Ammonia Synthesis Loops Variables Investigated by Steady-state Simulation

7/30/2019 Ammonia Synthesis Loops Variables Investigated by Steady-state Simulation

1/14

AMMONIA SYNTHES IS LOOP VARIABLESINVESTIGATED BY STEADY-STATE SIMULATION

LARRY D GAINESResearch Development and Engmeenng I)lvlslon, Apphed Automation, Inc. Pawluska Road, Bartlesvtlle, OK74004,USA

(Received 16June 1977, received for pubhcatmn 11May 1978)Abstract-A steady-state model of an ammoma synthesis loop contauung two parallel quench-type converters ISdeveloped The system modeled Includes converters, compressors, separator, purge and recycle and the model ISused to determme the effects of syntbesls loop vanables upon ammoma productIon Model slmulatlon mdxates thatlow purge rates lead to locreased production, however, rnefficlent converter operation may prohibit operation mthis regon due to certam process constramts A low H2/N2 ratto also results m Increased production The optundratio, however, 1s pnmanly dependent on synthesis compressor power costs

UWROXJUCTIONAmmoma IS produced by passmg a hydrogen-mtrogenmixture over a promoted uon catalyst at elevatedtemperatures and pressures High temperatures areemployed to achieve a sticlent reaction rate Since thereaction IS exothermlc. the gas leavmg the catalystsectlon IS at a high temperature and heat exchangebetween this stream and the feed stream IS employed tobring the feed gas to reactlon temperature Through thisheat exchange the heat of reactlon supphes sufficientenergy to mamtam the reactlon zone at desired reactiontemperature Such processes are termed autothermlcprocesses and have been discussed m detd by vanHeerden [ 11The extent of reaction m the converter IS pnmardy afunction of the synthesis reactlon eqtuhbrmm and am-monia content m the reacted gas 1s normally m the rangeof ll-13% for a process operating at 150 atmospheresBecause the per pass converslon IS low, a recycle loop 1srequired to make the process economically feasible

Synttmsls gosmmpnnrsorr

Rgure 1 dlustrates how recycle IS accomphshed m theprocess modeled Part of the reactor effluent I S mlxedwith feed gas, cooled, any condensed ammonia ISremoved, and the remamder Is subsequently recycledback to the converterThe feed to the synthesis loop contams a near stol-chlometnc ratio of hydrogen and nitrogen Argon andmethane are Inert gases and trace quantities are presentm the feed gas As hydrogen and nitrogen react andammonia IS removed, the merts level budds These mertsddute the gas stream and tend to quench thereaction[2.3] At high pressures, the merts level maystabthze at an acceptable value due to merts being da-charged from the synthesis loop m the hquld ammomaproduct stream When lower pressures are employed,part of the recycled gas must be purged to mamtamadequate reactlonConverter stabdlty and the effects of process variablesupon converter efficiency have been dlscussed bynumerous authors [4-6] Ammonia converters have

&-ramnla aepamtor

Fig 1 Synthesis loop flow dtagram37


2/14

38 L D GAINESfrequently been modeled to obtam a better understandingof the effects of each variable upon converteroperation [2,7-91This work extends a previous converter model[lOl toinclude additional mayor synthesis loop processingcomponents, (compressors and separators) and m-vestlgates the mdlvldual and collective effects of processvanables on the synthesis loop Process variables consl-dered are feed rate, merts level, merts level in the feed,converter temperature, separator efficiency, and HJ N+ratio Important constramts affecting synthesis loopbehavior are compressor speed and compressor driverpower, merts level, heat exchange, and synthesis looppressure These constramts are discussed in hght ofsimulation results

AMMONIA WIYTHESISOOPDESCRIFIIONTh e process modeled 1s a low pressure, 600 TPD plantas shown m Fig 1 Synthesis loop feed consists prunardy

of a near stolchlometnc ratio (3 1) of hydrogen tomtrogen with trace quantities of argon and methane Thefeed IS compressed to the synthesis loop pressure by twocentrtfugal compressors m series The compressors areon a common shaft and are dnven by steam turbinesRecycle gas, containing approximately 1 -13% ammoma,IS introduced mto the last stage of the second compres-sor The compressor discharge IS cooled m an array ofheat exchangers m which several flow paths may existsunultaneously to achieve the desired temperature Thechdled gas enters a separator where hquld product am-moma IS removed

The remammg gas passes through an exchanger whereIt IS heated by a process stream before being divided andfed to two identical parallel converters as shown in Fig1 Each of these two streams must be further dividedinto five portions, the first of which enters the top of theconverter to be heated, and the remammg four streamsare used as quench as shown m Fig 2 The feed gasentering the top of the converter flows downward aroundthe outside of the insulated catalyst section At thebottom of the converter, the gas IS heated by countercurrent heat exchange with the reacted gas and enters acentral riser The riser empties into the top quench zonewhere cold feed gas 1s introduced The gas at the resul-ting temperature enters the first catalyst bed where reac-tion between hydrogen and nitrogen forms ammomaaccording to eqn (1)

The temperature of the gas mcreases as it reacts in thebed and must be cooled in the next quench zone beforeentering the second bedGas from the fourth bed has a near equlllbnum

concentration of ammonia and supplies heat to the feedgas before leaving the converterThe converter efluents are mixed, cooled, and a rela-tively small purge stream IS drawn off The purge gas ISfurther cooled and lighter gases enter a fuel systemLiquid product ammoma IS drawn from the purge

JJ ouenchOuench#I -

Quench 962

Quench # 3

Quench#4

----lr---- IL: -* FToduct gosFig 2 Ammonia syntheslr converter

separator The major portion of the converter effluent ISrecycled to the last stage of the second compressor, thuscompleting the synthesis loop

CompressorsMODELDEWUWMENT

Synthesis loop feed gas IS compressed m two nme-stage centiugal compressors operatmg in seriesRecycle gas IS introduced on the runth stage of thesecond compressor m addltlon to the feed gas Suchcompressor configuratlons are commonly found m am-moma plants and are discussed by [1 1]The molecular weight, MW, of the stream IS computedfrom the component molecular weights, m, andstream composltlons, F, (lb mole/hr)

and suction volume, V (acfm), I Scomputed from

(3)


