specific recommendations can be made:

      (1)  Continue to use the analytical model as a design tool
           for the  controls as the component selection is finalized.
           A  blend  of analytical and experimental programs, such as
           the one  carried out during the completed phase of the
           program, should be continued for the purpose of maintaining
           a  high confidence level in the design of the control system.

      (2)  Optimize the control concept for the selected hardware
           configuration and extend operation to meet the 20:1
           operating  range.

      (3)  Expand the initial facsimile control work and develop
           the required wide-range fuel-air ratio control for an
           external combustor system.
      (4)  In conjunction with development of the operating controls,
           develop  the automatic startup and shutdown control system.
      (5)  Investigate alternate feedwater control concepts in
           addition to the present feedwater control concept.
 1-4

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(6)  Investigate alternate control systems  that would not
     require the use of the variable-ratio  drive for the
     auxiliaries.
(7)  Develop an electro-pneumatic servo for in-car combustion-
     air damper operation.
                                                              1-5

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                               SECTION 2

                        PROTOTYPE CONTROL DESIGN
      Examination of a steam-car bibliography indicates that the on-off
type burner and feedwater controls dominated the early designs.   Some
references can be found describing proportional control.  Valuable exper-
imental performance evaluation reports were not available.  On the basis
of the adaptation of controls used on modern low-water-inventory steam
generators, a predictive metering control is proposed in this program
for the modern steam car.  This control concept generates a measure of
the instantaneous steam demand of the expander.  On the basis of the
steam flow information available, a strong predictive controller adjusts
the burner firing rate to the known steady-state requirement of  the steam
generator plant.  The feedwater flow input is also matched by the con-
troller to the steam rate of the expander.  Deviation in steam-generator
output temperatures and pressures from the desired set points are inde-
pendently corrected through closed-loop controls in the burner firing
rate and feedwater supply systems respectively.  Because the large var-
iations in the variables are corrected by the open-loop predictive con-
troller, the lower gains available in the closed-loop trim-control circuits
can satisfactorily correct for deviation in plant characteristics and
dynamic responses of the temperature and pressure control systems.  De-
tailed implementation on the computer were evaluated using the expander
as the flowmeter and an engine-driven, fixed-displacement, feedwater
pump and combustion-air blower.  The predictive control, in this case,
was implemented through the use of a limited-range, variable-speed drive
between the expander and the auxiliaries.  Closed-loop pressure-trim
control was accomplished through the use of a feedwater bypass valve.
The temperature trim control was implemented through a combustion-air
throttle valve.  An independent fuel controller was used to maintain the
air-fuel ratio at the desired level.

2.1   HISTORICAL BACKGROUND

      The difficulty experienced by early steam car designers in automatic
control of the monotube boiler becomes apparent on an examination of ava-
ilable bibliographies, such as the references listed at the end  of this
section.  In addition to the slow thermal response from the boiler, the
strongly intercoupled thermodynamic relations between pressure and tem-
perature eliminate the possible use of the simple, independent,  pressure
and temperature controls.  In the following, three typical monotube
steam-generator control concepts will be reviewed.

      Late-Model Doble Steam-Car Controls

      Probably the most successful, and also most sophisticated, fully
automatic boiler controls of early steam cars were developed by Abner
                                                                    2-1

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Doble.  Several versions of controls can be found on the various models
of Doble cars.  A typical late control design uses an exhaust-steam-driven
combustion-air blower, a feedpump and a condenser fan.   The control is
accomplished through the use of solenoid valves operated by a combination
of pressure and temperature switches.  The burner control is operated
through a pressure sensor.  At the low pressure limit (typically 1400
psi), the burner is turned on.  As the steam pressure rises to the high
limit setting (typically 1700 psi) of the burner control switch, the
burner is turned off.  An additional thermostat switch is set at the high
temperature limit (such as 875ฐF) and the burner is turned off regardless
of the value of the steam generator pressure.  The steam-operated feedpump
is controlled through a solenoid valve.  The pump is turned off at the
minimum temperature setting, such as 840ฐF.  To improve the dynamic res-
ponse of the steam generator, Doble uses a normalizer which injects
feedwater near the entrance of the superheater at the time the feedwater
flow to the boiler is increased.  The flow through the normalizer helps
to reduce the response time of the boiler.  The fuel flow is controlled
automatically through a simple carburetor equipped with a venturi injec-
tion supplied from a level-regulated fuel bowl.  Further detailed descrip-
tions of this and other Doble steam-car control concepts may be found in
Reference 1.

      The General Motors SE-101 Steam-Car Control

      One of the most recent steam-car control designs can be found on
the GM SE-101 Steam Car as reported in Reference 2.  This control system
includes a fully automatic turn-key starter sequence control.  The operat-
ing control system is a three-level controller.  This system uses a fixed-
displacement engine-driven pump and solenoid bypass valves in the feed-
water control circuit.  The combustion air is supplied by an expander-
driven combustion-air blower.  The fuel system includes a fuel pump and
a bypass solenoid.  Air-fuel ratio essentially is controlled by the
common mechanical drive of combustion-air blower and fuel pump.  A three-
level control mode, consisting of off, low and high outputs, is employed
both in the combustion control and the feedwater control systems.  Depend-
ing on a decision matrix, considering both output pressure and temperature,
an electric decision network selects the appropriate level of operation
for both the burner and the feedwater control.  The use of a multilevel
controller offers a more continuous operation of the steam generator and
a better matching of steam output to the requirements of the vehicle.

      Late-Model White Steam-Car Control System

      One of the few attempts at proportional control of the monotube
steam-car boiler is that of the late White system.  The burner fuel supply
is set to be proportional to the feedwater input to the steam generator
by the "flowmotor."  The flowmotor is a piston-spring combination.  The
differential pressure to the piston actuator is supplied from a pressure
drop across  a metering orifice in the feedwater system and correction is
applied to the feedwater system.  The steam-temperature correction signal
2-2

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is applied to the flowmotor as a separate input.  Performance data, des-
cribing steady-state or dynamic operation of the White flowmotor system,
was not found in any of the references.

      These examples of steam-car control typify the development status
of steam car controls.  With emphasis placed on emissions, and also on
vehicle performance, the capabilities of the On-Off controls fall short
of the requirements.  A continuously variable proportional control for
the combustor becomes an essential part of the low-emission steam car.
An examination of the controls used on modern low-water-inventory, high-
response boilers indicates the use of highly sophisticated controls,
expecially tailored to the operating characteristics of the boiler.  In
general, the fast-response controllers often employ some form of a con-
trol system that utilizes information defining the instantaneous operating
level.  This is often done by defining the steam output of the boiler and
using this for at least one or both of the feedwater and thermal input
controls.  Using this background as the baseline, a predictive metering
concept was developed for steam car control.  This control concept is
described in the following section.

2.2   PREDICTIVE FLOW CONTROL CONCEPT
      A major requirement for the design of the steam-car control system
is the achievement of the operational simplicity, safety and convenience
that is currently available in present-day spark-ignition automobiles.
In order to achieve this goal for the Rankine-cycle power plant, three
types of controls will be essential:
      (1)  The operating control,
      (2)  Safety controls, and
      (3)  Automatic start and shutdown sequence controls.

It is apparent that the design of the automatic start and shutdown se-
quence controller is contingent on a near complete definition of the
operating controls if maximum economy is to be realized.   Safety controls
will again have to interface with the operating controls.  For this rea-
son the control design effort to date has been focused on the development
of the operating controls for the steam car.
      The accelerator pedal is to be the only normal input required of
the driver to control the Rankine-cycle power plant.  The operating con-
trol system then must operate not only the steam input to the expander
to produce the required accelerating torque from the power plant, but
also automatically control the feedwater supply and heat input to the
steam generator to satisfy the operator's demand.  In order to accomplish
this operating convenience, the control system must:

      (1)  Provide proper steam pressure, temperature and flow
           output from the boiler at all times.
                                                                     2-3

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        (2)  Regulate feedwater supply to the boiler and properly
            modulate the burner output.
        (3)  Maintain the burner air-fuel ratio  at all times^at the
            desired value to reduce exhaust emissions and maintain
            burner efficiency.
        (4)  Control the function of the condenser to keep up with
            continually varying steam output flow rejected by the
            expander.
        (5)  Satisfy the safety requirements at all times during
            the operation of the power plant.
        (6)  Have the capability for automation of the startup and
            shutdown functions to the extent that only turn-key
            operation is required of the vehicle operator.
        The general concept of the metering control system is shown in
  the  schematic of Figure 2-1.  The essentials of the control concept
  include the use of the expander to determine the steam flow consumption.
  The  steam flow signal thus generated is used to control the feedwater
  flow to the steam generator and also the combustion air flow to the
  burner at the proper level.  The burner fuel flow is controlled indep-
  endently on the basis of combustion air flow, thus maintaining an air-
  fuel ratio dictated by the combined requirements of low emissions and
  high burner efficiency.   Several mechanizations of this basic concept
  are  possible.  The one shown in Figure 2-1 is an early concept of mech-
  anization and utilizes an expander-driven feedpump and electrically
  driven combustion blower and combustion-air atomizing pump.  The basic
  control concept readily can accommodate changes in auxiliary drives.
  The  schematic of Figure 2-1 includes all functions necessary for the
  complete automatic control of the steam generator plant.   A more detailed
 discussion of a selected implementation of the predictive flow metering
 concept is described in  the next section.
       During the program period, operating characteristics of the pre-
 dictive steam-generator  control system were evaluated both through ana-
 lytical and  experimental  means.   Details of the experiments are described
 in Section 5  under Control  Mode Simulation Experiments.   Typical results
 from this  program are  compared  with the analytical predictions in Fig-
 ure 2-2 to show the  feasibility of the  control concept just described.
 The figure shows  closed-loop results of the predictive control concept
 when  a  step demand is made  in throttle  setting.   The results  indicate
 that  successful  control  of  the  monotube steam generator is possible in
 spite of the  relatively  slow thermal response of the steam generator.
 The experimental  setup used the same gain  setting for the corrective
 trim  temperature  control  loop that was  predicted through  the  prior com-
 puter simulation  of  the  system.   The results shown are typical of the
 performance capabilities  of the predictive controller.  The controller
 is  providing an approximate temperature control  band of 75ฐF.
2-4

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                                                                    EXHAUST  STEAM
                                                                                                    CRANKCASE
                                                                                                      VENT
                                                                                        VARIABLE
                                                                                         CUTOFF
                                                                                        CONTROL
                                                                                  FEED I (FIXED
                                                                                  PUMP I CAP'Y)
NJ

Ul
                                            Figure 2-1  - Metering Control  System Schematic

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ho
                                                                     100
                                                               TIME, SECONDS
                                                                                           150
                                                                                                                 200
                                                            Experimental Results
                                                  SES-4 Steam Generator Manual Pressure Control
                                                 Predictive Temperature Control Nominal Setting
+20U-
+ 100
AT 0-
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                                              50
      100
TIME, SECONDS
150
200
                                                           Hybrid  Computer Results
                                                SES-5 Steam Generator Bypass Pressure Control
                                               Predictive Temperature Control Nominal Setting

                            Figure  2-2  - Effect of  the Step Change From 39 to 56  Percent  Flow Level
                                          in Steam Throttle Setting on Superheater Output  Temperature

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2.3   PREDICTIVE-CONTROL CONCEPT IMPLEMENTATION WITH
      VARIABLE-SPEED AUXILIARY DRIVE

      The flow metering concept described functionally in the previous
section can be implemented in many ways.  The choice of power source
for the feedpump and combustion-air blower has a significant effect on
the type of controls that are practical.  In general, several horsepower
are required for auxiliary drives.  Thus, the use of electric drive, al-
though it would be desirable for controls, becomes bulky, because not
only the electric motors are needed, but a significant increase in the
size of the alternator used must be included in the package to provide
the electric power generation in the steam car.  Two other power sources
remain attractive:  mechanical power derived directly from the expander,
or the steam output derived directly from the steam generator or from
the exhaust of the expander.  The use of mechanical power from the ex-
pander was selected for the initial concept evaluation.  The use of steam-
driven auxiliaries may be an attractive choice; however, at the present,
the development status of these auxiliaries does not appear to fit the
overall program timetable.
      The schematic diagram of the predictive flow metering control con-
cept is shown in Figure 2-3.  The schematic includes only the operating
controls.  Both the feedpump and the combustion air blowers are driven
by the expander through a variable-ratio transmission.   The predictive
control is limited to setting of the transmission ratio.  The pressure
trim control is accomplished through the use of a pressure-operated
feedwater bypass valve.  The temperature trim control on the superheater
outlet temperature is accomplished through the use of an air damper lo-
cated at the air-blower inlet.  The fuel control for the concept main-
tains the desired air-fuel ratio independently.  Condenser drive control
and condenser shutter control are also independently operated.  The only
input from the driver is through the accelerator pedal.  A power boost
may be required to hold the oeprator's effort to a minimum level.  The
basic control unit is an electronic central controller.  This unit, on
the basis  of expander speed and expander inlet-valve position, computes
the instantaneous steam flow of the expander.  When a positive displace-
ment expander is used, this computation can be accomplished with sufficent
accuracy for the predictive control loop.  On the basis of this informa-
tion, the transmission ratio is set to provide the proper speed to the
auxiliaries.  The use of a common variable transmission for both pump
and blower drives compromises the prediction accuracy.   The flow output
of the pump is essentially proportional to pump speed.   On the other
hand, the combustion-air blower output is not a linear function of
blower speed because of boundary effects over a wide flow range.  This
mismatch between the two functions must be corrected by the closed-loop
trim controls.  As long as the design speed variation of the expander
is held within reasonable limits, such as the 3:1 to 4:1 speed range
anticipated for the expander operation, it will be compatible with the
corrective capabilities of the temperature and pressure trim controllers.
                                                                     2-7

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Figure 2-3 - General  Schematic Diagram of Predictive Flow Metering Concept

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The concept shown in Figure 2-3 has been evaluated using an analytical
model of the entire power plant.  Details of the analytical model, and
further results obtained through this study are given in Section 4.  In
the following, examples of the results obtained from this analytical
control evaluation are given.  These examples also point out operational
peculiarities that appear at the two extremes of the steam-generator
power range:  the very low-level operation for idle and the very high-
level operation near the maximum power point.

      Typical results of the analysis are shown in Figures 2-4 and 2-5.
The figures are reproductions of strip chart recordings showing simultan-
eous recordings of sixteen selected variables describing the responses
of the power plant to a step input in throttle.  The reduction of throt-
tle position, a, from 38 degrees to 8 degrees initiates rapid response
in the ratio  setting, Rp, of the variable-speed auxiliary drive.  This
setting changes from an initial value of approximately 1.4 to 0.35.  An
examination of the feedwater input traces, Wfe, and the combustion-gas
flow rates, Wg, indicates that the prediction circuit was properly pre-
dicting the new gas flow rate to the burner.  Because of the singular
prediction signal operating only the common variable-transmission ratio,
the feedpump output was somewhat in excess of the requirements of the
expander.  The closed-loop pressure trim control, through operation of
the bypass valve opening, x, reduced the bypass valve opening in order
to maintain the steam-generator outlet pressure at the desired level.
Examination of the additional traces shown in Figures 2-4 and 2-5 indicate
satisfactory operation of the control concept as long as reasonable pre-
diction accuracy is maintained.  If, however, the system is operated be-
yond the range of the predictive system, the burden of control is shifted
to the closed-loop trim controller.

      An example of insufficient predictive control is shown in the ana-
lytical results given in Figures 2-6 and 2-7.  A large error in the pre-
diction signal causes excessive temperature oscillations.  The error can
be best visualized through examination of the combustion-gas flow rates,
Wg.  At the initiation of the transient, the initial, relatively rapid,
reduction of the gas flow is the result of the prediction signal.  Because
of physical limitations in the simulated control system, this flow is
almost four times the flow rate required for the steady-state operating
point of the steam generator.  The transient described was initiated at
the beginning of the trace with a step input from high to low power by
changing the throttle from the intake valve position, a, corresponding
to 47 degrees to 23 degrees.  The fluid flow level corresponding to the
high intake-valve opening is 0.365 Ibs/sec while the level corresponding
to the lower intake-valve opening is 0.05 Ibs/sec.   Significant differences
in the performance characteristics of the steam generator can be seen for
the large transients.  The first transient, taking place from the high
power level operation is dominated by the slower thermal response of the
boiler at the low-level operation.   At the initiation of the transient,
the predictor circuit rapidly reduces the speed ratio  of the variable-
speed drive.  The reduction in auxiliary speeds can be noted as a rapid
                                                                    2-9

-------
  LB/SEC
    W,
  LB/SEC
   VI.
    rfe
  LB/SEC
   X
 BYPASS
OPENING   0
  AD
 T     100% •
  max
DAMPER
OPENING   0 —
        800 —

       1500 —
        500 —

 LB/SEC  0.8 —
   Wg

        0.4 —
            STEAM FLOW
            FEED WATER FLOW

              tttttt
                                                               - FEED PUMP OUTPUT "TT
                                                                BYPASS VALVE OPENING
           •AIR DAMPER OPENING


                                                                SUPERHEATER OUTLET PRESSURE
                                                                SUPERHEATER OUTLET TEMPERATURE
        +X COMBUSTION GAS FLOW
~'~" I  ' :  P
                10 SEC
              TIME MARK
                           Figure  2-4 - Hybrld-Coiiputer Results;  Fast  Equal-Percentage Bypass Valve

-------
                          ECONOMIZER METAL TEMPERATURE I
                          SUPERHEATER METAL TEMPERATURE
0-
              •Figure 2-5  - Hybrid-Computer Results; Fast Equal-Percentage Bypass Valve
                                                                                                                         2-13

-------
                3RUSH INSTRUMENTS DIVISION. GOULD INC.
   m
STEAM FLOW
                                              :FEEDWATER FLOW
                                                                           I
                                                                                           t
                                              FEED PUMP OUTPUT-
                                                            tฑ
                                                                     tt
                      qzn
                                              BYPASS VALVE OPENING -
                                              AIR DAMPER OPENING
    i
Psia rt
 750
1500
                                              SUPERHEATER OUTLET PRESSURE

                                                                                           rf-H-frr++-i
                                              SUPERHEATER OUTLET TEMPERATURE
                                                  COMBUSTION GAS FLOW
10SEC f
TIME MARKS
                                                 FigurB 2-6 - Closed-Loop Engine-Driven Accessories
                                                                                                                                               2-15

-------
10 SEC TIME MARKS
Figure 2-7 - Hybrid-Computer Resales; Step Transients in
             Intake-Valve Position
                                                                                                                                                 2-17

-------
decrease in both  the feedpump output and also in  the combustion gas  flow-
Because of limitations of  the minimum ratio  setting for the particular
variable-speed drive employed in  the analysis, the lower speed-ratio limit
of the drive is reached.   Thus, at the point of operation, both feedpump
capacity and the  combustion-air blower capacity exceeds the requirements
of the steam generator, thereby calling for significant corrections by
both  the pressure and temperature trim controller.  The slow thermal res-
ponse at the low-level operating point of the boiler results in a larger
temperature overshoot which  is eventually corrected by the closed-loop
temperature control operating through actuation of the air damper.  As
the temperature exceeds the  trim control setting, the damper is closed
and the superheater outlet temperature is corrected.  Gain of the trim
controller is near the stability limit as defined by the longer phase
shift at the low-level operation of the boiler.  For the second transient,
going from this lower level  back to the higher power plant operating level,
the performance is significantly different.  The prediction point for the
variable-speed drive of the  auxiliaries corresponds more closely to the
required output for both the feedpump and the combustion gas flow.  In
addition, a significantly  faster response time is evident for the steam
generator at the  higher operating level.  For this reason, both the mag-
nitude of the temperature  and pressure errors and the time required to
reach the final equilibrium  condition is significantly less than for the
down  transient.

