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Calhoun: The NPS Institutional Archive
Theses and Dissertations Thesis Collection
1980
An interactive code for a pressurized water reactor
incorporating temperature and Xenon feedback.
Heath, Gregory Garver
Monterey, California. Naval Postgraduate School
http://hdl.handle.net/10945/17616
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MONTEREY, CALIF 339*0
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Monterey, California
THESISAN INTERACTIVE CODE FOR A PRESSURIZED WATER REACTOR
INCORPORATING TEMPERATURE AND XENON FEEDBACK
by
Gregory Garver Heath
June 1980
Thesis Advisor P. J. Marto
Approved for public release; distribution unlimited
T196S81
UnclassifiedSECURITY CLASSIFICATION Of THIS *tSt fW*»m* Dmia EnffH)
REPORT DOCUMENTATION PAGErcrort nuuK*
READ INSTRUCTIONSBEFORE COMPLETING FORM
2. OOVT ACCKMIOM NO. » »ICi»iInT'$ CATALOG NUMIE*
4 TITLE rantf SuA«(((«)
An Interactive Code for a Pressurized WaterReactor Incorporating Temperature and XenonFeedback
»• TYPE OF REPORT * PERIOO COVERED
Master's Thesis: June 1980
• PERFORMING OMC. REPORT NuMIIN
7. AuTMORru i. CONTRACT OH GRANT NUMltMn
Gregory Garver Heath
# »(MO*MINd ORGANIZATION NAME anO ADDRESS
Naval Postgraduate SchoolMonterey, California 93940
10. PROGRAM ELEMENT. PROJECT taskAREA * WORK UNIT NUMIIM
II CONTROLLING OFFICE NAME ANO ADDRESS
Naval Postgraduate SchoolMonterey, California 93940
12. REPORT DATE
June 1980II. NUMBER OF PAGES
93
IT MONITORING AGENCY NAME » » <• CSSCI » (it Itarmnt Irom Contrnlltnt OMIe«l
Naval Postgraduate SchoolMonterey, California 93940
I*, security class. «h >m ( report)
Unclassified
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IE. DISTRIBUTION STATEMENT (ot t*i» *»»«/ii
Approved for public release; distribution unlimited
17. DISTRIBUTION STATEMENT fat r^. f.-rmr, ~(m~4 In !•<>* 20. II dllfrmni Kmmon)
IS. SUPPLEMENTARY NOTES
IS KEY WORDS (Continue em r»w»r»» ><*• II n»c»*»arr and i4i\itty my bloc* nsm»mmr>
Pressurized Water Reactor, Xenon Feedback, Moderator Feedback
20. ABSTRACT (Conllnttm on rmvmrmm «/<»• II n»c»*»+rr an* Idrntitllr my mlmcM •»«j
An interactive computer model of a highly enriched pressurized water reactor
was developed, using the applicable plant parameters from Shippingport AtomicPower Station. The point reactor kinetics equations for one delayed neutronprecursor group were linearized using small perturbation theory. The model
included both moderator and Xenon-135 reactivity feedback effects, as well as
an automatic reactor protection and average reactor coolant temperature control
system. The thermal response of the model plant was simulated for normal
operating transients induced either by control rod or turbine load changes. The
nn womm*» I JAN 71 1473 EDITION O* ' MOV SS is OBSOLETE
S/N 0102-014- 4401:
Unclassified
Unclassified<"*>•• «•«• *««•*•*
post-shutdown Xenon transient response was also modeled. The interactive
program was coded in FORTRAN-IV language, and the simulation program was
coded in IBM CSMP-III language.
DD Form 1473 Unclassified,,J <la.1j3 .. . i i
Approved for public release; distribution unlimited
An Interactive Code for a
Pressurized Water ReactorIncorporating
Temperature and Xenon Feedback
by
Gregory Garver HeathLieutenant Commander, United States Navy
B.S.M.E., United States Naval Academy, 1971
Submitted in partial fulfillment of therequirements for the degree of
MASTER OF SCIENCE IN ENGINEERING SCIENCE
from the
NAVAL POSTGRADUATE SCHOOLJune, 1980
ABSTRACT
An interactive computer model of a highly enriched pressurized water
reactor was developed, using the applicable plant parameters from the
Shippingport Atomic Power Station. The point reactor kinetics equations
for one delayed neutron precursor group were linearized using small
perturbation theory. The model included both moderator and Xenon-135
reactivity feedback effects, as well as an automatic reactor protection
and average reactor coolant temperature control system. The thermal
response of the model plant was simulated for normal operating transients
induced either by control rod or turbine load changes. The post shutdown
Xenon transient response was also modeled. The interactive program was
coded in FORTRAN-IV language, and the simulation program was coded in
IBM CSMP-III language.
TABLE OF CONTENTS
I, INTRODUCTION - 9
II. MODEL DESIGN CONSIDERATIONS - 11
A. PRESSURIZED WATER REACTOR - - 11
B. MODEL AND PROGRAM FLEXIBILITY 13
III. MODEL DEVELOPMENT 15
A. POINT REACTOR KINETICS EQUATIONS - 15
1. Zero Power Reactor Transfer Function 17
2. Limitations 18
B. REACTIVITY FEEDBACK MECHANISMS — 19
1. Xenon-135 Transfer Function 21
2. Moderator Temperature Feedback 24
C. THERMAL ANALYSIS - 26
1. Reactor Heat Transfer Function 26
2. Heat Exchanger Transfer Function 29
3. Primary Piping Transfer Functions 35
D. OVERALL PLANT BLOCK DIAGRAM - - 37
E. MODEL CONSTANTS 38
IV. RESULTS 42
A. INTERACTIVE PROGRAM 42
B. SIMULATION PROGRAM - 48
V. CONCLUSIONS AND RECOMMENDATIONS - 54
APPENDIX A: DEVELOPMENT OF THE XENON FEEDBACK TRANSFERFUNCTION COMPONENTS 56
APPENDIX B: DEVELOPMENT OF THE REACTOR HEAT TRANSFERFUNCTION 61
APPENDIX C: DEVELOPMENT OF THE HEAT EXCHANGER TRANSFERFUNCTION 65
APPENDIX D: DEVELOPMENT OF THE PRIMARY PIPING FUNCTIONS -- 67
APPENDIX E: INTERACTIVE PROGRAM LISTING - 71
APPENDIX F: SIMULATION PROGRAM LISTING 78
APPENDIX G: INTERFACE PROGRAM LISTING 91
LIST OF REFERENCES 92
INITIAL DISTRIBUTION LIST 93
LIST OF FIGURES
1. Schematic of the Model Plant 12
2. Block Diagram of the Zero Power Reactor TransferFunction 18
3. Block Diagram of the Reactor Transfer Function withXenon-135 and Moderator Temperature Feedback Loops,and External Reactivity 20
4. Xe-135 Decay Scheme - - 21
5. Simplified Xenon-135 Decay Scheme 22
6. Block Diagram of the Reactor Transfer Functionwith Xenon-135 Feedback and External Reactivity 24
7. Block Diagram of the Reactor Heat Transfer Function 29
8. Block Diagram of the Heat Exchanger Transfer Function 34
9. Block Diagram of the Transport and Mixing DelayFunctions 38
10. Overall Plant Block Diagram 39
11. Interactive Program, Introduction and Instructions 43
12. Interactive Program, Initial Power Level andSimulation Choice Inputs 44
13„ Post-shutdown Xenon-135 Behavior 45
14. Interactive Program, Control Rod Movement Inputs 46
15. Interactive Program, Turbine Load Change Inputs 47
16„ Interactive Program, Logic Routine Example 49
17. Power, Reactivity, and Temperature Response to a
20 to 40% Turbine Load Change - 50
18. Power, Reactivity, and Temperature Response to a
60 to 40% Turbine Load Change - 51
19o Power, Reactivity, and Temperature Response to a
10 second Inward Control Rod Movement 52
20. Power, Reactivity, and Temperature Response to a
10 second Outward Control Rod Movement 53
D-l . Mixing Volume 69
I. INTRODUCTION
Computers have been used in the design and analysis of nuclear
reactors since the inception of reactor technology, and many codes have
been developed. More recently, digital computer programs have been used
as learning devices for nuclear engineering students. The formulation
of the computational problem of predicting the behavior of a nuclear re-
actor has been greatly facilitated with the use of high level computer
languages.