3/14

Ammoma synthesis loop var iables mvestlgated by steady-state sunulatl on 392 IS the compresslbll lty factor, T IS suction temperature(OR), P, IS suction pressure (atm), and R IS the gasconstant With these data and compressor speed, S , thecompressor performance curves may be used to deter-mme discharge pressure and the horsepower requued bythe compressor

T he performance curves are computeri zed usingequahons stmll ar to those dlscussed by Wh ite1121 forcentr ifugal compressor surge contr ol T he surge flow, AS(acfm) may be computed as a lmear function of speedfor the compressor operatmg r ange from

AS = Cl,S+ C2, (4)Where Cl , and C2, are constants The head at mmlmumflow, HM (ft) I S a functi on of speed squared and 1scomputed from

HM = C3,S2 f C4, (5)and the head at operatmg condltl ons I S obtamed from

H = HM+C5,(V-AS)2 (6)This value of head 1s used to calculate compressor

discharge pressure, Pd (atm), fromPd = Ps(CQHMWI1545 O/T + 1 0)c7, (7)

Constants Cl ,-C7, m the above equations were deter-mmed by regresslon with data from performance curves

T he first compressor discharge I S cooled and anycondensate I S removed This stream becomes the feedfor the second compressor whi ch 1s computed in twoparts The performance of the first section (compnsmgthe first eight stages) of this compressor I S computedfrom eqns (2)-(7) using a new set of constants and P, setequal to the discharge pressure of the fir st compressorBefore repeatmg this procedure for the mnth stage of thesecond compressor, the recycle gas and feed gas must becombined

T he suction temperatu res and first stage suction pr es-sure are taken from design dataHeat exchange

T he compressor discharge I S split and cooled m anarr ay of exchangers H eat I S removed by exchange withprocess water, a process stream and by refweratlonsupphed from a refrl geratl on umt not shown m F ig I Amodel of the refrl geratl on unit and complex exchangernetwork I S beyond the scope of this study T he only heatexchange equipment modeled I S that physicall y locatedm the converter I t I S assumed that all other processstreams may be mamtamed at design temperaturesF actors affecting heat exchange requirements are,however, discussed since heat exchange limit&tons maybe encounteredAmmonca separator

T he compressor discharge, af ter bemg cooled, entersan ammonia separator where hqmd ammoma I S removed

The separator operates at -10F and at a pressureseveral atmospheres below the discharge pressure At140 atmospheres, only minor quantl tles of reactants andmerts are present m the liquid ammonia product T heammonia m the overhead gas, however, I S cntical sincethis stream I S fed to the converter Guerren [ 131computed per cent ammonia as a function of tempera-ture, pressure, and merts level H I S results show that theper cent ammoma leaving the separator m the gas phasewil l change by less than 0 2% over the pressur e andmerts range encountered m the separator Smce theseparator operates at a specifi ed temperature, constantvalues may be used for a hqmd-vapor eqmhbrmmconstants, K , T he liquid fraction, L f, of the separatorfeed I S computed using the component molar flow rates,E (lb mole/hr) such that

(8)

The component vapor flows, F,, (lb mole/hr) from thetop of the separator may then be computed from the inl etflows

F = KE(l-Lf)a Lf + K ,(l - 4)L iquid-vapor equthbrmm constants,Table 1Ammoma syntheses converters

(9)K,, are grven m

The gas leaving the ammonia separator 1s heated andthe stream is split between the two converters E qualquanti ties of gas are assumed to enter two tdenttcalconverters, thus, calculati ons for only one converterneed be performed

The converter catalyst section consists of an outerpressur e shell and a catalyst basket w ith cold feed gasflowing m the annulus between the two (Fig 2) T hecatalyst basket I S coated with msulatl on to mlmmlze heattransfer from the hot catalyst to the cold feed gas Anunmsulated ri ser tube transports gas from the heat ex-changer section thr ough the catalyst sectlon to the firstbed M mmtal heat transfer to the riser gas IS due to asmall heat transfer area and because the ri ser gas hasbeen heated to near reactlon temperatu re The catalystsection 1s nearl y adtabatlc and an assumption of umformradml temperatu re I S made A umform velocity profile mthe bed may also be assumed smce the bed dmmeter I Smuch larger than the catalyst dmmeter T he assumptionsof undorm radial temperature and velocity reduce theTable 1 L lquld-vapor eqtnhbnum constants at 140 atm and-10F

Nitrogen l69.ocoAnlnolua 0.023Argon 1oO.ooOMethane 66,Qxl


4/14

40 L D GAINESequations t o an easily usable form while retammg thedesued accuracy of the modelMa r en a l b a l a n c eThe hydrogen consumed m a dlfferentlal element of acatalyst bed IS aven by

F,dx=rAdz (10)where A I S he cross-sectional area of the bed in ft and x1s the fraction Hz converted F , IS the flow rate of Hzmto the bed m lb mole/hr Units are temperature, R,pressure, atm and flow, lb mole/hr Figure 3 IS asunphfied diagram of the synthesis converter whichshows flows and temperatures used m the followmgequations The component flows at any point m the bedmay be computed from mlet flows by

F; = F,(l -x)F; = F2 F,x/3F; = F 3 + 2F,x/3 (11)FL= F4F:=Fs

where the order of components IS hydrogen, mtrogen.ammoma, argon and methaneThe rate of reaction r , for eqn (10) IS found from the

Temkm [ 141 equatronr= 3+.~cl,(~)- - ($)-aJy (12)

where K, IS the equdlbnum constant, k z I S the rateFru ,

Fig 3 Block diagram of synthesis converter

constant of the reverse reaction, y IS the catalystactlvlty, 5 IS the effectiveness factor and a l , UZ ,and u3are the actlvltles of hydrogen, nitrogen and ammoniarespectively K , I S calculated from the equation ofGlllesple and Beattle [15 1log,,K, = -26911221ogjo T -S 519265x lo-- T

