Atomic Energy of Canada Limited

PROSPECTS FOR 44% CYCLE EFFICIENCY

FROM 400°C (750°Fj STEAM

IN LARGE ELECTRIC GENERATING STATIONS

DM-96

LM-3A CYCLE

ORGANIC LIQUID PRIMARY HEAT EXCHANGE

DM-96 Supplement

by

W. BENNETT LEWIS

Chalk River, Ontario

December 1968

Reprinted with Supplement

K

December 1969 AECL-3221

DM-96

PROSPECTS FOR 44% CYCLE EFFICIENCY FROM 4Q0°C (750cF)STEAM IN LARGE ELECTRIC GENERATING STATIONS

by

W. Bennett Lewis

S U M M A R Y

Current trends towards large capacity nuclear powerstations point towards high steam pressure and thereby highsteam cycle efficiency. By putting together the best ofcurrent practice without any technical break-through, a grosssteam-cycle efficiency of 44% is estimated to be practicablefrom 172 bar (2500 psi) steam at 400°C <750°F). Steam ateven higher pressure 240 bar (3500 psi) has in the past fewyears come to predominate in large fossil fuelled generatingplants at 540°C (1000°F). At lower temperatures, however,due to the special characteristics of supercritical steam,such higher pressures confer no advantage in conventionalturbines. The optimum pressure has not been establishedbut probably lies in the range 170 to 200 bar for steamlimited to 400°C as the maximum temperature. It is believedthat such steam can be generated from organic-cooled heavy-water-moderated reactors. The suggested cycle is developedfrom that used in the Douglas Point nuclear generating station.Suggestions are made for turbine characteristics and feed-water heating systems that could lead to still higher efficiency

Chalk River, OntarioDecember, 1968

AECL-322

Possibilités d'un rendement de cycle de 44$ pourïïïie vapeur à 400°C (750°F) dans les grandescentrales électronucléaires

par W. Bennett Lewis

Résumé -Les tendances actuelles de l'implantationdes grandes centrales ëlectronucléaires montrentque l'on s'achemine vers de hautes pressions devapeur et par conséquent vers de hauts rende-ments de cycle pour la vapeur. En réunissant lesmeilleures données de la pratique actuelle etsans innovation technique spectaculaire, unrendement brut de cycle de vapeur de 44% semblepouvoir être obtenu à partir d'une vapeur de 172bars (2 500 psi) à 400OC (750°F). De la vapeura une pression encore plus élevée de 240 bars(3 500 psi) a pu être obtenue au cours desrécentes années principalement dans les grandescentrales alimentées par un combustible fossilea 540°C (1 000°F). Aux températures inférieures,cependant, par suite des caractéristiques spé-ciales de la vapeur surcritique, les pressionsplus élevées ne confèrent aucun avantage guxturbines classiques. La pression optimale n'apas été établie mais elle se situe probablementdans l'intervalle de 170 à 200 bars pour lavapeur limitée à 400°C comme température maximale.On croit que cette vapeur peut être produite pardes réacteurs dont le modérateur est de l'eaulourde et le caloporteur un fluide organique. Lecycle suggéré est développé à partir de celuiutilisé dans la centrale nucléaire.de DouglasPoint. Des suggestions sont faites pour lescaractéristiques de la turbine et les systèmesde chauffage de l'ëau d'alimentation lesquellespourraient conduire à un rendement encore plusélevé.

Chalk River. OntarioDécembre 1968

AECL-3221

CONTENTS

PAGE

1. Introduction 1

2. Special Nuclear Reactor Considerations 2

3. Review of Reactor Steam Cycles 3

4. Optimum Pressure and Cost Differential Analysis 3

5. Evaluation of Gross Steam Cycle Efficiency 6

6. Douglas Point Steam Cycle 9

7. LM-3A Steam Cycle 9

8. Comparison of Douglas Point and LM-3A EfficiencyCalculations 10

9. Discussion 10

10. Acknowledgements 11

Table I Reactor Steam Cycle Characteristics 12

Figure 1. Review of Gross Steam Cycle Efficiencies 13

2. Data for the Douglas Point Steam Cycle

at 100% Load 14

3. LM-3A Steam Cycle (revised 13/1/69) 15

4. Mollier Chart LM-3 Cycle • 16

APPENDIX: Sheet 1 Douglas Point Steam Cycle Data 17

2. Douglas Point Staged EfficiencyAnalysis 18

3. LM-3A Steam Cycle Heat Balance 19

4. LM-3A Additional Data 20

5. Comparison of Efficiency Calculations 21

6. More Detailed Comparison and Analysis 22

CONTENTS (Continued)

PAGE

Supplement: Summary 23,24

Table IA LM-3A Primary Heat Exchange 25

IIA LM-3A Organic System for 340 to 410°C 26

Figure 1A LM-3A Primary Heat Exchange 27

2A LM-3A Steam Cycle " 28

3A LM-3A Division of Primary Heat Exchangersfor 340°C to 410°C Organic 29

PROSPECTS FOR 44% CYCLE EFFICIENCY FROM 400°C (750°F) STEAMIN LARGE ELECTRIC GENERATING STATIONS

by

W. Bennett Lewis

1. Introduction

The advent of the large nuclear power station with water orboiling water coolant has led to the practical realization of steamcycle efficiencies up to 80% of the thermodynamic maximum or Carnotcycle efficiency for the given temperature range. At higher steamconditions lower standards of 70% or less of Carnot have been acceptedin units generating about 100 MWe. With the prospect of unitsgenerating 1000 or 1500 MWe it seems worthwhile to assess the effi-ciency practically attainable without demanding excessive temperaturesor pressures in the steam generators, superheaters and reheaters.

Canadian climatic conditions have led to condensing at 1"Hg absolute pressure 26.1°C (79°F) for which the Carnot cycle effi-ciency from 399°C (750°F) is 55.5% so 80% of Carnot, which should bepracticable for the gross steam cycle efficiency would be 4 4.4%.Expressed in the customary units of the power industry the grosscycle "heat rate" would be 76 86 Btu/kWh, a level hitherto associatedonly with 1000°F reheat cycles.

The principles are now well established by which a steamcycle may approach the Carnot cycle efficiency and it is these princi-ples that have been applied in the successful attainment of 80%of Carnot in modern water-cooled nuclear power units. The greaterpart of the losses is still in the turbine stage efficiency despitethe offsetting reheat factor. It has proved necessary and economicto adopt a large number, six to ten, of regenerative feedwater heatersto deliver the boiler feedwater at a high temperature. Ideally allthe supplied primary heat should enter the working substance at thetop temperature. This condition is naturally approached in saturatedsteam cycles, especially at temperatures where the heat of evaporationis large and so well below the critical temperature 374.1°C (705.4°F).To achieve the same result in the superheat temperature range it isnecessary to apply reheat to the turbine steam as the work performedprogressively reduces the pressure and cools the steam. Continuouslyreheating turbines have not yet been developed and in their absencethe practice is to adopt two reheaters in succession where the highestefficiency is desired. In the large units contemplated the accompany-ing increase in capital cost for the same heat capacity should not be

- 2 -

large enough to make any net penalty in unit power cost, even when thefuel cost is negligible.

A second basic principle of steam cycle design for high effi-ciency is to keep to a minimum the temperature differences in the manyheat exchange processes involved.

Since the removal of all the unavailable heat at the lowesttemperature of the cycle is essential for high efficiency, as muchattention must be paid to the steam exhaust and moisture separation inthe low pressure stages as is usual now for saturated steam cycles.