      The maximum steam consumption of the expander must be limited by
the controls to avoid the  possibility of swamping the boiler.  In vehicle
operations, it is possible to demand a continuous power level from the
expander that exceeds the  maximum output of the steam generator.  The
need  for limiting the maximum intake valve or cutoff angle as a function
of RPM, is illustrated in  the traces of Figures 2-8 and 2-9.  Again, the
same  16 selected  variables are shown as a function of time for a step
input in cutoff angle.  When the steam demand of the expander becomes
larger than the output capability of the steam generator, additional
demand for power  results in  a drop of both steam pressure and steam tem-
perature.  In the particular transient shown, the steam pressure drops to
850 psi and the steam temperature drops to 650ฐF from the nominal 1,000
psi and 1000ฐF operating points.  At the same time, in spite of the in-
crease in intake  valve settings, the expander speed remains at essentially
the same level.   The figures illustrate the definite need for a speed-
sensitive limit on the maximum intake-valve position.  It is believed
that  for practical reasons,  this limit must be kept slightly under the
maximum capacity  of the steam-generator output.  An alternate input to
the limit controller could be the steam temperature.  In other words,
anytime steam temperature  is below some acceptable minimum value, the
cutoff angle must be reduced to avoid the possibility of demanding power
in excess of the  steam generator output.
      The examples shown above indicate that proportional control of a
tnonotube boiler for steam  car application is feasible.   However, it must
be noted that care must be taken in component selections.  For example,
                                                                    2-19

-------
           0.4- -^ I^I+ff
  LB/SEC      =
    Ws     0.2 -
                    ttt
           0.4-
   LB/SEC
    w
     fe
   LB/SEC
              in
            o^httL
                    H-
           100---
           50
              ^ฑ
           100 -
    X
  BYPASS
 OPENING
     D
   Amax
  DAMPER
 OPENING
              -t-M-f
              —I—  ! !
5ฐ-ฑฑH
           i.o-irr-i
   its
            0
          1250
          1000
    PSIA
           750
          2000
      T.  1000
       S
            0
   LB/SEC
   10 SEC
TIME MARKS-
           0.8-
           o.4
            0— -L
                         HT"
                                       -f

                                           —n—
                                   w
                                   E
                                                               jjcammt t.--rmซ-
           IOIOIG-4~M
                T-f-H-H-
              :S
                                                        fflbttt
                                                     m
3
                                                      ฃiฑ3
                                                      cnxif
                                                                                 IE
                                                                              ฑฑ
                                                                         5S
                                                                                          c1
                                                                              in
                                                                              H
                                                                            ~t~ *
                                                                              CO
                                                                              a.
                    Figure 2-8 - Hybrid-Computer Results; Operation above  100 Percent
                                Capacity
              2-20

-------
     a.
    DEG
 RAD/SEC
     2.5x10
  IN/LB
iTORQUE
      2.5x10J —
    PSIA
    T    500
     me
    'mb
    T   1000
     ms
  10 SEC
ME MARKS
                Figure 2-9  -  Hybrid-Computer
                              Capacity
Results; Operation above  100  Percent
                                                                                  2-21

-------
if range limitations of the variable-speed drive can not be eliminated,
a more elaborate predictive controller must be developed and the predic-
tive control of both the combustion air damper and the feedwater bypass
valve may become essential.  A somewhat more elaborate predictive con-
troller for the air damper is described in Section 3 for the facsimile
control system.  In addition, it is apparent that some excess capacity
of both feedpump and combustion air delivery system, above the steady-
state power plant requirement, is needed to allow this system to recover
from transient disturbances.  This recovery capability of the power plant
will affect the responsiveness or feel of the steam car power plant to
accelerator command inputs from the operator.  Additional results describ-
ing the effects of component selection on steam generator and power plant
performances are given in Section 4 of the report.
2-22

-------
                              REFERENCES
1.  Walton, J. N., "Doble Steam Cars, Buses, Lorries and Railcars,"
    Light Steam Power. Kirk Michael, Isle of Man, U.K.,  1965.

2.  Vickers, P. T., C. A. Amann, H. R.  Mitchell,  and W.  Cornelius,
    "The Design Features of the GM SE-101 - A Vapor Cycle Powerplant,"
    SAE Paper No. 700163, SAE Automotive Engineering Congress,
    Detroit, Michigan, January 1970.

3.  Hoess, J. A., et al, "Study of Unconventional Thermal, Mechanical,
    and Nuclear Low-Pollution-Potential Power Sources for Urban Vehi-
    cles," Summary Report, U.S. Department HEW Contract  No.  PH-86-67-109.
    Battelle Memorial Institute, March  1968.

4.  Garner, H. D., "Control of the Monotube Boiler," 1st Technical  Meet-
    ing, Steam Automobile Club of America, Oak Ridge, Tennessee,
    September 1971.

5.  Skinner, J. H., R. P. Shah, and W.  A. Boothe, "Modeling, Analysis
    and Evaluation of Rankine Cycle Propulsion Systems," EPA Contract
    No. EHS-70-111, General Electric Co., February 1972.
                                                                   2-23

-------
                                SECTION  3

                        FACSIMILE  CONTROL  DESIGN
      The purpose  of  the  facsimile  hardware  is  the  demonstration  of  the
overall control  concept at  a minimum  cost.   In  order  to minimize  develop-
ment costs, maximum utilization  of  available commercial components was
made in implementing  the  predictive control  concept.   Electronic  analog
signal processing  was selected as it  offers  maximum flexibility in the
testing program  at the lowest possible cost.  The facsimile hardware
design is an extension of the already tested temperature  control  concept.
The results of this testing is reported  in Section  5,  Control Mode Simu-
lation Experiments.   The  facsimile  predictive control  again relies on a
laboratory-type  feedwater flowmeter to monitor  the  power  level of the
steam generator.   On  the  basis of feedwater  flow and instantaneous blower
speed, the analog  signal  processor  predicts  the desired air damper posi-
tion.  This prediction concept is sufficiently  flexible to accommodate
both the expander-driven  auxiliary  configuration and also independently
driven auxiliary configurations  currently planned for  the initial testing
phase.  For an in-car installation, the  feedwater flow would not be used
and the expander speed and  cut-off  valve position would be used to pro-
duce the predictive damper  signal.  A commercially  available Bendix RSA
fuel controller  was modified through  the addition of an electric servo
positioner for the idle-mixture  control.  This combination then becomes
the forerunner of  a two-stage fuel  control system believed to be essential
for the wide 20:1, possibly 40:1, operating  range required for the proto-
type fuel controller.  The  command  signal to  the idle-mixture position
servo is again obtained through  the use  of the electronic signal processor
by combining air damper position with air-damper differential pressure to
obtain an air flow signal in the idle range.  An additional piece of hard-
ware is a direct-pressure-operated  bypass valve for the facsimile steam-
generator feedwater control.  The bypass valve does not include a predic-
tive signal.  The  feedwater pump controller must bear  the burden of the
predictive control to establish  stable operation of the feedwater con-
trol.  In the following,  a  more  detailed description of the individual
component design and design verification test results  of the major com-
ponents will be  presented.

3.1.  FACSIMILE  CONTROL CONCEPT

      The schematic of the  control circuit is shown in Figure 3-1.  The
feedwater supply to the vapor generator  is measured with a turbine-type
flowmeter.   The  flowmeter is designed to simulate the  steam flow signal
that normally would be obtained  on the basis of expander characteristics.
Using blower speed and, calculated or experimentally established,  flow
characteristics  of the throttle valve, the throttle position corresponding
                                                                     3-1

-------
 to  the  required  steady-state  firing  rate is  predicted.   A linear electric
 positioner  loop  is  used  to  actuate the throttle plate and position it to
 the predicted  position.   The  circuit also corrects the prediction signal
 for the actual blower  speed.   The closed-loop temperature trim-control
 signal  modifies  the position  command to the  throttle valve.   Superheater
 outlet  temperature  in  excess  of  the  desired  set point will introduce a
 corrective  signal at this point  into this circuit  that results in a re-
 duction of  combustion  air flow by commanding a reduction of  valve opening.
 The use of  the tachometer signal allows the  use of a simple  linear servo
 positioning actuator in  the throttle-plate control loop.   Although blower
 speed is not a precise indicator of  combustion air flow,  it  will give an
 accuracy that  is acceptable for  the  predictive  temperature  control loop.

       The RSA  fuel  control  unit  operates approximately over  a 4:1 fuel
 flow range. In  order  to extend  automatic air-fuel ratio  control over
 the entire  burner operating range, the use of the  idle-mixture control
 valve was selected  to  control air-fuel ratio below 600 Ibs/hr combustion-
 air flow rate.  In  normal aircraft application,  this control valve is
 directly operated from a throttle lever linkage in the idle  range.  In
 an  aircraft, sufficiently accurate relationship exists  between air flow
 and throttle plate  position near the idle range to use  this  information
 for the fuel control signal.   However,  in the case of an  external combus-
 tor, large  variations  in blower  speed are anticipated.   Thus, the throttle
 position alone is not  sufficient to  be a direct indicator of combustion
 air flow.  An  additional parameter is required to  calculate  the air flow.

       The present concept,  shown in  Figure 3-1,  for the idle-mixture
 control signal was  selected after considering other methods.   Some of the
 more promising concepts  and reasons  for discarding them are  given in the
 following.   The  use of a laboratory-type inlet-air flow meter was discarded
 because nearly all  available  flow meters have pressure  losses that would
 be  excessive for the present  application.  In addition,  the  use of this
 concept appears  impractical for  an on-road vehicle application.   A second
 concept of  combining throttle position signal with blower speed,  similar
 to  the  manner  used  in  the predictive circuit,  was  not believed to be suffi-
 ciently accurate for fuel metering,  particularly when combustor emissions
 are likely  to  require  stringent  air-fuel ratio control.   The use  of throttle
 plaue position with throttle-plate differential pressure  combines simple
 mechanization  using available components with a predicted accuracy that
 should  be satisfactory for  burner fuel  flow  control.   In  addition, this
 concept will allow  a mechanization compatible with the  fuel  controller
 and may be  the forerunner of  a two-stage fuel control system.  It is
 believed that  this  type  of  a  system  will be  essential when combustion-gas
 flow variations  on  the order  of  25:1,  to possibly  as high as 40:1, will
 be  considered  in the final  vehicle application.  An additional feature is
 the relative ease of addition of other  information such as a trim control
 signal  based on  actual combustion gas temperature.
3-2

-------
                                                              VAPOR GENERATOR
FEEDWATER
  INLET
                                                                              AIR INLET
                              Figure 3-1  - Predictive-Control Schematic

-------
3.2   BURNER CONTROL
      In order to quantitatively examine different mechanization  of  the
predictive burner control concept, a simplified model of  the burner  air
flow system was developed.  The model includes the flexibility of handl-
ing any blower nonlinearities because the model uses a three-dimensional
table of any suitable number of points to define the effects of both
blower speed variations and combustion system flow resistance on  the flow-
pressure characteristics of the blower.  As this model is of general
use it is described in some detail.  Following the general description
of the burner simulation model, specific results as applicable to the
facsimile burner control design are discussed.
      3.2.1   Computer Simulation of the Burner System

              In the development of models of any system, a first step
is the derivation of the system equations.  In general, the equations
are based, in part, on certain assumptions which are made not only to
facilitate the modeling task but also are those assumptions which result
in the least sacrifice of accuracy.  Some of the basic assumptions made
concerning the present system model are:
              •  All gas flow through restrictors is proportional to
                 the 1/2 power of the pressure drop except for the
                 exhaust restrictor. Several exponents, including 1,  1/2,
                 and 1/1.8, were evaluated for the exhaust restrictor.

              •  The specific heats of the recirculated combustion gas
                 (CGR) and the inlet air-CGR mixture are assumed  equal
                 to the specific heat of the inlet air.
              •  The temperature rise in the combustor is fixed (3000 F).

              •  The air/fuel ratio is constant (20:1).

              •  Heat transfer from the burner system to the surroundings
                 is neglected.

              •  Injected fuel is completely vaporized.

              A schematic block diagram of the burner system is given in
Figure 3-2.  Gage pressures shown in the figure are those for maximum
exhaust flow and were used to compute the orifice coefficients for the
various restrictors.  The symbols used in the block diagram, along with
the variables and constants used in the equations to be presented, are
all to be found in the nomenclature at the end of this section.   For  each
symbol,  the nomenclature includes a brief description of the symbol,
the units,  and,  where applicable, the numerical value.

              Basically,  in the system model, incoming air, WA, at atmos-
pheric temperature  and pressure,  is mixed with a prescheduled amount  of
recirculated combustion gas,  WR.   This gas then passes through the com-
bustion air blower.   Fuel,  WF, is mixed with the air which then passes
on to the combustor and is burned, producing a temperature rise,   and  is
exhausted to atmosphere as WE.  The exhaust restrictor, RC, represents
 3-4

-------
               ' GAGE PRESSURE IN INCHES OF H2O AT MAXIMUM EXHAUST FLOW.
          0.0
                     -1.5
                                          -2.0
                                                        WF
                                                             PF
                                                                         3.0
         PAB
         TAB
WA
                RI
PI     WG    pG  s~*s:	*
TI     —   TG f   \ pB
      ~      v   yTB
      RG        V	/
                                                               TO
                                                                   WM
                                                         PCI
                                                         TCI
                                                                    RD
                                                              COMBUSTOR
                                                             WR
                                                                        3.0
                                                             RR
                                   Figure 3-2 - Schematic Block Diagram of  Burner  System
                                                                                               0.0
                                                               PCO
                                                               TCO
WE      PAB
                                                                                                      RC
                                                                                                                CM
                                                                                                                in
Ln

-------
  the heat exchanger resistance.   SES experimental evaluation indicates  that
  an exponent of 1/1.8 is most representative for the design.  Other designs
  may deviate from this value, with theoretical limits of 1.0 to 1/2.  The
  system equations are presented  below.
         1.    WG = WA + WR


         2.    WM = WG + WF


         3.    WF = 0.05 WA


         4.    WM = WE + WR


         5     UA - (PAB - PI)
         5.    WA
                       R].
                      RD
        8.  WE =  
        9.  WR •
       10.  PB = PO = PF
       11.  PCO = PCI
       12.  PB = PG + f (SP, VG)
3-6

-------
            TT _ WA x CPA x TAB + ^ x CPR x TCO
                       WA x CPA + WR x CPR
      14.    TB = TI


                 WG x CPG x TB + WF x CPF x TF - HLG x WF
      15.    TO =
                           WG x CPG + WF x CPF


      16.    TCI = TO


      17.    TCO = TCI + 3000


      18.    VG = WG/D


                PG x 0.03613 x 1728
      19.   D =
                 640 x (460 + TG)
              The effects of this throttling scheme on exhaust gas
flow versus blower speed is also shown in Figure 3-3 with the throttle
angle given in parentheses.  Examination of the results shows outflow
to be quite linear with respect to blower speed for a given inlet area
over the usable speed range of the blower.
              From the results shown in Figure 3-3, a new set of curves
may be derived.  They are exhaust outflow versus blower inlet valve
position, with blower speed as a parameter.  These curves are the curves
of constant blower speed shown in Figure 3-4.  It can be inferred from
the nonlinearity of these curves that some nonlinear response or curve
shaping might be required for a shutter or throttle butterfly-valve
inlet control.
              The effect of changing the flow exponent to 1/1.8  was
evaluated.  This figure corresponds more closely to SES experimental
results of steam generator pressure-flow tests.  The calculation results,
as far as pressure-flow relations are concerned, remained nearly unchanged,
as shown in Figure 3-5.
              The block diagram indicates that no active CGR valve exists
and CGR flow is determined by system pressure distribution.  Assuming no
heat transfer from the gases, the 1/2 flow exponent model indicates a
constant 218ฐF combustion inlet temperature.  The 1/1.8 flow exponent
model gives a blower inlet temperature of 247ฐF at minimum inlet flow and
216ฐF blower inlet temperature at the maximum inlet flow.  At the extreme,
if laminar flow conditions are assumed in the heat exchanger and a flow
                                                                    3-7

-------
GJ
00
                                      1000
                                                           2000                  3000
                                                              SP - BLOWER SPEED - RPM
4000
                     5000
                                        Figure 3-3 - Exhaust Gas  Outflow versus Blower Speed

-------
      8.6
THROTTLE PLATE-TYPE VALVE POSITION, DEGREES

 17.1         25.7         34.3         42.9
                                                               51.4
                                                                           60
                                                  - BLOWER SPEED = 1000 RPM
                  10          15         20          25

                   SHUTTER-TYPE VALVE POSITION, DEGREES
                                                          35
Figure 3-4 -  Exhaust  Gas Outflow versus  Burner  Inlet-Valve Position -
               Flow Exponent =1.0
                                                                           3-9

-------
(9

O
    LB/SlcC
    0.7-
    0.6
8.6
          THROTTLE PLATE-TYPE VALVE POSITION, DEGREES

           17.1          25.7         34.3         42.9
                                                                           51.4
                                                                                       60
                             10          15          20          25

                                 SHUTTER-TYPE VALVE POSITION, DEGREES
                                                                      35
       Figure  3-5 - Exhaust Gas Outflow versus  Blower  Inlet-Valve Position
                     Flow Exponent =1.8
     3-10

-------
      PRTGEN
                 9!18   04/18/72  TUE.
BU3WER SUING FACT3R
RIซRDซRR 2.3146
RC= 3.3136 D BFLY=

.
CFM
.
87.
175.
262.
350.
437.
525.
612.
700.
787.
875.
962.
1050.
1137.
1225.
S C
00

00
50
00
50
00
50
00
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00
50
00
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00
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00
A L E D
.00
RPM AT
.00
325.00
651.00
976.00
1302.00
1627.00
1953.00
2278.00
2604.00
2929.00
3255.00
3580.00
3906.00
4231.00
4557.00
B Y :
.50
1.75
3.1174 76.3614
5.00

1.00
1 • 6002
1.50

2.00
IN idF H20
1043.00
1075.00
1 189.00
1375.00
1603.00
1882.00
2202.00
2527.00
2856.00
3188-00
3522.00
3858.00
4194.00
4532.00
4871.00
1489.00
1514.00
1564.00
1710.00
1894.00
2108.00
2351.00
2640.00
2958.00
3280.00
3606.00
3935.00
4266.00
4599.00
4933.00
1853.00
1854.00
1876.00
1986.00
2152.00
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3368.00
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4336.00
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2232.00
2379.00
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2749.00
2965.00
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.
CFM
.
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612.
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787.
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962.
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1 137.
1225.
00

00
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3
RPM
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2627
2624
2662
2772
2930
•

.
.
•
.
•
.
00
AT
00
00
00
00
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3104.00
3296
3502
3733
3975
4223
4533
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.
.
.
.
.
.
.
.
00
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4.00
5.
00
6.00
7.00
IN 3F H2d
3034.00
3036.00
3027.00
3050.00
3129.00
3259.00
3420.00
3597.00
3789.00
3994.00
4217.00
4457.00
4702.00
4966.00
5230.00
3402.
3396.
3383.
3392.
3444.
3559.
3709.
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4453.
4672.
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5389.
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3708.00
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3843.00
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4132.00
4303.00
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4681.00
4887.00
5106.00
5325.00
5544.00
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4007.00
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4037.00
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422 1 . 00
4373.00
4536.00
4712.00
4895.00
5097-00
5299.00
5501.00
5703.00
           .00
      CFM
           .00
         87.50
        175.00
        262.50
        350.00
        437.50
        525.00
        612.50
        700.00
        787.50
        875.00
        962.50
       1050.00
       1137.50
       1225.00
   8.00
 RPM AT
4331.00
4304.00
4286.00
4283.00
4301.00
4353.00
4464.00
4593.00
4757.00
4927.00
5105.00
5283.00
5461.00
5639.00
581 7.00
   9.00   10-00
IN 0F H20
4587.00 4835-00
4567.00 4816.00
4547.00 4795.00
4541.00 4786.00
4550-00 4792.00
4601.00 4836.00
4693-00 4910.00
4811.00 5025.00
4968.00 5153.00
5125.00 5281-00
5282.00 5409.00
5439.00 5537.00
5596-00 5665-00
5753-00 5793-00
5910.00 5921-00
  11 -00

5076-00
5053.00
5030.00
5018.00
5025.00
5060-00
5166.00
5272.00
5378.00
5484.00
5590.00
5696.00
5802.00
5908.00
6014.00
Figure 3-6 -  Pressure-Flow Characteristics of  Blower
                                                                     3-11

-------
 exponent of 1.0 is used, the blower inlet temperature varies  from 502 F
 at minimum flow conditions to 212ฐF at maximum flow condition.   The
 temperature calculation was made with no heat transfer  to  the surround-
 ings.  The effect of heat transfer, however, is believed to be  significant
 and will affect the results.  The prime purpose of the  present  study  was
 the determination of pressure-flow control characteristics for  the system
 studies, and did not require accurate temperature predictions.