+1848863x10-T2+20016/T+26899 (13)where T L Sm K Activity coefficrents used to calculatethe actlvltles for eqn (4) are computed fromRTln Yr =[(B ,-A,/RT -C,/T)+(A, -S,)* /RT] P

(14)where T I Sm K and R has the value 0 08206 ConstantsA,, B, and C, are gven by Nlelsen[5] and appear inTable 2 S, IS computed as

where rnI refers to the mole fraction of component I Inthe mixture and the summation includes all componentsm the mixtureNlelsen[S] discusses the possrble values of a andconcludes that LI should have a value between 0 5 and0 75 A value of 0 5 was used m this study to obtam thereverse reaction rate constant k , as outlmed by Dysonand Slmon[l6], but usmg actlvltles as dtscussed m thispaper Data from Nlelsen[5] for an mdustrlal catalystwere fitted and the resultmg equation

k 2 = 0 146578 x 10 ec--nr) (16)has units of lb molesh ft3 and T I Sm R R IS the umversalgas constant and has a value of 1987 Btu/lb mole RThe effectiveness factor ,$ IS derived by Dyson andSunon[l6] for Isothermal spherical catalyst particlesThey assumed the ddfuslon coefficients were mdepen-dent of posltlon m the parucle and that Knudsendiffusion does not occur The resulting equation

~=bo+b,T+b2q+bjT2+b.+qf+b5T3+b6$(17)

gwes the effectiveness factor as a functton of T),temperature, K and pressure for a 3 to 1 H2/N2 ratio and12 7% merts Constants for use wzth eqn (17) are gvenby Dyson and Sunon[ 161 and appear m Table 3 for

Table 2 Constants for calculation of actlvlty coefficrents1 A1 BI CplO-4

%J 1 0.1975 0.02096 0.0504N2 2 1.3445 0.05oLb 0.4209 3 2.3930 0.3w5 476.87A 4 1.2907ml4 5 2.2769


5/14

Amm oma synthesis loop variables Investigated by steady-state stmulattonTable 3 Constants for cakulatlon of effectweness factor

Pressure, ati b. bl b2 b3 x 104 bq b5 x 10' b6150.0 -17.5391 0.07698 6.90055 -1.08279 -26.4247 4.92765 38.9373225.0 - 8.2126 0.03774 6.19On -0.53547 -20.8696 2.37914 27.8840

41

300.0 - 4-6757 0.02355 4.68735 0.346331 Il.2803 1.54088 10.4663

pressures of 150, 225 an d 300atm The fracttonal con-verston of nttrogen IS 7 and may be obtamed fromv = molar flow of NH&molar flow of NH3

+2 molar flow of Nz) (18)Catalyst actlvlty, Y deteriorates with catalyst useFactors affecting deterloratton are discussed byNielsen [5]Ene r g y ba l a n ce

The converter may be divided mto three majorsections, catalyst beds, quench zones, and heat ex-changer The energy balance for the converter may beobtamed from the energy balance equatrons for the threesections by combining them m the proper order Theenergy balance equations for each major section aredeveloped The energy balance for a honzontal dlfferen-teal element of catalyst bed mvolves four heat terms

dQrcact,en + dQ,.s + dQ r ,scr dQannu,uo = 0 (19)of which the last two are small as already mdlcated butare included m the model The heat generated by theformation of ammonia IS dQractlon and IS expressed as

dQ,,,,t,,, = - f AH,F, dx (20)A& IS the heat of reaction corrected for the heat ofmlxmg m Btu/lb mole NH3 formed, and 1s computedfromHR = - 23840 57 + (P - 300)(1 08 + (P - 300)[0 01305

+ (P - 300)(0 83502 x lo- + (P - 300)x (0 65934 x lo-))]} + 4 5(1391- T) (21)

This equation was derived from the heat of reaction datapresented by Kazarnovskuf17] for the formation of amixture contammg 17 6% NH3 in a 3 1 HJ Nz mixtureMost of the heat liberated by reaction IS expended mralsmg the temperature of the reaction mixture asexpressed by

dQ ea. = - MCJ T , P)dT (22)where M 1s the moles of reaction gas C,(T , P) I S afunctton of pressure, temperature, and composition andIS calculated from the BWR equation Calculation ofmixture heat capacltles, enthalpies, viscosities, and

thermal conductlvltles are explained elsewhere m thispaperThe remammg two heat terms refer to heat transferredfrom the bed to either the gas flowing in the annulus onthe outside of the catalyst basket or the gas in the riserHeat transferred to the gas m the annulus IS computed by

dQ ann u l u s k C ( T - TAM dz = f&fL.(Ta. I+) dT,(23)

stnce the primary resistance to heat transfer IS the cata-lyst basket msulatlon Neglectmg all other heat transferresistances reduces computational effort wlthoutsubstantially affectmg the accuracy of the sunulationFor converters without insulation, heat transfer to thefeed gas would be conslderable and an overall heattransfer coefficient would be necessary Smce the riser 1sunmsulated, an overall heat transfer coefficient U ISconsidered m calculatrng thus heat term

dQnsc r - l J C (T - T R ) dz = f ,M& (T ,z , PF) dTR(24)where C IS the outslde circumference of the riser Theoverall heat transfer coefficient IS mven by

l/U = l/ho+ l/h,,,+ A,/(;?& In (L&,/D))+ E (25)ho IS calculated for heat transfer mslde packed tubes asdescrtbed by J _eva[lB]

ho=YAW (26)where W I S he mass flow m lb/hr The geometllc factorIS calculated from

# = 3 5 etA4 6~01(Do/A) /D (27)where D, 1s the catalyst diameter m feet and D I S hebed diameterThe bed heat transfer factor A IS a function of reacttongas thermal conductlvtty K and gas viscosity and is givenby

A= ICI/LO7 (28)The Colburn equation IS used to calculate the heat

transfer coefficient mslde the nserh ,,, = 0 023 GC,, (Tn . PF)N ; ;213 2 DJ Do (29)

