Cl~ernical Cngr,~rrrirrg Sciwzw. Vol. 13, No. 8. pp. lXOlL1806. 19X8. Printed in Great Bntain.

m9-2sn9/xx $3.oo+o.w Pergamon Press plc

Simulation of Steam Reformers for Methane

M.A. Soliman, S.S.E.H. El-Nashaie*. AS. Al-Ubaid and A. Adris

Chemical Reaction Engineering Group (CREG), Chemical Engineering Department King Saud University, P-0. Box 800, Riyadh 11421, Saudi Arabia

*Author to whom correspondence should be addressed.

Abstract A model is developed for industrial steam reformers for both-top fired and side fired furnaces. The catalyst tube model is a one- dimensional heterogeneous model with intra- particle diffusional resistances. The two point boundary value differential equations of the catalyst pellets are solved using a modi- fied novel orthogonal collocation technique to obtain the effectiveness factor variation along the lenkth of the reactor. The side fired furnace equations are algebraic equat- ions, the top fired furnace equations are two- point boundary value differential equations which are solved using the orthogonal colloca- tion technique. A recently developed more general rate expression is used. The model performance is checked against industrial steam reformers. The model is used to invest- igate the effect of various parameters on the behaviour of the catalyst tubes and the fur- nace. The effectiveness factor variation along the length of the catalyst tube is also analysed.

Keywords: steam Reforming. Reactor modeling, Digital Simulation, effectiveness factor

Introduction Most sy*gas used today is produced by steam reforming. In Saudi Arabia and Egypt this percenta&e is almost 100%. and therefore economic comparison -with other methods is not available, however based on a plant starting up in the U.S.A. in 1987 steam reforming Is the most economic choice (Gaff and Wan& 1987. Table 1). Lfficient steam reformers need to be designed on kinetic basis rather than thermodynamic equilibrium basis, (Rose 1977), and therefore kinetic SXp~C2SSiOn and diffusion-reaction models which are reliable over a wide range of parameters are needed. Extensive work dealt with the kinetics of steam reforming of methane over nickel supported catalyst, and some of the review papers offer a good survey of the subject (Rostrup--Nielsen, 1984, Elnashaie et. al.. 1988). Some investigators obtained negative effective order for the reaction with respect to steam (e.g. Bodrov, 1964, Al-Ubaid et. al.. 1987). while others obtained positive effective order (e-g. De- Deken et. al.. 1982). The more general rate SXpt-SSsiOn which shows this non-monotonic dependence of the rate of reaction upon the steam partial preseure obtained recently by Froment and co-workers (Xu and Froment 1988) is used in this investigation.

The performance of steam reformers is strongly affected by the heat transfer from the furnace to the catalyst in the tubes (Hyman, 1968). The developed model for the whole steam refor- mer including both the catalyst tubes and the furnace is checked successfully against a number of industrial reformers, with both top and side fired furnaces in the open literature (Singh and Saraf 1979) and in the ammonia and methonal industries in Egypt and Saudi Arabia. The model was then used in a detailed para- metric study that elucidates the COmplSX

interactions between the different processes taking place in the catalyst tubes, and the complex interaction between the catalyst tube and the furnace. Catalyst Tube Model Development: Reaction Kinetics Used: The kinetic rate expression used in this work is the one developed by (Xu and Froment 1988) with the following equations: cn4 + Hz0 co + 3n2 (I), co + H-JO cop + Hp (II), CH,, + 2H20 co2 + 4H2 (III> based upon adsorption-desorption mechanism model consisting of 13 steps, three of them are rate controlling and thus arrived at the following kinetic rate equations:

PCH4 0.5

=1 = kl ( . PH20 PH2 - PC0

P2.5 - Kl )/DEN*,

H2

PC0 r2 = k2 (

* PH20 PC0 - ~)/DEN~

pH2 K2

po.5 * PC0

r3 = kg ( PCH4 . Pi,, Hz

P3.5 - Kl . K2 2)/DEN.2 (2)