2) Special Nuclear Reactor Considerations

The high levels of radiation experienced around nuclearreactors, even when shut down, slow up and increase the cost ofmaintenance operations. The larger the generating unit the morecostly is a period for maintenance. Even when the cost of powermay have been reduced to 2 mill/kWh this is $2000/MkWh so a 100 hourmaintenance operation on a 1000 MW unit would face a base penalty of$200,000.

Reviewing the different possible coolants, maintenanceoperations, though difficult, may be carried out while a coolantgas is still at high pressure, or a sodium coolant is still highlyradioactive. Mass transport of radioactive corrosion products canalso produce high radiation fields, even with water and steam cool-ants. Experience has shown that for rapid access the organic-cooledsystems are very good. The fact that they operate only at temperaturesbelow about 430 C (800°F) is also an advantage in reducing plant cool-down time. Moreover with steam as the working substance the organiccoolant has two significant advantages,first that its operating tempera-ture is high enough to raise steam at any desired high pressure andsecond that its own operating pressure can be reasonably low so thatthe costly and difficult-to-maintain high steam pressure equipment isremote from the high radiation fields of the reactor core. The samecharacteristics can be claimed for sodium but as already mentionedsodium in the primary circuit is highly radioactive and even that ina secondary circuit needs protection from both water and the atmosphereand to be kept warm to circulate. Apart, however, from any comparisons,the organic coolants are very premising for convenient association withboth nuclear reactors and high pressure steam.

Exposure to radiation and in particular to fast neutronsincreases the number of ways in which structural materials fail andchanges the relative importance of the mechanical properties. Failureis liable to occur from fracture by stress rupture, slow strain ratefatigue (including tertiary creep) embrittlement and loss of ductility,anisotropic growth, over-ageing (growth in size of precipitates),incomplete recovery from radiation damage as well as the many mechanismsof corrosion, including stress-corrosion-cracking and hydrogen embrittle-ment. Although an increase in temperature promotes recovery fromîn.-.tial or primary radiation damage, it is generally true that inpractical operating ranges increasing temperature accelerates mostmechanisms of ultimate failure, including residual radiation damaae.

- 3 -

Generally speaking there seems more prospect of engineering forhigher temperatures within reactor cores with ceramic materials, includinggraphite and silicon carbide than with metals. Residual radiation damagemay, nevertheless, limit the use of graphite for extended life of severalyears to below 800°C.

The shortening of lifetime of structural materials in reactorcores presents a technical challenge to achieve economic designs. Ingeneral permissible temperatures for structural solids, under tensilestress, in reactor cores are lower than for their conventional use.These considerations point strongly towards maximising efficiency inthe lower range of steam temperatures below 500°C (932°F) .

3) Review of Reactor Steam Cycles

Fig. 1 shows the Carnot steam cycle efficiency as a functionof the top steam temperature for exhaust at 1" Hg 26.1°C (79°F).Similarly plotted is the line for 80% of Carnot efficiency. Thefigure shows the gross steam cycle efficiencies for actual or proposednuclear plants, details of which are given in Table I. For those inreferences 1 and 2 of the table the exhaust conditions were 1.5" Hg0.0508 bar 33.2°C (91.7°F). The effect of this is to lower the Carnotcycle efficiency by AT/Ti = 7.1/673,16 = 0.0105 at 400°C or 7.1/523.16 =0.0136 at 250°C, just sufficient to explain the lower efficiency shewnfor the PWP, but not a major factor in the low efficiencies for theHWOCR, SGR and HTGR. These lower efficiencies are attributable tothe acceptance of less efficient cycles than should be achievable.The efficiencies, however, are typical of current practice in the super-heat range and justified by economics under the conditions establishedin the past.

Detailed discussion follows on the Douglas Point and theproposed LM-3 cycles included on Fig. 1.

4) Optimum Pressure and Cost Differential Analysis

It remains to examine the forecast in the introduction thateven when the fuel cost is negligible, the increased capital cost forthe higher pressure and the same heat capacity will lead to no netpenalty in unit power cost. At the inlet pressure of 172 bar adoptedin LM-3 the expansion line in the high pressure turbine terminating at320°C is somewhat short, Fig. 4, but because the flow is relatively highit nevertheless contributes about 10% of the total power. In a smallsize plant the complication of this high pressure stage would not beeconomically justifiable and in fact the use of double reheat in cyclesfor organic-cooled heavy-water reactors has been ruled out by previousstudies. It is only when the size is large and the high pressure stagecontributes 60 to 150 MWe that the extra cost is expected to be repaid.Little if any further gain seems likely to be attainable from pressureshigher than about 200 bar, unless the continuously reheated turbine isdeveloped. The reason is apparent from the Mollier chart. Theenthalpy of the steam drops and, recognising that entropy must increase

- 4 -

in any practical turbine, the expansion line whether limited by tempera-ture or saturation becomes very short. While this is compensated bythe higher temperature at which most of the heat enters the workingsubstance and by the greater enthalpy rise of the first reheat, any netchange in cycle efficiency is small.

The recent history of the trend to higher pressure in fossilfired plants has some relevance. A study made on the basis of offeredprices was presented by H.J. Petersen to the American Power Conferencein 1963 (Proceedings p.444). He compared costs and performance ofplants of 300, 400, 500 and 600 MWe to operate on reheat cycles charac-terized as 2400 psig 1000/1000°F, 3500 psig 1000/1000°F, and the doublereheat cycle 3500 psig 1000/1000/1000°F. For low cost fuel at 20 cents/million Btu and load factors over 50% the last showed a very smalleconomic advantage only for the high outputs 500 and 600 MWe. Heconcluded "Economics of 2400 paig versus 3500 psig for large capacityunits i.s still not clear-cut". Subsequently, however, the trend to3500 psi for fossil fired plants has accelerated and it was reportedto the 1.967 American Power Conference by Davis, Griffin and Wybenga ofBabcock & Wilcox (Proceedings p.333) that over the previous three years52% of such boilers purchased had been for 3500 psi. It must be notedthat at 1000°F (540°C) the Mollier chart does not show any penaltyagainst the higher pressure comparable with that at 400°C (750°F) .The inference may, however, be drawn that the cost penalty against thehigher pressure components has dropped with experience.

It is of interest to present the cost differential analysisin a manner relevant to nuclear power plants. The total capital costof the plant is divided between two components X and Y for which Y isproportional to the thermal power (Q ) and X is independent of thethermal power. This division is always valid over a limited rangeof the thermal power. The cost of power in mill/kWh

871T6ÏÏ

where X + Y = capital cost of plant in $

QE = electrical power kWe (net)

e = station net efficiency

QT = thermal power .W(th) (= 0.,/e)

note X/QE + Y/eQT = specific capital cost ($/kWe)

A = annual charge rate on capital in percent

8766 = hr/yr; 24 = hr/day

- 5 -

u = utilization (fraction)

c^ = cost of operating (other than fuel andannual interest-dependent charges)

P = net cost of fuel in $/kgU

B = fuel burn-up in MWd/kgU

Tne-,

_ A_ 5(X + Y) 5e i . Ys A . P \ , . - ,876.6~ïï QE e2 V Q T

; 8 7 6 T F i ï + 2 T B / • " ( 4 - ~ }

T h e r e i s a c o s t s a v i n g i f 6c i s n e g a t i v e , i . e . i f

6e_ ) _Y_ P . 8 7 6 . 6 U ) 6 (X + Y) . . . ( 4 . 3 )e 2 { Q T

+ 24BA 1 QE

For convenience of relating to familiar estimates of the componentsthis may be rewritten

Y _,_ P . 876.6u > ,/X + Y, ... (4.4)\ 2 4Be A

For example:

if u = 0.8 and A = 7%, 876.6u/A = 100

P/24Be is the net fuel supply cost which may be 0.5 mill/kWh

Y/Q- may be $4 0/kWe

there would then be a cost saving if

~{40 + 50} > 6(* + Y) ... (4.5)e QE

Y + VAs the size of the plant increases 5(—-—-) becomes smaller, and, evenif the fuelling cost is very small, E the contribution fromY/Q to the left hand side of the inequality (4.4) can make it economicalto increase the efficiency.