G I S the mass velocity lb mole/hr ftZ and the Prandtl


6/14

42 L D GAINESnumber 14 computed

&r= c , ( T R. PF ) d K (30)D, and Da a r e the inside and outslde diameters of theriser respectively m feet and the Reynolds number 1scomputed

NR C = 0, Glcl (31)The thermal conductlvlty of the riser material IS k and lthe fouling factor, was assumed to have a value of 0 001

The energy balance for the quench zones above eachbed IS developed assummg the zone to be adlabatlc andperfectly mlxed Equatmg the energy flow from theprevious bed (Fig 3), (or the riser m the case of the firstzone) and the energy m the quench stream to the energyflow mto the next bed yields for the ith quench zone

M 1s the molar flow from the previous bed or riser and HIS the enthalpy, Btu/lb mole The enthalpy of the quenchIS obtamed by multlplymg the fraction of feed used forquench m the ith zone by the total enthalpy of the feedFlow rate of a component out of the zone IS equal to thesum of the flows m smce no reactlon occursThe energy balance for the heat exchanger located atthe bottom of the converter may be obtained from theheat transfer equations The cold gas flowing from theannulus passes upward through the exchanger tubes mtothe riser and reacted gas flows countercurrent and IScooled on the shell side The energy balance for thereacted gas m the shell IS

M&,(T,, Pp) dTs - UA(T, - TT ) dz (33)where A IS the heat transfer area m ft* per ft of heatexchanger height and Ts and TT a r e shell side and tubeside temperatures respectively U, the overall heattransfer coefficient for the heat exchanger IS computedaccordmg to eqn (25) ho for the heat exchanger 1scomputed by the method of Donohue[l9] for dak-and-doughnut baffles

where D, 1s he equivalent diameter m inchesD, = 4(f l ow area/wetted perimeter) (35)

and G, IS computedG, = W/S, (36)

S, IS the geometric mean of the crossflow and thebaffle-hole areas, I eS, = (cross flow area x baffle-hole area)05 (37)

The energy balance for the gas m the heat exchangertubes yields

f ,M& ,(T=, PF)~TT = - UA(Ts - Tr ) dz (38)The mdlvlduat heat transfer coefficient for the tube side,!I,~, 1s computed according to eqn (29)

Equation (33) and (37) may be solved by numerlcalintegration or analytically as suggested by Shah[2] IfC,( Tap ) and C,(T,, P,) a r e assumed to be constantThe solution may be obtamed by letting

andB =~,MFC~(TT.PF) (39)

4 = MFCJTS, P,)Equatmg eqns (33) and (38) and integrating

(40)

Ts = Ts+PI+(TT - TRB) (411where Ts and TRB result from evatuatmg the constant ofintegration Substitution of eqns (37), (40) and (41) into(38) and rearranging

(42)Integration over the length of the exchanger yields afterrearrangmg

TRB = TA(1- p/t@) - Ts(1 e(-P*UA ue)e(--89uA WB- PI* (43)C,(TS, P,) and Cp(TT , P,) ar e evaluated at averageconditions Tp may be obtamed from eqn (41) since atthe bottom of the exchanger T ,- and becomes TA and Tsbecomes Tp

CAUXJIATIONAL PROCRDUREComponent fresh feed rates and compressor speed, S,

must be available to begm compressor calculations Aseach compressor computation (eqns 2-7) 1s performed, acheck IS made to determine whether the compressor ISabove surge flow rate, AS The computation of the laststage of the second compressor requires an estunate ofthe recycled component flows to be madeThe compressor discharge flow IS separated mto con-verter feed and product according to eqns (8) and (9)The separator gas IS split m half and each half entersseparate converters at the same predetermined tempera-tureThe basic converter model eqns (lo), (111, (19), (331,(41) and (43) cannot be solved analytically for componentflows and converter temperatures since they are non-hnear and coupled due to heat transfer between the feedand reacted gas The method of solution consrsts ofassuming a value of TRT and T, IS hen found such that


7/14

Ammonia syntheses loop variables mvestlgated by steady-state sunulahon 43eqn (32) IS sattsfied to within 0 1F on T, At th1.s pomt,values of TR, TN, TA and F exist and an integrationthrough the first bed IS performed usmg a fourth orderRunge-Kutte method The fraction HZ converted, x ISset to zero to start the integration Molar flows at anypomt in the bed may then be found by eqn (11) Thepressure IS assumed to vary linearly with the fractioncatalyst traversed

P = P+AP(B-z)/B (44)where P IS the pressure at the bottom of the bed, AP ISthe pressure drop through the bed, B 1s bed depth and zIS posItIon measured from the top The pressure,temperature, and mole fractions are used to calculate arate for eqn (10) and a value of dx/dz IS obtamed Valuesof dT/dz, dTJ dz and dTA/dz for the element are foundfrom eqns (19), (23) and (24) New values for TA, TR, T,and x result at the end of each mtegratlon step Ten stepsare normally required to obtam desuable accuracy

When the bed mtegratlon IS completed, composltlonsare adjusted according to the value of x and thetemperature and composltlon m the next quench zone IScomputed, which leads to tnihd values for the next bedintegration

This procedure IS repeated for each bed unti the endof the catalyst section IS reached The temperatures ofthe reacted gas and the cold annulus gas are entered mtothe exchanger equations along with estimates of Tn, andTp Heat exchanger calculations are repeated ustng thenew values of TR and Tp until convergence on the twotemperatures ts obtamed

Solution of the heat exchanger equahons completesone pass through the converter To check for con-vergence, the value of TRe from heat exchanger cal-culatlons 1s compared with the value obtamed after m-tegratlon of the last catalyst bed If they differ by morethan 0 25F, a new value of TN IScomputed