H2

(1)

where, DEN = ' + kCOPCO + kU2 'H2 -+ 'CH4 'CH4

+ kH20 PH20/PH2 and rl, r2 and r3 are the

rates of reactions I. II and III respectively. The rate of disappearance and appearance (formation) of CHq and CO, are given by: rCH4 =rl+r3, rco2 = r2 + r3 (3)

For this reaction system the number of moles of the different components can be expressed in terms of number of moles in feed and the

CH4 and CO2 conversions, which are defined as:

XCH4 = (FCH4 - nCH4)jFCH4 ' XC02

= (n CO2 - FC02)'FCU4 (4)

1801

1802 M. A. SOLIMAN et al. A2

Detailed analysis of the non-monotonic behavi- our of these rate equations is given elsewhere (Elnashaie et al., 1988). Also all kinetics and equilibrium constants and their dependence upon temperature are given elsewhere (Xu and Froment. 19881. The Mathematl&l Model: Mass, Energy and Momentum Balances: Single reformer tube perf0rmsIIce is assumed to be representative of snY other tube in the furnace. The material, energy and nxxnentum balance equations can be written as:

dXcH 4=A.p

dll B ' 'GE4 - rCH4/FCH4 '

d%u 2_ 7 - -A - PB - nco

2 - fC02/FCH4

[PB AH1 "m 4 (rL+r3) + PB

(5)

AH2 nC02(r2+r3) + 4 & (Ts-T)] (6)

heat transfer coefficient U (Kcal/m2.hr.0K) is calculated by the relation (Xu and Froment, 1988):

(7)

where a'i (Kcal/m.hr.°K) is the convective heat transfer coefficient in the packed bed and is obtained from the correlation of (Leva et. al., 1951). The differential pressure drop Is given by the equation of (Fanning, 1954) and the friction factor is computed using (Hicks 1970) correla- tion together with the modification of (Singh and Saraf 1979). Catalyst Partifles Equations: The catalyst pellets are assumed to be slabs with a char&teristic length, a,. Material balance equations take the following form:

d2% B Q2

- = ai 10 -5

dw2 RT-rf, ai'l,

Di,e for i E en4 and -1 for i R CO2 (8)

with the boundary conditions at w = 0

d pi 8 -=o ; w-1 .+ti Fi,s = Fi _-

The slab is assumed isothermal and external msss and heat transfer resistances are negli- gible. All physico-chemical parameters formu- lae are taken from (Reid and Sherwod 1966) and (Froment and Bischoff 1979). Also all hydrocarbons higher than methane in feed are assumed to instantaneously crack into CH,,, H2, CO3 and Co. Radiative Heat Transfer Two tvnes ot turnaces firinp: are considered: side gired furnaces and top-fired furnaces. - s- Modeling of side fired furnaces: Side fired turnaces are uf3UGZill y modeled well- stirred enclosure having a mean tema&rature which is different from the exit temperature. The radiative heat transfer equation takes the form (Singh and Saraf, 1979):

1 (p 't>R

1 (L-l)+ a A,

P Et (' 't)R,black

(G Nt)R.black = 'g (a Ac + P

AR 1 + c,/l(l-~~) FR 1 ) , Eg - TV T; s

t

Et 4

- a Tt,o

(G StjR is the total exchange eree between

flue gas and reformer tube, and (c '&)R black is that for a black tube surface, A, = ' n . dc . L. P a is a factor accounting for the part of the radiation that falls on the cold plane and which is observed by the tubes. FR -aA /

(AR+aAc) P

where 0 c AR < 0.5a 5 =P

=P- Assuming 2% of the net heat release Qn lost to the surroundings and that the heat transfer by convection is negligible, the following heat balance equation csn be written as shown in equation (11) where QG is the sensible heat of the exit flue gas:

0 fL QR de = 0.98 Qn - e; (11)

b- Modeling ~of top fired furnaces: ROesler model as modified by (Filla 1984) takes the form:

d2 EL - - al [6(E1 - df2.