        3.2.2   Facsimile Burner Control System
               Detailed design evaluation of the circuit described  was
 carried out for the facsimile system.  The blower characteristics  of  the
 combustion air blower used in the breadboard circuit were scaled to
 provide a blower input definition for the analysis.  The flow-pressure
 characteristics at various speeds are shown in Figure 3-6.  It  is  believed
 that for the purposes of the present design analysis, these curves are
 satisfactory and the design of the system is sufficiently flexible to
 accommodate any deviations that are expected in the performance character-
 istics of the blower to be selected by SES.

               The calculations assumed the use of a simple butterfly-type
 valve  for the throttle.  System pressure drops specified by SES were  used.
 Figure 3-7 shows the blower speed correction curve required for the pre-
 dicted air flow calculation.  The results indicate that for the specified
 blower characteristics a 4:1 blower speed range can be accommodated with-
 out the need for blower curve fitting.  Also a simple two-segment curve
 can be used to accommodate a blower speed range in excess of 10:1 for the
 predicted position control of the air damper.   The required air damper
 position for the blower speed corrected input  signal is shown in Figure
 3-8.  This curve may be fitted with four straight segments to provide
 an accuracy better than that required for the  predictive throttle-plate
 control system.
3-12

-------
                Computer Program Listing  (1 of 3)

GASCK
100  CJMMdtM  DAC 14)*OBC 15)*DC< 14* 15)
110  READ*CU*J)*J=1* 15)* 1=1* 14)
140  READ*TAB*PAB*TF
150  READ*CPA*CPF*HLG
160  KEAD*RI*RD*RC*tfR
162 ARC=l.b
164 RC=3**<1/XRC)
165 RC=RC/.65
166 PRIiMT"RC- "*RC
170  THE1=35.
180 122  C3NTINUE
190  PRINT*IT*/-WR*CPK*VซG*CPG)
400 PG=PA8-**XRC+CRD*WM)**2
420 B=PB-PG
430 TJ=CWG*CPG*TB+WF*CPF*TF-HLG*WF)/A
440 TCJ=TJ+3000.
450 D=C.03613*1728ซ*PG)/<640.*(460 • + TG) )
460 C=WG*60./D
470 IFC1 1.-8)322/312*312
480 312 CONTINUE
490 CALL  IiNPJLCB*C*SP)
bOO GJ Td 329
510 322 SP=0.
520 Gซ3 TJ 329
530 329 CONTINUE
540 PRINT* WA* WG* WM* VIE* v,'R*PVvR* Its* TC-4*PG* PB* B* SP*G*D
550 PRINT
                                                               3-13

-------
                     Computer Program Listing (2 of 3)

       GASCK   CJtMTINUED
560
570
580
590
600
610
620
630
640
650
660
670
680
690
700
710
720
730
740
750
760
770
780
790
800
810
820
830
840
850
860
870
880
890
900
910
920
930
940
950
960
970
980
990
1000
1010
1020
1030
1040
1050
WA=WA+.05
IFC.7-WA>230*230*50
230 CONTINUE
 IF<0.-THE1>372*379*379
 372 THEl=THEl-7.
G*J TJ  122
379 WNTIdUE
END
SUBKdUTINE  INPdL
CJMMdN AA< 14)* BBC 15>*C(X 14* 15)
1=1
40 IFCXA-AACI>>70*50*50
5U 1=1+1
GJ T0  40
70 Ml =1-1
M2=I
1=1
100 IFCXB-BBCI)) 130* 1 10* 1 10
110 1=1+1
63 TJ  100
130 N 1=1-1
iM2=I
GC=(XB-BBJ2)-BB -CCC.vi 1 *
CHI = CC<
            1 ) + < CC< M2* N2 ) - CC C M2ป
                                             1 ) ) *G6
                                             1 ) ) *GG
           RETURN
           END
           SDATA
           0>*.5*1>*1ซ5*2>*3>*4.*5ซ*6.*7.*8>*9.*10**11>
           O./ 50.* 100.* 150.* 200.* 2 50.* 300.* 350.* 400.* 4,50.* 500.*
           550.*600.*650.* 700.
           0.* 325.* 651 .* 9 76.* 1302.* 1 627. * 1 953. * 2278. ป
           2604.* 2929.* 3255.* 3580.* 3906.* 4231.* 4557.*
           1043.* 1075.* 1 189.* 1375.* 1603-* 1882. * 2202. * 252 7. *
           2856.* 3188>* 3522.* 3858.* 4 194.* 4532.* 4871.*
           1489.* 1514.* 1564.* 1710.* 1894.* 2108. * 235 1 .*2640>*
           29 58.* 3280** 3606.* 39 35.* 4266.* 4599.* 49 33.*
           1853.* 1854.* 1876.* 198 6.* 2 1 52. * 2341 .* 2553. * 2793. *
           3053.* 3368.* 368 7.* 40 10.* 4336.* 4664.* 499 4.*
           2125.*2143.*2151.*2232.*2379.*2553.*2749.*2965.*
           3205.* 3452.* 3765.* 4082.* 4403-* 4 72 7.* 5054.*
           261 7. * 262 7. * 2624. * 2662. * 2 772.* 29 3U.* 31 04.* 329 6.*
           3502.* 3 733.* 3975.* 4223.* 4533.* 4849.* 5 165.*
            3034.* 3036>* 3027.* 3050.* 3129.* 3259.* 3420.* 359 7.*
            3 789.* 3994.* 42 17.* 445 7.* 4 702.* 4966ซ* 5230.*
            3402.* 339 6ซ* 3383.* 3392.* 3444. * 3559 .* 3709 •* 3375. *
            4053.* 424 7.* 4453.* 46 72.* 491 1 . * 51 50.* 5389.*
            3736.* 3 723ซ* 3708.* 3 71 2.* 3753.* 3843.* 39 72.* 41 32.*
            4303ซ* 4483.* 4681 .*4837ซ* 5106.* 5 32 5.* 55 44.*
3-14

-------
            Computer Program Listing  (3 of 3)

GASCK  CONTINUED
1060 4044.*4024.*4007.*4007.*4037ซ*4105.*4221. *4373ซ*
1070 4536.ป4712ป*4895.*5097.*5299.*5501ป*5703ซ*
1080 4331• *4304.* 4286.* 4283.* 4301• *4353ซ* 4464.* 4593ซ*
1090 4757.,4927.*5105.*5283.*5461.*5639.*5817.*
1100 4587.*4567.*4547**4541.*4550ซ*4601.*4693ซ*4311.,
1110 4968.* 5125.* 5282.* 5439.* 5596.* 5753.* 59 10.*
1120 48 35.* 48 16.*4 79 5.*4 73 6.*4 79 2ซ* 48 36.* 49 10.* 5025.*
1130 5153.* 5281.* 5409.* 5537.* 5665-* 5793.* 5921.*
1140 5076.* 5053-* 5030.* 5018.*502bซ* 5060.* 5166.* 5272.*
1150 5378.*5484.*5590.*5696.* 5802.* 5908•*6014ซ
1160 70.*407.*70.
1170 .243* *4*156.
1180 1.96*2.54*4.62*68.4
                                                        3-15

-------
                                  NOMENCLATURE
Symbol
CPA
GPP
CPG
CFR
D
HLG
PAB
PB
PCI
PCO
PF
PG
PI
PO
RC
RD
RG
RI
RR
SP
TAB
Units Value
btu/lbฐR
btu/lbฐR
btu/lbฐR
btu/lbฐR
lb/ft3
btu/lb
psia
in H20
in H20
in H20
in H20
in H20
in H20
in H20
(in)171'8 (sec/lb)
Yin" sec/lb
Yin" sec/lb
"Y in sec/lb
Y in sec/lb
rev/min
ฐF
0.243
0.4
	
	
	
156
14.7
	
	
	
	
	
	
	
2.83
2.54
1.09
1.96
68.4
	
70
Description
Specific heat of air
Specific heat of gasoline
Specific heat of inlet air and
EGR mixture
Specific heat of EGR
Density of blower inlet gas
Heat of vaporization of gasoline
Atmospheric pressure
Blower outlet pressure
Combustor inlet pressure
Combustor outlet pressure
Fuel pressure
Blower inlet pressure
Pressure of inlet air + EGR
mixture
Pressure of PI + fuel mixture
Combustor restriction
Blower + duct outlet restriction
Blower inlet restriction
Inlet restriction
EGR duct restriction
Blower speed
Inlet air temperature
3-16

-------
Symbol
TB
TCI
TCO
TG
TI
TO
WA
WE
WF
WG
WM
WR
VG
Units Value
ฐF
ฐF
ฐF
ฐF
ฐF
ฐF
Ib/sec
Ib/sec
Ib/sec
Ib/sec
Ib/sec
Ib/sec
ft3/min
	
	
	
	
	
	
	
	
	
	
	
	
	
Description
Blower outlet temperature
Combustor inlet temperature
Combustor outlet temperature
Blower inlet temperature
Temperature of inlet air + EGR
mixture
Temperature of TI + fuel mixture
Inlet air flow
Exhaust gas flow
Fuel
Inlet air + EGR flow
WG + fuel mixture flow
EGR flow
Blower volumetric flow rate
3-17

-------
u>
00
                    0.3n
                    0.2-
                    0.1-
                                 ,f(0B>
                                              TWO SEGMENT
                                              APPROXIMATION
                                                                 TWO SEGMENT OPERATING RANGE
                                                                         400-4000 RPM
1000
                                                                             2000

                                                                   ACTUAL BLOWER SPEED, d, RPM
                                                        3000
                                                                                    4000
                                                    Figure 3-7  - Blower  Speed Correction  Curve

-------
                             100
                              80
                         ซ/>
                         ui
                         UJ


                       oc
                       <

                                                                                                                     op
                                                                                                                     oL
                                             0.2
0.4           0.6           0.8

     NORMALIZED AIR FLOW
1.0
1.2
CO
                                             Figure   3-8 - Air Damper  Command  Signal Curve

-------
 3.3   FUEL CONTROL

       The wide flow range of the external combustor falls beyond the
 reach of available fuel control systems.  It is believed that an open-
 loop fuel metering scheme that relies on precise scheduling between air
 valve position and flow area opening of a fuel control valve would
 not be a practical device for mass fabrication techniques.  In addition,
 when flow ranges in excess of 20:1 are considered, even with the ulti-
 mate in practical fuel filtering, erosion  of the metering surfaces over
 an extended period of operation - such as a 50,000 mile vehicle opera-
 tion - would alter the metering characteristics of this valve to a
 degree that is unacceptable for a low-emission combustor.  On the basis
 of a brief evaluation of various fuel control concepts, a commercially
 available aircraft-type fuel injection system was selected to provide
 the primary fuel control function.  One of the major factors influenc-
 ing this decision again was the constraint of the time and funding
 schedule of the present program.  It is quite possible that an alternate
 fuel control system could be the result of a more extended fuel control
 system program aimed at optimization for the external combustion system.

       Carburetor-Type Fuel Control
       A fuel control system relying on a constant-level fuel bowl for the
 fuel supply system and the air differential pressure developed in a
 venturi section has a limited range between maximum and minimum flow.
 If only inches of water is available as the maximum flow air differential
 pressure developed across the metering venturi, a specially designed
 variable-area fuel metering jet would become an essential part of this
 type of a metering system.  At the present time, additional constraints,
 such as the use of an air atomizing nozzle believed to be essential for
 proper fuel vaporization, eliminate the carburetor-type fuel control
 system from consideration.

       Pulsed-Type Fuel Metering

       An intermittent or time-modulated fuel metering system, such as
 the electronic fuel injection system currently used on the VW 1600 cars,
 will offer a very flexible fuel metering system.  The degree of sophis-
 tication will be limited only by cost considerations - primarily of the
 various sensors needed to mechanize a particular control scheme.  It is
 believed that this concept may evolve to offer the ultimate in precision.
 The intermittent nature of the fuel flow may affect the emission charac-
 teristics of a continuous combustor.  It was felt that combustion experi-
 ments would be essential to answer this question.  Because neither funds
 not time were available to answer this uncertainty, the electronic fuel
 injection concept was not selected.
3-20

-------
      Continuous Fuel Injection

      Continuous fuel injection has been in extensive use in the aircraft
industry.  This system offers the remote location of the fuel injection
nozzle from the air metering and fuel control section.  In addition,
commercially available continuous fuel injection systems - often called
pressure carburetors - include two independent mixture adjustments:
the manual mixture control and the idle mixture control.  An independent
mechanization of either one of these two with a controller, such as an
electric servo positioner, can easily convert the unit to a two-stage
fuel metering system.  Because of the ready availability of fuel injec-
tion units and the ease of modification that was found to be essential
for thepresent application, this fuel control concept was selected for
the facsimile fuel controller.  A Bendix RSA fuel control device was
selected as the prime component.  In the following, the principle opera-
tion of this device and the operating characteristics of the particular
unit incorporated in the facsimile fuel control system will be described
in greater detail.

      3.3.1   Operating Principles of the Bendix RSA Fuel Control Unit

              A functional block diagram of the RSA fuel control system
is shown in Figure 3-9.  Briefly, the system operates as follows:

      (1)   Pressurized fuel from the fuel pump flows through the
            manual mixture control and idle cut-off valve to one side of
            both the enrichment valve and fuel diaphragms,  and also to
            the enrichment and cruise jets.
      (2)   If the sum of the spring force and the pressure of the
            fuel from the enrichment jet exceeds the fuel pressure
            from the manual mixture control valve,  the enrichment
            diaphragm moves to open the enrichment valve.   Thus in
            the power range (manual idle valve open greater than idle
            setting), fuel may flow through the series combination
            of the fixed enrichment jet and variable enrichment
            valve and in parallel with the flow through the fixed cruise
            jet.

      (3)   Air is drawn into the unit through the inlet and boost
            venturi which causes a pressure drop.   The differential
            pressure across the venturi is applied to either side of
            the air diaphragm.  Thus the net force on the air diaphragm
            is a function of the air flow through  the unit.

      (4)   Fuel leaving the manual idle valve is  applied  to the remain-
            ing side of the fuel diaphragm and the  variable ball valve.
            The air and fuel diaphragms are constructed so  that the
            fuel diaphragm opposes the force of the air diaphragm.   Since
            the net force on the air diaphragm is  totally  dependent on
            air flow, and since the pressure on one side of the fuel
                                                                    3-21

-------
U)
                PRESSURIZED ,
                INLET FUEL
                             MANUAL MIXTURE CONTROL
                              & IDLE CUT-OFF VALVE
                                        ENRICHMENT   _
                                      VALVE DIAPHRAGM
                                                                                                             TO BLOWER INLET
                                                                                                                                 TO FUEL
                                                                                                                                 NOZZLE
                                                                                                                              P-84-257-2
                                              Figure 3-9 - Functional  Block Diagram of RSA Unit

-------
             diaphragm is  fixed at any point in time,  the two diaphragms
             will  balance  only when the remaining fuel pressure,  that to
             and through the ball valve, is of the proper valve,  thus
             ensuring  that fuel flow is a function of  air flow.

               From the above, it should be evident that the fuel control
 system is  a  closed-loop system with input air flow as an input  control
 parameter.   It should further be evident that by varying the sizes  and
 taper  of the enrichment jet and the spring which bears against  the
 enrichment valve  control  diaphragm, the unit can be tailored to  track
 nearly any fuel schedule.

       3.3.2    RSA Fuel-Control Test Results

               The Bendix  RSA fuel control device was  fabricated  by
 Bendix Energy Controls Division in South Bend, Indiana.   The unit initially
 was  calibrated by ECD personnel and final tests, which will be described
 below,  were  performed in  the presence and under the direction of Bendix
 Research Laboratories personnel at ECD.  The overall  performance of the
 unit was deemed satisfactory.  A photograph of the unit is  shown in
 Figure 3-10.

               In  order to test the effectiveness of the unit's regulation
 of fuel flow,  the fuel outlet line was terminated with a variable orifice.
 Fuel flow  and fuel pressure were then measured for constant values  of
 air  flow while the variable load resistance was varied from full open
 to a minimum opening.   The data so obtained is shown  in Figure 3-11.
 In all cases,  the fuel inlet pressure was 30 psig,  the test fluid was
 naphtha (specific gravity 0.735),  and the ambient temperature was 70 F.
 It may be  seen from the data that  the operating characteristic of the
 unit is such as to provide a fuel  flow appropriate to  a  given air flow,
 and  to maintain this  air/fuel ratio essentially constant over an extremely
 wide range of fuel outlet pressures.   Thus,  within reasonable bounds,
 the  proper operation  of the unit is unaffected by the  choice of  fuel
 atomizing  nozzle.

               Prior to the performance of the  BRL-directed  tests at ECD,
 preliminary  data  on the actual-unit fuel-schedule performance was sub-
 mitted  to  BRL.  This  data is shown as the data points  designated by the
 symbol  "X" on Figure  3-12.   Also shown on this figure  is the desired
 fuel schedule (heavy  line)  and the + air/fuel  tolerance  zone (area
 between the  dashed lines).  Data for the fuel schedule  was taken  by  BRL
 personnel both with increasing air flow and decreasing air  flow  so  that
 any  hysteresis in  the  fuel schedule would be noted.  The data points for
 increasing air flow are denoted by the symbol  "o" in the figure,  and the
 data points  for decreasing air flow are denoted by the symbol "A".  The
 fuel schedule  data was taken with  the manual idle valve  in  the full open
position.   No  hysteresis  was noted in the fuel schedule  for  air  flow
 greater than  1000  Ib/hr.   Some hysteresis in the  fuel  schedule is noted
for air flow  of 1000 Ib/hr  and less.   The maximum pressure  drop  across the
RSA unit was  3.7  inches of  1^0 at  2000 Ib/hr air  flow.
                                                                    3-23

-------
N3
                                                                   BOOST VEfMTURI
                                                                                     FUEL INLET
                                                                                  (FROM FUEL PUMP)
                                                                                         IDLE MIXTURE
                                                                                        CONTROL VALVE
                           MANUAL MIXTURE
                              CONTROL
                                                                                    FUEL OUTLET
                                                                                    (TO NOZZLE)
                                  Figure 3-10  -  Modified  Bendix  RSA-11 Fuel Control Device

-------
                        INLET AIR FLOW 2000 LB/HR
                         INLET AIR FLOW 1000 LB/HR
                                                       INLET FUEL PRESSURE

                                                            30 LB/IN2
                        INLET AIR FLOW 600 LB/HR
30
                            10           15



                            OUTLET FUEL PRESSURE (LB/IN'
         Figure  3-11 - Fuel  Flow Versus Outlet Fuel Pressure for

                     Constaftt Values of Inlet Air Flow
                                                                                   op
                                                                                   CL
                                                                       3-25

-------
to
                        5
                        1C


                        Ul
                        3
                        u.