Suppose, for example, the net station efficiency is increasedfrom 38 to 40% so that 6e/e = 0.02/0.4 = 0.05, the inequality (4.5)would show a cost saving if . /X + Ys is less than $4.5/kWe.

QE

- 6 -

For the fossil fired plants studied by Petersen the capitalcost difference between a 2400 psig 1000/100Q°F and a 3500 psig 1000/1000/1000°F plant was between $4 and $5 per kWe. The difference mthe optimum net plant heat rate was close to 4%. For a fuel cost of25$/million Btu and a net plant heat rate of 8800 Btu/kWh* the fuellingcost is 250 * 8800 * 10"6 =2.2 mill/kWh. The fixed charga rate heused is 12.4% so for 80% load factor 876.6 u/A = 56.5. There is acost saving if

0.04{(Y/Q_.) + 56.5} > 4 to 5 .

i.e. if Y/Q_, > $43.5 to $68.5 per kWe

and in the result the saving appeared very small.

It should be noted that both in the above analysis and inPetersen1s the operating cost (c ) has been kept constant in mill/kWh.If the increase in efficiency is applied to increase the outputof the plant, co is likely to fall giving an added advantage to thehigher efficiency.

5. Evaluation of Gross Steam Cycle Efficiency

Because of the complexity of any practical cycle a great dealof arithmetic is involved in any procedure for calculating the cycleefficiency. There can be errors in identifying the exact cycle becauseit has to satisfy many conditions of practicality such as heat transferrates, turbine stage efficiencies and pressure drops, and there can bearithmetical errors. Arriving at a correct and exact formulation ofa cycle is tedious, especially the correction of errors that may involvethe reworking of many subsidiary cycles that must be satisfactorilyclosed. Moreover in the exploratory design of cycles the effect of areal change is often desired in explicit form. Current practice hasestablished procedures for such analyses, but it seems worthwhile topresent an exposition which incorporates this practice but is set interms that make clear the interrelations and special virtues of the,"heat balance", "condenser check" and "unavailable energy (or entropy)balance" and also show how errors may be allowed to remain and theireffect on other quantities evaluated approximately. In the basic heatcycle the necessary quantities are collected into five sums, eachaccumulated from around the whole cycle which is then written

*This heat rate allows somewhat over 10% loss betweenfuel and steam.

- 7 -

EQs = Z Work + T2EcFAs - LQ^ + ZQ^ ... (5.1)

(?) (2) (3) (4) (5)

where Z represents summation around the whole cycle with exceptionsexplained below.

(1) IQ = heat supplied from primary source via boiler andreheaters

(2) E Work = sum of work output from all turbine stages

(3) T2XcFAs is a quantity of great analytical utility andsignificance as discussed below. At its simplest,however, it may be equated to

Q = heat removed from condenser (usually bythe "circulating water")

T2 = condenser temperature (absolute)

s = entropy

F = flow (of steam or water)

As = net change of entropy between any two identifiedpoints in the cycle or over any regenerativeoperation where heat is transferred from onepart of the cycle to another

Any Q orWork = EFAh where Ah is the enthalpy change

Any real flow in the main or a subsidiary cycle is positive. Theassociated entropy change may be positive or negative. In anyregenerative heat exchange IFAh = o but the net ZFAs is positive.

EcFAs may be evaluated in several ways. On the one handEcAs = EFs flowing into the condenser - ZFs flowing out from thecondenser and is positive. On the other hand, because the overallentropy cycle must close Ec

FAs is also the net change along th&main cycle (or any loop of it) from the exit to the entrance intothe condenser. This particular method of evaluation is used to showexplicitly the losses at different parts of the cycle.

2 c is conveniently considered as the "unavailableenergy" of the cycle. If heat (Q) is contributed at temperature (T)at any point of the cycle, Carnot's principle indicates that the

- 8 -

portion unavailable as work is QT2/T = T2EFAs.

(4) EQ = sum of any extra heat from an external source,x e.g. boiler feed pump, generator cooling or by

error.

(5) EQ = sum of heat wasted unnecessarily, by error, orw not otherwise counted.

By definition:

,.-. . £ WorkGross Steam Cycle Efficiency = E G - ^

where E_ = generator efficiencyG

by equation (5.1) the gross steam cycle efficiency

^ WOrK _ _ I , C X w | , c -, >EG £Q " EG I 1 EQ /T2 ' •" (5"i)

W)/T21JThis equation (5.2) relates the four methods used to evaluate thecycle efficiency. The expression on the left represents the "heatbalance" method, while that on the right represents the methods usedin

(a) the heat balance condenser check in which Q replaces T2E FAs

(b) the unavailable energy balance in which E FAs is taken fromaround the main cycle, and c

(c) the unavailable energy check in which EcFAs is the net (entropychange x flow) between the input and output of the condenser.

To ensure satisfactory consistency in evaluations it isnecessary to assign flows, entropies and enthalpies to a completeseries of flow junctions and stage ends throughout the cycle. Moreinformation may be derived if required by breaking up stages intosmaller sections across which entropy changes are evaluated. Whenthe information is not required, however, no error is introduced byomitting such extra subdivisions.

If enthalpy and entropy values are assigned so that allmain and subsidiary cycles are closed, then any of the four methodsof evaluating the cycle efficiency will give the same result.Whether the result is a valid practical estimate depends on whetherthe entropy changes are correct. Even in the evaluation by heatbalance, where no entropy values appear, they have in effect beenused in setting exit enthalpies for the turbine stages, and boilerfeed pumps and the temperature differences across heat exchangers.

- 9 -

6. Douglas Point Steam Cycle

Data for the Douglas Point steam cycle at 100% load are shownin Fig. 2 taken from the heat balance drawing No. FSCN-40000-3 Rev.3 of11 November 1960. Converted to centigrade temperatures and joules/genthalpy units the data are given in the Appendix pp 1 and 2. Flowsappear precisely consistent and also enthalpies except for what appearsto be a small error in the heat balance across the moisture separator.No entropy values were given so these have been derived from tables andthe Mollier chart. Where redundant information on temperatures emdpressures does not accord precisely with the enthalpy values, the latterhave been taken as valid for the calculation of entropies and the temper-atures are modified. Within the numerical accuracy of the data andcomputations all the cycles are consistent.

It will be noted that entropy values are quoted, sometimesto a large number of significant figures. It is not because thesehave any absolute significance but in the use of the entropy valuesin assessing heat exchangers there is an exact heat balance and oftenonly a small entropy difference. In order that these small differencesmay be significant, it is necessary that the entropy values shouldcorrespond to the exact enthalpy values used in the heat balance. Inother cases entropy values to two decimal places are sufficient, as itis only necessary that the same value should be used when tracing entropyaround the cycle. Whatever the values are at these points they cancelin any complete cycle.