The purpose of this work was to develop a computer assisted learning
device to be used by students taking nuclear engineering courses at the
Naval Postgraduate School. The program graphically displays the simulated
kinetic and thermal transient responses of a pressurized water reactor
power plant. The reactivity feedback effects of Xenon-135 poisoning and
moderator temperature are incorporated into the model. Normal operating
transients, starting from a steady state critical condition, can be
induced by ramp changes in either control rod position or turbine load.
The initial reactor power level and reactivity change mechanism are
chosen by the program user. For user convenience, these inputs are
prompted and entered interactively, after which the simulation is run
with no further user action required.
A pressurized water reactor (PWR) was modeled as it is the most
common type of reactor used for power plant application. As a realistic
reference, the model design incorporated the applicable characteristics
of the Shippingport Atomic Power Station, whose nuclear parameters were
the most compatible, of available PWR core data, with the assumption of
a highly enriched core made in the model design,
9
The model was developed by treating the reactor kinetics and other
plant component thermodynamic relationships as transfer functions [1].
The point reactor kinetics equations were linearized using a Taylor ex-
pansion or small perturbation technique [2]. This same perturbation
technique and lumped parameter analysis were applied to the plant's
linear differential heat transfer and Xenon-135 equations [3]. The
individual transfer functions were then integrated into an overall plant
block diagram. The program also features a reactor protection and
average reactor coolant temperature control scheme.
IBM CSMP-III language was chosen to formulate the model simulation
as it has inherent routines to invert the transfer functions to the time
doma i n
,
10
II. MODEL DESIGN CONSIDERATIONS
Lumped parameter analysis and first order perturbation theory were
used throughout the model development. All the nuclear variables which
were used in the reactor kinetics and Xenon-135 decay equations were
considered to be averaged values of the variables over the neutron
energy spectrum. Similarly, the thermodynamic variables which were
used in the plant's heat transfer equations were considered to be
averaged values over the volume of the particular component.
Because of these approximations and other assumptions made in the
model development, the simulation should not be considered a design or
stability analysis. Instead, the simulation shows the model's large
scale reactor kinetic and thermal trends during normal operating
transients.
For similar reasons, although certain plant parameters of the
Shippingport Atomic Power Station were incorporated into the model, the
simulation cannot be considered to reflect the operating characteristics
of this power plant. The. Shippingport core is a seed and blanket type
whereas the model core is a uniform mixture of highly enriched fuel
and support materials.
A. PRESSURIZED WATER REACTOR
The pressurized water reactor (PWR) is the most widely used reactor
type in central power plant applications and the only type currently
used in naval propulsion.
A simplified schematic of the modeled PWR plant is shown in Figure
1. The modeled reactor contains a highly enriched core which is light
11
PRESSURIZER
COOLANTPUMP
THROTTLEVALVEa:
CONDENSATEPUMP
Figure 1. Schematic of the Model Plant
water moderated and cooled. The modeled plant contains two closed loop
thermodynamic systems coupled by a heat exchanger.
In the primary system, heat generated from thermal fission is trans-
ferred to the coolant as it passes through the reactor, raising the
coolant's temperature. The high temperature coolant leaving the reactor
is circulated by a pump through tubes inside of a heat exchanger where
it gives up some heat to a secondary fluid. A pressurizer maintains
the primary coolant at a sufficiently high pressure so that the bulk
temperature of the coolant is kept below the saturation temperature.
On the secondary side of the heat exchanger, the temperature of the
entering feedwater is raised to the boiling point and saturated steam
is produced. The steam is delivered to a turbine, is condensed and
returned to the heat exchanger, completing the second closed loop
12
B. MODEL AND PROGRAM FLEXIBILITY
The model was designed to simulate only normal operating transients
from an initial steady state condition. Reactor accidents and startup
were not considered in the model development. However, the model's re-
actor control module will simulate either a full or constant insertion
of control rods if certain parameters are exceeded. This feature was
not incorporated for accident analysis, but to keep parameters within
the limits of the assumptions made in the model development.
The applicable Shippingport plant characteristics are fixed in the
simulation program. User inputs are limited to choosing initial power
level, from which other initial parameters are adjusted, and the plant
perturbation mechanism, either control rod movement or turbine load
change. These perturbations occur at fixed rates which limit the amount
of reactivity that can be inserted during the simulation.
Originally it was envisioned that an overall program would feature
not only interactive input capability but also produce real time graphi-
cal displays of the transient for a user at a time sharing computer
terminal which had graphics capability. However, the computer language
(IBM CSMP-III) required to solve the model's algorithms was not avail-
able on the time sharing system, and furthermore, the programs long
execution time precluded any real time response.
In order to still provide as much user facility as possible with
these restrictions, the interactive feature was partially retained for
data input. Using a computer language available on the time sharing
system (FORTRAN-IV), a program which interactively prompts the user to
13
enter specified controlling parameters was developed. This program in-
corporates logic routines which permit inputs only within requested
ranges „ This feature allows for data reentry if a user error is made,
and ensures that only inputs compatible with the model development's
assumptions are used in the simulation.
After the input is completed, the initial transient conditions are
displayed at the terminal. A separate internal control program then
transfers the user supplied inputs into the simulation program. The
control program then transfers the simulation program to the batch
processing system for execution., The interactive program notifies the
user that this has been done. Hard copy plots of the transient are
subsequently produced. The post-shutdown Xenon behavior module is
located in the interactive program and produces real time graphical
displays at the terminal.
Thus there are three distinct programs:
1. An interactive program to prompt and receive user input,
and also simulate post-shutdown Xenon behavior.
2. A batch processed program which simulates the transientresponse to the user inputs.'
3. An internal control program which interfaces these two
programs.
14
III. MODEL DEVELOPMENT
A. POINT REACTOR KINETICS EQUATIONS
A lengthly and formal derivation of the point kinetics equations is
found in Reference 4 and will not be repeated here. More generally these
equations are obtained from one group neutron diffusion theory with the
assumption that the neutron flux is deparable in time and space, and
with the inclusion of delayed neutrons [5]. They are listed below.
djitl. £ltk£_ Ht)+ £ N C.(t) + Q(t) (1)
^.(t)iJ 1
0(t) - A,C,(t) 1 = 1,2,... (2)dt A y ^' Mi
where 0(t) = Flux amplitude function
/0(t) = Reactivity
C.(t) = Effective concentration of the ith delayedneutron precursor group
Q(t) = Extraneous delayed neutron source strength
a. = Effective delayed neutron fraction from
the ith group
*i = Decay constant of the ith group
and /3 » 2A s Total effective delayed neutroni fraction
These equations are referred to as the point reactor kinetics equa-
tions not because the reactor is considered as a single point in their
application, but since spatial variations are neglected, /3j,A;and A^
are assumed to be constant.
If the neutron flux spatial variation is assumed to be time invari-
ant, 0(t) can be considered to represent the number of neutrons in the
15
core. Equation (1) may then be seen as a neutron rate equation where
the three terms on the right hand side represent the rate of production
of prompt, delayed, and source neutrons respectively.