                        E

                        <

23.05
23.0-
22.05
22.0-
21.05
21.0-
20.0-
19.0-
18.0
17.0-
16.0-
15.0-



	 X























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X,

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100% AIR FLOW - 2000 LB/HR
MAX. FUEL FLOW - 113.38 LB/HR
TOTAL MA)




<. GAS FLOW




-211 3.38 LB




/HR




LEGE^
D:
[J ECDDATA
0 BRL DATA. INCREASIN
& BRL DATA
EDULE

"OLERANCE
^*X
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'X,






. DECREASE


ZONE

X
^

^X. '
X,
X,





3 AIR FLOW
G AIR FLOW






1 ^V
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-18.64
-17.64
-16.64

                                      200      400       600      800      1000      1200


                                                                 INLET AIR FLOW (LB/HR)
1400
         1600
                 1800
                         2000
                                   Figure  3-12- Fuel Schedule -  Air/Fuel Ratio Versus Inlet Air  Flow

-------
              It may be seen that  the  fuel  schedule  data  points  lie within
or near the tolerance zone for air flow in  the  range from 600  to 2000
Ib/hr.  It is also evident that the operation of  the unit is such as to
become increasingly fuel rich at lower values of  air flow.  This is be-
cause of the inability of the air  and  fuel  diaphragms to  respond to the
extremely low control signals generated at  low  air-flow rates.   The dia-
phragm materials used in this modified aircraft-type RSA  unit  are some-
what stiffer, to fulfill aircraft  reliability requirements, than would
be necessary in an automotive-type application.   In  addition,  there is
a small change in the effective area of the air and  fuel  diaphragms as
they move in response to control signals.   These  small changes in effec-
tive area also tend to degrade the tracking accuracy of the unit at low
air flow levels.  In discussions between BCD and  BRL personnel concern-
ing the possibility of extended control range,  ECD has assured BRL per-
sonnel that a modest diaphragm development  program would  most  probably
result in a unit which could track accurately down to approximately 350
Ib/hr air flow.

              The effect of the manual idle valve on fuel flow is  shown
in Figure 3-13.  The data was taken at zero air flow, and the valve was
more or less arbitrarily adjusted  for  5 Ib/hr minimum fuel flow.   This
minimum set point may be changed even  to full cutoff by a simple mechanical
adjustment without changing the maximum flow setting.  It may be seen
from the figure that the maximum fuel  flow  through the manual  idle valve
alone, as presently constructed, is 24.5 Ib/hr.   Referring again to the
fuel schedule, Figure 3-12, it may be  seen  that the  changeover to  idle
valve flow alone occurs at some value  of air flow slightly less  than 400
Ib/hr.  The control scheme for automatically operating the manual  idle
valve is described below.  It is apparent that  the RSA unit as presently
modified will perform adequately in all respects  and does conform  to
SES Specification SP-1041, dated 30 March 1972.

      3.3.3   Idle Mixture Control
              The air flow signal,  below 600 Ib/hr flow rates, is  generated
by combining the flow characteristics  of the throttle valve with the
square root of the pressure differential across the  valve.  Figure 3-14
gives a typical flow function for  the  throttle plate.  The function shown
is calculated; however, experimental corrections  due to possible changes
of C
-------
V
oo
                                                               INLET AIR FLOW 0.0 LB/HR
                       0

                     CLOSED
MANUAL IDLE VALVE ANGLE (DEGREES)
                                                                                                                OPEN   i
                                   Figure 3-13 -  Fuel Flow Versus Manual Idle Valve  Angle at
                                                       Zero Inlet Air Flow

-------
a
ฃ
cc
ง
    1.0
                          10                    20

                               DAMPER POS.TIONI. 0Q, DEGREES



                  Figure  3-14 - Airflow  Computer Function
                                                                           3-29

-------
I
CO
o
                    50-
                   40-
                                     W  =f
                   30-
                                                                                               7
               ui
               cc
               (9
               111
               Q
                   20-
                   10-
                               0.2         0.4
                                                    0.6         0.8        1.0


                                                                      *


                                                                    WA
                                                                                    1.2         1.4
                                                                                                         1.6
                                           Figure  3^15 -  Idle  Fuel Control  Valve Position Command

-------
3.4   BYPASS VALVE

      A single-stage, direct pressure-operated bypass valve  is a separate
piece of hardware designed for  facsimile  steam generator  control.  This
bypass valve does not include a predictive  signal.   It  is operated in the
feedwater circuit by direct application of  superheater  outlet pressure.
The use of the superheater pressure,  instead  of  the  inlet pressure for
control, eliminates the shifting pressure at  the outlet that would result
from the AP variations caused by the  variations  in flow losses.  The unit
thus performs only the trim control function  in  the  feedwater circuit.
The feedwater pump controller must bear the burden of predictive control
to achieve stable control of the feedwater  circuit.  An additional func-
tion of the bypass valve is the function  of a safety device, as the full
flow  capability of the valve is 4 GPM, and exceeds  the flow capacity
of the feedwater pump.  Thus the bypass valve is capable of pressure
relief in the event of failure  in the feedwater  control circuit.

      The valve was assembled and initial adjustments, to establish proper
spring preloading and correct metering stem-to-seat  closure, were made.
A linear potentiometer, actuated by the valve control piston, has been
provided to facilitate the acquisition of valve  displacement data.  Ex-
ploded and assembled views of the bypass  valve are shown in Figure 3-16.
Also provided, but not shown in the photographs, is an overpressure limit
stop, the function of which is  to  limit, mechanically, the travel of the
control piston to prevent extrusion and resultant damage to the piston
seal in the event accidental overpressure of  the valve occurs.  In addi-
tion, two end caps, to accommodate springs  varying from 2.5 to 6.0 inches
in free length, have been fabricated.  The  test  set up and results are
described below.

      The pressurized water source consisted  of  a pressure vessel filled
with water and pressurized from a high-pressure  nitrogen source through
a constant-pressure regulator.  The valve control pressure was obtained
from the same nitrogen source through a second,  independent, constant-
pressure regulator.  Pressure transducers were installed in the water and
control pressure lines.  A laboratory water flow meter was installed at
the valve water inlet port.  A  voltage source was connected to the linear
potentiometer.  Thus, all control and operating  parameters were available
as electrical signals which were monitored  simultaneously on a strip chart
recorder.  In addition, the water flow and  control pressure signals were
monitored, with expanded scale, on an X-Y plotter.
      With a water supply pressure of 800 psi applied to the bottom port
of the valve, that is, so that  water  flowed in through the seat and out
of the side port, the control pressure was  cycled slowly up to approxi-
mately 1060 psi and then down to 700  psi.   The results of this test are
shown in Figure 3-17.  The data points denoted by the symbol '0' are
those obtained from increasing  pressures; the data points from decreasing
pressure readings are denoted by the  symbol 'A'.  The best straight line
approximations (graphically determined) to  the increasing and decreasing
                                                                   3-31

-------
        (a)  Exploded  View,  Direct Pressure Operated Bypass Valve
       (b) Assembled View,  Direct Pressure  Operated Bypass  Valve




          Figure 3-16 - Direct Pressure Operated Bypass Valve
3-32

-------
                                                                              100% FLOW = 3.0 GPM
                                                                              FLOW DIRECTION - IN THRU SEAT
                                                                              WATER SUPPLY PRESSURE * 800 PSI
                                                                                                                   to

                                                                                                                   =9
                    700
                                           800
900
1000
                                                                                                               1100
                                                        CONTROL PRESSURE (LB/IN2)

                              Figure3-17 - Percent Flow  Versus Control Pressure Showing Hysteresis
u>

-------
pressure points are also shown.  A semilog scale  (log  percent  flow versus
linear pressure) was chosen because a perfect equal-percentage valve
would yield a log percentage flow versus pressure plot which would be  a
perfectly straight line in these coordinates.  Three gal/min was  chosen
as the 100 percent flow-point.  Good linearity is -noted at  flows  above the
10 percent (0.3 gpm) point.  Repetitions of this test  showed some varia-
tion in the valve cracking pressure because of stiction of  the metering
stem-control piston assembly.  The valve displayed good repeatability,
but the hysteresis caused by moving part friction, which is  evident in
Figure 2-17, was noted in these and all subsequent tests.   The waive
was then cycled as described above at constant water supply pressures
of 600 and 1000 psi.  The results of these tests are the mean  value curves
 (hysteresis averaged out) shown in Figure 3-18.  This  data  is  repeated
in Figure 3-19, but plotted as actual water flow versus control pressure
in rectangular coordinates.  The valve was then connected so that water
flow was reversed, that is, in through the side port and out through the
seat,  and flow tests identical to those previously described were conducted.
The results were the same as those previously obtained except  in  two
respects:
       (1)   The valve cracking pressure was essentially constant
            at all supply pressures.
       (2)   Lower flow rates were obtained for a given supply
            and control pressure.
These  effects are due primarily to the effect of the differential areas
between the metering stem and control piston.  The difference  in  flow
rates  are shown as the mean-value curves in Figure 3-20.
       With the valve still connected so that the flow  was out  through
the seat, the valve was tested dynamically by applying a control  pressure
of approximately 1050 psi, and then dropping the control pressure approxi-
mately 200 psi in 100 milliseconds.  The results of the test are  shown
in Figure 3-21.  Some delay in initial response is apparent, but  subse-
quent  control-pressure tracking and lack of overshoot  or oscillation is
noteworthy.

       Closed-valve leakage tests were performed with 700 psi control
pressure applied, and a water supply pressure of 1000  psi applied in both
flow directions.  With flow in through the seat, a leakage  flow of 0.0324
gpm was obtained; with flow out through the seat, the  leakage  was 0.0306
gpm.

      From the foregoing, it is believed that the valve will perform
adequately and safely in the S.E.S. system.  The safe  operation of the
system is enhanced by the fact that the maximum flow capacity  of  the by-
pass valve (approximately 4.1 gpm at 1000 psi pressure drop) exceeds the
expected capacity of the feedwater pump.  Therefore the valve, equipped
with the spring and adjusted as in the tests reported  above, has  been
delivered to S.E.S.  A set of instructions for operation and adjustment
has also been submitted.  It is anticipated that any additional tests
will be performed on the facsimile hardware at S.E.S.
3-34

-------
    100% FLOW = 3.0 GPM
    FLOW DIRECTION - IN THRU SEAT
                        SUPPLY PRESSURE = 1000 PSI
                                                                  SUPPLY PRESSURE = 800 PSI
                                                                 SUPPLY PRESSURE = 600 PSI
700
800
900
                                                                    1000
                                                                    1100
                                   CONTROL PRESSURE (LB/IN2)
           Figure 3-18- Mean Values of Percent Flow Versus Control Pressure
                     for Constant Values of Water  Supply Pressure

-------
V
CO
                                             FLOW DIRECTION - IN THRU SEAT
                                                     SUPPLY PRESSURE = 1000 PSI
                                                 SUPPLY PRESSURE = 800 PSI
                                                                   >— SUPPLY PRESSURE = 600 PSI
                            700
800               900               1000



       CONTROL PRESSURE (LB/IN2)
                                                                                                        


-------
                  100
               I-

               iu
               u
               oc
               UJ
               a.



               Q
                         100% FLOW = 3.0 GPM

                         WATER SUPPLY PRESSURE = 800 PSI
                                              FLOW DIRECTION - IN THRU SEAT
                           FLOW DIRECTION - OUT THRU SEAT
                     700
800
          900



CONTROL PRESSURE (LB/IN2)
                                                                                          1000
                                                                     1100
                                   Figure 3-^20- Effect of Flow Direction on Flow Versus Control

                                               Pressure at Constant  Supply Pressure
OJ


u>

-------
                          600
WATER SUPPLY PRESSURE (PSI)
                         2000
            FLOW (GAL/MIN)
   CONTROL PRESSURE (PSI)  1000-
       DISPLACEMENT (IN)
  CONTROL PRESSURE (PSI)  900 —
                         800
                         700
         Figure  3-21 - Valve Response  to  a Step Change in Control  Pressure
      3-38

-------
3.5   FACSIMILE DESIGN VERIFICATION TEST

      A series of tests were conducted at  the  Bendix Research Laboratories
to verify the operation of the facsimile burner  control hardware.  All
input signals were activated either directly or  through a simulated signal.
All functional parts of the predictive controller and the fuel controller
were tested.  In addition, calibration curves  for both the air flow cir-
cuit and the fuel control circuit were established.  This section de-
scribes the test conditions and presents the results of the testing.
      Figure 3-22 shows the complete  test  schematic.  The RSA fuel
control unit was connected through a  transition  section to the butterfly
throttle valve.  A suitable fuel source, using a Stoddard-T solvent as the
working fluid, was used in the fuel circuit.   An air-box vacuum system was
used to provide the source for the air flow during the test program.
Both the air-damper servo actuator system  and  the fuel trim-control servo
actuator were fully energized and operated with  both circuits simultane-
ously connected to the 809 analog signal processor.  All signals, including
both the throttle differential pressure input  and the water flow signal,
were also connected to the 809 analog signal processor.  For the purposes
of testing, low-pressure cold water was used to  operate the water flow-
meter.  In order to simplify testing, a voltage  source was substituted
for the blower tachometer signal.  A  photograph  of the test configuration
is shown in Figure 3-23.  A closeup of the throttle valve, the RSA fuel
control, and the two electromechanical servo actuators are shown in
Figure 3-24.

      The calibration of the RSA unit, using the Stoddard-T solvent, was
performed to verify the operating characteristics of the unit.  The re-
sults, shown in Figure 3-25 verify the operating characteristics already
established at an earlier date and reported in Section 3.3.2.  The cali-
bration curve for the fuel control trim servo  at zero air flow was also
established.  The results are shown in Figure  3-26.  The range of control
and general smoothness of the curve indicates  that the fuel control trim
servo has the capability of controlling over the required fuel flow
range.
      The  throttle-plate valve-position signal  and differential pressure
across the butterfly valve are used to compute the air flow signal for
the low-range fuel controller.  The calibration  curve in terms of these
two variables are shown in Figures 3-27 and 3-28.  The first figure gives
the overall calibration over the entire operating range.  After the
blower selection is finalized by SES, the  operating characteristics of the
blower can be combined with these curves to define the expected operating
range of the fuel control system in terms  of AP  range and throttle-angle
range.
      Dynamic operating characteristics of the two servo actuators is
shown in Figure 3-29.  The response to a large step change in water flow
signal is shown.  It is believed that the  signal rates indicated in the
figure exceed the requirements of the SES  facsimile plant.
                                                                    3-39

-------
WATER

SOURCE
FLOW
METER
                           O- 3GPM
                           2.40 at 100%
                 FLOW
                 METER
              ELECTRONICS
                  809 ANALOG
                SIGNAL PROCESSOR
                       0-6000 RPM
                                                                                                   O-0.31 GPM
                                Figure  3-22 -  Burner-Control Test Setup

-------
4S-
                                  INSTRUMENTATION

                                     RECORDER
                                        I        ป~ซ~*!
                                                SERVO AMPLIFIER

                                                 ELECTRONICS
                                                      Figure  3-23  - Facsimile Test Setup

-------
         FUEL OUTLET

        TO FUEL NOZZLE
          FUEL TRIM

       SERVO POSITIONER
                  •   Kg! BUTTERFLY SERVO

                        If   POS1T1OWER
                         1
        Figure  3-24
RSA Fuel-Control Unit with Idle  Circuit  and

Throttle-Plate Servoactuators
3-42

-------
20-
19'
18-
17-
16-
AIR/FUEL 15.
RATIO
14-
13-
12-
11-
10;
4

00ฐ
A ฐ 0ฐ
0ฐ 00
o
o


SPECIFIC GRAVITY 0.767 AT 60ฐF
O


                200 400 600 800  1000 1200 1400 1600 1800 2000 2200 i

                              AIR FLOW LB/HR


Figure 3-25  - RSA Unit Calibration Using  Stoddard-T  Solvent
   FUEL FLOW
     LB/HR
             25-
             20-
             15 -
             10-
             5-
SPECIFIC GRAVITY - 0.767 AT 60ฐF
                               468

                              INPUT COMMAND  VOLTS
                                                       10
      Figure 3-26  - Fuel-Control  Trim-Servo Calibration
                                                                     3-43

-------
                                                 THROTTLE POSITION
                                                  SIGNAL - VOLTS
                                                                      10
                                      AP IN. OF H2O
         Figure 3-27 - High  Flow Calibration  of Throttle-Plate Valve
           600-1
           500-
           400-
    AIR FLOW     .
     LB/HR  30ฐ"
                                                 2.5
                   12345678910
                                       Ap IN.OF H20
            Figure 3-28 - Low  Flow Calibration of Throttle Valve
3-44

-------
CO
                       o
  10

   8 •

   6 •

   4

   2 •
                          0 —
s
                          0 •
                          2 -

                          "
                          6-
                          8 •
                         10

                         10 •
                          8
                       12  6.
                       5  4.
                          0 —
                         10 •

                          8 '

                       J2  6
                       P  „.
                                                                 1 SEC
                                                               TIME MARKS
 COMPUTER
 AIR FLOW
  SIGNAL
WATERFLOW
   SIGNAL
FUEL SERVO
  POSITION
AIR DAMPER
  SERVO
 POSITION
                                                                                                                      en
                                                                                                                      CN
                                                                                                                      
-------
                                SECTION 4

                     ANALYTICAL CONTROL EVALUATION
      The purpose of  the  analytical  evaluation of closed-loop  control
concepts was the optimization  of  the operating controls  for  the  steam-
car vapor generator.   Because  no  suitable analytical description of
this class of monotube steam generators was available at the start of
the program, a new  analytical  model  of  the boiler,  to fit the  time and
cost frame of the project,  had to be developed.   It was  assumed  from the
onset of the program  that simplifying assumptions would  be an  essential
part of this model  to keep  computer  costs within reasonable  limits.
For this reason, a  continued examination and comparison  of analytical
results with experimental data available from the ongoing burner-boiler
experimental program  conducted by SES on their model SES-4 vapor genera-
tor system was made a planned  and significant part  of the analytical pro-
gram.

      The analytical  model  of  the entire power plant system  was  defined
in terms of physical  interfaces.   This  definition of individual  components
and subsystems offered maximum flexibility for evaluation of various con-
trol concepts including new evolutionary concepts that were  not  a part
of the original planning.   The model was designed to handle  all  signifi-
cant nonlinearities in order to allow evaluations of both the  system and
the control concepts  during large transients which  are believed  to be
a significant part  of automotive  power  plant operations.
      The most significant  part of the  analytical model  development was
that of the vapor generator.   A three-section model,  characterizing fluid
properties in the preheat,  boiling,  and superheat range,  was selected.
In each of these sections,  heat transfer calculations from gas to metal
and from metal to fluid were independently calculated on the basis of
selected average temperatures.  A simplified form of  the equation of state
was used in each of the three  sections.   In order to reduce  inaccuracies
resulting from shifts of  the phase transition points, instead  of  using
the fluid temperatures, the enthalpy of the fluid was used as  the output
variable denoting the thermal  energy of the fluid at  the outlet  of each
of the sections.  In  order  to  maintain  a computer model  offering  simple
and fast computations,  both the expander and also engine-driven  auxiliaries,
such as the combustion blower,  were  described in terms of  partial deriva-
tives.  The use of  this model  required  some initial  processing of the data
for the main simulation.  The  amount of time and cost for  this processing
was minimal and repaid itself  in  the form of a greatly simplified system
model.
      The initial hybrid-computer simulation results  were  compared with
the open-loop static  and  dynamic  experimental results developed  at SES
                                                                      4-1

-------
 for  the  quarter-scale model  SES-4  vapor  generator.   This model was then
 used to  develop  the  predicted  closed-loop temperature control for this
 vapor generator.   The resulting closed-loop  experimental program is report-
 ed in further detail in Section 5.   After the model was thus experimentally
 verified,  two prototype vapor  generators, the cross counterflow and also
 the  counterflow  vapor generator models,  were simulated on the hybrid com-
 puter.   The use  of engine-driven auxiliaries operated through a variable-
 speed drive was  also evaluated in  this portion of  the program.

 4.1    HYBRID-COMPUTER MODEL
       In the following,  the  final  form of the hybrid-computer model is
 presented.   This model,  when used  without the expander,  and with simple
 models for  the auxiliaries,  such as  the  simplified  gas-generator model
 described at the very beginning, was used during the initial parts of
 the  program dealing  with the model SES-4 vapor generator system.  The
 model presented  in the  following thus describes the hybrid simulation
 including the steam  generator,  the expander,  vehicle load and also the
 control  concept  based on the use of  expander-driven auxiliaries through
 a variable-ratio drive.   Following the general description of the model,
 results  of  the SES-4 unit are  given.  The section is then concluded by
 hybrid-computer  results  based  on the prototype SES-5 design.