7. LM-3A Steam Cycle

-The LM-3A steam cycle is shown on the Mollier chart, Fig. 4and Sheets 3 to 6 of the Appendix. The designation LM-3A may beinterpreted as Lewis and Meikle, Cycle 3, amended. The cycle isconstructed by using the same output flow as the Douglas Point cycleand beyond bleed #3 and feedwatsr heat exchanger #3 the cycles areidentical. In particular, the expansion line in the LP turbine ismerely a prolongation upwards of the expansion line for the LP turbineof the Douglas Point cycle. Above that the cycle is designed on theprinciples explained above, namely that the heat should enter theworking substance at the highest possible temperature and that tempera -ture drops over heat exchangers should be kept small. In connectionwith the latter point it will be noticed that there are quite largetemperature differences in the first stage of reheater #2 and inFWHEx #9. The reason is that the bleeds are the only high temperaturesteam available that is close to the saturation line. Taking highertemperature steam from other points in the cycle is not satisfactorybecause1 the condensation temperature is too low. These large tempera-ture drops lead to the suggestion that the cycle could be improved bytaking the bleeds for regenerative feed water heating from one or moresmall auxiliary turbines, starting from the exit point of the HP turbine.This would replace the IP-2 turbine which now does this duty for bleeds5 and 6. Although the IP-2 turbine is shown in Fig. 3 as on the same

- 10 -

shaft as the main turbines, this is unlikely to prove practical, insteadthe IP-2 turbine may very well be two turbines on separate shafts withthe duty of driving the boiler feed pumps and the organic coolant.These duties would be taken over by the suggested auxiliary turbinesproviding bleeds to all the feedwater heat exchangers above #4. OnSheet 3 values for turbine stage efficiency are entered. It will benoted that these cover quite a wide range. They have been derivedfollowing the guide lines given in General Electric Report GER-2007A"A method for predicting the performance of steam turbine generators . . .165Q0 kW and larger" by R.C. Spencer, K.C. Cotton and C.N. Cannon. Inexplanation of the variation, these are not strictly stage efficienciesbut overall efficiencies and include, where appropriate, entry andexit losses. In the case of the HP turbine the efficiency is low,73.96%, largely because the expansion line is short and entry and exitlosses are significant. In the case of IP-1 the steam is taken overa satisfactorily large pressure range from 91.1 bar to 23 bar and stageefficiencies can be high. All the turbine stage efficiencies areextremely important to the overall cycle efficiency and it will benecessary to obtain confirmation from, manufacturers that the indicatedvalues are indeed practical, it is probable that in some cases they maybe improved and in other cases not achieved. It is hoped that thevalues are at least attainable and may be raised.

8. Comparison of Douglas Point and LM-3A Efficiency Calculations

The comparison is set out on Sheet 5 from which the distribu-tion of losses may be seen in the column headed Net PAs. A furtherbreakdown of these values is given on Sheet- 6, it is necessary toexplain the first two entries. At the head of the table the value ofFAs representing the entropy entering the system from the organic cool-ant is shown. Dividing Qg by the top temperature on the absolute scalegives the item entered at the top of the Net FAs column. The nextitem is the difference between this top item and the value of FAs givenat the top of the table and represents the extra unavailable energyfrom taking in heat below the top temperature. It will be noticedthat for Douglas Point this extra is very small, whereas it is signifi-cant for LM-3A.

9. Discussion

It has been suggested above that the HP and IP-1 turbinescould be improved if a continuously reheated design were developed.On the Mollier chart the expansion line would then follow the isothermalline for the top temperature and this should be carried on until itmeets the chosen final expansion line. Whereas the Douglas Point finalexpansion line is quite efficient by modern standards, it is possiblethat improved efficiency would result from starting the expansion at alower pressure and thereby limiting the wetness of the exhaust steam.Improvement is also being sought in current turbine designs fromcontinuous moisture removal in the wet stages.

- 11 -

Also suggested above is the idea of using an auxiliary tur-bine to provide the bleed steam for regenerative feedwater heating.This scheme may have further advantages and might be continued throughthe whole feedwater chain. If this were done the main turbine wouldbe simplified. One further advantage of this would occur if nuclearreneat were incorporated. Once the steam pressure is sufficientlylow it would become practical to heat the steam directly in thereactor to temperatures above 400°C, thereby raising the cycleefficiency. On the other hand, there is a danger in such dry steamthat in the event of a fuel rupture fission products would be carriedover Into the turbine. The situation Would be very much complicatedif this turbine had many bleed points to the feedwater heat exchangers,as decontamination might prove an elaborate operation. In the boilingwater reactors already designed this problem is almost eliminated,because of the considerable decontamination achieved in the steamseparator. With nuclear superheat there is no steam separator andit might prove desirable to provide some device such as a centrifugalseparator as an insurance against contamination of the turbine.However, such devices are lively to introduce losses. It is clearthat the advantages of nuclear superheat cannot be obtained in fullmeasure from any current designs. It remains to pursue the alter-native line emphasized in this paper of optimising an indirect cyclewith only organic coolant in the reactor.

Concerning the top temperature of the organic coolant,Fig. 3 indicates 425°C. It is not really necessary that this temper-ature should be above 410°C in order to provide steam at the requiredoutput. However, organic coolant is not a good heat transfer mediumto solid surfaces and the use of the lower temperature would involvemuch enlarged heat exchangers. No study has been made of the econo-mic optimum, probably the top temperature would lie in the range 410to 425°C. To achieve higher temperatures from the reactor some othercoolant such as liquid metal or gas would be necessary, both havetheir own disadvantages. The opacity and radioactivity of liquidmetal has proved troublesome and the relatively large temperature dropsin gas make a gas coolant less suitable for tube- type heavy-water-moderated reactors.

10. Acknowledgement

I am very much indebted to contributions by A.B. Meiklewho has checked many of the numbers involved and eliminated severalerrors in the course of arriving at the LM-3A cycle. He has alsocontributed several pertinent references to literature that helpedto keep the suggestions in line with recent practice.

WBL/g

- 12 -

TABLE I

REACTOR STEAM CYCLE CHARACTERISTICS

Net Station EfficiencyThrottle Steam Temp.Throttle PressureFeedwatar Temperature

CycleReheat TemperatureThermal Power to SteamGross Generator OutputCross Cycle EfficiencyEstimated Fuel- Cycle CostEstimated Power Cost

t•cB a r' C

"CMUHHu

tm i l l.'kWhm i l l kWh

pwa'

31.12 5 7 . I t

4 4 . 6 22 2 8 . 1 7

1 Reheat2 4 1 . 4

122310JS

i : . a i1 . 7 - 1 . 84 . 3 - 4 . 4

SSCH'

3 1 . J2 5 4 . 4

4 1 . 3 72 2 3 . 9

7-

32051051

12 .741 . f..- 1 - 94 . 1 - 4 . 4

KWB-U'

.'i, . 8241.1

3 4 . 1 31 1 7 , ;

7-

3562, •>"1081

30 .401 . 0 - 1 . 23 . 1 - 4 . 1

KIVII-Tn1

26. t241 .1

Î 4 . 1 31^7.2

No rehc-at-

3 6 7 6 . 5 »1080

29. 18l . J - 1 . 44 . 1 - 4 . 6

«TOP1

4 4 . 456 5 , li2 4 2 . )S2 8 0 . 8 9

1 R e h e a t5 6 5 . 6

2 2 701029

4'. . 331 . 3 - 1 . 43 . i - 3 . ?