Equation (2) is a rate equation for the ith delayed neutron precursor
group. The two terms on the right hand side represent production and
decay rates respectively. The precursors are comprised of approximately
thirty isotopes which historically have been divided into six groups
-1 235with decay constants ranging from 0.0124 to 3.0 seconds for U. In
the model, the precursors were considered to be represented by one
effective group characterized by an averaged decay constant [2],
A. Z
-1
and an effective delayed neutron fraction
1*1 1
In equations (1) and (2), reactor power may be substituted for neutron
flux if the precursor concentration is modified by C = E. S*C . .
since
P-Ef Z,»
where P
Ef
If
= Reactor power
= Energy released per fission
= Macroscopic fission cross section
= Integrated one group neutron flux
For a reactor operating at power, the source term contribution to
the overall neutron population is negligible, i.e., Q « 0.
16
Incorporating these assumptions, equations (1) and (2) reduce to:
dPM = S&-Z* P(t) + A C (t) (3)
dCM = -£p(t) - XC(t) (4)
1 . Zero Power Reactor Transfer Function
Consider an initially critical steady state reactor at some power
level P at t £ 0. Equations (3) and (4) each yield:
C =-TX^ P (5)A.A ov
'
Now let P(t) = P + </Wt)
C(t) = CQ
+ </c(t)(6)
/>(t) =/Q
+ <#(t)
where the zero subscript denotes the initial steady state valueand the delta prefix a small perturbation imposed at t = aboutthis value.
Noting that /& =0 for a critical reactor and substituting equa-
tions (5) and (6) into equations (3) and (4) yields for t >
fa Mt) - % S/xt)- 4r Ht)+ x/c(t) (7)
fa <A(t) = -£<A(t) - \</t(t) (8)
where the —-^ term in equation (7) has been neglected.
Taking the Laplace transforms of equations (7) and (8):
s ^>( s ) = A.^($) - JjL<fp(s) +- A</C(s) (9)
SC(s) = —cPPts) - \<Pc& (10)
17
Solving equation (10) for C(s), substituting this into equation
(9), and rearranging, yields:
c/Vcs) _ ?o
tyw s(a* S+A -)
(ID
This is the zero power reactor transfer function, so called because
the reactor is assumed to be operating at a sufficiently low enough
power that no feedback effects are realized. These effects are examined
in the following sections.
Equation (11) may be represented in block diagram form as shown in
Figure 2.
e/>(s>
PGZ(s)
</*P<s)
Figure 2. Block Diagram of the Zero Power ReactorTransfer Function
where Z(s) =
S (a /3
S-hX -)
2. Limitations of the Point Reactor Kinetics Equations
The point reactor kinetics equations derivation is based on
the assumption that the spatial dependence of the neutron flux is
negligible. This assumption limits the validity of these equations to
transients where this remains a reasonable approximation such as those
which result in only small changes in reactivity. The small perturba-
tion technique used in the development of the zero power reactor
18
transfer function also requires only small changes in reactivity if the
first order approximation is to hold. Specifically, /> must be less
than 0.5/3 . Physically, when ^/Jis greater than /3 the reactor is
critical on prompt neutons alone. The simulated transients imposed on
the model were limited to ensure that excessive amounts of reactivity
were not introduced.
Bo REACTIVITY FEEDBACK MECHANISMS
In the zero power reactor point reactor, the power level is assumed
to be so low that it does not affect the reactivity, thus there are no
feedback effects. However, for a reactor operating at a useful power
level, feedback effects do exist and the reactivity becomes an implicit
function of the reactor power level (or neutron flux). This dependence
arises since reactivity depends on macroscopic cross sections which in-
volve the atomic number densities of material in the core:
X = nct
where S = Macroscopic cross section (cm""*)
N = Atomic number density (atoms/cm 3 )
C = Microscopic cross section (cm z)
The atomic number density can depend upon the reactor power level
since the concentrations of certain nuclei are constantly changing due
to neutron interactions. Material densities also depend upon tempera-
ture which is a function of reactor power level and hence the flux.
Duderstadt, J. J. and Hamilton, L. J., Nuclear Reactor Analysis ,
1st ed., Wiley, 1976.
19
The two feedback mechanisms considered in the model were Xenon-let and
moderator temperature.
Reactivity can also be changed directly by an external source such
as control rods containing a neutron absorbing material. Thus the over-
all reactivity can be written as:
where "fix = F P
<#>T= FT
P
cfo = Overall core reactivity
ofee = Externally added reactivity
<//Ox = Xenon-135 feedback reactivity
<fer = Moderator temperature feedback reactivity
Fx = Xeonon transfer function
Fj = Moderator temperature transfer function
Equation (12) with the previously developed zero power reactor
transfer function, given by equation (11), are incorporated in the
block diagram below:
(12)
<&>&)
u*
'*
\ <fal* .P Z(»)
</P<s)
J\*
FT(s)
Figure 3. Block Diagram of the Reactor Transfer Function
with Xenon-135 and Moderator TemperatureFeedback Loops, and External Reactivity
20
The next step in the model development was to derive the Xenon-135
and moderator temperature transfer functions.
1 . Xenon-135 Transfer Function
Xenon-135 is the most significant fission product poidon because
of its enormous thermal neutron absorption cross section and relatively
large fission yield. This isotope is not only produced directly from
fission but also from the decay of other fission products as shown in
Figure 4.
Figure 4. Xe-135 Decay Scheme
Since the /3>~ decay of Iodine-135 (
135I) and Xenon-135 (
135Xe), with
the largest half-lives, are the controlling steps in this decay scheme,
it was simplified by making the following assumptions:
1) All I is produced directly from fission (the production of
Antimony-135 ( Sb) and subsequent decay to I is consideredinstantaneous).
135m,2) The short lived metastable Xe is ignored.
1 353) The removal of I by neutron absorption is negligible for
the neuton flux levels used in the model (10 neutrons/cm sec)
21
4) All the135
I decays to135
Xe,
With these assumptions the effective decay scheme is shown in Figure
5.
Ax 135Cs
Figure 5. Simplified Xe-135 Decay Scheme
135
135Using this decay scheme, the resulting rate equations for I and
Xe are:
c/X(t)
(13)
where
5^= VxZ^tfCt) + AiI(*)-Ax X(t)-<3a x 0<:-t)X(t) (14)
135 3I = I number density (atoms/cm )
135 3X = Xe number density (atoms/cm )
2f = Macroscopic fission cross section (cm" )
= Integrated one group neutron flux (neutrons/cm sec)
1 35dr = Effective I fission yield
1 35^
x= Effective Xe fission yield
A_ = I decay constant (sec" )
1 35 -1
A x= Xe decay constant (sec )
<3" = Xe microscopic thermal neutron absorption crossa
section (cm?)
22
Again, using the first order perturbation technique, let
I(t) = IQ
+ cTl(t)
X(t) = XQ
+ o^X(t) (15)
0(t) = jB
Q+ c/fe(t)
where the zero subscript denotes an initial steady state value andthe delta prefix denotes a small perturbation about this value.
Upon substituting equations (15) into equations (13) and (14), and
taking the Laplace transform, the relationship between the perturbation
1 35in Xe and is derived (See Appendix A):
— = ; : —— = <=»x (S) (16)</*0<s} S 2 + (03X^o tV*Ai)st Ar(<3a
x ^ +
A
x )
135The change in reactivity caused by a small perturbation in Xe
concentration is also derived in Appendix A:
<7^X(S> 1? = *x (17)^<s > X + SL U
<5*
1 35where X = Equilibrium Xe number density before the
perturbation (atoms/cm^)
U = Uranium-235 ( U) number density (atoms/cm )
<% = Microscopic thermal neutron absorption cross
section of U (cm )
A perturbation in neuton flux ( o^0) is directly proportional to a
perturbation in reactor power ( cfp) as shown by:
23
where Ef
= Energy released per fission
O O C 1
Hf= Macroscopic fission cross section of U (cm )
or
c/<Zks)= k
The product of equations (17), (16), and (18) yields
(18)
^x(s) ^X(S) </fr(s) <^x(s) _= —I — ^X^X (S) K - rV(SJ</>X(s) ^(s) c^PCS) ^P(s)
x xwi* x (19)
Equation (19) is the transfer function relating reactor power and
135Xe feedback reactivity. This equation is shown in block diagram
form in Figure 6 where it has been incorporated with the previously de-
rived zero power transfer function given by equation (11).