Q 11 T;) + Et s,(EL -

Q II, T:,& + Er *,(<l - ‘4’) El - 4~ Ez) 1 (12) d2 E2 - - u2["t dL2

st(E2 - ~(1 - $) T:,,,) + Er sr

(‘4 E2 - (1 - 9) El>1 (13)

1 dE1 With the boundary conditions at * = 0, --

1 dE a1 df

---2 a2 dlL

- & [(I - Q) El - 'IJ E2)1

atn-L,-1~=--=- dE1 1 dE2 er df

[(l - +) El - ;&

a2 de 2-s=

The differential heat balance on the flue gas stream will take the form

Gg %

d[Tg - F(T* - Tg,J]

df -28 [E1-WEgl

(14) where T* is the adiabatic flame temperature, and F is the fraction of the fuel burned along the reformer the flue gas Inlet temperature T g o is given by

,

Tg,o _ [El.0]1/4

00 (15)

and the heat transfer by radiation to the reformer tubes is given by:

QR - 2 Et st(Er - Et) V where E, = CJ T: and V

is the free volume of the reformer. Numerical Solution: Differential equations (5,6) are solved using 4th order Runge-Kutta subroutine with automatic step size to ensure accuracy of the solution. The furnace equat- ion for side fired becomes a single non-linear algebraic equation in the flue gas temperature which is solved by an IMSL (International Mathematical and Statistical Library) subrout- ine celled %SPOW based on a discretized Newton method. The differential equations describing the heat transfer in the top fired furnace are discre- tized by the orthogonal collocation method using a cubic polynomial leading to 12 simul- taneous non-linear algebraic equations which are solved by ZSPOW. At every location along the reactor the effectiveness factors have tc~ be calculated and for this purpose a modiffed collocation method is used. Using this method there Is no need to solve four equations

A2 Simulation of steam reformers for methane 1803

simultaneously as in the standard two point collocation method. In the present method two equations are solved at the first colloca- tion point followed by the solution of two equations in the second collocation point. Details of the numerical algorithm IS given elsewhere, (Elnashaie et al 1987b). Numerical experience with this approach indicated a saving of 314 the time required to solve the particle equations using a three-splines accelerated collocation method developed by (Xu and Froment 1988) with a comparable accu- racy. The overall computer program is equipp- ed with the facility to compute the change of physical properties along the length of the the catalyst tube and for the change in feed compositions. Results and Discussion: The mDde1 developed in the previous parts is characterized by including the essential phenomena taking place in the catalytic Steam reforming tubes. Model Verifications: The developed model has been checked anainst a number of industrial reactors with both side fired and top fired furnaces in EgyPt. Saudi Arabia and in the open literature. A sample of the results is given in Tables 1, 2 for both side fired as well as top fired furnaces. The agreement is obviously quite Eood. Effect of -Operaring and Design Variables on the Performance of Steam Reformers: A detailed parametric study using this more general and industrially verified model has been carried out (Elnashaie et al 1987b). A sample of the parametric study performed is presented in this paper. 1. Effect of molar steam to methane ratio: Table 3 shows the results for different values of molar steam to methane ratio which are chosen around the usual industrial values. It is shown that the methane conversion decreases slightly with the increase In steam to methane ratio, this decrease in the conversion is not due to a decrease in the rate of reaction because the system is in the positive reaction order region with respect to steam where the rate of reaction increases with the increase in steam partial pressure, in fact it is due to the decrease in residence time with in- crease in steam flow rate. The conversion to CO2 increases due to the increase in the shift reaction and the methane conversion to CO2 with the increase in steam partial pressure. The increase in methane conversion decreases the temperatures beC&Use the reactions are endothermic. For the same methane conversion the increase of CO2 on the expense of CO causes an increase in the temperature because the CO2 reaction is less endothermic than that for Co. Obviously the increase of the heat in the exit (due to higher flowrates) decreases the temperature. In Table 3, although the methane conversion decreases and 002 conver- sion increases with the increase in steam feed, the temperature decreases because of the heat increase at exit with the excess steam. Fig. 1 shows the efffectiveness factor profile along the length of the catalyst tubes. The behaviour of n depends upon the balance between different processes taking place inside the catalyst pellet. The increase in the intrinsic rate of reaction tends to cause n to decrease, while the increase in the diffusion coefficients tends to CBUQO 17 to

increase and the increase in the equilibrium constants tend to cause n to increase. The final behaviour of n depends upon the

interaction between these factors along the length of the tube. The n in Fig. lshows a complex non-monotonic be Y& viour, uhile nCH

monotonically decreases. Similar non-monotolf nit behaviour of nCo2 has been observed by