SGH'

4 3 . 65 3 7 . 8242 .35266 .6

1 Reheat537.1)

23361050.5

44 .971 . 4 - 1 . 53 . 9 - 4 , 0

HWÛC1C( M - C E I U C

3 4 . 1 1 3 4 . 0 1385

6 3 . 0 92 l c . , 6

Sup.Ht * Nue.Reheat385

29431068

37.01.0-1 .53 . 4 6 - 3 . 5 4

Net station efficiencyThrottle Staam Temp.Throttle PressureFeedwater Temperature

CycleReheat TemperatureThermal Power to SteamGross Generator OutputGross Cycle Et £xciency» .timatod Fuol Cycle CoatCan routed Fewer Cost

CANUU- !'HW ' BLW-l^O* SHHWv

Doublas P t . Cent I i l y Winf r1 th5QIW* Nwochih IIWH-P V e s s e l ' CANDU-PHW-500' C 2 I

Marvikan P ick«nnq(SM99/29 l

1

"cBar"(_'

"cMW

im i l l 'kwhm i l l 'kh>.

29250

40171

1 Hehi2 2 16 6 02 20

31I5

.1

b M t

. 3

.0

. l>

J l . l2 6 6 . 5

5 1 . 7228

1 Reheat23B.27772 6 6 . 3

34 .3o.a5 . 2

3 0 .2 7 8

62199

2 Sépara-

289.100

3 4 .

.

J

turs

5

50

3 1 . 1 5278

622 0 0

-1082.2

37734. T>-

28.51254

47120

Saturated-

4 6 3138

29.81-_

32.55472

41126

Superheated-

593200

31.73--

2 9 a2 5 0 . 8 3

4 0 . 3171

1 R e h e a t2 2 4

1 6 1 2 . 75 4 3 . 1 2

3 3 . 70 . 63 . 4

95» of Reactor • •rr-'W Powoi

' M . W . P a s c n t h a l • < a i . J I M I - 3 6 B I . , T i t l - s ".. •• k 1 "i . 1 . . I a n . I ' ) 6 ' .' P . I I . K u a t e i l e t a l . O R N L - 1 ' 1 2 1 . u » e n . i l o t x - 1 *. I , J a n . 1 9 6 7' " N u c l e a r P o w e r i n C i n . i d . i " A l : C ! . - r » S 0 , J u n e 1 9 6 4

a n d A £ C L - 2 0 ù : , J u n e 1 9 6 4" G . A . P o n , C A N O U - D l . w - 2 5 0 , S M - 9 9 / 3 2 P r u c . p . 2 0 1 , 1 9 6 7l H . C a r t w r i g h t , D e s i g n o f - h e S Q I W R , S M 9 9 / 6 4 , P r o u . p . 1 8 1' R . N l l s . i n , T h e M a r v i k e n P e a c t o r S t a t i o n . S M 5 1 / 1 7 , P r o f . ] . 1 I) I' L . R . H a y u o o d a n d A . ' ! . A i k i n S M I I i n . P r - J C . p . 4 6 9

Procet?dinc|S of u Symposium onHeavy Water Power R e a c t o r s h e l dIn Vienna , Sep tember 1967,p u b l i s h e d by I n t p r n a t i or.alAtomic Energy Agency, Vienna 1968

The a n a l y s i s f i g u r e s for thû AI-CK HWOCR in ORN'L-3921 do riGt a p p e a r s e l f - c o n s i s t e n t .

F l q . 6.10 shows g r o s s g e n e r a t o r o u t p u t 1088.•> MWe, which checks w i t h Tab io B .2 . (1088 MWe I . Subt r a c t inq 12 MM forp l a n t a u x i l i a r y e l e c t r i c a l power l eaves 1076 MW, but t h i s cannot be the ne t s t a t i o n o u t p u t for Tab lo 6 .1 s h o » s 38 KHr e q u i r e d as pumping power (o r the r e a c t o r c o ' - ' U n t , l e a v i n n on ly 101H MWe n e t .

Taking t h e r e a c t o r t he rma l ou tpu t from Tabu- Ci. 2 an 1091 MU i n c l u d i n g i h o c o o l a n t pumping pownr l e a v e s t he n e ta ' . a t l o n « f f i c i c n c y a t 1O3S/3O55 -• 13.98% i n s t e a d of 34.8% quoted in T a b l e s 6 .2 and 3 . J .

Reactor tho rmal power t o c o o l a n t i s g iven i n T a h l o fi.2 as 2921 MW b u t in T a b l e 6.1 as 2943 MW and from F i g . 6 . 1 0I t i s 10.06184 • 1 0 ' B t u / h r - 2948.7 MW ! ?>93 kKh/10* B t u .

Since the definitions are not clear

1 take Reactor Power JO1)! MWPowor t o Steam 294 3 MWGro53 Generator Output 10B8.5 MW<sNet S'.jtion " tput 1 0 J Û MWuNot Stati t . r , E f f i c i ency 34.0%Gros»; St*'a- Cycle E f f i c i e n c y 37.01

GROSS STEAM CYCLE EFFY. %

Ci

TOmy—t

m

0

Ci500COCO

COHtn^

nnm

mT )

•nnmzn»~4

-tO•0

</>

HmZ

zmX

-4ca>m

00

OO

POOO

OO

OO

mOO

en00

00

rn

- £T -

- 14 -

<o

oo

U4

u

H5/3

O

5/5

OQ

ceop-

<H

CM

I—I

308.11 "C

BOILER

3 I 3 C ^ ^

307°C

320°C#9

30 7 "C

172.5BAR

307°C J

399°C

HP

340°C

320°C

322°C

399.8°C

REHEATER# 1

96BAR

399°C

307.1°CREHEATER # 2

22BAR

L P O.O24BAR

CONDENSER

0.034 BAR26 I2°C

^42OC

* 8

2ll7eC

* 7 .147 9°C

G2# 4

IO5°C

OEAERATOR

8 5 6°C 65S°C 33.2°C

2I7OC I8O°C

FIG. 3 LM-3A STEAM CYCLE fRevised 13/1/69)

GENERATORCOOLING ETC

- 16 -

«i «a <• <•

e

ENTRWt IN WATT-SEC / QRAM-'C

SI 34 it 1» »O 12 «• M H 70 rt '« '6 '» «3 t2 • • • • • ' 90

4- T- n î r t -T f f |8' î- is

ENTHALPY- ENTROPY DIAGRAMFOR WATER m METRIC UNITS

ocuw orrici

«NIK3M1.L miuimm antaiKH ( I T M L I V W < V

mm caw

IN W*TT-SEC/GR*U-"C

FIG. 4 MOLLIER CHART LM-3 CYCLE

APPENDIX

SHEET 1DOUGLAS POINT STEAM CYCLE DATA

F1 ow ( F ) p ( ba r ) Wet » Fh

Cond'ate from turb.Steam at exhaustSt.Ej.Drain to Cond'rSp.Lk. Dram to Cond'rDrain to Cond'rCcmd.Ext.P. outGenerator Cooler ouiFW from E ] . St. CoolerFW for Sp. Lk. CoolerFW from Drain Cooler to #1D r a m from #1 FWHExEj . k Spindle Lk. St.: Fh (FW.Drains i passing St.;St.after bleed to #1Bleed to *1 FWHExFW to 12 FWHExDrain from »2LFh (FW,Drains t. passing StlSt.after bleed to »2Bleed to (2 FWHExFW to *3 FWHExD r a m from #3ZFh (FW,Drains i passing st.!it. after bleed to #3Bleed to »J FWHExGland St. to «3FK to »4 FWHExLFhiFW, D r a m s i. passing StlSt. after bleed i-.o *4Bleed to t4 FWHExDrain to *4FW to B1r.F. Pump ,FW to 15 FWHExFh (FW,Drams & passing St.'.