<fp#)
d/Ogs)
«x ^ </X($)
QJ»^«/b(s)
K
c*
sP Zls)
</>(s)
)
Figure 6. Block Diagram of the Reactor TransferFunction with Xenon-135 Feedback andExternal Reactivity
2. Moderator Temperature Feedback
Reactivity feedback from changes in moderator temperature occurs
as a result of changes with moderator density. The density is also a
24
function of pressure, however, the pressure coefficient of reactivity is
typically two orders of magnitude smaller than the temperature coeffi-
cient, and was therefore not considered in the model development. The
moderator density affects the moderator number density and hence the
macroscopic scattering cross section of the moderator.
The primary mechanism for the thermal ization of the prompt and de-
layed neutrons is by elastic scattering interactions with the moderator
nuclei. Hence, variations in the macroscopic scattering cross section
will affect the rate at which neutrons become thermal ized. Changes in
the thermal ization rate affect the fission rate, or reactor power level.
Power level changes affect the moderator temperature, thus a feedback
loop is created.
Because the moderator is also the coolant in the model, its temper-
ature is not only a function of reactor power but also a function of the
heat transfer process occurring in the heat exchanger, thus complicating
the feedback loop.
Therefore, in order to develop this feedback mechanism analytically,
it is first necessary to model the plant's heat transfer processes.
This thermal analysis is done in the following section.
Another temperature feedback mechanism is the broadening of the
Uranium-238 resonance absorption cross section for neutrons with increas-
ing temperature. Because a highly enriched core was assumed in the
model, with little Uranium-238, this effect was not considered.
25
C. THERMAL ANALYSIS
1 . Reactor Heat Transfer Function
A lumped parameter model was assumed. This simplification pro-
vided a set of ordinary differential equation which were sufficiently
accurate for the simulated normal operating transients. In the lumped
parameter model, heat transfer in the reactor was assumed to occur at a
single point. Thus, spatial variations were neglected. The core was
considered to be a homogenized mixture of the uranium alloy fuel, the
fuel cladding, and other structural materials, with a constant thermal
capacity,,
The equation for the heat flow from the core was obtained from a
basic heat balance. The heat generated from fission equals the heat
required to change the temperature of the core materials plus the heat
transferred to the coolant. On a per unit time basis, the heat balance
is:
P(t) = C,
dTF(t)
dt+ h
FMTF Ct)-
TAV
(t) (20)
where P(t) = Total Power generated in the core (Btu/sec)
T-(t)= Average temperature of the core materials (°F)
T.v(t) = Average reactor coolant temperature (°F)
CF
= Total thermal capacity of the core materials (Btu/°F)
hF„ = Total heat transfer coefficient (Btu/° F sec)
A similar heat transfer balance must also hold for the heat being
transferred to the coolant and transported out of the core. This heat
is the last term of equation (20) and is transferred to the coolant as:
26
hF«[TFCt)-TAv(*)] = CM s!2*<S + mmc [V^-Tc,^] (21)
where CM= Total thermal capacity of coolant in the core
M(Btu/°F)
m = Coolant mass flow rate (lbm/sec)
TH
= Average reactor coolant outlet temperature (°F)
o
Tp
= Average reactor coolant inlet temperature (°F)L
i
C = Specific heat of reactor coolant (Btu/lbm°F)
For simplification, the average reactor coolant temperature is
assumed to be given by:
TAv(t) = [T^(0+ THo (t)] /2 (22)
Rearranging equation (20) yields:
TfC-O -ffi dl££l = TAvCO + li PCt) (23)dt Cp
where • l - tt—"fm
Substituting equation (22) into equation (21) to eliminate T"
Hyields:
TAv(t)[lt2|l *Ya <%£! = TF (t) + 2|Tc,(t) (24)L To J dt 'o 1
where Jo = —— , and
^FM
27
Now 1 et
P(t) = P + c/,
p(t)
Tp(t)= T
p+ cTT
F(t)
o
TAV
(t)= TAV
+ ^ TAV
(t) (25)
o
TH f*)- T
Ho+ ^T
Ho(t)
Tc.
(t)= Tci
+ ^c (t)
1
where the zero subscript denotes a steady state value and the
prefix a small perturbation about this value.
Substituting equations (25) into equations (22), (23), and (24),
eliminating the average temperatures TF(t) and T.„(t), and taking Laplace
transforms, the following relationship between reactor coolant inlet and
outlet temperature perturbations is derived (see Appendix B):
<^TH<> <s) = —(26)
where
tf = lj-'° =
This equation is the reactor heat transfer function for the reactor
coolant outlet temperature as a function of the reactor coolant inlet
temperature and reactor power.
A block diagram representation of this transfer function is shown
in Figure 7.
28
</*PU)
Figure 7. Block Diagram of the Reactor HeatTransfer Function
In Figure 7
BC(S) = {?-HH«)-<H
GH(S) =
{^-HfihHH2. Heat Exchanger Transfer Function
As in the derivation of the reactor heat transfer function, a
lumped parameter model was assumed for the heat exchanger . Two points
of energy storage were assumed, the primary coolant water and the water
on the secondary side of the heat exchanger. The thermal capacity of
the heat exchanger metal was included with that for the secondary water,
since the thermal resistance on the primary side is predominant. A
further assumption was made that the time spent by the primary coolant
29
while it passed through the heat exchanger was negligible in comparison
with the time spent in the primary piping.
The state of the steam produced on the secondary side of the heat
exchanger was assumed to always be dry and saturated and that the
secondary water is always at the saturation temperature for the existing
pressure. These assumptions were justified because moisture separators
and recirculation can provide high quality steam and preheating of the
feedwater.
With these assumptions, the following equations were obtained from
a heat balance per unit time:
%c [^Ctl-Tcfel] = Z» £E^> + h™[TAvecO-TsOo] (27)
hmffaeta -TsCtl ] = Cg sO|£} + pl(o (28)
where Tu .(t) = Heat exchanger coolant inlet temperature (°F)HI
Tr
(t) = Heat exchanger coolant outlet temperature (°F)
T.w (t) = Heat exchanger average coolant temperature (°F)AV
B
Ts(t) = Saturated steam temperature (°F)
h TM = Total heat transfer coefficient (BTU/°F.sec)In
C,. = Total thermal capacity of coolant in the heatn
exchanger (BTU/°F)
Cs
= Total thermal capacity of heat exchanger metal
and secondary water and steam (Btu/°F)
P.(t) = Power delivered by the heat exchanger (BTU/sec)
m = Coolant mass flow rate (lbm/sec)m
30
The total heat transfer coefficient, a function of the primary coolant
flow rate (a constant in the model) and the heat transfer characteristics
of the heat exchanger, was assumed to be constant. Also, the time delay
in transferring heat across the heat exchanger tubes was neglected.
This changed the shape of the initial thermal transient but had little
effect on the basic dynamics of the secondary loop.