       4.1.1   Steam-Generator Analytical Model

               All variables used  in the analytical models are described
 in alphabetical  order by symbol in the nomenclature at  the end of this
 section.  In addition to a short description of the variable itself,  the
 specific units required to be  compatible with the other  variables are
 listed.   Also, when  applicable,  typical  values of the variables describ-
 ing  the  SES-4 and SES-5 steam  generators are entered.   It must be noted,
 however, that these  values describe  current  design  and  some of the co-
 efficients  may be revised to reflect the final SES  design.
               Burner

               The simplest  burner model assumes constant air-fuel ratio:


                           W  -  Wp  (1 + A/F)                         (1)
                            T  = constant                           (2)
                             O


 Gas  residence  time in the burner and boiler is assumed  to be negligible
 compared  to the boiler dynamics.  In a later section, gas flow prediction
 using  engine-driven auxiliaries will be described.
4-2

-------
               Feedwater Pump

               Feed-pump characteristics may be elaborated during the
program.  For the present, description of a controllable, variable-dis-
placement pump may be sufficient.  A more detailed analytical description
of an engine-driven, fixed-displacement pump is given in a later section.
                           T.  . = constant
                            fei
                                            (3)
                          W.  . = D  p,  .  6
                           fei    p  fei  p
                                            (4)
Other feedwater supply system characteristics such as the condenser model
can be readily incorporated in the analysis.

               Steam Generator

               The steam-generator model is subdivided into three fixed-
length sections: economizer, boiler and superheater sections.  The infor-
mation flow diagram representing  the steam generator is shown in Figure
4-1.

               Economizer

               The flow through the economizer is nearly incompressible.
                          W,  . = W,   = W.
                            fei     feo    fe
                                            (5)
The energy equation for  the economizer section yields
mr  — h  =
 fe dt  e
                                *
                                nfe
               fe
                                           ei
                                              - h
                          eo
                                            (6)
Pressure losses through  the  length of  the economizer section are described
in terms of viscous flow losses.
        2 f  L
- P
                               pfei 8o ฐe
                                             w
                                                                   (7)
The metal-to-fluid heat flow is a function of the temperature difference
between the metal and the fluid.
                         = Ufe AHf
                me
                                          - T
                      .
                      fe
                                            (8)
                                                                     4-3

-------
Figure 4-1 ~ Information Flow Diagram - Steam Generator
                                                                                                                   4-5

-------
                          u
                           fe
                      _
                   Wfe*
                                           0.8
                                                                     (9)
 The enthalpy of the inlet fluid Is a nearly linear function of the fluid
 inlet  temperature.
           4
          ei
                                 *   , ,
                                 fe  fei
                                            oe
                                                                    (10)
       The gas-to-metal heat flow is based on the temperature  difference
 between the inlet gas to the economizer and the metal  temperature at  the
 inlet.   However,  for the particular heat exchanger  design,  the  inlet
 metal  temperature is sufficiently close to the fluid temperature  at the
 inlet  so that the fluid inlet temperature, Tfei, may  be  substituted for the
 heat transfer calculations.
                  T    - T    = e  (T   .  - T-  .)  = AT
                   gei    geo    e  gei     fei      ge
                                                  (11)
                       N T U   =1.1
                                       -N T U
1
1
- e
1
2
2
-N T U
e
2
-0.4
                                     w  *
                                                                    (12)
                                                  (13)
For computation simplicity  the above coefficients may be included into
a single function describing the gas-to-metal heat conductance, f .
                           f   (W ) - C   e
                            e   g'    ge  e
                                                  (14)
Wg Ahge
Cge (ATge>
                                      = f
Metal temperature calculation is based on the difference between the heat
flow from gas to metal and metal to fluid in addition to the mass and
specific heat of the metal.
                                                                    4-7

-------
                      m    C   4-T    =Q     -  Q_*                    <16)
                       me  me dt  me    gme    inre
  The fluid output  temperature  at  the  economizer  outlet can be computed
  on the basis  of outlet  enthalpy.
                          T     = — (h    -  h  )                      (17)
                          feo   Ce   v eo     oe'
                                  re

        The model presented  so  far accomodates  static  heat transfer charac-
  teristics and primary dynamics of  the  economizer  section.   The distrib-
  uted nature of the thermal wave  is not sufficiently  represented for the
  economizer section.   An approximate description of the  thermal wave
  propagation time was  used.
                                    m   C
                                     me  me                          (18)
                                eh   WfeCfe
  Equation (18)  indicates  that  the  thermal-wave propagation  time  is  a
  function of fluid flow.   During the  anticipated operating  range of the
  steam generator,  the changes  in thermal-wave propagation must be accom-
  modated.  The  most convenient inclusion of the variable thermal wave
  propagation time  was through  the  use of the digital portion  of  the hybrid
  computer.  The outlet enthalpy of the economizer was thus  delayed  prior
  to use in the  boiler energy equation.
                          h*    =  h     * -  e"eh                       (19)
                           feo    feo   T  , s
                                        eh
        Inclusion of the thermal wave  propagation time reuses  the metal
  thermal inertia in addition to the account for m^  in  equation (16).
  The repeated use of the metal mass resulted  in a  slower  computer result
  than the experimental results indicated.  Reducing  the metal mass to
  0.75 fflme in equation (16)  offered a  simple empirical solution to this
  problem.  The same correction was also made  in the  boiler  and superheater
  sections of the steam generator  model.
4-8

-------
      Boiling Section

      Conservation of mass in  the  fixed  length  representation  of  the
boiling section may be stated  as
                        V  —— n   = W     - W
                        Vb dt pfb   Wfeo   Wfbo
                                                                    (20)
The energy equation  for  the boiling  section defines  the  time  rate  of
change of average  fluid  density  and  internal energy  for  the volume of
the boiling section.
Vb dF (pfb Ufb>
                        Wfeo hfeo  "  Wfbo hfbo  = Vb  (pfb  Ufb + Ufb  pfb>

                                                                   (21)
 Thus  the  rate of  change of internal energy  in the  boiling  section is
 defined as
        U
         fb  "  p..  V.  (0 _,  + Wฃ    hc    -  W,,   h,.    -p_ U   V  )    (22)
               fb   b   Tnfb    feo  feo     fbo   fbo     fb  fb  b
       The outlet  flow from the boiler  section  is based  on viscous  flow
 calculations  for  the given length of pipe  and  the  square root  of pressure
 difference between boiler and superheater  pressures.
                 W
                  fbo
pfb 8o ฐb
2 f, L,
b b
1/2
A , Y
xb v

rP, - P
b s
                                                                   (23)
 The  gas-to-metal heat flow rate is  again defined  as  a  function  of  tem-
 perature difference between the gas inlet temperature  and  the boiler
 metal  temperature.
                                           -NTU,
                W  C ,   (T ,. - T ,)  \l-e
                 g  gb   gbi    mb'
                                                 W
                                                  W*,
                                                    i
                                                      -0.4
                                                                   (24)
                                                                   4-9

-------
 This equation, again, may be simplified by  representing the heat conduct-
 ance by a single function, f. .



                     V ' 'b  
                                fb   fb   mb '   fb
4-10

-------
       The basic definition of enthalpy, as the sum of internal energy
plus  a PV term, is used to calculate enthalpy of the boiler fluid.
                            hfb - ufb
       The definition of specific volume and specific density of the
 fluid  is  used to calculate the specific volume of the fluid.
                                     l/Pfb                         (32)
       A simplified set  of  equations  are used to  describe  the  equation  of
 state  for the fluid in  the boiling section.   This  simplified  relation-
 ship was found to be satisfactory to evaluate the  control characteristics
 of  the boiler.  Comparison with  the  experimental data obtained  from  the
 SES-4  boiler was used to establish the  validity  of this assumption.
 The pressure in the boiler is  described as a function of  the  temperature.



                        Pb ' A + B Tfb  + f Tfb2                    (33)
On the basis of  the  Clapeyron  relationship for the boiler


                          dPb    (Ufb - UL) J




The internal energy  of the liquid is a linear function of temperature.
                            UL = D Tfb - E                        (35)
Using the Clapeyron relationship, the temperature of the fluid in the
boiler can be defined as follows:
                       T_.  =  (-b +Yb2 - * ac)/2a                 (36)
                        rb
                                                                4-11

-------
.
f   ft
                                     (V   - V)                         (37)
                    b = 460C  (V_ - V.) +  DJ  =  920a + DJ              (38)
                                ID     L






                   c = -J  (E + Ufb) -  (Vfb  - VL)  (A - 460 B)            (39)







          The  thermal wave propagation time is again included to improve the

   model's  representation of  the distributed  thermal capacitance of the


   monotube steam generator.
                                T.
                                      l ,  \s  ,

                                      mb  mo                           (40)
                                bh    W_   C,.,
                                       fbo   fb
    The resulting thermal  delay  is again  represented  as  a variable distributed

    delay time as a function  of  boiler  outlet  flow
                           h*    -h_     "  e                           (41)
                             fbo     fbo    T,,  s
                                          bn
          Superheater


          Conservation of mass  defines  the  change  of  density in the fixed

    section.
                             V   r P*   • w*u   ~  w
                              s  dt  fs     fbo     s
    The energy equation in the superheater may be  simplified because the

    contribution through the rate of change,  pU, is negligible.
                Vs dT (pfs Ufs}  = Qmfs + Wfbo  h*fbo  - W
4-12

-------
The superheater pressure loss  is again calculated on the basis of viscous

flow and the time constant relating  the volumetric charge time of the

superheater volume.
?fs-
2 f  L  W
   s  s  s	

  *        2
P   g  D  A
 SOS   XS
S + 1
                                                                    (44)
The heat flow rate  from  the  gas to metal again is defined as a function

of gas flow rate and  temperature difference between the superheater inlet

gas temperature and metal temperature.
         V = Wg Cgs  (Tgsi - Tg.o> - Wg Cgs ฃs       (45)
This equation may be simplified as follows
                    Q    = f   (W )  (T   . - T  ) W
                    xgms    s   g7   gsi    ms   g
                                               (46)
f  flow function in  the  superheater section is defined as follows
 S
                           f   (W ) = C   e
                            s   g     gs  s
                                               (47)
                                                  -0.4
E = 0.7
s
/ w
-NTU rrr2-
g I W"
1 - 0.96 e V gs /

                                                                   (48)
Outlet temperature of the gas leaving the superheater is calculated on

the basis of heat loss.
                              _.      _


                          gso ~  gsi   W
                                                                   (49)
                                                                 4-13

-------
    The metal-to-fluld heat  flow rate is  again calculated on the basis of
    temperature difference.
nfs  =  Ufs
                                          (Tms " V
                                            (50)
    The heat transfer rate is defined as  a 0.8 power of the fluid flow rate.
                              i.  " C..
                               fs    Is
                                         W
                                         W
                                        V  s
                   0.8
                                                                       (51)
          The metal temperature is calculated on the basis of heat flow
    difference between the gas-to-metal and metal-to-fluid flow rates.  In
    addition, a term describing the equivalent mass of the superheater steam
    and its specific heat is included in the equation.  This correction,
    however, is small.
                   (m.C   +m   C  )  -5- T    = Q
                     fs  ps    ms  ms  dt ms
                                            (52)
    The temperature of the working fluid in the superheater section is
    calculated on the basis of enthalpy and specific  heat.
                               T  = (h  - h  )/C
                                s     s    os  ps
                                            (53)
    The superheater pressure is a nearly linear function of the superheater
    temperature.
                              Pfs ~ CTps Tfs pfs
                                            (54)
    The superheater outlet temperature,  TS',  may be solved by substituting
    the expression for Qmfs, from equation (50), and the enthalpy, hs, from
    equation (53), into the energy balance equation (43).
4-14

-------
    C   T  ' + h   -  h*
    I ps  s     os     fbo
                      Ws - Wfbo hrbo + Ufs AHfs
T   - T
 ms    s
(55)
Equation  (55) now may  be solved for the superheater outlet temperature
    I        + TT  r
s   ufs AHfg + W  C
                               *    U*  T   + w*u  h*   - W  h
                               fs  Hfs  ms    fbo  fbo    s  os
                                                              (56)
The simplified  equation of state used in the superheater section allows
the insertion of  the  distributed effect of thermal propagation  in series
with  the  superheater  outlet temperature.
                           T  =  T'	
                                        -T  , S
                                         sh
                            S     S   T  ,S
                                     sh
                                                              (57)
 and similar  to  the  economizer  and boiler  section
                                  m   C
                                 _   ms  ms
                             rsh ~ W C,
                                    s   fs
                                                              (58)
      The steam  flow is  normally calculated  in  the  expander section,
however, a choked-flow,  variable-area  throttle  is used  to  compare steam
generator operation  independent  from the  expander to verify the analyti-
cal model with the experimental  results.   For the comparison, steam flow
was calculated on the basis  of throttle admittance  and  steam conditions.
                          Ws=KE
                                   yr  +  460
                                                              (59)
      4.1.2    Expander and Vehicle
               For the purposes of control  system evaluation, the expander
may be characterized as a  continuous  flow device.  This assumption is
valid because the steam-generator dynamics  are much slower than the in-
dividual flow pulsations that result  from the opening and the closing of
the intake valves on the four-cylinder expander.  Engine output torque
and mass flow consumed by  an expansion device can be accurately predicted
                                                                     4-15

-------
using the analytical model described by Taplin  and  Gregory*.   Using this
analytical model, the engine torque output,  in  general,  is defined by a
torque coefficient D_._.
                    Mi


                           Torque = D   x P                         (60)
The volumetric flow consumed by the expander  is defined  by  a  volumetric
coefficient D._T.
             MV

                                              •
                        Weight flow = D._7 p   6                      (61)
                                       MV  s
               In a variable-admission device, such as the  SES  expander,
both DMT and Djjy are also a function of the admission angle and engine
RPM.  It will be shown that a modified continuous flow model will  simply
and accurately describe the SES expander on the basis of available flow
and indicated horsepower data.  The information diagram for the combined
expander and load dynamics and engine auxiliaries is shown  in Figure  4-2.

               The engine torque, Tg, is calculated on the  basis of the
    model,
                        T,, = D.^   P  - P                           (62)
                         E    MT   I s    c


where D   is a function of both engine speed and valve position
                         DMT = DMT  (a) ฐ  <ฐ>                        C63)
Similarly, the flow demand of the engine is computed
                     W = D._.  |T  + 4601  R
                          MV  [  s      Is
(64)
*L. B. Taplin and A. J. Gregory, "Rotary Pneumatic Actuators,"  Control
 Engineering, December 1963.
 4-16

-------
                                                 EXPANDER + VEHICLE
                                                                                   ฐMV
(T. + 460)
                               TEMPERATURE CONTROL & BLOWER
                                       A
                                                                               VARIABLE SPEED CONTROL
               Figure 4-2 - Expander Vehicle and Auxiliaries  - Information Flow Diagram

-------
                          DMV = DMV
                The vehicle and accessory  loads  can  be  combined into a
 single second-order representation for preliminary  control studies.  A
 more elaborate model may be included at a future  time  as  engine perfor-
 mance and auxiliary loads become more firmly  defined.

                                         •
                          TT = K. 6 + K_ 6 + F                      (66)
                           LI      *•
 The vehicle  and  engine  inertia are  combined  as  reflected  at the engine
 RPM.
                        e =
T  _ T    	                       (67)
 1    L   IE+IL
        4.1.3    Auxiliaries and Flow Control Mechanization
                The present simulation describes an  engine-driven,  com-
 bustion-air blower and  feedpump operated  through a  variable-ratio  trans-
 mission.  The  transmission ratio  is controlled by a central  controller.
 A simple, steam-pressure-operated, bypass valve is  used  to implement a
 closed-loop pressure  control.  The closed-loop temperature controller
 actuates an air damper  valve to regulate  the combustion  gas  flow.   The
 separate fuel  controller is assumed to maintain fuel-air ratio,  and is
 not  included in the present system simulation.

                The pump speed is  controlled by the  variable-speed  trans-
 mission, represented  as a second-order system.
                                    Rp  e                             (68)
DMV
-4-

*
R T
s s
ms 0
""~~~"— o
*
p
S
pfei
+ 1

D

                      n  _  "•ป	a  a	J.KJ.  y                    fฃn\
                      Rp - —2	24	                     (69)
                           WRP   "RP
 Blower speed is directly proportional  to  the pump  speed.
4-18

-------
                              9B = RB
Pump delivery is defined by pump speed
                           WP ' ฐp pfei  6P
Steam-generator inlet flow is defined by  conservation of mass
                             Wฃe = W  - WR                          (72)
The bypass flow is calculated on  the basis of pressure, fluid properties
and valve opening
                      WR ' \          pfei Pfei
The  valve opening is defined by the geometry of the design and instantan-
eous valve position


                             AR = KB2 x2                            (74)
Valve position calculation  is based on  the solution of a second-order
system


                               i = &-                              05)
                     AF = P  A   -b  x-F  - K  x                  (76)
                           s  A    o      ps


The combustion gas flow is defined by blower speed, blower characteristics
and air damper opening

                                        AD
                       W  = W
                        g    S
                           (77)
"D max
                                                                     4-19

-------
 where damper opening is controlled by the temperature error signal.  The
 rapid time response of the damper system may be neglected for the initial
 evaluation.
                                    T    - T                         (78)
                                     set    s
       4.1.4    SES-5 Steam Generator and Expander Representation

                In order to increase the solution speed of the overall
 hybrid-system simulation, several relationships were precomputed and the
 results of these calculations were used to program analog, nonlinear, func-
 tion generators.  In the following, results of these calculations, as
 applicable to the SES-5 steam generator, are presented.
                The gas-to-metal heat transfer equations (15), (25), and
 (46) require the functions fe (W ) , f^ (Wg) , and fs (Wg) as a function
 of gas flow, Wg.  Figure 4-3 presents this information in the form cur-
 rently used in the hybrid simulation.

                The metal-to-fluid heat transfer coefficients, Ue and Us,
 as a function of fluid flow, Wp, are shown in Figure 4-4.  The functional
 relations are used in equations (3) and (50) to calculate metal-to-fluid
 heat flow in the economizer and the superheater.  In the boiling section,
 a fixed heat transfer coefficient is used.
                The expander D^x and DMV curves were generated on the basis
 of SES data describing expander performance at various operating condi-
 tions defined by valve angles from 20 to 70 degrees at 500, 1500, and 2500
 RPM engine speeds.  The 500 RPM data set was used to represent both the
 torque and volumetric coefficient maps of the engine in the hybrid simu-
 lation.  Results are given in Figure 4-5.  The speed correction curve,
 as used in equations (63) and (65), is given in Figure 4-6.

                Engine losses, auxiliary power consumption, and vehicle
 load were combined into a single load-torque curve for the initial studies.
 A second-order least squares fit to this curve was used to represent
 loads to the engine in high or low gear.  The following expressions were
 used to calculate total load.