St. after reheaterSt. after bleed to «5Bleed *.u (5Gland St. to »i>Separator DrainSt. to SeparatorFW to 16 FWHExlirai n f rorr. #6St. after bleed to tCBleed to i6 FWHExFW to Bol1erSt . after G.ar.d to #j

aîter G-and to «3t^ t-rtir.e f'-r W L T (..; •- -ri. k e "ect or

er.eat»»." r.-r.der.sate7 - t a l Water te Be i 1erT-. tal i t e x - i r x . B, l i e r

00000iîiî

n

u

00I0

1010

1Û01

10011

1

00011

101111

.97564

.97564

.0004563

.0003042

.14 7157

.123558

.123558

.123558

.12J558

.12 3 5 5 8

. 147157

.0007606

.97564

.06 784 0

.12355b

. 079317

.T4348

.01970

.12 3 5 5 8

.039*17

.0b3id

. .J3T6 7

.001947

.123556

.120850

.045237

.273488

.442333

.442 333

.16 6 13 7

.166137

."34522

.012169

. 143396

. 349055

.442333

.r'-S401

. 349055

.078401

. 442333

.4^7456

.4 i 9 i> 2 5

.4418-7

. J 4 - 1 3 S

262687100

26282B2832fa7250

70706571

8888d5 .

91 .

Ill ,111 .14 i.104 .

166.16b .134.128.12S.

2 2 1.

i 'J 1

151 .\ 1 56 .1 5 1 .

-147.153.175.175.

235.2 35.2 (i ''t .

- 5 '",

.1172

. 1172

.7314• J.

.2520

. 4032

. b364

. 7«, 19

.15 5 5

.7015

.0

.9B

.98

.52

.OB

.49

.49

.49

.1 3

. 1

.1

.9

.98

.89892501.69

li1212~"4 )Ob

97^ 2

515 1h j22

a

0.033840.0 3 3840.642871.01330.24.04 .04.04.04 . 0Q.2639.97

0.32540.3254•2.544

0.6447

0.66190.6619'2.5440.6447

1.3691. 3691.3692 .544

2.6 202.620•2.5442.544• 4 0

4. 7574.90914.909](4.909114.90915.171

4 0

9.0 J-

30.75

39.97

11.2

5.5

2.B

Just dry

Dry

DryWetWet

9.75

2 . 3

Just dry

1092259367

419144110119120120139283

2794

25012501274297

25932593338381

269426942760439.

2S0O.2800.564 .537.543.

2901 .2 72? .2~25 .

2^60.636 .

2494.625 .647.

2574.2574 .-JC.

2-60.-•60.2"94 .2^94.

.44

.01

. 392

.06

.142

. 333

. 3255

. 30

.95

.19

. 377

.92

. 15

.15

.07

.33

.26

.26

.16

.49

.90

.90

.96

.85

.2727.19.54.35

4561.619684406009424 236969É929 2

071100000006

7

00

+ 27711

•* 2

77

1 .+ 2.7.7

1 .

i .

7 _6.6.6 .1 .6.

1 .1 .6.b" .

-' ,

b

b .t-

.3B263

.5655

.16616

.3069

.49656

.JB353-•41554

.418770

.42092

.4 7B91

.92663

.05866

.377

.377

.89809

.96749

.126

.323

.323

.13924

.2039

.126

.2663

.2683

.435B

. 3623

.126 =

.22834

.228146 88.6131.6170502176 —-—,- 74 ^""^774058665292124ei"iB~661" 3817 39Dé 3586 84 Î5 6Ï5 86^ 5 t

(See f

-156412

-112

-30723

-282

-40215

-385

? .-494.

-Tinr

154.

-783.-621 .

neet 2

. 3BS

.701

.126

.561

.934

.583

.414

.115

.376,197,695

299

,692T9T

F.'.h 205.054

33.

902 .50.

1053.

598

324732

4 2-'

' J 9 . 9 "5 6 2 4 :=t 2 5 C . . 9 j <". .y-: hi

APPENDIXSHEET 2 DOUGLAS POINT STAGED EFFICIENCY ANALYSIS

Read as additional to Sheet 1

Ah(from point above) Fh

WorkAFh

Cumula-tiveWork

FhCorrec-tion (FromSheet 1)

NetFh

r*

E Cum WkG-Net Fh

Turbine exit loss 18.6Steam at exhaust 0St. after bleed to #1 242.14St. after bleed to #2 92.11St. after bleed to #3 101.64St. after bleed to #4 105.37St. after reheater 101.IB

Reheat 175.84St. after bleed to #5Separator correction + 1-45St. to separator 0St. after bleed to #6 60.02St. after glands 186.54St. to turbine for work ' 33.96Total steam from boiler -

" corrected for Sep.error

18.2203.2440.2706.2919.3138.3383.

147981222015062683488

3178.435+ 1.963

3473.0343941.1494029.9314366.8084366.808

-18.147

236.24196.115

110.094118.104117.990(205.054)

107.951266.27848.966

-18.147

218.094314.209424.303542.407660.397

768.3481034.6261083.5921083.592LOS1 .592

-112.561-282.225-385.173-486.695-621.891

A-605 .438

B-847.093

2327.6612423.7902533.8892651.9882761.597

2867.596

3182.83831B2.8453184 .809

0.09190.12710.16420.20050.2345

0.2627

0.33380.33380.3336

oo

#6 FWHEx DrainEjector SteamSpindle Leak

Glana St. to #3Gland St. to #5

ReheatFW Heat

AFin

50.7321.276Û.850

5.37633.598

205.054-902.324

-605.438

HF-Mi

1.275

205.054-1Q53.4J2

-847.093

* E r = Generator Effy. = 0,9805

Cumulative Workn = E. Net Fh

APPENDIXSHEET 3

LM i-P rEAM . . ÏCLE Hh/ HALA-CCE

Total Water to Boiler

HP Turbine Throttle

HP Turb. Exit TotalBleed to *2 RHTRBleed to #9 FWHExSteam Leaks (used)#1 ReheaterIP-1 Inlet

IP-1 Bleed IB

IP-1 to RHTR i, il

T°C

308 .11

399 .048

322.370

p b a r

175

172. 5

95.84

1381 .41.h*1525.59

2907M i - 9 0 . 9 J

2 3 1 6 . 0 7

• ! i 4 2 9 4 . ' j 7

399

271

' 2 "

. 7 7 7

. 4 2 B

04 2

9 1

34

21

. 1 0 S . \ ' t

.'": - 2 1 22MIJ i

! i - ^ a2S24

. 14

. 00

. 1 4

. 31

. 8 3

J . 2 8 9 7 9

t 5 . 7 3 2 4 e

5 . " 8 4 5 9

b . 2 76 76

t . J 2 4 5 -

h . ! 3 3 4

T u r b .S ta . iuK f f y .

89 .'I

t

0 .0 .J .

J .L .

r 1- l . ' " d

17 14 4 4i ) 6 7 1 J• 1 4 6 7 7

'/ 3 6 11• • 0 9 6 4

L i. B ' J O 4

owPass i rv j on

1.

1 .

0 .

998424

998424

998424

475390475 192

371"79

10 2 811

WorkF i

181 .

312.