Again, as a simplification, the average temperature of the coolant
in the heat exchanger was assumed to be given by
X - \ {\ +V (29)
The power delivered by the steam generator is proportional to the
product of the steam flow rate and the difference in enthalpy between
the steam and feedwater
Ktm [Tavb (0 - TsCt)] = m3 (HS - Hpv, ) * CS s!|£!
where m = Steam flow rate (lbm/sec)
H = Saturated steam enthalpy (Btu/lbm)
HF
= Feedwater enthalpy (Btu/lbm)
This equation assumes the steam and feedwater flow rates are always
equal and thus neglects any instabilities in the secondary steam.
As discussed in Reference 4, the enthalpy of saturated steam is
nearly constant over a wide range of pressure, varying from 1198 Btu/lbm
to 1204 Btu/lbm over a pressure range from 200 to 800 psig. The feed-
water enthalpy,which depends on condenser pressure, is usually between
50 and 100 Btu/lbm. Thus the enthalpy difference may be regarded as a
31
constant, and the power delivered by the heat exchanger is directly
proportional to the steam flow rate.
The impedance to steam flow caused by the turbine is nearly indepen-
dent of turbine speed. If constant backpressure is assumed, the steam
flow rate is directly proportional to the throttle opening at a given
pressure. Thus,
ms
= y*sA
2where /& = Saturated steam pressure (Ibf/in )
A = Proportionality factor which is a functionof the throttle setting 1 bm in^
Ibf sec
The numerous assumptions made in the heat exchanger model development
limit the accuracy of the resulting equations. However, the errors in-
volved in these assumptions are usually less than the amount of uncertainty
in the engineering valves of the coefficients used in the equations.
Recalling the previous assumption that
dTAV
B
(t)
-at— =° • and
Substituting equation (29) into equation (27), yields:
mMc[THl (t)- Tco(t)] = WT„r^[TH^t)+ TcoOo] - Ts(o]
Solving equation (30) for Tr , gives:Lo
(30)
Tc C*> = **(.)- tmi-O. (31)
fl+Kl)
32
where K, =
Ktm
Substituting equation (29) into equation (28) gives:
k™ (l [T̂ + Tci(^
J- Ts co . =r cs i^i1
+ put)
Rearranging this expression,
cs cJTsct) _ - r i i „(32)
Let THi
(t) = TH>
+ <AH<(t)
TC
f*)= T
Co °,
T e (t) = T v
PL(t) = P
L+
o
<As(t)
^Ts(t)
c/V,(t)
(33)
where the lowest zero subscript denotes a steady state value and
the delta prefix a small perturbation about this value.
By substituting equations (33) into equations (31) and (32), and
taking the Laplace transform, the following set of equations is derived
(see Appendix C)
:
<ATs (s) =Is + 1
(34)
^CoCS) = 2</Ts(s) - <AT^(s)[l-K±l
1+Ki
where T L = -r-*-
(35)
33
These equations are represented in a block diagram in Figure S.
JTC (s)
;rs*<LJ (
1
I
1 i
J
•Kl
<JTs(s>1
-
1*
Kl
Kl
1
' i
Jtc Is)
C- ;
v £+
J ^ ' \ji \
<*TH (rt
1
•
M
<S(!)
Figure 8. Block Diagram of the Heat ExchangerTransfer Function
With </tu and <P\>. as inputs, equations (34) and (35) form anH
iL
algebraic loop in </*Tr and c/t~. These two equations were solved
with the use of an inherent functional routine available in the CSMP-III
language which was used to formulate the model.
34
3. Primary Piping Transfer Functions
While circulating through the primary loop, the coolant under-
goes mixing and transport delay effects. For example, a transient in
the coolant temperature at the heat exchanger outlet does not appear at
the reactor inlet until sometime later. Where there are volume or flow
direction changes, as in the reactor coolant inlet plenum, mixing occurs
causing a smearing of a temperature transient. Both of these effects
were approximated in the model by combinations of two types of time
delays: a pure transport delay and a simple mixing delay.
Assuming no mixing in the primary piping and no heat loss with
perfectly insulated pipes, pure transport delays are encountered in the
piping runs between the reactor and heat exchanger. With these assump-
tions, the temperatures involved in the transfer of heat between the re-
actor and heat exchanger can be given as
TH
(t) = TH
(t - T3
)
where the inlet coolant temperature to the heat exchanger J u has theM
i
same form as the outlet temperature of the reactor T,, after a fixed
o
transport delay T-. This transport delay can be approximated by the
following differential equation which is derived in Appendix D.
i dt o
Similarly, the transport delay from the heat exchanger outlet to
the reactor inlet plenum is:
Y (t) r4
«c(t)=
'pdt
35
(37)
where Tp (t) = Reactor inlet plenum coolant temperature (°F)
Tr (t) = Heat exchanger outlet coolant temperature (°F)Lo
This can be approximated by
Tc .
p(t) 4-M ^ = fcoCt) (37)
As in the previous development, let
TH
(t) = TH
+ <AH
(t)
o o
TH
(t) = TH>
+ AH .(t)
(38)i p i po i p
v '
TC
(t) = TC
+ ^TC
(t)
o oQ
o
Tc
(t) = Tc
+ <AC
(t)
ip ipo ip
where again the lowest zero subscript denotes a steady state value
and the delta prefix a small perturbation about this value.
After substituting equations (38) into equations (36) and (37),
taking the Laplace transform, the following transfer functions are
derived (see in Appendix D).
A^ (5) = § (39)
</Tc,pC5) _ l(40)
</Yca (s) 14- HS
36
Combined mixing and transport effects are encountered at the reactor
and heat exchanger inlet and outlet coolant plenums. For simplification,
only the mixing effects at the inlet plenums were considered in the
model. Assuming perfect mixing and after performing a heat balance on
the plenum concerned, the combined mixing and transport effects are ex-
pressed by differential equations developed in Appendix D of the form
Y-£lil T <t)-T1<t)
whereT. (t) = Plenum inlet temperature (°F)
T (t) = Plenum outlet temperature (°F)
I= Mixing delay
As before, by using a small perturbation technique, and taking the
Laplace transform, the following transfer function is derived (see
Appendix D):
^2i!L = ,.., 1 (41)c/Th^(s) l tfs
Block diagram representations of the coolant piping transport and
mixing transfer functions is shown in Figure 9.
D. OVERALL PLANT BLOCK DIAGRAM
The transfer functions developed previously for the zero power point
1 35reactor, Xe feedback, reactor and heat exchanger heat transfer, and
the primary piping were interconnected resulting in the overall model
block diagram shown in Figure 10. The moderator temperature feedback
37
REACTOR<fT (s) / ^W s >
/•&TS
Jf
/ /
/^s /+%s1 1
i
HEATEXCHANGER
/
^o(s> / +fs Vtc
.csi
> r~
Figure 9, Block Diagram of the Transportand Mixing Delay Functions
loop is seen to consist of the heat transfer process and primary piping
transfer functions which yield v~[-. and <7T,, . These temperaturesi o
are summed, then halved, yielding T«v
, which is then multiplied by
the negative temperature coefficient °*T, generating the moderator
temperature feedback reactivity /^j.
E. MODEL CONSTANTS
The following plant parameters from the Shippingport Atomic Power
Station were obtained from Reference 6 and used in the model development,
38
/***
/*/&
hrn
<#htt)
t/Tc^s)
<*T
1
J
777T
—I
—
cT^ts]
'-
/-/ct
—*—
QH±K> c/*rHiis)
J+Ts
l+fts
Jtu#
GhOJ
<^>e<s)
O1
<&«s>V> + o^tts)
c^b(3)
A?fc)
cft^l
GhU)
</P(5)
<#<s)
G^S)
c^zka
Figure 10. Overall Plant Block Diagram
39
1
.