(DeDskan et al 1982) also this phenomenon has been observed for ammonia synthesis by Elnashaie et al (1987a). 2. Effect of catalyst activity: Table 4 shows the effect of decreasing the catalyst activity from 1.0 to 0.02. As a result the conversion decreases by about 24% and the catalyst tube temperatures increase by about 5-6%. while the furnace temperature increases by lo-11%. Such relatively limited changes compared with the very large change in cata- lyst activity is due to the very large diffu- sional limitation and thus the considerable increase in n with the decrease in catalyst activity as shown in Fig. 2. At this low catalyst activity the non-monotonic behaviour of n is not confined to also to nCH,. (Fig. 2).

nC02 but it extends

3. The effect of catalyst particle size: Fig. 2 shows the effect of catalyst particle size on n for catalyt activity = 0.02. It is clear that n increases considerably with the decrease in particle size. It is also noticed that as the particle size increases nCo

2' 'CHq ;pLa;aches each other and the sharpness in the

and minima of n profiles decreases. Table 5 shows the effect of pellet size on conversions and temperatures, it is clear that the conversions increases as the pellet sire decreases due to the increase in the n and the temperatures decreases, while the pressure drop increases. 4. Effect of feed temperature: Tbe increase in feed temperature causes considerable increase in the conversion of CR,, as it is clear from the comparison between Table 3 and Table 6. Holowever the increase of CO2 conver- sion is limited in comparison with that of CR, due to the fact that high temperatures favours the production of CO rather than COP. The high feed temperature cases shown in Table 6 correspond to feed conditions of negative reaction order with respect to steam. The implication of the non-monotonic kinetics of methane steam reforming is discussed in some details elsewhere (Elnashaie et al, 1988). It is noticed that in this region the effect of steam to methane ratio is more pronounced than in the positive effective order cases in Table 3 and therefore the conversion increases with the increase in steam to methane ratio despite the decrease in the residence time. Although the rate of reaction in these cases is much higher than that in Table 3. the effectiveness factor is larger. This is a manifestation that the effect of the increase in diffusion coefficients and equilibrium constants on effectiveness factor is stronger than the effect of the rate of reaction.

Acknowledgement: The authors would like to express their deep appreciation to Prof. G. Froment for providing them with his more general kinetic rate equations, and the sub- routine for the physical properties and for very stimulating dicussions in Riyadh and Gent during exchange of visits between the two departments. This work is supported by the King Abdulasiz City for Science and Technology grant no. AR-7-19.

M. A. SOLIMAN et al. A2 1804

Notation

&AC #AR P

a?.+

Cp*Cg

Ui,e

dc.dp

dte,dti

E1*E2

Ey.Et

F,f

reformer tube cross sectional area. cold plane and refractory surface ar as re ractory and tube side surface % L::1

to furnace free volume half ratio heat capacity of process gas and of flue gas [~cal/~g..~g.~]

effective diffusivity of component i

[m'/br] center to center distance between

tubes and catalyst particle diameter [m]

external and internal diameters of the reformer tube [ml grey and clear gas component heat

flux [Kcal/m2.hr] gas and tube emiesive powers (in side fired furnaces) [Kcal/t&.hr]

fraction of fuel burned along the reformer and friction factor

FH2*FCH4sFC0, Fco; FH20S molar flow rate of

H2, CH4, CO, CO2* H20 in feed[Kmol/hr]

FT total molar flow rate [Kmol/hr]

“.G& process gas and flue mass ve ocities t Kg/hr.m2]

AHi enthalpy change of reaction i

[Kcal/mol]

KlSK2 equilibrium constant for reaction I II [bar23 > c-1

klsk3. rate coefficients of reactions I, III

respectively [Kmol.bar/Kgcat.hr]

k2 rate coefficient of reaction II

[Kmol/K&cat.hr.bar]

kCH4.kC0.kt12 adsorption constants for CK4, CO

and H2 [bar-l]

kH20 dissociative adsorption constant for

Ii20 t-1 L.fi SEC reformer tube heated length, coordi-

nate, and characteristic length of pellets [ml

"i number of moles of component i in the

reacting mixture [Kmol/hr]