107.

h

7 1 7

7 8 3

4 24

HeatF i h

3048 .775

441 .245

IP-- Bleed »6

IP-2 I : M I #5

175.3b

151 .85

• 2 Sfi.ejtiT

Re.'ieater St.iqe

Pthedt"! l.x i tL5- Turb. 11. lft

LP Bleed »4

LV Bleed #3

LT b l . - f i #<-

LI [<ifi'J • 1

1. P 7 - i L . L A 11

3 0 7 . 0 8

i h - 1 5 J . 5 J9 . 0 2 6 " 1 . 0 h . 4 1

• h - 8 0 . •5 . 0 2 5 9 1 . 0 tj . 4 b

•.h*2 1 t i . l "2 2 . 2 3 0 3 5

3 9 9 . 0 0. 2JJ . ~J

1 1 6 . 8 9

111 .1

8 8 . 4 4

7 0 . 9 8

2 b . 12

2 6 . 12

224 . 80

2 .62

1. 369

0.662

0. 325

0.034

0. 3J4

• h - 2 0 b . S3 2 4 1 . b2 9 0 1

'. h - 4 4 1 . 5 )2 t i O H . 2 7

Mi - 1 0 r. . j -2 b 9 4 . 4 0

M i - 1 0 1 . 6 42 5 9 3 . 2 6

' h - 4 2 . 1 12 501 . I S

• h - 2 4 2 . 142 2 b 9 . 0 1

1 6 . b1 0 ' ) . 44

-.h 2 1 6 8 . 17

K'n 'Ht

6 . 7 4 4 3 6

" . . : 7 5 "

- . 2 2 8 3 4

" . 2 6 8 3

" . 3 2 3

" . j " 7

7 . 5 6 5 5

0 . i 6 2 6 Ï

:»8» = Gn. ss

Û. 0 ^ 6 5 4 8 ' J . 0 3 6 2 6 3

0 . G 1 6 2 6 3

1 . 1 6 8 0 0 4

1 . 1 fi B 0 0 4

0 . 0 4 7 1 5 4

0 . 0 3 7 6 7

0 . 0 3 9 7 0

C . 0 6 7 B 4

0 . 9 ^ 5 6 4

1 5 . 8 1 5

2 . 9 0 1

1.12 0 8 5

1 .08318

1. :434S

0 . y 7 51 4

0 . <* 7 5 fc 4Cur . Wi.i '•

T o t a l H t . t r o r ' r a a n ; ;Wk H t

l ' v c l d E f f i c i e n c y

5 1 5 . 7 0 9

118 .104

110 .094

96 . :15

2 16 . .'4 1

- I S . , 4 "

( 2 4 5 . 4 7 9 '

2 4 1 . 5 4 3

1 5 . . ' • >

i ; . s b i'j . 4 4 9 8 8 :.0 . 4 4 1 "8

APPENDIXSHEET 4

- 20 -

LM-3A ADDITIONAL DATA

Note: For Condenser to PWHEx #3 see Douglas Point Data Sheet 1

Flow(F) T°C p(bar)

#6 Dram to #5 Dr.#5 Drain to #4

FW from #4 to BFP #1

FW to #5 from BFP #1

FW to #6 from #5

FW f*:om #5 to BFP #2a

#7 Drain to BFP #2b

BFP #2(a) out

BFP #2(b) out

BFP #2(a) & (b) mixed

#8 Drain to #7 Dr.

FW from #7 to «8

FW from #8 to «9

FW from #9

Bleed #9 to «9 FWHEx

Bleed #9 to Rehtr

BFP #3 in

BFP #3 out

FW from #9 + BFP #3 mixed

0.0665480.114980

1.285692

1.285692

1.285692

1.285692

0.204575

1.285692

0.2045746

1.490267

0.103611

1.490267

1.490267

1.490267

0.336713

0.171444

0.508157

0.508157

1.998424

127

128

18C

184

241

307

308

.958

.7146

.0

.536

.98

.04

.11

2.540

40

21

200

200

200

200

200

647.09564.19

537.54

543.35

625.60

730.36

763.7

756.06

792.83

761.11

930.19

912.32

1049.49

1373

2816.07

2816.07

1384.25

1406,09

1381.41

1.8781.6788

1.61255

1.61725

1.8171

2.0590

2.138

2.C761

2.15709

2.08736

2.4905

2.41151

2.68563

3.2751

5.78459

5.78459

3.32142

3.3321

3.28979

APPENDIXSHEET 5

- 21 - REVISED 16/12/68

COMPARISON OF EFFICIENCY CALCULATIONS

Qs Boiler1st Reheater2nd Reheater

Total EQ

Qx Boiler Feed PumpsGenerator Cooling

ErrorTotal Q x

Qw FWHEx DrainError

Qx-Qw Total

FAh

3182

3182

8111il

5

16

EFAs Boiler)Top T.Carnot& Reheaters) Extra

H.P. TurbineIP-1 TurbineIP-2 TurbineLP Turbine

Total Turb.

Separator & Regen RHTRBoiler Feed Pump(s)FW Ht ExchangersMisc.Drains etc.

Generator Cooling£netFAs Total

Qc or IcFAs (check)

Qc - (Qx - Qw)

[Qc - (Qx - Qw)]/^Qs

Work/EQg

EG

E (Work/EQS)

E G [l+fQx-oC-Qci/^Qs1

FAh423.

~

660.1083.

2115.

2099.

DOUGLAS POINT

.83--.83

.3799

.1069

.963

.4498

.3187

.1131

FAh/T j

10.6350--

10.6350

0.037112

0.053900

= WORK19

39 el58 e*

35

219

0.340445

0.9805

EG{l+[ (Qx-Qw)/T2-!:netFAs]/(j;Qs/T

EG{l+[(Qx-Qw)/T2-5;cFAs }/(tQ,

t incl.cit loss

7.06813

7.01423

0.659542

(0.340458)

0.9805

1

FAs

6.13590

Net

+6.+0.

+0.

+0.

+0.+0.+0.-0.+ 0+ 7

7

7

0

(0

0

--

FAs

08386070044

216516

465991

10450500562509549101C622037112068523

06852

01462

.659579

340421)

.9805

Gross SteamCycle Effy.

0

0

0

0

.3338

.3338

.3338

.3338

LM-3A

FAh

3048.775441.245241.543

3731.563

57.569611.1069

-68.€765

5.31870.794

62.56 38

FAh/T2

4.0.0.

12.4684 5.

0.03712

0.209048

FAs

881530726143387006994679

Net FAs

+ 5.+0.

FAh = WORK181.717420.20718.716

1058.1161678.756

2115.35

2052.786

0.-449880 (

0.982

+0.+0.

net incl ."

+0.+0+0-0+0+ 7

7.06813 7

6.85908 6

0.550117 0

0.449883)(0

0.962 0

54335451329

104138082651009688583209779687

057707037964171261009885037112068525

.068525

.85948

.550149

.449851)

.982

Gross SteamCycle Effy.

0

0

0

0

.4418

.4418

.4418

.4418

APPENDIXSHEET 6

- 22 -

MORE DETAILED COMPARISON AND ANALYSIS

Douglas P o i n t LM-3A

FAS FAh FAs Fûh

Primary JEFAS Boilerheat \ 1st Reheatersupply 2nd Reheater

FAh/T j

H.P. Turb.to BleedBleed to exitH.P. Turb.IP-1 Turb.to Bleed 8Bleed 8 to Exit (7)IP-1 Turb.IP-2 Turb.to Bleed 6Bleed 6 to ExitIP-2 Turb.LP Turb.to Bleed 4Bleed 4 to Bleed 3Bleed 3 to Bleed 2Bleed 2 to Bleed 1Bleed 1 to ExitExit Loss

LP Turb.

All Turb.

Total

Extra

Total

Total

Total

Total

TotalSeparator net to Sep.DrainRegen.Re.ieaterBoiler Feed Pump 11

#2 and Mixing13 and Mixing

•9 Feed Water18 Heat#7 Exchangers• 6• 5#4• 3• 2#1

Total FWHExGenerator CoolingMisc.Drains etc.