General Parameters
a. Reactor thermal power
b. Reactor coolant system pressure
c. Reactor coolant average temperature
d. Steam pressure at full load
2. Reactor Coolant System
a. Reactor coolant flow rate
b. Reactor coolant outlet temperature
c. Reactor coolant inlet temperature
d. Coolant volume in core
3. Reactor Core
a. Configuration
b. Size
c. Fuel load U (seed)
d. Composition (seed)
1
)
Water
2) Fuel alloy
3) Zircalloy
e. Control Rods
1
)
Total rod worth
2) Scram time
231 MW
2000 psia
523° F
600 psia
6280 lbm/sec
538° F
508° F
103 ft3
Right cyl inder
6.8 ft. dia X 6 ft. high
75 kg
43.5 v/o
30.0 v/o
34.1 v/o
0.256 k
0.35 sec time delay
1 .0 sec rod drop
4. Nuclear Data
a. Thermal neutron flux
b. Prompt neutron lifetime
14 22 X 10 n/cm sec
5.6 X 10" 5sec
40
c. Effective delayed neutron fraction 0.0077
-4d. Temperature coefficient of reactivity -3.1 X 10 a k/°F
5. Reactor Protection Setpoints
a. Scram
1) High reactor power 138%
2) High reactor coolant outlet temperature 550°F
b. Cutback
1) High reactor power 114%
2) High Startup rate 1.74 Decades per min,
41
IV. RESULTS
A. INTERACTIVE PROGRAM
Figures 10-16 are examples of the interactive program output. As
shown in Figure 10, the user is first given a description of the purpose
of the overall program and general instructions for input entry. As
shown in Figure 11, the user is then prompted to enter an initial steady
state power level, after which the corresponding major plant parameters
are displayed. The user is then prompted to choose the type of simula-
tion, either plant transient at power, or post-shutdown Xenon-135
behavior. The user must be at the Tektronix 4012 console in the computer
center if the latter is chosen.
If the post-shutdown Xenon-135 behavior is chosen to be examined,
graphs such as those shown in Figure 13 are generated and displayed on
the Tektronix' s screen. In addition to showing the time response of the
Xenon, the time of its peak and associated maximum reactivity are dis-
played.
If the plant transient simulation is chosen, the user is then prompted
to choose the mechanism for initiating the transient, either control rod
movement or turbine load change. In Figure 14, the user has chosen the
former. The program then prompts the user to enter the direction and
time length of the control rod movement. In Figure 15, the turbine load
change was chosen and the program request the final turbine load. In
either case, a summary of the transient inputs is displayed with a
notice that the interactive portion is complete.
42
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xiia POST SHUTDOWN XENON TRANSIENT
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Figure 13. Post-shutdown Xenon-1 35 Behavior
45
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Figure 16 is an example of the interactive program's logic routines
which will only permit inputs within the requested ranges to be accepted
for the simulation.
B. SIMULATION PROGRAM
The program generates Versatec plots showing the plant's power, re-
activity, and temperature transient responses for the user's inputs.
Figures 17-20 are composites of these plots.
In Figure 17, the transient was initiated by a simulated 20 to 40
percent ramp change in turbine load at 1/2 percent per second. In
Figure 18, the transient was initiated by a simulated 60 to 40 percent
change in turbine load in the same manner. Both of these figures show
the characteristic power demand following and inherent stability response
typical of a PWR plant with a constant average temperature program and
negative temperature coefficient.
The inherent stability feature is further displayed in Figures 19
and 20. In Figure 19, the transient is initiated by a simulated 10 second
inward movement of the control rods at a reactivity insertion rate of
-4 n1.25 x 10 ok per second from an initial 50 percent power level. In
Figure 20, the transient response to an outward control rod movement
with the same parameters is shown. As seen in both cases, the reactor
power stabilizes about the initial turbine load and the average reactor
coolant temperature stabilizes at a new level to compensate, via the
negative temperature coefficient, for the reactivity inserted by the
control rods.
48
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SIMULATION OF R PRESSURIZED WATER REACTORPONER AND LOAD RESPONSE leseno
POWER
LOAD O
REACTIVITY RESPONSE
/\
TEMPERATURE RESPONSE
LEGEND
rho a
RHOE <D
COOERNAL")
RHffT A(temperature:)
LEGEND
TRY m
TH O
TC A
00 10.00 20.00 30.00TIME (SEC)
40.00 50.00*10 l
60.00 70.00
Figure 17. Power, Reactivity, and Temperature Response
to a 20 to 40% Turbine Load Change
50
*
(To
u.
SIMULATION OF A PRESSURIZED WATER REACTORPOWER AND LOAD RESPONSE iegcnq
pokeh
Lam o
-g a a a a » a a a a a e g
REACTIVITY RESPONSE LEGEND
RH8
RttE O(EXTERNAL)
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^fw^^— 6B ~^> '' ^- ©
TEMPERATURE RESPONSE LEGEND
TOT
TH O
TC A
"b7oo 10. 00 20.00 30.00 40.00 50.00 SO. 00 70.00TIME (SEC) xlO'
Figure 18. Power, Reactivity, and Temperature Response
to a 60 to 40% Turbine Load Change
51
o
c_3 <\j
SIMULATION OF fl PRESSURIZED WATER REACTORPOWER AND LOAD RESPONSE legeno
POWER E
LOflO O
REACTIVITY RESPONSE.LEGEND
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RHOE C(EXTHRKJAL)
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j Q-— o —a a g o agseesgoosse
TEMPERATURE RESPONSE LEGEND
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TH O
TC A
o a e a a 9 e
* a g a n a a a-
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240.00(SEC)
300.00 360.00 420.00
Figure 19, Power, Reactivity, and Temperature Response
to a 10 second Inward Control Rod Movement
I
52
SIMULATION OF P PRESSURIZED WATER REACTORPOWER AND LOAD RESPONSE legend
o
CO
GL
a a n a b b b
REACTIVITY RESPONSE
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TEMPERATURE RESPONSE
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POWER
LORD
LEGEND
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RHOE(EXTERNAL)
RHOT A(TEMPERATURE)
.LEGENO
TflV
TH O
TC A
Figure 20. Power, Reactivity, and Temperature Response
to a 10 second Outward Control Rod Movement
53
V. CONCLUSIONS AND RECOMMENDATIONS
The model responds to normal operating transients with the charac-
teristics of a PWR plant. The interactive program provides the user with
the facility to initiate the simulation or examine a real time display
of post-shutdown Xenon-1 35 buildup and decay.
As previously discussed, the purpose of this work was to develop a
learning device for nuclear engineering students at the Naval Postgrad-
uate School. The result was a relatively simple model of a PWR power
plant and a simulation program that suffers from a long execution time.
The model's simplicity resulted mainly from the consideration of
an effective single group of delayed neutron precursors in the reactor
kinetic equations and the use of lumped parameter analysis in the plant's
heat transfer processes. The model's sophistication could be increased
by:
1. Expanding the reactor kinetic equations to consider six
delayed neutron precursor groups.
2. Developing a multi-section heat transfer model for the
reactor and the heat exchanger.
3. Incorporating variable reactor coolant mass flow rate and
heat transfer coefficients.
These additional features will also increase the complexity of the
simulation program, thus aggravating the long run time problem. This
concern might be eased by taking a different approach to the model's
formulation than the transfer function method* While this method was
readily coded using the CSMP-III language, physical insight to the plant
dynamic processes was lost when the Laplace transform and subsequent
grouping of constant terms was performed. The use of state variable
54
theory might not only allow this physical appreciation to be retained
but also result in shorter program run times.
Regardless of the formulation method, a real time response to the
simulation can probably only be achieved by using an analog computer.
The wide range of time constants associated with the equations (prompt
_5neutron life-times on the order of 10 seconds to Xenon decay half
lives on the order of hours) require a small numerical integration time
interval over a large time period. The result is a prohibitively long
run time on a digital computer for an interactive program.