TRnkw = T~~~xcm.w + (TRW, - T=mwJ (45)and used to start another pass through the converterFour or five passes are usually reqmred to obtam con-vergence

mxture heat capacihes and enthalples are computedusmg the BW R equahon of state Table 4 contams pure

component constants whch are used to compute mixtureconstants to be used m the BW R equahon Mxtureconstants are computed according to Reid andSherwood[20] except for Bo, which IS computed

n nBc,, = 2 2 m,m,(B:,33+ B: ,333)3 (46),-a j-L

Mxture vlscosltles and thermal conductrvltres are cal-culated by methods suggested by Reid andSherwood[20] Pure component viscoslhes are obtamedusmg the theorehcal treatment of Chapman and EnskogLow pressure mixture vlscoslhes are then computed bythe method of Wtike and corrected for pressure by theequatton proposed by Dean and She1 Pure componentthermal conductlvltles are computed as suggested byBromley and mixture thermal conduchvltles arecomputed accordmg to the method of Lmdsay andBromley, the mixture value IS then corrected for pres-sure by the equations of Steel and Thodos

The fraction of the converter effluent to be purgedmust be specified The remamder of the convertereffluent IS recycled and the component flows arecompared with the estunated values used to start thecycle Strrct convergence on these ilows 1s necessary anda convergence accelerator ts applied on odd loop Itera-tlon cycles This IS done by computmg a new estnnate ofeach recycle component flow, F,, from the componentflows of the previous two cycles, F: and F;

F, = F: +E(F:-FynThe convergence cntenon 1s (47)

Gwand must be sunultaneously sahsfied for all componentflows

At the same tune, synthesis loop pressure drop mustsum to zero with the pressure nse across the last stage ofthe second compressor Pressure drop, AP, is assumed tobe proportlonal to the square of the flow mto the con-verter and mversely proportional to the density. p, of thestream

AP = G 2 (FiMwi)*/pt-1Table 4 BW R constants

Nitroaen Ax-non Methane Amnda mmnenA .03l2319 .02sS3% .0494 .10354l29 .ooo8108534A0 .8720e6 .=3.U7 1.855 3.784282 .18267058b .0032351 .002l5289 .OQ338004 .ooo7195852 .01092708% .0281066 -0222852597 .0426 .05l646121 .026%120974C X 10 &X547364 .0007982437 .a2545 .ooO15753298 .ooooo355307co X 10 6 .0078X375 .Ol314J25 -02257 .17J35709 .oooo99967II0 .OUDO709232 .oooo3558895 .oool24359 SWXCU652178 .oooo893277Y .0045 .00233827ll ,006 .019805156 .(X2576974


8/14

44 LDwhere G was determmed from operating data The con-verter pressure drop for the plant data was API3 with l/6of the drop occurrmg between the converter outlet andthe compressor suction The pressure drop computed byeqn (49) IS assumed to remam dtstrtbuted m this mannerto provide mlet and outlet pressures for converter cal-culatlons Loop p ressure drop ts determined by flow rateand must be matched by pressure rise across thecompressor which 1s a function of flow and compressorspeedThe compressor speed 1s adlusted each loop in aneffort to match pressures The new speed, S, 1s computedfrom the previously computed speed, S as follows

S = S - ((Pd - Ps) - AP)Cm, (50)where Cl0 was determmed by trtai Equation (50) pro-vldes an excellent esttmate of the speed for the nextcalculattonal cycle Convergence on pressure drop (atm)IS obtamed when eqn (51) 1s satisfied

ols ( (51)

DISCU~EON OF -TsThe synthesis loop model equations developed in theprevious sectlons were solved on an Interdata 7132 to

mvestlgate numerous modes of steady-state operationConvergence was dependent upon operatmg cond ttlonsan d mitml estimates of component recycle rates Thutyto sucty seconds of CPU tune was normally requtred forruns whtch converged to a steady-state soluttonThe prevtous study used the model presented m thiswork to determme effects of specific process vartablesupon converter operatmg effictency and stabthty In thatwork, a sunple method of determrmng near optlmd con-verter temperature was presented Tlus new method hassmce been unplemented on the converter modeled resul-tmg m an Increase m converslon and conformation of the

GAINES

model This paper 1s Intended to show the mteractlon ofthe converter with other components of the synthesisloop and variables investigated Include feed rate, mertslevel, methane leakage, separator efficiency, convertertemperatures and hydrogen-nitrogen ratlo The combmedeffects of several of these vanables are also included andconstramts are dlscussedA typlcal sunulatlon result 1s shown m Table 5. wherethe feed rate, combmed converter flow rate, recycle flowand a complete temperature profile are shown Forty-eight such runs were made to Investigate variableseffects upon ammom a production and synthesis loopefficiency The results of these runs are shown in Rgs4-17 and are dlscussed m the followmg text

Figure 4 mdlcates that ammonia productlon andsynthesis com pressor speed are almost hnear functionsof feed rate at a constant merts level and H/N ratio tnthe synthesis loop Figure 5 shows that recuculatlon rateand compressor dtscharge pressure also Increase withfeed rate The synthesis compressors are major utilityconsumers m the process, compresston of the feed gas tosynthesis loop pressure accounts for about 70% of thesynthesis compressors power requtrements whrle recur-culation accounts for the remamder A s expected, m-creased feed and product rates necessitate an increase mpower input, as mdtcated by compressor speed

The synthesis compressors are often the bottleneck ofthe synthesis loop and there are several related con -stramts which may become active as feed rate IS m-creased As already dacussed, compressor speed,horsepower requvements and dtscharge pressure all m-crease with feed rate and may become hmltmgCompressors are destgned to operate wtth a hmltedspeed range, usually 70-105% of design speed Operationoutside of thts range may result tn compressor damageand speed controls are deslgned so that the maximumrated speed cannot be exceeded

Synthesis gas compressors are normally &ven by oneor more steam turbmes and as horsepower reqmrementsIncrease, a restncted steam supply may hnnt compressor

Table 5 Typtcal model resultsConverter Temperat.uFe Profile (OF) Process Flow Rates (++=,Feed 3co.cORiser Outlet 751.14Bed 1 Inlet 751.14Bed 1 Outlet 875.03Eed2Inlet 816.82Bed 2 outlet 907.09&d 3 Inlet 832.69Bed 3 Cutlet 897.36BedLInlet 832.00