                Load = Engine Friction + Auxiliaries + Road Load
                Auxiliaries = Feedpump + Burner + Fan Power

 The engine friction horsepower load was defined as

                Engine Friction = 4 x 4.41 x 10~5 RPM (11 + RPM/100)
 Auxiliary horsepower loads were defined as

                Feedpump Power = 2.42 x 10~3 x Steam Flow  (Ib/hr)
4-20

-------
                  HX-GAS
1 1 :O1
                                       O2/O4/72
                                  537
                  INPUT 4  SCALE  FACTJHS
r
1*1*1*1
WG
.0268
• 0537
.0805
. 1074
. 1 342
• 161 1
.1879
.2148
• 2416
.2665
.2953
.3222
.3490
.3759
.4027
.4296
• 4564
.4833
.5101
• 5370
.5638
.5907
.6175
• 6444
.6712
.6981
. 7249
. 7518
.7786
.8055
CGE
.2932

FGE*WG
•007181
•013483
.019303
.024787
•030014
•035034
.039880
•044576
•049142
•053593
• 0 5 79 39
•062192
•066359
•070448
•074464
•0764lฃ
•062297
•086124
•089895
.093614 ,
.097283
. 100906
. 104485
. 108021
. 1 1 1518
. 1 14975
. 1 18396
.121 782
. 125134
• 128454
CGB
.326300

FGB*WG
.006533
•01641 7
.023775
.030727
.037353
.043707
.049830
•055753
.061500
.067090
•072b40
• 077862
•083067
•066165
•093164
•098071
• 102893
• 107634
• 1 12301
• 1 16896
• 121425
• 125891
. 130297
• 134646
. 136940
. 143183
. 147377
. 151523
. 155624
. 159682
CGS
•355400

FGS*WG
•002461
.004070
•005461
•006732
•007923
•009056
•010142
•01 1 191
•012210
•013203
•0141 73
•Ul 51 23
• 01 6O56
• 016973
•01 7676
•018766
•019644
•02051 1
•021368
•022216
•023055
•023887
•024710
•025527
•026336
•027140
•027937
•028728
•029515
•030296



FGE
• 2674
• 2511
• 2396
• 2306
• 2236
• 2175
• 2122
• 2075
• 2034
• 1996
• 1962
• 1930
• 19U1
• 1 6 7 4
• 164*
• 182S
• 1603
• 1762
• 1 7 62
. 1 743
• 1725
• 1 706
. 1692
. 1676
. 1661
• 1647
• 1633
• 1620
• 1607
• 1595



FGB
• 3178
.3057
.2952
.2661
.2782
.2713
.2651
.2596
.2545
.2499-
.2456
.2417
.2380
• 2345
• 2313
• 22d J
• 2254
• 2227
• 2201
.2177
.2154
• 2131
.2110
• 2O69
.2070
.2051
.2033
.2015
• 1999
• 1982



FGS
. 09 1 V
.0758
• 0676
.0627
• 0590
• 0562
.0540
• 0521
• 0505
• 0492
• 0480
• 046*
• O460
• 04b2
• 0444
• 0437
• 0430
• 0424
.041*
• 0414
• 0409
• 0404
.0400
• 03*6
• 0392
• 0389
• 0365
• 0382
• 0379
• O376


                                Figure 4-3 -  Heat Transfer Coefficients for Gas-to-Metal  Side

-------
               >: 52    02/03/72   i'MUซ
V. HMbE          .3330   Clii
 I.^HUT 3  SCALE
.' 1 • \ * 1
                                                Cl
                                                     • 149UOOฃ-02
vvF
• 0166
.0333
.0499
• 0666
• Od32
.0999
. 165
. 332
• 49d
• 665
• d31
• 99d
• 2164
• 2331
.249 7
*2664
• 2d30
.ฃ997
.3163
. 3330
• 3496
. 3663
• 3d 29
• 3996
• 4162
• 4329
.4495
. 466^
. 4d2d
.4995
Vr'ijT 3 oCALc.
• 333* ป0039 D* •
Us.
.000360
• 000626
•000do6
•001090
.001303
•OOl bOd
.001 70 b
ซ001d9d
.002085
•002269
• 00244rf
•00262b
•002799
•OU2969
• uo3 i 3d
•Ou3304
• oOjAri;
-c^3631
.O0379 1
. OU39:>U
. 0 tJ 4 1 0 7
- OJ4eio3
• UU'4'll 7
. 00 4. -3 70
.004722
•004d72
.005022
ซ u 0 5 1 70
•OO 531 7
.005463
H ACi J/36
.00ld94
•001950
.0020O6
.U02U61


Ut/ wK
•021b95
•0 18500
•01 7335
•016366
• 01565P-
•U15091
.014633
•Ol424d
.013916
.013626
.013368
. 0 1 3 1 38
.012929
.012739
•012564
•0124O3
• 0 12234
•012114
• Ul 1 9cS4
• Olid 6
-------
PRTEXP
9139
11/15/71  MJN.
INPUT FJR  500*1500*2500*XX*XX*
5 DMT FACTJRS
? 6.7878*5.8597*5.1582*1*1
5 DMV FACTORS
? 3. 360 7* 2. 8611* 2.4893* 1* 1
                     RPMS
TS
HPF
a
LEAD
20.
30.
40.
50.
60.
70.
a
LEAD
20.
30.
40.
45.
50.
60.
a
LEAD
20.
30*
35.
40.
45.
1460.00 RS 1027.20 PS
30.7493 TF 63025*358 DMVF
500*00 RPM
IHP
10.13
21.63
36.86
53*85
70.83
86.90
1500*00
IHP
24.91
54.37
94.21
116*25
139.46
185.09
2500.00
IHP
35.46
78.51
105.95
137*16
169.78
T
1277.
2726.
4646.
6788.
8928.
10954.
RPM
T
1047.
2284.
3958.
4884.
5860.
7777.
RPM
T
894.
1979.
2671.
3458*
4280.
DMT
1.2769
2.7265
4.6462
6.7878
8.9282
10.9538

DMT
1.0466
2.2845
3.9584
4.8845
5.8597
7.7769

DMT
• 894U
1.9792
2*6710
3*4578
4.2802
DMV
.5170
1.1663
2.1291
3.3607
4.8079
6.4231

DMV
• 4281
• 9765
1.8235
2.2988
2.8611
4.0969

DMV
• 3655
• 8421
1 • 1 68 1
1.5596
1.9948
ws
.0180
.0407
• 0743
• 1 173
• 1679
• 2242

WS
• 0448
• 1023
• 1910
• 2408
• 2997
• 4291

WS
.0638
* 1470
• 2039
• 2722
• 3482
1000.00 UP 10U
238*6866
DMT /DMV NUMT NDMV
2.47 •
2.34 .
2.18 .
2*02 1
1 .86 1
1-71

DM I /DMV
2.45 •
2.34 .
2. 17 .
2*12 •
2.05 1
1.90 1

DMT/DMV
2.45 •
2.35 .
2.29 .
2.22 •
2.15 •
1881 •
4017 •
6845 .
• 0000
• 3153
1538
3470
6335
) .0000
1 .4306
1.6)37 1.9112

1 iMDMT
1786 •
3899 .
6755 .
8336 •
• 0000
• 3272

NDMT
1 733 •
3837 •
5178 •
6704 •
8298 •

NDMV
1496
3413
6373
8035
1.0000
1 .4319

NDMV
1468
3333
4693
6265
8013
 50.  204*61   5158* 5*1582  2*4893  ซ4346  2*07  1.0000 LOOOO
          Figure 4-5  - Expander D
                   .„, and D..,, Calculation Results
                   MT     MV
                                                              4-23

-------
DW)  0.6
                                                 NORMALIZING
                                                  FACTOR AT
                                                    500 RPM
     0.4
     0.2
                                                                   261.80
                                                                    2500
RAD/SEC
  RPM
                   Figure 4-6 - Engine  Speed Correction  Curves
   4-24

-------
                Burner Power = 0.68 + 1.18 (Steam flow (lb/hr)/1290)3
                Fan Power = 2.084 - 5.94 x 10~3 RPM + 4.36  x 10"6  RPM2
Road Load was  defined in terms of grade and vehicle speed  for  a 4600
pound vehicle.
  Road Load  = V (ft/sec)/550   [4600/  y 1 + G2  (%)  (G(%)  (G(%) + 0.0154)


          +  9.91  x 10~2  V(ft/sec)  + 1.44  x 10~2  V2  (ft/sec)]


The  conversion factors between engine RPM and vehicle  speed were  given  as

      RPM =  20.50 V(ft/sec)         Second Gear

      RPM =  13.16 V(ft/sec)         High Gear

                The composite load-torque  is  shown as a function of engine
speed for the high-gear  condition  in Figure  4-7, and for the  second-gear
engagement in Figure 4-8.  Least squares  fits of both  curves  are  also
shown in Figures  4-9 and  4-10 respectively.   The least squares fit repre-
senting the  high-gear conditions is used  at  the  present for the load tor-
que  in the current control system  studies as required  by equation (66).

                Both the maximum and minimum  values of  the cut-off valve
were limited as a function of engine speed as shown in Figure 4-11.  The
maximum limit is  required to hold  the expander steam consumption  within
the  maximum output of the steam generator.   The minimum setting  provides
the  idle-speed regulation for the  expander.
                The vehicle ratio-drive settings are similarly controlled
to avoid over-speeding of the auxiliaries and are shown in Figure 4-12.
Maximum available ratio  from the drive is 0.9 and the  minimum is  0.3 for
the  design currently considered.

4.2   EXPERIMENTAL VERIFICATION FOR MODEL SES-4

      During the  early part  of the program,  experimental transient results
of the SES Model  4 steam generator were compared with  hybrid-computer
results.  The comparison indicates that the  hybrid-computer program repre-
sents the dynamic characteristics  of the  SES vapor generator.  The accuracy
of the hybrid-computer simulation  is satisfactory to serve as a design
tool in the  program.  Figure 4-13  presents steam-generator, open-loop
characteristics at 100 percent and 40 percent steady-state levels.

      Figure  4-14 compares step-transient results in gas flow at  the
0.0819 Ib/sec fluid  flow rate.  The gas step change was adjusted  to pro-
duce an outlet  temperature change  from 800ฐF to 1040ฐF to duplicate the
experimental  conditions designated as SES-1.   During the change,  both in
the experimental  and  in the  simulation, the  throttle was manually controlled
                                                                    4-25

-------
    OKAOE
FK?CII
b
.20
10
Ib
. 71
20
1 .01
2b
l-3b
30 '
3b
d.- 12
40
2-55
43
3.02
bO
3. 32
4-05
60
4.61
6b
3.21
7u
3.64
75
6. SI
60
7.20
65
7.93
90
6.69
9b
9.49
100
10.32
KK.'i
Jol KJAD
96.51
.96
193.01
2.01
269.52
3.20
366.03
4.60
462.53
6.26
b79.04
6.25
673.55
10.64
7 72 -Ob
13.47
666. 56
16.63
965.07
20. 76
1061 .b7
2b.33
1 1 56.06
30.61
36.63
43. bl
1447.6U
b 1 . 2 7
Ib44. 1 1
59 . 1 7
1640.61
69.69
1737. 12
60.46
1633.63
92.41
1930-13
105.54
KAO/iEC WbT
FA.M FEEO P bUKNKK [.) f ttt_
10. 106
l.bb
20.212
1.10
30.316
. 73
40. 425
.44
50O31
.23
6U-637
. 10
70. 743
.06
6U.B49
.09
.21
101 -U61
.41
1 1 1 . 1 66
. bo
121.274
1 .04
131 . 360
1.49
141 .466
2.01
1 31.392
2.61
161 . 69'i
3.30
1 71 .604
4>O6
161-911
4.91
192.01 7
202. 123
6.64
.00706
.05
.00662
• 06
.01 109
• 06
•01404
. 1 1
•01 776
• 14
. lo
.02616
• 23
.03509
.04333
.36
•03301
• 44
•06427
.54
•07724
• 6b
.09205
• 76
• 10664
. 12774
1-06
• 1 4U69
1 .26
. 1 7242
1.46
• 19646
1 -66
. 22 72 1
1.93
•2b677
2.20 1
.66
. 66
• 66
• 66
.66
.66
• 66
• 66
• 60
• 66
• 69
• 69
• 70
• 71
•73
• 76
• 6O
• 67
• 96
• 09
3.45
4-29
b-40
6.63
6.66
10.93
1 3 • V ฃ
1 7-09
21 .09
25-61
31-29
37.61
44.62
52.99
62- 19
72.49
63.95
96.63
1 10.62
125.99
r J.-iiiOt
22b,.41ซ
1 401 .633
1 173.712
1 1 Ib. 775
1 130.366
1 169. 606
1279.630
1394.710
1 330- 626
1 66 S • 4 1 6
1637. 70o
2046* hU 3
22bl • 3U4
2471.9*6
27U7.M6
2956.602
3224.672
3506*030
3602.349
41 13.974
     kST    907.2911     122.703
          Figure  4~7 -  Expander Load Tabulation - High Gear
4-26

-------
GHAOE
             •00    PEKCENT
FKICTl
5
• 33
10
.74
15
1.23
20
1 -
-------
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        Figure 4-9 - ฃ>econd-0rder Load Curve  Fitting - High Gear
4-28

-------
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Figure 4-10 - Second Order Load  Curve  Fitting - Second Gear
                                                               4-29

-------
    60-
                                              POWER LIMIT LINE
    50
    40-
a   30-
    10
                              OPERATING ZONE
                      IDLE STOP
                  500
1000         1500
       EXPANDER RPM
2000
2500
                  Figure 4-11 - Power Limiter and Idle Stop
   4-30

-------
 1.0-
                                                                                 00
                                                                                 a.
 0.8-
                                   UPPER LIMIT OF
                                VARIABLE RATIO DRIVE
• 0.6-
 0.4-
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                          LOWER LIMIT OF
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                500
           1000         1500
                   EXPANDER RPM
2000
2500
                 Figure  4-12 - Variable-Ratio Drive Limits
                                                                        4-31

-------
UJ
NJ

Super Heater
Gas Outlet
Economizer
Gas Outlet
Fluid Flow
Fluid Outlet
Gas Inlet
Gas Flow Rate
FLOW LEVEL
100%
SES
EXPERIMENT
1663
830
0.211
1000
3200
0.358
HYBRID
RESULT
1764
851
0.211
1000
3200
0.400
40%
SES
EXPERIMENT
1320
542
0.0835
1000
3200
0.130
HYBRID
RESULT
1426
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0.0835
1000
3200
0.137
UNITS
ฐF
OF
Ib/sec
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Ib/sec
                                                                                                           s
                                                                                                           2
                                   Figure 4-13 -  Steam Generator Open Loop Characteristics

-------
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                  1000 	
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                  700
                 10 SEC
               TIME MARKS
                                Figure 4-14 - Experimental and Hybrid  Computer Transient Results
                                     Duplicating Experiment SES-1  at 0.0819 Ib/sec Flow  Rate

-------
to maintain the steam outlet pressure at 800 psia.  Figure 4-14 indicates
excellent correlation between hybrid-computer results and the supporting
SES experimental data points.
      Figure 4-15 shows results of a similar transient at a higher steam
flow rate of 0.1722 Ib/sec.  Again the throttle was manually controlled
to maintain the steam-generator outlet pressure at 800 psia.  Superposi-
tion of SES experimental data from the experiment designated as SES-9
again indicates good correlation.  It must be noted that, during the
experimental tests, a slight droop in the temperature occurred from the
initial value.  It is difficult to maintain a constant fluid-flow rate
in the experimental test setup as pump delivery is sensitive to extraneous
disturbances.
      Figure 4-16 compares the hybrid-computer results with experimental
data obtained in SES-10 experiment for a constant throttle setting.  The
agreement between the outlet steam temperature traces of the hybrid com-
puter and the experiment is good.  The system pressure for the step
varied from 800 to 700 psia in the experiments and from 800 to about 745
psi in the computer prediction.  This discrepancy between hybrid computer
results and experimental data was resolved as a result of thermal drift
in the steam flow control valve opening.  In the experimental setup, ther-
mal expansion of the valve body affects the valve opening and it is diffi-
cult to maintain a constant valve opening for temperature changes in
excess of 200ฐF.  Thermodynamic calculations verified the predictions of
the  hybrid computer, provided the throttle area was held constant.

      The results of this direct comparison between experimental and
analytical transient results verify the validity of the computer model.
General shapes and physical relationships between temperature and pressure
correspond closely between the analytical and experimental model.

      One of the nonlinear characteristics of the steam generator is shown
in Figure 4-17.  In open-loop operation, the computed outlet steam tempera-
ture is shown as a function of combustion gas flow at 10, 30 and 100 per-
cent fluid flow levels.  The significant change in system gain over a 10
to 1 operating range is one of the contributing factors to the difficulty
encountered when simple feedback control is applied to the control of the
monotube steam generator.

4.3   PROTOTYPE RESULTS

      The hybrid-computer model of the SES-5 steam generator was inter-
grated with the vehicle and engine-driven auxiliaries.  The general sche-
matic of the integrated system is shown in Figure 4-18.  The information
flow  diagram includes a very simple predictive controller that sets the
variable-speed auxiliary drive ratio on the basis of throttle position
and engine speed.  The pressure trim controller is a simple direct-acting
bypass valve.  The temperature trim control is a position-servo-actuated
damper in series with the combustion air blower.  The model was used to
establish preliminary gain settings that provide stable operation in the
 4-34

-------
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                               Figure  4-15 - Experimental  and Hybrid Computer  Transient Results
                                    Duplicating Experiment SES-9 at 0.1722  Ib/sec Flow Rate

-------
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                 Figure  4-16 -  Experimental and Hybrid Computer Transient Results
                     Duplicating Experiment SES-10  at 0.1722 Ib/sec  Flow Rate

-------
  1200
             10%
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                        Figure 4-17-  Steam Temperature Versus  Gas  Flow Rate
0.4
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-------
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                                     Figure 4-18 - Block Diagram of Hybrid Simulation;

                                      Engine Driven Auxiliaries and Closed-Loop Control

-------
pressure and  temperature  trim control  loops  and  also  to  evaluate  the
operating characteristics of  several bypass  feedwater control  designs.

      Two prototype  steam generator  configurations were  evaluated.  The
initial SES concept  had a single  boiler-section  pass  nearest to the burner.
General evaluations  of several control concepts  were  conducted using  this
model.  The results  of these  investigations  are  reported under the desig-
nation of cross-counter flow model.   The later  design  included  the inser-
tion of an additional pass of a bare-tube  superheater pass directly located
at  the burner outlet.  These  results are described under the counterflow
configuration.

      4-3.1     Hybrid-Computer Results of  Cross-Counter  Flow Steam
                Generator  Configuration Effect  of Prediction Accuracy
                Initial evaluations were conducted using  an idealized  pre-
dictive flow  control system.   In  this  idealized  case  it  was assumed that
a rapid-response  (2  to 5  hertz bandpass) fuel  and feedwater controller
would be able to  match the demanded  fuel flow  and feedwater flow commanded
by  the predictive controller.   Using this  idealized model of the predic-
tive control  system, the  required prediction accuracy was evaluated.
Results indicated that an 80  percent accuracy  in the  prediction is suffi-
cient, and trim control operation was  able to  handle  as  much as 50 percent
error in the  predicted signal.  With zero  prediction,  in other words  with
straight closed-loop control,  pressure transients were controlled suffi-
ciently by a  bypass  controller.   However,  the  temperature errors in the
non-predictive case  become excessively large.  This initial evaluation of
the control concept  also  indicated   the need of  combustion air and feed-
water pump capacity  in excess  of  the steady-state requirements.  This
excess is mainly  needed during acceleration.   In the  feedpump  circuit, a
sudden increase in the accelerator pedal may open the  inlet valve to  the
expander at a rate faster than the feedwater system can  follow.  If no
excess combustion gas and pumping capacity is  available,  then  the initial
steam pressure droop exists for an extended period of  time.  This reduces
the maximum torque capability  of  the expander  and limits the acceleration
of the car.

                On-Off Pressure Control
                Operation  of an on-off  pressure controller was  also
evaluated.  This  pressure concept offers a very  simple mechanization  of
the bypass concept.  Over the  required flow range of  20  to 1,  excessive
limit cycling was  present.  Pressure switch deadband was  varied from  990
to 1010 psi,  for  a narrow-range design,  to 900 to 1100 psi for a wide-
range deadband  evaluation.  The speed  of the bypass valve was  slow and
was varied from two  to four seconds  end-to-end travel  time.  Under these
operating conditions, excessive pressure variations were  present.  The
relatively slow-response  bypass valve  was  to simulate  an electrically
operated bypass valve.   The extreme  limit  cycle  was evident at all condi-
tions evaluated.
                                                                     4-39

-------
               Proportional Pressure Control

               An optimization of the bypass feedwater control, when com-
bined with the simple predictive control of engine-driven auxiliaries,
was carried out.  This combination may offer the simplest and least expen-
sive control configuration of the feedwater control system.  Typical
results describing this optimization study are given in Figures 4-19
through 4-24.  The examples are given in order to illustrate the use of
the hybrid-computer model for the control system optimization.  In
this type of control system, interaction is present between feedwater
and fuel control.  In addition, the relative capacity of combustion
air blower and feedpump, as compared with steady-state engine demand,
also influences the design and the performance of the power plant controls.
A wide range of choices are practical in these areas.  At the present
time, the choices were not optimized.  However, results presented are
typical of the type of controls proposed for the prototype control mode.