It is hoped that refinement of the model will be continued. While
the existing model does reflect the general characteristics of a PWR
plant to normal operating transients, it does exhibit relatively large
power over/under shoots in response to turbine load changes as seen in
Figures 17 and 18.
55
APPENDIX A
DEVELOPMENT OF THE XENON FEEDBACK TRANSFER FUNCTION COMPONENTS
1 . Derivation of G (S)A
1 o f" IOCThe previously developed rate equations for I and Xe from the
135effective Xe decay scheme shown in Figure 5 are:
-— = *xZ£<z$<t) - ax ret) (A 1)
j^ = #x Z{£M + AxlCt) -Ax xa)-CaX^^)X(t) (A 2)
135 3where I(t) = I number density (atoms/cm )
135 3X(t) = Xe number density (atoms/cm )
_ 235 -121 f = Macroscopic fission cross section of U (cm )
2= Average integrated one group flux (neutrons/cm sec)
*I = Effective135
I fission yield
*X = Effective135
Xe fission yield
1 = I decay constant (sec" )
A X = Xe decay constant (sec" )
135 2g-X = Xe thermal neutron absorption cross section (cm )
Let I(t) = I + </l(t)
X(t) = XQ
+ (fX(t) (A 3)
0(t) =Q
+ ofy(t)
where the zero subscript denotes an initial steady state value and
the delta prefix denotes a small perturbation about this value.
56
With the reactor at an initial steady state with a flux level , the
135 135equilibrium values of I and Xe are found by setting their time
dependence in equations (A 1) and (A 2) equal to zero. Thus,
X = -i—2. (A 4)Ai
X. = hl£*l±2*ll (A 5)
*x + cr£ do
Substituting equation (A 4) into equation (A 5)
x.=^jMZpj.
(A5)
Substituting equations (A 3), (A 4), and (A 6) into equation (A 1)
and (A 2), neglecting the o0 J*X term (first order approximation)
^ cPr = a'rSfo^-Aio^: (A 7)
sL cPx = hxJl f (XxZp-q^Xoi^-CAx+QaV^^X (A 8)d-fc
Taking the Laplace transform of equations (A 7) and (A 8) and solving
for cA(S) and cfx(S)
^<s) - ski (A 9)
^X(s) = ^^ tUf-rfXo)^(aio)
57
Substituting equation (A 9) into equation (A10)
</X(s)
Expanding, collecting terms, and solving for ^g7s>
—— a . — ^xtS)
This is equation (16) on page 23.
2. Derivation of °^x
The effective core neutron multiplication factor k is given by
the familiar "six factor formula" found in the literature
k - fl*f>f PpNL PTNL
(A11)
where '{ = Thermal fission factor of the fule
€ = Fast fission factor of the fuel
p = Resonance escape probability
f = Thermal utilization factor in the core
PFNL
= Fast neutron non-leakage probability
PjNL
= Thermal neutron non-leakage probability
The core reactivity is defined as
^ 1 (A 12)
Consider a reactor at an initial critical steady state condition
1 35with an equilibrium Xe concentration Xo. By the definition of
critical ity
58
ko = 1 (A 13)
/^o =
In this condition
£-yCOKi - p v
( A 14)
where r-r F r , . . , . ..2»a
= Fuel macroscopic thermal neutron absorption crosssection
21 = Initial core macroscopic thermal neutron absorptiona° cross section
2, = Initial Xe macroscopic thermal neutron absorptioncross section
and the macroscopic thermal neutron absorption cross sections of the
moderator and other core materials has been neglected, a reasonable ap-
proximation in a highly enriched core.
1 35Now impose a small perturbation in the Xe concentration on the
1 35core. Neglecting the effect of the Xe perturbation on neutron leakage,
a reasonable assumption for a large reactor,
X = Xo + </x
k = ko + <fk
f - fb + cff
/>=/>o + <fy = d)C since /> =
The remainder of the terms in equation (A 11) are unchanged, thus
cA = <ff
and
k = k = * since ko - 1
59
where
r - *-*3 = =ti (a 15)
from equation (A 12) and (A 13)
(A 16)^= ^x= ^i= l--|
substituting equations (A 14) and (A 15) into equation (A 16)
&* = 1 = *x
This is equation (17) on page 23. Since fuel depletion effects are
not considered during the short simulated transients, the fuel number
density U is assumed constant and °^cis invariant in the time and the
Laplace domains.
60
APPENDIX B
DEVELOPMENT OF THE REACTOR HEAT TRANSFER FUNCTIONS
The previously developed equations describing heat transfer from the
reactor to the coolant are
TF(fc) * ti ^f£* = TAv(t) + 3k p(t ) (B 1)
TAv (t)[lM |]tf2^ = TF (t) + 2%TCl (t) (B 2)
TAvCtl = ICTHoC^ t Tc^Cfc)] (B 3)
where Tp(t) = Average core material temperature (°F)
T»v(t) = Average reactor coolant temperature (°F)
T„ (t) = Reactor coolant outlet temperature (°F)
T . (t) = Reactor coolant inlet temperature (°F)
P(t) = Total power generated in the core (Btu/sec)
Cr = Total thermal capacity of core materials (Btu/°F)
C., = Total thermal capacity of coolant in the core (Btu/°F)
C = Specific heat of coolant (Btu/blm°F)
m = Coolant mass flow rate (lbm/sec)m
and
hFM
= Total heat transfer coefficient (Btu/sec°F)
Jo =
r\pH
f2 = £m_Hfm
61
Let P(t) = ?n + <A(t)o
TF(t) = T
f(J+ c/*T
F(t)
TAV
(f) = TAV
+ <AAV
(t) (B 4)
THo
(t)=
THo
+ ^Ho (t)
o
Tc.(t) = Tcio
+ JVc^t)
where the lowest zero subscript denotes a steady state value and
the delta prefix a small perturbation about this value.
From substitution of equations (B 4) into equations (B 1), (B 2),
and (B 3), with the reactor in a steady state at t ^ 0,
TAV*(i + 2 \) = TF^2^Tc id(B 5)
TAV = -|- ("ft*.* Tcio )
and
(B 6)
(/l>(o) » <^7av(o) b J%c io) a c/'Tc. (O) a C
Impose the perturbations at t=0. Substituting equations (B 4) and
(B 5) into equations (B 1), (B 2), and (B 3) with the initial conditions
given in equations (B 6),
</TP (.h) + Yi & <ftpfc) =* c/T/wCt) *•~ cPPdt) (B 7)
^TavCO * I jVrHo (fc) + o^ (t)] (B 9)
62
Taking the Laplace transform of equations (B 7), (B 8), and (B 9)
with the initial conditions given by equations (B 6)
Tf(5) [sfi+i] = Tav(s) *- ^rPCs) (B 10)
Tav(s) [sfz + i+z ±|] = TF( s) t a~Tc(s) (B 11)
TAV(S) = \ [ TN (S) + Tc (5)] (B 12)
where the delta prefix and lowest subscript have been deleted for
readability
Substituting equation (B 12) into equations (B 10) and (B 11)
TF (sfiti) = \ ( TH t Tc) + % P (B 13)
|(TH +TC ) (sTz + | + 2 £) = Tp + a^Tc (B 14)
where the s domain dependence notation has been deleted for
readability.
Solving equation (B 13) for TV and substituting into equation (B 14)
j<Jh + Tc) {sti+t+Z ^)(sfifl) = I (Th+Tc) + A p + 2 J| (sfi+ l)Tc
After expanding the products and collecting terms
After multiplying through by and solving for T„( <Pln (s))
o
63
where 5 = —-y and all notation has been restored^"2
This is equation (26) on page 28.