Feed 6W.33H2 4793.80N2 1602.10AR 19.19wa 25.25Converter Feed 35791.o9H2 22383.22N7 7695.94Bed 4 OutletConverter Outlet 883.29 G3 818.U566.28 2161.96

ccmpressor TimHorsepowerDischarge Pressure (atm)Recycle Pressure (atm)fk NH 3 into cowetir$ NH3 out of ConverterEQ/N2 RatioFraction Feed HeatedQuench Fractxon Bed 1Ciuench F'rectxon Bed 2&ench Fracticn Bed 3Quench Fractum Bed 4

10197.90lO826.00m?2:29Il.982.910.700.000.080.l.l0.11

a4 2731.85R.xxcle 32431.80

63ARr&ed Purged 2710.00k Inerts III Converter 4.06Feed 13.47

NH3 Produced 630.53 Tcm&'w


9/14

650

640

6302ts' 620E3Ii 6102=600

590

60

Ammoma synthesis loop variables mvestlgated by steady-state stmulatlon 45- fO.400

- 1OW

E

;-lO.200*8E9- 10.100

l3g 4 Production and compressor speed vs feed rate at constant l3g 5 Recirculauon rate and &charge pressure vs feed rate atmert level constants merts level

operation If this occurs, compressor suction pressurewill rise reqmrmg that part of the feed be vented Thecompressor discharge pressure may also become ab-normally high as feed rate IS increased Since thesynthesis loop IS designed with relief valves, this pres-sure must be mamtamed below some predetermmedlevel

The maximum feed rate may also be determined byreclrculatlon rate Recuculation rate affects heat ex-change and refrtgerabon requuements and converterstability At high feed rates, coolmg capacity may be

148

146

I44

I42 E6_IWO r3t5i.3601

362

34

32

insufficient or the converter may lose reactlon as HrllI bediscussed laterThe effects of feed rate upon each of the variablesdiscussed above are shown m Table 6 This table shows

the process vanables which were held at constantvalues and the dlrectlon of change of other pertmentvmables Vmables which may reach constramts are alsomdlcatedThe slmulatlon results shown in Fus 4 and 5 reflectfeed rate vanatlons at one specific merts level Figure 6indicates that pressure and recuculation rate also m-

Table 6 Effect of variable Increase on other process varrables

Vanable mcreased PI-C- com- Fkc1rcu- Pressure Horse- Temp.duct1on pressor 1ataon power 9% ,$

IIErtSLevelSpeed Ftate Ratio

Feed rate t -t 'T 'T -7-- - - t -

Inerts level t-T-f-I--f--- -LCf--

Separator efficxency -11 il--------

Convertertemp. I-I-T-I-T-

%/N2 ratio

Nitrogen feed rate t7-t

w-7-17-+-l-T_(_t I1 I ta Ccmverter w-ill become unstable and temperature cannot be held constant at high inerteb IZcreases tdl ratio of 2.2, then increases_ -.-c Decreases tiLl ratio Of about 2-q 18 reachedVariable held constant or did not changeVanable increased/decreasedVariable mcreasedandmay becomelznutang


10/14

L D GAINEZS

32poo

- 144- 143 E

ij_- 142 f

tk-I41 s

$- 140 s

- 139

L I I I _ 113613 14 I5 16 17 ,ePercmt herts an converter feed

Fig 6 Reclrculatton rate and discharge pressure vs merts levelat constant feed ratecrease wtth tnerts concentratton at a constant feed rateChanges m recuculatlon rate are partially due to per-passconverston as shown tn Ftg 7 A t tnerts concentrattonsabove 19%. the converston ts so low that the converterbecomes unstable, t e heat produced IS not sufficient tomamtam the reaction zone at an elevated temperatureThe merts concentration IS controlled by ventmg p art ofthe recycle gas as shown m Fig 1 Since reactants arealso vented tn the purge stream, a loss of productronresults from operation at low tnerts levels A high mertslevel affects the synthests loop tn much the same way ashigh feed rates and the same synthests loop constramtsmay be encountered as mdlcated in Table 6 Stnce thesynthesis loop 1s normally operated agamst one of thementloned constraints, a balance between feed and purgerates must be mamtam ed For example, d the synthesisloop pressure constraint ts active and feed rate ts tn-creased, there must also be a corresponding tncrease mpurge rateFigure 8 shows that methane leakage has a detrunentaleffect on productton at a constant merts concentration Arise in methane concentration tn the feed of 0 06% resul-ted m a loss of production of 0 32% (ZTPD ) Substantiallosses tn production may result from meficlent reformeroperation which leads to increased methane leakage

Converter temperature profile affects converterefficiency and stability Figure 9 shows a reduction mreclrculatlon rate and compressor speed as the outlettemperature of the catalyst section 1s reduced Smcecatalyst actlvtty also rapidly detertorates at Hughtemperatures, tt IS desuable to mamtam low convertertemperatures Lower temperatures, how ever, decreasethe reaction rate and tend to lead to stablltty problemsFigure 10 tndlcates that synthests pressure also decreaseswith converter temperature Except for stabllrty relatedproblems, lower temperatures wh tend to movesynthesis loop operation away from compressor, pres-sure, and flow related constramts and increase synthesisloop capacity as shown m Table 6

Fig 7 ProductIon and cowerston vs merts level at constant feedrate

62Y80-

6260 -

23M*,ho:, &~5mo*26 27 2e

Fig 8 Production as a function of methane in feed

Figure 11 shows the effect of ammonia separatorefficiency Runs were made with 2 0, 2 3 and 2 6% am-moma tn the converter feed Faure 12 shows thatsynthesis loop pressure and reclrculatlon rate increaseswith ammon ia concentration tn the converter feedHigher ammonta concentrations also tend to reduceconverter stab&y, as discussed by Shah[2]Converter efficiency and synthests loop operation areknown to be affected by the hydrogen nitrogen ratio tnthe synthests loop A value somewhat lower than thestotchometrlc ratio of 3 ts usually considered to beoptunal Since the molecular wetght of the synthesis gaschanges with H,/N, ratio, there are several factors whichmust be considered Fust, the pressure drop m thesynthesis loop WIIIbe altered as shown by eqn (49) and atthe same tune, the molecular wetght change will affectthe compression ratio as predicted by eqn (7)Component activities wrll change with HZ/N2 ratio and