               Slow Bypass Valve
               Figures 4-19 and 4-20 describe the operation of the by-
pass pressure control using a relatively slow bypass valve, requiring
two seconds to travel from the fully opened to fully closed position.
The two figures represent simultaneous recordings of 16 selected variables
in the hybrid-computer simulation.  The designations of the variables can
be correlated with the schematic (Figure 4-18).  Pressure control charact-
eristics of the feedwater system, using a slow-response bypass valve,
are described in Figures 4-19 and 4-20.  The traces describe the response
of the steam generator and the control system to a step change in throttle
setting corresponding to approximately 40-70 mph vehicle speed range with
the corresponding steam flow rates from 0.315 to 0.415 pounds per second.
Wide excursion in steam pressure occurs during the transient.  The slowly
responding bypass valve is not able to cope with the rapidly varying
steam flow.  Pressure excursions of plus and minus 500 psi occur in this
case.  A slow temperature transient follows the pressure transient.
Steady-state steam pressure varies from 1200 to 1020 psi.  The offset
is due to the fact that pump capacity is significantly larger than steady-
state engine requirements.  The bypass valve opening changes from 40 per-
cent to 1 percent opening.  In general, this indicates that at a minimum
engine speed, low excess feedwater and combustion air flow is available.
Figure 4-20 shows the corresponding metal temperature transients and
boiler pressure transients.  In addition, it may be seen that the pre-
dictive circuit changes the variable-speed drive ratio from 0.64 to 1.20
during the transient.

               Fast Linear-Area Bypass Valve

               A sufficient reduction of the pressure transient is achieved
through the use of a faster bypass valve.  Results are shown in Figures
4-21 and 4-22 for the new bypass valve design.  The details of this de-
sign are shown in Section 3 under bypass valve.  The traces indicate a
 4-40

-------
  Amax
DAMPER 100%
OPENING
              10SEC
            TIME MARK
                                                                Figure 4-19 - Hybrid-Computer Results; Slow Bypass Valve

-------
Figure 4-20 - Hybrid-*Comput4r Results; Slow  Bypass Valve
                                                                                                                4-43

-------
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           0-

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    P  100%-
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                10 SEC
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                                       Figure  4-21 - Hybrid-Computer Results; Fast Linear  Bypass Valve
                                                                                                                                                             4-45

-------
 RAD/SEC
  IN/LB
ATORQUE
                                   Figure 4-22 - Hybrid-Computer Results; Fast Linear Bypass Valve
                                                                                                                                                     4-47

-------
     -mm
LB/SEC
              Figure 4-23 - Hybrid-Computer Results; Fast Equal-Percentage Bypass Valve
                                                                                  4-49

-------
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ATORQUE
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                        Figure 4-24 - Hybrid-Computer Results;  Fast  Equal-Percentage Bypass Valve
                                                                                                                                   4-51

-------
linear scheduling between valve  position and valve  opening  area.   The
results indicated that  this  type of  a valve operates  satisfactorily  over
a range of feedwater  flow levels.  However, at  the  very  low flows, at
the order of  3-5 percent of  the  full steam generator  capacity,  the gain
of this type  of a feedwater  bypass valve becomes  too  high and unstable
operation results.  The steam pressure is controlled  between 1000 and
960 psi during the  transient.  The maximum pressure excursion is 1010
psi and indicates a significant  improvement over  the  previous design.
The bypass opening  varies from 50 to 10 percent opening.  The variable-
drive transmission  ratio varies  from 1.37 to 1.55 during the transient
shown.  The steam-flow  steady-state  levels vary from  0.16 to 0.015
for this transient.

               Fast Equal-Percentage Bypass Valve

               The  performance of the feedwater control  system  can be
improved at the low flow levels  through the use of  an equal-percentage
valve opening.  In  this configuration the area  of the bypass valve varies
approximately as the  square  of the valve position.  This can be achieved
by the use of a taper-plug  design as already described in Section  3.  This
is the configuration  that was selected for experimental  evaluation and
for the facsimile system.   Typical results are  again  shown  in Figures
4-23 and 4-24.  Typical operation is shown for  a  steam flow change from
0.195 to 0.007 pound  per second.  The minimum flow  level is  l/50th of the
rated 100 percent flow  capacity  of the SES-5 boiler.  The steam pressure
is controlled between 1020  to 960 psi.   The variable  transmission drive
ratio is changed during the  transient from 1.42 to  0.45.  It appears that
satisfactory  steam-generator pressure control can be  achieved through the
use of the equal-percentage  bypass flow controller.

      The above results are  shown mainly to illustrate the use of the
hybrid computer in  the  optimization  of  control.   However, in the selected
traces, the steam temperature is too high.  This  was  due to  the relatively
low proportional gain used  in the trim temperature  control.  The high-
excess, combustion-air-blower capacity required a nearly 50  percent  clos-
ing of the damper.  It  became apparent  during the evaluations that capa-
city matches  between  the combustion  air blower  delivery and  the feed-
water pump delivery must be  carefully observed.   The  speed-flow charact-
eristic of the feedwater pump is  essentially linear, while  the blower
output-speed  curve  is highly nonlinear  in the operating range.   This sug-
gests that for the  blower characteristics currently used, it may be
necessary to  add a  predictive signal to the air  damper control in order
to maintain an improved steam temperature control.

      4.3.2    Hybrid-Computer Results  of Counterflow Steam  Generator
               Configuration
               The hybrid-computer simulation was later modified to
accommodate the changes due  to the new  SES-5 design of the steam generator.
The effect of the additional fifth pass,  which  is a bare-tube superheater
                                                                     4-53

-------
section, and  the  change of  the  gas  flow path  that  locates  the new  section
directly at the burner phase, were  evaluated.   Perhaps  the most  significant
effect  stems  from the differences between  blower and  pump  output charact-
eristics as a function of auxiliary drive  shaft speeds.  In  general,  com-
bustion air blowers  of the  type considered for  the present application
have a  strong exponential characteristic of flow versus shaft speed.
On the  otherhand,  positive-displacement pumps have a  nearly  linear out-
put versus speed  characteristic.  When additional  restraints are imposed
on the  control design, then the design of  the temperature  and pressure
control for the steam generator  becomes increasingly more complex.   An
example of such an additional restraint  is the need  for a minimum
differential-pressure drop  across the combustion air  circuit in  order to
reduce  its sensitivity to ambient pressure disturbances  at the inlet
or exhaust of the combustor.  A large number of parametric evaluations
becomes essential to optimize the temperature and  pressure control
systems.
               The interaction  between the temperature  and pressure con-
trol system is illustrated  in Figures 4-25 through 4-28.   In both  in-
stances the steam generator is  operated at a low level.  At  this point
the system time constants,  both in  the pressure and temperature  control
loop, become  the  longest.   Closed-loop control  under  these conditions
becomes the most  difficult  due  to the slow but  strong interaction  be-
tween the temperature and pressure  control system.  In  both  cases  a step
change  in engine  demand is  initiated by changing the  cut-off position
alpha.  In both cases, the  feedwater flow  input is manually  controlled to
maintain the  steam pressure at  1,000 psi,  with  a lower  gain  in the tempera-
ture trim control.   As shown in Figures 4-25 and 4-26,  the control of the
pressure becomes  very difficult as  the changes  in  feedwater  flow affect
both the outlet pressure and temperature.   When the temperature  control
is improved as shown in Figures 4-27 and 4-2 8,  pressure  control  becomes
less affected by  temperature changes, thus significantly improved  control
over the outlet pressure can be achieved even at very low  operating
levels. The  increased level of air damper activity is  a direct  result
of the  increased  temperature gain.   It is  apparent that  good prediction
accuracy will be  required at the low power levels  to  reduce  the  air
damper  activity and  the resultant modulation of the combustion gas flow
at the  minimum operating levels.  The examples  shown  are characteristic
of the  general results of the hybrid-computer simulation which appears to
offer a very  realistic tool for the optimization of the  steam generator
control system.
4-54

-------
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Figure 4-25 - Hybrid-Computer Results; Revised Boiler-Manual Feedwater
              Control - Low-Gain Temperature Control
                                                                     4-55

-------
  a
 DEC
 RAD/SEC
   IN/LB
ATORQUE
PSIA
  me
 'mb
 ฐF
Tms

   10 SEC
TIME MARKS-
I S.LEiELt
u
30 	 1

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200 — *
o —
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            Figure 4-26 - Hybrid-Computer Results; Revised Boiler-Manual Feedwater
                          Control - Low-Gain Temperature Control
        4-56

-------
   LB/SEC
     W,
   LB/SEC
     W
      fe
    LB/SEC
     W_
     X
  BYPASS
  OPENING
    max
  DAMPER
  OPENING
       s

     PSIA
     'FT,
   LB/SEC
     Wg
   10 SEC
TIME MARKS
    Figure 4-27 - Hybrid-Computer Results, Revised Boiler-Manual  Feedwater
                  Control - High-Gain Temperature Control
                                                                         4-57

-------
      a
     DEG
     6
 RAD/SEC
   IN/LB
ATORQUE
      RP
     rb
     PSIA
      me
     'mb
      ms

20 —



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          10 SEC
       TIME MARKS

       Figure  4-28 - Hybrid-Computer Results; Revised Boiler-Manual Feedwater
                     Control - High-Gain Temperature Control
        4-58

-------
Symbol

a

A
A^max

A/F
 A

AT

Axb

A
 xe

A
 xs

b

b
c

C
Units



lb/in'


in2

in2



in2


in2


in2


in2

in2

in2

in2

in2


in2
                              NOMENCLATURE

                               Value
                           SES-5    SES-4



                           4267.5   4267.5
         19.0

625      324


1150     312


373      288


3.518 x  10~3

0.169

0.05373

0.192    0.116

0.107    0.075


0.192    0.116
              Ib-sec/in    0.444

              lb/in-ฐF     -20.35   -20.35
              lb/in2-ฐF2   0.0527   0.0527
               R/sec
                           0.4049
      Description

Equation-of-state constant

Equation-of-state co-
efficients, boiler

Air damper opening

Maximum damper opening

Air-fuel ratio

Heat transfer area, fluid
side, economizer

Heat transfer area,
fluid side, boiler

Heat transfer area,
fluid side, superheater

Bypass flow area

Bypass differential area

Flow area, throttle

Flow area, boiler tube

Flow area, economizer
tubes

Flow area, superheater

Equation-of-state constant

Bypass damping coefficient

Equation-of-state co-
efficients, boiler

Equation-of-state constant

Equation-of-state co-
efficients, boiler

Superheated steam, gas
coefficient
                                                                     4-59

-------
                                Value
Symbol
Cle
Cls
Cd
Cfe
V
C
ge
C
gs
mb
C
me
C
ms
C
ps
T*
Units
btu/
(in2-sec-ฐF)
btuX
(in -sec-ฐF)

btu/lb-ฐF
btu/lb-ฐF
btu/lb-ฐF
btu/lb-ฐF
btu/lb-ฐF
btu/lb-ฐF
btu/lb-ฐF
btu/lb-ฐF
in/ฐF
SES-5
3.95 x
10-3
1.49 x
10-3
0.80
1.09
0.3263
0.2932
0.3554
0.11
0.11
0.11
0.697
1437
SES-4
3.86 :
1.736
0.80
1.09
0.434
0.287
0.310
0.11
0.11
0.11
0.68
1437
D
btu/lb-ฐF    1.26      1.26
              in.

              in.
             0.495     0.385

             0.37     0.31
                                                      Description
                                                mizer, fluid side
heater, fluid side

Orifice coefficient,
throttle

Specific heat, fluid,
economizer

Specific heat, gas,
boiler

Specific heat, gas,
economizer

Specific heat, gasป
superheater

Specific heat, metal*
boiler

Specific heat, metal,
economizer

Specific heat, metal,
superheater

Specific heat, fluid,
superheater

Superheater conversion
coefficient, temperature
pressure

Equation-of-state
coefficients, boiler

Boiler tube I.D.

Diameter, economizer
fluid tube
4-60

-------
Symbol


DMT (a>

DMV
Units
     Value
SES-5    SES-4
D
D (6)
f  -
 1Ps
in -sec/rad

in -sec/rad
              in /rad
in.
              btu/lb
              in-lb
              Ib
             0.0872   0.385
0.495
             148      148
             1399.2
                           4 x 10~3 4 x 10~3
                           4 x 10"3 4 x 10~3
             160.55

             4 x 10~3 4 x 10~3
      Description

Expander torque constant

Expander volumetric
constant

Water pump displacement
coefficient

Tube diameter,
superheater

Velocity coefficient,
expander constants

Liquid, internal
energy equation
coefficient

Vehicle rolling
resistance revised

Friction factor, boiler

Boiler heat transfer
function

Friction factor,
economizer

Economizer gas-to-metal
heat transfer function


Flow function
                      Bypass spring preload

                      Friction factor,
                      superheater

                      Superheater heat
                      transfer function
              in/sec'
              btu/lb
             386
         386
Gravitational constant
for mass conversion

Enthalpy - average
economizer
                                                                    4-61

-------
 Symbol
  ei
  eo
  lfb
  fbo
  oe
  OS
  'fsi
Unit
               btu/lb
               btu/lb
               btu/lb
               btu/lb
               btu/lb
               btu/lb
               btu/lb
               btu/lb
               btu/lb
                         2
                in-lb-sec /
                rad
    Value
SES-5    SES-4

190.4    219
             -51
         -51
             864.35   825
             1187.8   1187.8
      Description

Enthalpy - economizer,
inlet

Enthalpy - economizer,
outlet

Enthalpy - fluid,
boiler

Enthalpy - fluid,
boiler outlet

Enthalpy - baseline,
economizer

Enthalpy - baseline,
superheater

Enthalpy, superheater
output

Enthalpy, superheater
inlet

Enthalpy, vapor
                                   Engine inertia
 K.
                in-lb-sec  /
                rad           901

                             371
                in-lb/btu     9335     9335
                             1.273    1.273
II F         3.5 x
             ID'3

in2-ฐF/      0.3605
sec
                                   Reflected load
                                   inertia - High Gear

                                   Reflected load
                                   inertia - Second Gear

                                   Thermal conversion
                                   coefficient

                                   Ratio of specific
                                   heats, superheated
                                   steam
                                                  Damper positioner  gain

                                                  Simulated throttle
                                                  admittance
4-62

-------
Symbol
Unit
    Value
SES-5    SES-4
k
V
K
s
Kl
K2
Lb
L
e
L
s
"DP
mfb
mfe
m.
f s
m
mb
m
me
m
ms

Ib/in
in-lb-
in-lb-
sec/rad
in.
in.
in.
Ib-sec2/in
Ib
Ib
Ib
Ib
Ib
Ib

228
7.626 x
10~2
-0.0189
739 265
538 336
240 240
2.163 x

1.7 0.28
4.29 x
10 2.08 x 10
39.3 x 20
0.75
28.4 x 32.5
0.75
7.8 x 22
0.75
NTU,
NTU
NTU
             1.10
                           1.10
                           0.12
              psia
      Description

Expander velocity
constant

Bypass spring constant

Load torque - second
order coefficient

Load torque - speed
coefficient

Length of boiler

Length of economizer

Length of superheater

Mass of bypass valve


Mass of fluid, boiler

Mass of fluid, economizer
                                                 Mass of fluid, super-
                                                 heater

                                                 Mass of metal, boiler
                      Mass of metal,  economizer

                      Mass of metal,  super-
                      heater

                      Heat transfer units,
                      boiler

                      Heat transfer units,
                      economizer

                      Heat transfer units,
                      superheater

                      Boiler pressure
                                                                     4-63

-------
Symbol
P
c
PDP
Pei
P
eo
Pfs
PM
P
P
P
s
P *
s
gmb
gme
gms
^mfb

-------
Symbol

R
 fb
 fe
 fei
 feo
 fs
 fsi
 gbi
 gbo
 gei
 geo
 gsi
 gso
Unit

in/ฐR
              in/ฐR
              in-lb
     Value
SES-5    SES-4

1027.4   1027.4
             1027.4   1024.4
                           220
                      220
                           3200     3060
       Description

 Gas  constant,  super-
 heated steam

 Gas  constant,  vapor,
 superheater

 Torque output,  expander

 Fluid  temperature,
 boiler,average

 Fluid  temperature,
 economizer,  average

 Fluid  temperature,
 economizer inlet

 Fluid  temperature,
 economizer outlet

 Fluid  temperature,
 superheater, average

 Fluid  temperature, super-
 heater  inlet

 Gas temperature,
 burner  output

 Gas temperature,
 boiler  inlet

 Gas temperature,
 boiler  outlet

Gas temperature, econo-
mizer  inlet

 Gas temperature,
 economizer outlet

Gas temperature,
 superheater inlet

 Gas temperature,
 superheater outlet
                                                                    4-65

-------
 Symbol


 TL

 T ,
  mb


 T
  me


 T
  ms
 u
  fb
 u
  fe
 U
  fs
 u
  fs
 u
  gb
 u
  gs
 U
  fb
Unit

in-lb

ฐF
     Value
SES-5    SES-4
                R
                R
btu/(sec
in -ฐF)
             1460
             1460
      Description

Load torque

Metal temperature,
boiler,average

Metal temperature,
economizer, average

Metal temperature,
superheater, average

Superheater outlet
temperature

Nominal superheater
outlet temperature
                                               -2
                            1.929 x  1.929 x  10   Heat  transfer
                            10-2
btu/(see-in -ฐF)
btu/lb
btu/sec-in -ฐF
          o
btu/sec-in -ฐF
               btu/sec-in -ฐF
btu/sec-in -ฐF
btu/lb
                     coefficient, fluid
                     side, boiler

                     Heat transfer
                     coefficient, fluid side,
                     economizer

                     Fluid internal
                     energy - superheater

                     Heat transfer
                     coefficient, fluid
                     side, superheater

                     Heat transfer units,
                     boiler, gas side

                     Heat transfer
                     coefficient, gas side,
                     economizer

                     Heat transfer
                     coefficient, gas side,
                     superheater

                     Internal energy, fluid,
                     boiler
4-66

-------
Value
Symbol
UL
Vb
Vfb
VL
VM
V
s
W
WF
W,,
fbo
Wfe
Wfe*
W. .
fei
Wr
feo
W.
fs
W. *
fs
W
g
W *
g
W *
gt>
Unit
btu/lb
in3
in3/lb
in3/lb
in3
in3
Ib/sec
Ib/sec
Ib/sec
Ib/sec
Ib/sec
Ib/sec
Ib/sec
Ib/sec
Ib/sec
Ib/sec
Ib/sec
Ib/sec
SES-5 SES-4

142. 1 35

37.67 37.67
40
46.1 30

0.0277 0.0174


0.333 0.2153



0.333
0.537 0.3472

0.537 0.3472
                       Description

                 Internal energy, liquid,
                 boiler

                 Volume of boiler

                 Specific volume,
                 fluid, boiler

                 Mean specific volume
                 of liquid,  boiler

                 Manifold volume

                 Superheater volume

                 Fluid flow, expander

                 Fuel flow rate

                 Fluid flow, boiler
                 outlet

                 Fluid flow, economizer

                 Nominal  fluid flow,
                 superheater

                 Fluid flow,
                 economizer  inlet

                 Fluid flow,
                 economizer  outlet

                 Average  fluid flow
                 rate, superheater

                 Nominal  fluid flow,
                 superheater

                 Gas  flow rate

                 Blower delivery
                 characteristics

                 Reference gas flow
                 rate - boiler
                                    4-67

-------
 Symbol

 W*
  ge
            Unit

            Ib/sec
                 Value
             SES-5    SES-4
             0.537
         0.3472
      Description