TOTAL

+ 6.1539045--

+ 6.1539045+ 6.08386+ 0.0700445+ 0.1644429+ 0.05207352+ 0.21651642

------

+ 0.06180526+ 0.04404940+ 0.05924995+ 0.05034792+ 0.18390814+ 0.06063+ 0.46599067

+ 0.68250709+ 0.02735773+ 0.07714719+ 0.00562510

-----

+ 0.01848607+ 0.01807309+ 0.00949293+ 0.01003360+ 0.00926641+ 0.03013905+ 0.09549115+ 0.0371122- 0.01062156

+ 7.068523

3182.--

3182.

315.107.423.

------

117.118.110.96.

236.- 18.660.1083.

94.205.8.----

151.118.109.91.94.151.718.11,

828

828

241950191

98810309311424147391

582504053799

10363760578485631351069

+4.88153032+0.7261427+0.38700644+5.99467946+5.54335+0.45132946

--

+0.1041379+0.0705384+0.0121128+0.0826512+0.0078753+0.0018131+0.0096884+0.17828413+0.0447891+0.0592499+0.0563479+0.1839081+0.06063+0.58320913

+0.77968663_

+0.05770722+0.0060428+0.0261025+0.0058190+0.0490863+0.0112602+0.0229627+0.0094134+0.0203452+0.00874601+0 ,01004139+0.00926641+0.03013905•+0.17126066+0.0371122-0.00988543

+7.068525

30484412413731

1813Ï2107420152

.775

.245

.543

.563

.717

.783

.424

.207

.815

.901

role51511811096

236- 181058

1678

245739114822042251341051099194151159911

.709

.104

.094

.115

.241

.147

.116

.756-.479.46987,.00155.09815.113.4193.3504.688.7475.7607.78.48.56.898-1.1069

17/12/68

- 23 - AECL-3221DM-96 (Supplement)

LM-3A CYCLE - ORGANIC LIQUID PRIMARY HEAT EXCHANGE

by

W.B. Lewis

Since it seems desirable to keep the top organictemperature not above 410°C, I have reviewed the heatexchange to boiler and reheaters. The overall pictureis given in Table IA and Fig. 1A.

For a top organic temperature of 420° to 425°Cthe heat exchangers may be fed simply in parallel and allexit the organic at 340°C, as shown in Fig. 3 (see original text)

For a top temperature of 410°C it seems necessaryto adopt a more complex circuit. An extra circuit isadded to each reheater taking organic at 360°C from theboiler and exit at 340°C, as shown in Fig. 2A. (Possibly356°C would be a better break point than 360°C).

The necessary flows are evaluated in Table IIAand the temperature differences are shown in Fiq. 3A.

This revision leaves the steam cycle unchanged.

Atomic Enerqy of Canada, Limited

Chalk River, Ontario

December, 19 68

- 24 - AECL-3221DM-96 Supplement

Cycle LM-3A - Echange thermique primaire par liquide organique

par W.B. Lewis

Résumé

Comme il semble souhaitable de maintenir à 410°Cla température maximale du caloporteur organique, l'auteura révisé l'échange thermique à la chaudière et aux réchauffeursLes résultats d'ensemble figurent dans le Tableau IA et sur laFigure IA.

Si le caloporteur organique a une temperaturemaximale de 420°C à 425°C, les échangeurs thermiques peuventêtre alimentés en parallèle et évacuer tous le caloporteurorganique à 340°C comme on peut le voir sur la Figure 3 (voirtexte original).

Si la température maximale est de 410°C il semblenécessaire d'adopter un circuit plus complexe. Un circuitsupplémentaire doit être ajouté à chaque réchauffeur lequelprend le liquide organique à 360°C dans la chaudière etl'évacué à 340°C comme on peut le voir sur la Figure 2A.(Il se peut que 356°C soit un meilleur point de rupture que360°C) .

Les écoulements nécessaires sont évalués dans leTableau IIA et les différences de température sont indiquéessur la Figure 3A.

Cette révision ne modifie pas le cycle de vapeur.

L'Energie Atomique du Canada, Limitée

Chalk River, Ontario

Décembre 1969

28/12/68

TABLE 1A

LM-3A PRIMARY HEAT EXCHANGE (DATA FPOM DM-96 AND NEL STEAM TABLES)

SECOND REHEATER (2ND STAGE)

T

307.308.310

320330340350354.

3603703S0390

399

°C

0811

6

p=17 5 bar

h(J/g)

1381.411392

1450151315811663171125302610271427892851

p=172.5

2907

F=l.998424

Fh

2760.6432781.806

2897.7153023.6163159.5083323.3793419.3035056.0135215.8875423.7235573.6055G97.507

5809.419

p=92 bar

h(J/g)

2816.072825287229142952

29883021305 33084

p=91.1 bar

3115.14

Ah

8.93r5.9397.93

135.93

171.93204.93236.93267.93

299.07

F=l.475390

F/.h

13.17582.519

144.485200.550

J224.983

253.664302.352349.564395.301

441.245

p=22 bar

h(J/g)

3035

3043

3066308931113134

3156317832003222

3241.8

Ah

8

31547699

121143165187

206.8

F=

9

366388

115

127

141167192218

241

1.168004

FAh

.344

.208

.072

.768

.632

.452 {

.328

.025

.721

.417

.543

In

Diff.

Total

I(Fh+FAh)

2760.6432791.150

2947.0983169.2073392.7613639.5613771.7375408.4475610.8795893.1006115.8906311.225

6492.207

2760.643

3731.564

1

toLn1

- 26 -

TABLE IIA

LM-3A Organic System for 340 to 4I0°C

To clear the pinch point in the boiler,at 380°C all the organicis taken from the reheaters to the boiler. Then some organicis bled back at 360°C to serve the reheaters and then exit at340°C.

From data of Table I and DM-96

Total heat to cycle EFAh = 3731.563 (= 100%)

Assume enthalpy h linear with T for organic then

FAh %Flow of organic FQ = - ?

= 1.87589 FAh/AT %

Reheaters

Boiler

Steam Temp.Range

350-399°C

310-350

OrganicTemp. Range

410-380

380-360

360-340

From

366

316

AT

30

20

20

FAh OrganicTable I Temp.Range

.606 410-380°C

.182 360-340°C

Fo(From above)

77.076%

100%

70.344%

Total FAh

AT

30°C

20°C

F =O

Water-SteamFAh = 0.

1232

1066

749

= 3048

.630

.160

.980

.770

1.87589 FAhAT

22.924%

29.656%

533080 F ATo

420

320

300

2802000 3000 4000 5000

(ARBITRARY ZERO)6000 6500

3oeii°c

i**T^>? CD

CONDENSER0.034 BAR

26.I2°C

6 I**7 I72.6°4 *k \\**j

G2

I47.9°C3 128.7° I28°d

IO5°Ci * j »* - ,

v^DEAERATOR [

Bi.€t

*&.

^655°C

S ^332°C

2I7°C I8O°CGENERATORCOOLING ETC.

FIG. 2 A L M - 3 A STEAM CYCLE

tooo

420

t

FIG.3A LM-3A DIVISION OF PRIMARY HEATEXCHANGERS FOR 340°C TO 4IO°C ORGANIC

toD

2802000 3000 4 000

Fûh5000 6000 6500

Additional copies of this documentmay be obtained from

Scientific Document Distribution OfficeAtomic Energy of Canada Limited

Chalk River, Ontario, Canada

Price - $1.00 per copy

3835-69