11/14

Ammoma synthesis loop variables mvestlgated by steady state slmulatlon32.600

32.400

32,200

3l@oO -&yy!&J.lo,1208-mTemperature of gas leaving cotalyd section. *F

Fig 9 Recmxlatlon rate and compressor speed vs temperature Fig 1 Reclrculatlon rate and discharge pressure vs separatorat constant feed rate efficiency

I4

I3

E I36fIg,3s5B 13

I3

OF

59 -

a-

7-

6-

5-85l

Temperature of gas Ieavlng catalyst sectlm.FFig 10 I)lscharge pressure and conversIon vs temperature atconstant feed rateexcept at composltlons very close to eqmhbnum. thereaction rate IS mcreased at ratios shghtly below stol-chlometnc

In order to determme the results of changmg theHz/N2 ratlo, several sunulatlon runs were made at aconstant total feed rate and constant merts level m thesynthesis loop The merts composlhon m the feed wasvaried m relation to the concentration of reactantsFigure 12 shows the effects of Hz/N2 ratlo on horse-power requirements and hydrogen consumption at threelevels of merts and two temperature levels m the con-verter Each horsepower curve represents a constantammoma production rate Horsepower requuementdecreases W&I mcreasmg HJ N2 ratio und a ratio ofabout 3 IS reached At this pomt, horsepower IS nearlyconstant with further Hz/N2 ratio changes Hydrogen

32.500

32,300-0Ei0 I_

f 32.00

bl-

,-

I20 23 26Percent NH, I converter feed

47I41

140

Effwa

ghf0

36

37

4.m

4;160

YI!iP I

4.750_

$f 4,740

I4=0

4.720

,

I-

/

I5

if\

- 12.ooo

- ll,5Qo

20 25 30 35 40y IN, ratioRg 12 Hydrogen feed and power Input at &fferent Hz/N2 ratiosand constant product ratesconsumption goes up with Hz/N2 ratlo over the rangeinvestigated This result IS expected since the hydrogenconcentration m the purge increases with Hz/N2 ratioDepending on the relative costs of power and hydrogen,an optional Hz/N2 ratio below 3 ~111 exist m each caseAs can be seen from Rg 12, operation above a ratio of 3requues an elevated hydrogen feed rate with no reduc-tion of power reqtured by the synthesis compressors

Synthesis compressor discharge pressure curves (FigI3), have the same general shape as the horsepowercurves Higher molecular weights of the gas at low ratiosresult m a larger pressure rise across the synthesiscompressors Since the feed rate Is constant and most ofthe power 1s required to compress the feed gas, horse-power IS greatly influenced by discharge pressureReclrculatlon rate affects horsepower and the curves


12/14

L D GAINES

,6295TPD1645% harts

6248TPD136D% lmrts

I I 1 420 25 30 35 49H2/N2 Rat10

>t L I I I

IS 20 25 30 35 40H/N, r-doRg 13 lhscharge pressure at dlfferenl HJ N, ratios and constant

product ratesAg 15 Compressor speed at different HZ/N2 ratios and constant

product ratesshown m Fig 14 have a somewhat different shape and 15 reveals that compressor speed and reclrculatlonRecuculatlon rate IS a functron of per pass converslon m rate curves are reI ated and compressor speed exhibits athe converter and as already mentloned, component muumum at a Hz/N2 ratio of around 2 2 Thus indicatesactlvltles result m higher con* erslons at low Hz/N2 that d mexpenslve compressor driver power 1s avadable,ratios, although conversion WIII also mcrease with a reduced Hz/N2 ratlo wdl not only reduce hydrogensynthesis pressure consumption but will also decrease compressor speedCompressor speed IS another unportant variable whichIS affected by Hz/N2 ratio, flow, and compression ratioand results are shown m Fig 15 Comparison of Figs 14

Figures 12-15 also illustrate the importance ofcontroJ hng merts level and converter temperatures Ap-proximately 5 tons/day increase m production resultsfrom operation at an merts level of 16 45% rather than

6295 TPD16 45% Inats

6248 TPD1360% lna?s

20 2.5 3D 35 40

Fig 14 Recuculauon rate at different Hz/N2 ratios and constantproduct rates

6295 TP DI6 45% lnerts

6248 TPD/ 1360% lnerts

625.9 TPD

/1395% Inerts

I+ m,-3530

E

I

Fig 16 Production vs merts level at dtierent Hz/N2 ratros


13/14

Ammoma synthests loop vanables mvestlgated by steady-state skmulauon 49

-1 I, 36 H/N, Ratlo2, 303. 26E 4, 22c mo - 5. 17eiii2Z8@5 BID-Bfbt2g --EI-

Fig 17 Temperature vs merts level at different H2/Nz atios13 6% Figure 12 also shows that an increase of 1 ton/daymay be achteved at an merts level of 13 95% and tn thiscase, lower converter temperatures also resulted m areduction m power requvements The synthesis loopvmables discussed are shown to be of sizable economtcconsequence and illustrate the need for a basic under-standmg of process vmable effects on the processSlmulatlon runs wtth a constant hydrogen feed rategive further Insight into synthesis loop operation If theprocess 1s feed limited These runs were made by chang-mg the mtrogen and argon feed rates and purge rateFigure 16 shows that production mcreases with nttrogenfeed rate (reduced Hz/N2 ratlo) at a constant hydrogen feedrate and merts level In these runs, no attempt was made tohold converter temperatures constant and Fig 17 showsthat the converter may become unstable d large changesare made m synthesis loop variables wlthout changmgquench flows Curve 1 shows that operation at high Hz/N2ratios and merts levels ts not possible with the quench flowsused Curve 5 shows that very low HJ N2 ratios also tend toreduce converter stabrllty Frgure 17 also shows that highertnerts levels tend to reduce converter temperatures andconsequently reduce converter stab&y