Reference gas flow
rate - economizer
(to be set until
Tฃ   = 1000ฐF)
 fso
w *
V
W
w,
 R
W
a

AF


Ah


Ah
  3


AP.
fe
 Jfeฑ
Jfs
Ib/sec


Ib/sec

Ib/sec

Ib/sec


in.

degrees

Ib


btu/lb


btu/lb


lb/in2


lb/in3
              lb/in3
              lb/in'
              Ib/in'
0.537    0.3472       Reference gas  flow
                      rate - superheater

                      Pump output flow

                      Bypass flow

                      Steam flow,
                      superheater outlet

0-2.250               Bypass valve travel

0-60                  Admission angle

                      Sum of bypass valve
                      forces

                      Enthalpy difference,
                      fluid economizer

                      Enthalpy change,
                      economizer gas

                      Pressure drop,
                      boiler

                      Average fluid
                      density,  boiler

1.389 x  1.389 x 10~3 Mean density of
10~3                  fluid
             0-0346
6.944 x
               94
               ~4
                                              fluid in boiler

                                              Fluid density,
                                              economizer inlet

                                              Fluid density,
                                              superheater

                                     6.944 x  Mean density,
                                     10-4     superheater
4-68

-------
Symbol

6
ป
e
 eh
 lbh
 ah
 sp
 wb
'RP
"RP
Unit

rad

rad/sec
       f
rad/sec

rad/sec

rad/sec
    Value
SES-5    SES-4
              sec
              sec
              sec
              sec
                            8.7
                           4.0
                      24
                            15.0     19.0
                      27
              rad/sec
                           0.068    0.068
                           0.008    0.008
                           0.707
             3.14
      Description

Expander  rotation

Expander  speed

Expander  acceleration

Blower speed

Water pump speed

Heat exchanger
effectiveness coefficient,
boiler
Heat exchanger
effectiveness co-
efficient, economizer

Heat exchanger
effectivness
coefficient, superheater

Economic enthalpy
time constant at
nominal flow

Boiler enthalpy time
constant at nominal
flow

Superheater enthalpy
time constant, at
nominal flow

Superheater, pressure
buildup time constant

Boiler, flow inertial
time constant

Variable drive
damping ratio

Variable drive
frequency response
                                                                    4-69

-------
                               SECTION 5

                  CONTROL-MODE SIMULATION EXPERIMENTS


      In October 1971  the flow control mode concept was evaluated using
the SES Model 4 burner and vapor-generator test setup.  The experiments
were designed to simulate, as much as practical, the closed-loop, hybrid-
computer, design studies.  Because of the limitations of the available
experimental, vapor-generator, test bed, only the temperature control
loop was implemented on  the  experimental test setup.  The feedwater sys-
tem and the steam throttle were controlled manually during the tests to
maintain the vapor-generator pressure at the desired level.  The tempera-
ture control of the vapor-generator outlet is the most critical control
problem, because of the  large thermal lag presented by the vapor-genera-
tor thermal inertia.   The experimental results indicated an approximate
66ฐF control band for  large-step fluid transients from 40 to 100 percent
flow levels.  The previously mentioned computer predictions indicated a
75ฐF control band for  the same gain settings.  This is considered to be
good agreement between the hybrid-computer prediction and the experi-
mental observations and  indicates that the computer is an established
tool for the control design  studies.

      In the following section, first the control-mode simulation experi-
mental test setup is described in detail.  Prior to the closed-loop
steam-generator control  tests, preliminary hardware design tests were
conducted.  These results are included in the section prior to the closed-
loop steam-generator test results.  In addition to the closed-loop testing,
open-loop dynamic-response tests were conducted on the steam generator
during the test period.  The results of these transients tests are com-
pared with hybrid-computer results and indicate good agreement between
the analytical model and the experimental results, thus again verifying
the accuracy of the hybrid computer model.

5.1   EXPERIMENTAL-SYSTEM DESCRIPTION
      The general schematic  of the control-mode simulation test setup is
shown in Figure 5-1.   The setup simulates the predictive flow-control
mode of the vapor-generator outlet-temperature control.  The basic system
operation is as follows.
      The feedwater flow is  measured and this information is used to
control the operating  level  of the burner.  Relying on the known plant
characteristics, a prediction curve is established for the open-loop
control of the burner.   At any flow level, the heat input to the vapor
generator is adjusted  to the appropriate steady-state level necessary
to produce the desired steam-generator outlet temperature.  The response
of this predictive control loop can be made rapid compared to the thermal
inertia of the vapor generator.  Thus, the predictive control loop re-
                                                                    5-1

-------
                                          VAPOR
                                        GENERATOR^
PRESSURE
 GAUGE
                                                                          THROTTLE
                                                                            VALVE
WATER
SUPPLY
           FEED PUMP
                                                         PREMIXED
                                                        COMBUSTION
                                                           GAS
     CURVE SHARER
            Figure 5-1 - Control-Mode Simulation  Block Diagram

-------
spends to rapid large-scale changes in power levels as anticipated to
occur during the operation of an automotive power plant.  In order to
correct for errors in predictions that may be caused by initial manu-
facturing tolerances of the plant, or because of changes in plant operat-
ing characteristics with time, a closed-loop temperature controller is
added to the control scheme.  The temperature controller measures the
actual outlet temperature of the vapor generator, compares it to a desired
set point and appropriately adjusts the combustion gas delivery.  This
controller may operate in a proportional control mode at the relatively
low gain that is tolerable with the slow thermal response of the vapor
generator.  The controller will increase the burner heat output when the
outlet temperature is below the desired set point and conversely decrease
the burner output when the outlet temperature exceeds the set point.
This control is secondary and operates at a much slower rate than the pre-
dictive control system.

      The test setup, as shown in Figure 5-1, utilized the SES Model 4
vapor-generator test unit to the maximum extent.  The feedwater is
supplied through a variable-speed pump which is manually operated through
a speed control.  The flow delivered to the vapor generator was measured
using a laboratory-type flowmeter.  The flow information then was pro-
cessed through a curve shaper and used to adjust the blower inlet valve
position using an electric positioner.  The blower motor operated at a
constant speed.  A separate fuel control, which is a part of the SES
Model 4 for test setup, maintained fuel/air ratio.  In order to maintain
operating safety at or near critical gains, the intake valve was designed
with a leakage path to prevent a flameout.  The temperature controller
was a commercially available process instrument and operated in a pro-
portional mode only.

5.2   SIMULATION-HARDWARE-DESIGN TEST RESULTS
      Prior to installation of the control hardware and mating with the
SES Model 4 vapor generator, a number of design verification tests were
conducted at the Bendix Research Laboratories to assure proper function-
ing of the control hardware and thus minimize the time requirements placed
on the already heavily scheduled SES experimental test setup.

      The general block diagram of this test setup is shown in Figure 5-2.
The main combustion-gas control element is the inlet shutter assembly
attached to the blower inlet.  An electric positioning loop is used to
set the shutter position to the desired level.  In order to improve the
overall linearity of the control system, the shutter was designed to pro-
duce an outlet area proportional to the square of the shutter position.
The shutter and blower assembly is shown in Figure 5-3.  The valve has
10 inlet orifices that are equilateral triangles at all times.  As the
inner orifices are rotated with relationship to the outer rectangular
orifices, the equilateral triangular shape is maintained at all shutter
positions.
                                                                     5-3

-------
Ui
I
                 Q.
AMPLIFIER
VOLTAGE
 SUPPLY
                      ฑ15
 MOTOR
CURRENT
 SHUNT
                                                               +28
PRIMARY
VOLTAGE
 SUPPLY
                                                                                          DIFFERENTIAL INPUT
                                                                                         OR FLOATING GROUND
                                    FEEDBACK
                                I
                           SERVO
                          AMPLIFIER
                             I
  WATER
FLOWMETER
                            i
                         STRIP CHART
                          RECORDER
                          FUNCTION
                         GENERATOR
            SIGNAL
                  ACTUATOR
               ROTARY
               SHUTTER
              ASSEMBLY
            BLOWER
                         BLOWER
                          INPUT
                       MANOMETER
   AIR
FLOWMETER
                      BLOWER
                      OUTPUT
                    MANOMETER
                                                                AIR FLOW
                                                            ACTUATOR POSITION
                                                                                                                       in

                                                                                                                       8
              INPUT FOR HYSTERSIS &
            FREQUENCY RESPONSE TESTS
                                       Figure 5-2  - Schematic of Rotary-Shutter  Test Setup

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                                           Figure 5-3 - Shutter  and Blower  Assembly

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      Blower outflow characteristics, as a function of static pressure,
are shown in Figure 5-4.  This curve was used for the combustion-air
control design.
      Experimental results describing blower outflow versus shutter posi-
tions are shown in Figure 5-5.  The figure indicates a good linearity
of the air control system in the operating range.  Dynamic tests were
also conducted to verify the predicted stability and frequency response
of the shutter positioner loop.  The block diagram of the shutter-posi-
tion stability analysis is shown in Figure 5-6.   Using this linearized
model, the analytical prediction of the frequency response for the shutter
position and loop is shown in Figure 5-7 indicating approximately 14 Hz
at the -3 db point.  A frequency-response test of the actual positioner
servo was subsequently made.  The results are given as the amplitude
ratio versus frequency response plot in Figure 5-8.  From this plot it
is readily seen that the -3 db point occurs at approximately 13.5 Hz and
is in excellent agreement with the prior calculations.  The system response
to a step input was also tested.  The test data was recorded and the chart
is shown in Figure 5-9.  The rotary shutter was commanded to close and
then to open, in each case, through an angle of 23 degrees.  The slew rate
achieved was 320 degrees per second with very good linearity and little
overshoot.  A 30 millisecond delay in the response of the air flow to the
shutter movement may also be seen in the figure.  The small reversal in
motor current to correct the slight overshoot is also apparent.

      To simulate anticipated input from the feedwater transducer, a tur-
bine-type flowmeter was connected to a tap-water source and its output
connected to the control-servo amplifier.  The action of the shutter with
this input was evaluated only qualitatively.  The shutter actions observed
were deemed satisfactory.  Additional adjustments were deferred until
such time as tests could be made at SES in conjunction with the burner
and steam generator.

5.3   CLOSED-LOOP STEAM-GENERATOR TEST RESULTS

      Following the design test evaluations, the control-mode simulation
hardware was delivered to SES and installed to control the SES Model 4
vapor generator.  The experimental test program included initial calibra-
tion tests to establish the shutter position required to maintain the
desired outlet steam temperature as a function of fluid flow rate.  It
was found that a straight line approximation of the relationship between
feedwater flow rate and intake valve position resulted in a small tempera-
ture error for the operating range tested. It was the intention of the
test to have such an inaccuracy in the prediction signal and thus be able
to test the closed-loop corrective temperature control system.  The super-
heater outlet temperature was monitored as a metal temperature at the
superheater outlet using a commercial temperature controller.  The output
of the temperature controller was summed with the position command signal
of the combustion-air controller.  The control circuit diagram is shown
in Figure 5-10, and shows that the temperature error summing was accomp-
 5-6

-------
100
                               BLOWER OUTFLOW VS. STATIC PRESSURE*
        "FROM CINCINNATI FAN & VENTILATOR CO. TABLES

        FOR PB-9 BLOWER AT 3450 RPM.
                1234


                         PRESSURE (IN. H2O) —ป•





       Figure 5-4  - Blower Outflow versus Static Pressure
n

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a.
                                                                   5-7

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                   250
I
00
                   200-
                 2 150
                   50
                                                                                                              s
                    CLOSED
                                              10           15           20

                                                       SHUTTER OPENING (DEC.)
25
30
      35


FULL OPEN
                                       Figure 5-5 - Blower Outflow versus Shutter Position

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                                          MOTOR
                                         ARMATURE
          IN-LB
         v/v     i
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                         JTS
                         T2S+1
                            V/RAD
                                          J.
                                          70
                                       GEAR RATIO
                                        OF MOTOR
   1
   7
GEAR RATIO
 TO LOAD
                                       GEAR RATIO (MOTOR
                                       OUTPUT TO FEEDBACK
                                        POTENTIOMETER
Figure 5-6 - Block Diagram of Stability Analysis of Shutter Positioner

-------
(Jl
                soo-
             ui
             a
             3
             ฃ
               100ฐ-
               150ฐ-
               200ฐ •
                                                               PHASE
                                                                    7
                                                                         ^AMPLITUDE
CLOSED LOOP SHUTTER CONTROL
     INPUT AMPLITUDE
  ฑ2.28ฐ (SIMULATED DATA)
                                                                                           10
                                                                    FREQUENCY (HZ)
                                                                                                                \
                                        \
                                                                                                                    \
                                                                                                                         \
                                                                                                                                •20
                                                                                                                                -30
                                                                                                                                    ffl
                                                                                                                                .-40  Q
                                                        c
                                                        111
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                                                        H
                                                    ••50 j
                                                                                                                                --60
                                                                                                                                -70
                                                                                                                              100
                          Figure 5-7  - Analytical Prediction of  Frequency  Response of  Shutter Positioner

-------


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-------
            23 DEC. INPUT
                           i- I  I  i  !
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                                                 ,	i
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                    MOTOR CURRENT
                      0.1 AMPS/DIV
        j^O/TSAMPS
                      I
                             Irrl
           CLOSED -f
      23ฐ
                                 _ VEL. = 230DEG/SEC
        f        I   POSITION _
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                                           1-1—1-
                                                   !  i . I-
                   AIR FLOW
              210 CU. FT.
                        r
                                                                        If
                                EJT    L 10 CU-FT/DIV
                               110CU. FT.
                                    -I—I-
                                ฑ5ES
                                            CHART SPEED 100 MM/SEC
                                                                             00
   Figure 5-9  -  Dynamic Experimental Data Showing Response of  Shutter
                 Positioner  to  a  Step Input
5-12

-------
1000 PSI
                                              FLOW METER
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SHUTTER F.B.
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TEMPERATURE
CONTROLLER
BARBER
COLMAN #537

3.5 VOLTS
                                                                       100 ฐF
                              -15VDC
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         Figure  5-10 - Control Circuit  Diagram

-------
lished in such manner that if the superheater outlet temeprature exceeded
the temperature set point, the intake valve was to close to a degree pro-
portional to the temperature error.  No derivative or integral temperature
control signals were needed in the temperature control loop.  The propor-
tional control mode was selected in order to maintain the simplest type
of control to be used in an automotive power plant.
      Figure 5-11 shows the air flow/shutter position curve established
for the experimental setup*   Again, good linearity is indicated between
air flow and water flow.  The air flow was measured in arbitrary units
defined in terms of the flue-gas differential pressure, as no direct
measurement of combustion gases was practical.  The estimated air flow
control range is approximately 2:1 which is in good agreement with the
preliminary open-loop tests conducted at Bendix Research Laboratories.

      Following operating qualification tests of the control-mode simu-
lation hardware, open-loop hot testing was conducted to determine the
required shutter position for the predictor circuit.  The predictor was
selected to maintain the superheater outlet temperature at 800ฐF.  The
results of these tests are shown in Figure 5-12.  Using the open-loop
data, the straight-line shutter prediction curve, shown as a function
of feedwater flow, was programmed.  This simple prediction curve was
used in all of the following closed-loop tests.  Typical closed-loop
data points are superimposed for comparison in Figure 5-12 and indicate
slight deviation from the programmed straight line.  This is the amount
of prediction error that had to be corrected through the closed-loop
temperature controller at the various parts of the operating range.  As
a reference, one millimeter of water-flow error corresponds to approxi-
mately 75ฐF superheater temperature deviation.

      The control system, operating over a wider range, will require
curve fitting between the predicted shutter position and the feedwater
flow rate.  The calculated combustion gas flow required to maintain
1000ฐF steam temperature is shown as a function of steam flow rate in
Figure 5-13 for the entire anticipated operating rnage of the steam
generator.  It appears that the required performance could be achieved
by fitting the function generator with three straight lines.  This
straight-line approximation would not result in prediction errors in
excess of the correction  capability of a proportional temperature-con-
trol loop.

      Closed-loop stability of the steam temperature control was evaluated
first.  After the stability evaluation was completed, large step changes
in water flow, corresponding to the total gas-flow control range of the
control-mode simulation hardware, were conducted.  At the completion of
the testing, the proportional gain of the temperature loop was increased
until the gain margin of the experimental setup was established.  In the
following, typical examples of the experimental results are presented.
 5-U

-------
                      10
20           30
   WATER FLOW
                                                         40
                                                                     50 MM
Figure 5-11 - Experimental Results of Airflow-Shutter Position  Test
                                                                   5-15

-------
 34ฐ-50 MM
       40
SHUTTER
POSITION
                            Q SET UP CELL CHECK
                            O RUN
                            A OPEN LOOP CHECK RUN
                       10
20
                                                                 40
 50 MM
1.22GPM
      Figure 5-12 -  Shutter Position Required to Maintain  800ฐF Superheater
                     Outlet Temperature  as  a Function of Feedwater Flow

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                       COMBUSTION AIR INFLOW (W'G) VS.

                            WATER INFLOW (Ws)
       (LB/SEC)
               0.1
                                                         344
                                                         258
WG|



(FT3/MIN)
                                                         172
                                                                  CM
                                                                  in
                                                                  a>


                                                                  s
                                                 Ws (LB/SEC)




                                                 Ws (GPM) —
Figure 5-13  -  Calculated Combustion-Gas  Flow Required to Maintain  1000ฐF

               Steam Temperature as a Function of Steam Flow
                                                                      5-17

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      A small step change in water flow is shown in Figure 5-14 for a
reduction in water flow rate.  The predictive loop immediately commands
a reduction in air intake-valve position.  It may be seen that the area
of the combustion control valve closely follows the water flow changes.
A temperature undershoot of -8ฐF and a maximum overshoot of +20ฐF is
rapidly damped out.  The total shift in the outlet steam temperature is
approximately 4ฐF.
      Closed-loop control characteristics of the predictive flow controller,
in response to a large step change in water flow, is shown in Figure 5-15.
Water flow is increased from the minimum to near the maximum operating
level.  It again may be observed that the predictor loop commands the
combustion-air control area to correspond with water flow changes.  A
temperature excursion of +60ฐF overshoot and -6ฐF undershoot is again
rapidly corrected.  The total shift of the operating outlet temperature
is -6ฐF for this large step input.  Several other steps of input signals
were used to verify this stable operation of the control system.  In
general, the results were similar to the example shown above.

      The proportional gain of the temperature loop was increased in two
steps to find the critical gain level for the experimental test setup.
The critical point was reached at approximately three times the initial
gain setting.  The operating characteristics at this setting are shown
in Figure 5-16.  The temperature oscillations are slowly damped out,
indicating a closed-loop proportional temperature gain near to the criti-
cal system gain at the operating conditions.  The overshoot of the system
is approximately 80ฐF and the undershoot is approximately 3ฐF.  These
results establish that the operating gain margin for the system corresponds
to previous computer predictions.

      The experimental closed-loop results presented clearly establish
verification of the flow control mode as proposed for the temperature
control system, and also that the hybrid computer model is a valid tool
for the control system design of the steam generator.
5-18

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                         WATER  30 	
                          FLOW
                                        .,     .           .

                                     "1- -'-          -  V-'- ' Mf
                          AIR
                         VALVE  20
                        POSITION  Q
                          MM
                                40

                                50

                              1 SEC
                            TIME MARKS
Ul
                            Figure 5-14 - Experimental  Closed-Loop Temperature Control, Small  Step
                                           Input in Water  Flow at Design Gain

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NJ
O
            WATER  30 	
            FLOW
             MM   20 	
            AIR
           VALVE   20 ~
          POSITION
            MM
                1 SEC
             TIME MARKS
                              Figure 5-15 -  Experimental Closed-Loop  Temperature Control, Large
                                             Step  Input in Water Flow  at  Design Gain

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745— BH
                         Figure 5-16 - Experimental Closed-Loop Temperature  Control,  Large  Step
                                       from Low  to High Water  Flow near  Critical  Gain (3X Design)
                                                       •
                                                                                                                                                5-21

-------