64
APPENDIX C
DEVELOPMENT OF THE HEAT EXCHANGER TRANSFER FUNCTION
(
The previously derived heat exchanger heat transfer equations are
[t + Kl]Tc (t) = r .
— 1(C 1]
£l ill^+TsU) = Y[THi (t)-hTCoa)] + h^pL(t) (c 2)
where T„.(t) = Heat exchanger coolant inlet (°F)
Tp (t) = Heat exchanger coolant outlet (°F)
T-(t) = Saturated steam pressure (PSIQ)
P. (t) = Power delivered by heat exchanger (Btu/sec)
C<» = Total thermal capacity of heat exchangermetal and secondary water and steam (Btu/°F)
C = Specific heat of coolant (Btu/lbm°F)
\\ = Total heat transfer coefficient (Btu/°F.sec)
m = Coolant mass flow rate (lbm/sec)
and 2mM C*i s "u
—
Ktm
Let Tm (t) = THi
+ </TH1
(t)
o
TCo
(t) = TC
+ Ato^ (C 3)
Ts(t) = T
s+ <A
$(t)
PL(t) = p
l+ dP
L(t)
o
where the lowest zero subscript denotes a steady state value and the
delta prefix a small perturbation about this value.
65
With the plant in an initial steady state at t < 0, equations (C 1)
and (C 2) are
aTSo -TH io [i+ Ki ]Tco„ =
[n-^i](C 4)
Ts = f ( TH. tTc0o )+ ^-PL (C 5)
and
JTu.^o) = cPTq (o) = <^Ts (o) = o
Impose the perturbations substituting equations (C 3), (C 4), and
(C 5) into equations (C 1) and (C 2)
(C 6)
(C 7)
Cs d
[£ 57 <*i<0 + Ate) = 1 [jYHiW + Ac.tel] - £VPL te) (c 8)
Taking the Laplace transforms of equations (C 7) and (C 8) with the
initial conditions given by equation (C 6)
jTs(s) = *
|[Ac (5) + C^Th.(S)]-4i^PL^
I5S + 1
where<-f
- Cs
Ktm
<^TCoC3) = (2c^rs (s)-(y)
TH.(s) [l-Kl]}/[l+*il
These are equations (34) and (35) on page 33.
66
APPENDIX D
DEVELOPMENT OF THE PRIMARY PIPING TRANSFER FUNCTIONS
1 . Transport Delay Transfer Functions
The transport delay of the coolant between the outlet of the reactor
and the heat exchanger inlet plenum can be expressed as
TH
. (t) -THo
(t-r3
) (Dl)p
where T„. (t) = Heat exchanger inlet plenum coolantHi
p temperature (°F)
TH
(t) = Reactor coolant outlet temperature (°F)
T*3 = Transport time delay (sec)
Rearranging terms and expanding equation (D 1) in a Taylor series
r Ji- 2.' dtdt 3t (° 2 )
For slow temperature changes, the second order and higher terms
in equation (D 2) may be ignored.
Tv*)*f3 ^S.T,..Wdt HoW(° 3)
Let
THi
(t) = TH
. + <Pl (t) (D 4)
po 1p
THo
(t) - th
+ eAHo (t)
°o
where the lowest zero subscript denotes a steady state value and
the delta prefix a small perturbation about this value.
67
From substitution of equations (D 4) into equation (D 3) and for
the temperatures in a steady state at t^O
THi
= THo (D 5)1 po o
v ;
and
(D 6)
Impose the perturbations at t=0. Substituting equations (D 4) and
(D 5) into equation (D 3) with the initial conditions given in equation
(D 6)
Taking the Laplace transform of equation (D 7) and solving for
^THi / ^ T
Ho
C^THoCS) 1 + fzS
This is equation (39) on page 36,
By following a similar development, the transport delay between a
perturbation in the heat exchanger coolant outlet temperature and the
resulting perturbation in the reactor inlet plenum coolant temperature
can be expressed by the transfer function
tTTcoCs) 1-1- MS
where o *c^p = Reactor inlet plenum coolant temperatureperturbation (°F)
c^Tco = Heat exchanger coolant outlet temperatureperturbation (°F)
\. = Transport time delay (sec)
68
This is equation (40) on page
2 Mixing Delay Transfer Function
Let the reactor and heat exchanger coolant inlet plenums be represent-
ed by the control volume shown in Figure D-l
.
Tj,m *
m,T4
MT
M
T
o
m
= Mass of coolant in plenum (lbm/sec)
= Average plenum coolant temperature(°F)
= Plenum inlet coolant temperature(°F)
= Plenum outlet coolant temperature(°F)
= Coolant mass flow rate (lbm/sec)
Figure D-l. Mixing Volume
If perfect mixing is assumed, the coolant outlet temperature is equal
to the average coolant temperature in the plenum. At heat balance on
the volume requires that the net heat flowing into the plenum be equal
to the increase in stored energy. Assuming no ambient heat losses, and
assuming the temperature of the plenum structural material remains
constant, the heat balance per unit time over a time interval At is
mAtC [T\(t)-ToW] = MC ^gr^t = matedt
dTo(fc)
dt
where C = specific heat of coolant
The resulting equation is
m at dt(D 8)
69
where T =
Let
M— = mixing time delay.
m
V^ sTio
+ AwTo(t) = To
o+ A) (t)
(D 9)
where the lowest zero subscript denotes an initial steady state value
and the delta prefix a small perturbation about this value.
Substituting equations (D 8) into equation (D 9)
i ZJZ <^Tofc) + ofVoCt) * d%(±)dt
Taking the Laplace transform of equation (DlO)and solving for
(D10)
A/ AcTT (s)
This is equation (41) on page 37,
70
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91
LIST OF REFERENCES
1. Schultz, M. A., Control of Nuclear Reactors and Power Plants , 2nded., McGraw-Hill, 1961.
2. Bell, G. I. and Glasstone S., Nuclear Reactor Theory , 1st ed.,
Van Nostrand, 1963.
3. Radkowsky, A. (ed.), Naval Reactors Physics Handbook , vol. I, USAEC,1964.
4. Henry, A. F. , Nuclear Reactor Analysis , 1st ed., MIT, 1975.
5. Duderstadt, J. J. and Hamilton, L. J., Nuclear Reactor Analysis ,
1st ed., Wiley, 1976.
6. The Shippingport Pressurized Water Reactor , personnel of the Naval
Reactors Branch, AEC, and Westinghouse Electric, Addison-Wesl ey,1958.
92
INITIAL DISTRIBUTION LIST
No. Copies
1. Defense Technical Information Center 2
Cameron StationAlexandria, Virginia 22314
2. Library, Code 0142 2
Naval Postgraduate SchoolMonterey, California 93940
3. Department Chairman, Code 69 1
Department of Mechanical EngineeringNaval Postgraduate SchoolMonterey, California 93940
4. Professor Paul J. Marto, Code 69Mx 1
Department of Mechanical EngineeringNaval Postgraduate School
Monterey, California 93940
5. Professor Gil 1 es Cantin, Code 69Ci 1
Department of Mechanical EngineeringNaval Postgraduate SchoolMonterey, California 93940
6. CDR C. J. Garritson, USN, Code 34 1
Naval Engineering Curricular OfficerNaval Postgraduate SchoolMonterey, California 93940
7. LCDR G. G. Heath, USN 1
Department Head Class 67
SW0SC0LC0M, Bldg. 446Newport, Rhode Island 02840
93
Thesis
c.lHeath
109507
An interactive codefor a pressurized waterreactor incorporatingtemperature and Xenonfeedback.
Thesis 189507H^4l7^5 Heathc.l An interactive code
for a pressur : ze J water
reactor incorporating
temperature and Xenon
feedback.