CONCLUSIONSSunulatlon results should provtde a better understand-mg of process variable effects upon synthesis loop

operation and a basis for synthesis loop optmuzatlonLarge economrc benefits can be denved from goodoperatmg strategyDecreased purge rate (tncreased merts level) wdl rasethe production rate since a smaller quantity of reactants1s lost A hmltatlon to thus operatmg mode ts that Itresults m increased compressor reqmrements, synthesisloop pressure and recuculatron rate, and consequently aCES Vo l 3 4 N o I-D

process constramt may be reached Methane leakageleads to increased purge rates and lost productronDecreased separator efficiency results m lugheroperating costs and reduced converter stability Lower

converter temperatures also result m reduced stabilitybut increased synthesis loop efficiencyThere appears to be an optnnal HZ/N2 ratto at whichthe synthesis loop operates most efficiently Operationbelow this ratio results tn increased utdtty consumptionwhile operation above requires more hydrogen feed Theophmal ratio IS prunarlly dictated by the cost ofcompressor drtver power

AA,A'

AS;B,Bo,Bornb,C

Cl-C7CC,

(CP)D90.DODPE

E, FL l -7F ,f .GG,

wA.HR

hoh 01,K,K,kk

kzL

L f1

MM lMFMP

Mwm

NOTATIONcross-sectlonal area of catalyst bedconstants in eqn (14)heat transfer area per foot of exchangersurge flowactivity of component icatalyst bed depthconstants in eqn (14)BW R constant for component 1BW FZ mixture constantconstants in eqn (17)circumference of riserconstants used in compressor calculationscircumference of catalyst sectionconstants tn eqn (14)nuxture heat capacitybed diametermslde tube diameterequivalent diameteroutside tube diametercatalyst pticle diameterconvergence accelerator for component 1molar flow of component Imolar flow of component I from separatorfraction inlet flowmass velocityequivalent mass flow ratemolar enthalpy of mixtureheat of reactionoutside heat transfer coefficienttube side heat transfer coefficient1 refers to HZ, 2 refers to NZ, 3 refers to NH X,

4 refers to A, 5 refers to CH 4eqmhbnum constant m terms of acttvltleshquld-vapor eqmhbrmm constantthermal conductivtty of heat transfer surfacethermal conducttvlty of catalyst basket m-

sulattonrate constant for decomposltlon of ammonialength of heat exchangerfraction hqutdsubscnpt refers to converter zonemolar flow ratemolar flow rate in zone 1molar flow rate of feedmolar flow rate out of convertermolecular we&t of component 1mole fraction of component I


14/14

50

NRNRC

PP

APPdPFPPPSQR

s, BSe

S,TT IT *T FTPTR

TRBTSTTuW

xZ

L D GAINESPrandtl numberReynolds numberpressurepressure at bottom of bedpressure dlfferentml across bedcompressor discharge pressur epressure of feedpressure out of convertercompressor suction pressureheat termgas constantreaction ratecompressor speedgeometr ic mean of the crossflow and the

baffle-hole areasterm computed from eqn (15)temperaturetemperatu re m of I th zonetemperatu re in annulustemperature of feedtemperatu re out of convertertemperatu re in ri sertemperatu re at bottom of ri sertemperature on shell side of exchangertemperatu re on 2 J e side of exchangeroverall heat transfer coefficientmoss flow ratefraction H z convertedcompresslbrllty factordistance from top of zone

G r e e k sy m b o l sa kinetic parametersB defined by eqn (39)Y activity coefficientE fouling factor7 conversl on based on mtr ogen

1:;[3 1I41

(51

van H eerden C , In d Ena n a Chem 1953 45 1242-Shah M J , Ind Engng Chem 1%7 59( 1) 73Baddour R F, Bnan P L T , Loneats B A and Evmerv JP , C h em Engng Sci 1965 20 281- _ -Slack A V , Aiigood N Y and Maune H E , Chem EngngPrvg 1953 49 393Nielsen A, An hoestrgatron on Promoted I r on Ca ta l ysts o rt h e Synrhesrs o f Ammon ra , 3rd Edn Jul G]eilerups,Copenhagen 1968161Annable D , Chem Engng SCI 1952 1 145171 Kubec J . Burtanova J . Bunanec Z Int Chem EnPnn 1974

I8 1[9 114(4) 629 I I

Hay J J and Pailar I M , Br C h em Engng 1%3 S(3) 171Kiaer J , Calculatron of Ammonra Converters on an Elec-tromc Dtgual Computer Akademtsk Forlag, Copenhage n1% 3

thermal conductivi ty of gas mixtur eviscosity of gas mixtur ecatalyst activitygas densityeffectiveness factorbed heat transfer factordefined by eqn (40)geometnc factorthickness of msulatlon

_CES

ii01 Games L D . I&EC Proc Des Dev 1977 16 381riijii21n3 1[I41I151i i 6 3[I71[181r191I201

Sultan R F , Br Chem Engng Equtp Sup I%8 81Whrte M H , Chem Engng 1972 79(24) 54Guerreu Cl , Hydrocarbon Proc 1970 49(12) 74Temkm M , J Phys Chem (USSR) 1950 24 1312Gtiiespre L J and BeattIe J A , Ph ys Rev 1930 36 743Dyson D C and Stmon J M, In d Engng Chem Fundis,1968 7(4) 605

Kazarnovakn Ya S , Zhur Fiz Khrm 1945 19 392Leva M , Wemtratib M , Grummer M and Clark E L I n dEn gng Chem 1948 40(4) 747Donohye D A In d En gng Chem Fu nd ls 1949 41(11) 2499Rerd R C and Sherwood T K , he Proper t r es of Gasesand L l qm ds, 2nd Edn M cGraw-Htll, New York 1966

Documents

Ammonia Synthesis Loops Variables Investigated by Steady-state Simulation