Fi.Fi.S partial pressure of component i at the

gas bulk and in catalyst particle [bar ] QG’Qr’QR sensible heat of exit flue gas, net

heat release, rate of heat transfer by radiation [Kcal/hr]

R.Re gas constant, Reynold's number [~calj~m~l.~]

r1*r2.r3 rates of reactions I, II and III res-

pectively [Kmol/Kgcat.hr] T.Tg,Ts temperatures of the reacting mixture,

furnace gas and inner tube skin [K] T r.Tg.0 refractory and fuel gas inlet temper-

ature* [K] Ti,o.Tt,i.Tcot outer and inner tube surface

temperature and catalyst temperature [K]

Us.V superficial mass velocity [kg/m*.hr],

free volume of reformer [m31 z dimensionless distance along the

tube, L/L a radiation distribution factor Eg,Er.Et,Es emissivity of the flue gas, ref-

ractory slab and tube wall. And solid void fraction

x g.%t process gas and tube metal thermal

conductivities [Kcal/m.hr.K] Y viscosity of the reacting gas mixture

[Kg/m.hr] oB.og catalyst bed bulk density and process

gas density [Kg/m31 nCH

4 ,nCO effectiveness factor of CHA dis-

2 appearance rate and CO2 formation rate

o.w stefan-Boltsman constant and dimen- sionless coordinate of the catalyst pellet

-

References

Al-Ubaid. A-S.; Elnashaie, S.S.E.H.; Abba- shar, M.E.E., 1987, "The Influence of the Support on the Effective Order of the Steam Reforming Reactions", Methane Conversion SYmP-. April 27-May 1, Auckland, New Zealand. Bodrov, N-M.; Apel'baum, L-0.; Temkin, M-I., 1964, "Kinetics of the Reaction of - Methane with Steam on the Surface of Nickel", Kinet. Catal., 5, 614. DeUeken, J-C.; Uevos, E-F.; Froment, G-F., 1982, "Steam Reforming of Natural Gas". Chemical Reaction Engineering ACS Symposium Series 196 (BOSTON), 1982. Elnashaie, S.S.E.H., Al-Ubaid, A-S., Soli- man, M-A., and Adris. A.M., 1988. "On the Kinetics and Reactor Modelling for the Steam Reforming of Methane - A Review", J. Ensw Sciences, King Saud Univ.. Vol. 14, 2. Elnashaie. S.S.E.H.; Abbashar, M.E.E.; Al- Ubaid, A-S., 1987a. "Simulation and Optimi- zation of An Industrial Ammonia Reactor". submitted for publication 1987. Elnashale, S.S.E.H., Al-Ubaid, A-S.. Adris, A.M., and Soliman, M-A., 1987b, "Parametric Investigation of Industrial Steam Reformers in Egypt and Saudi Arabia", unpublished report. Elnashaie, S.S.E.H., Al-Ubaid, A-S., Adris, A.M., and Soliman, M-A.. 1988, "On the Non- monotonic Behaviour of the Methane steam Reforming Kinetics", submitted for publica- tion. Filla. M., 1984, "An Improved Roesler-type flux method for radiative heat transfer in one-dimensional furnaces", Chem. Eng. Sci., 39, 1. 159-161. Froment. G., and Bischoff, K., 1979, "Chemi- cal Reactor Analysis and Design", Wiley and Sons, 1979. Gaff. S-P.; Wang, S-1.. 1987, "Syngas Pro- duction by Reforming". Chem. Eng. Frog.. 38, 8. 46-53. Hicks, E-E., 1970, -Pressure drop in packed bed of spheres", IEC Fund.. K, 500. Hyman, M-H., 1968, "Simulate Methane Reform- er Reactions", Hydr.Carb.Froc. 49. 131. Leva, M., Winstraub, M., Grummer, M., Poll- chlk, M. and H.H. Starch. "Fluid Flow through Packed and Fluidised Systems". 1951, U.S. Bur. Mines, Bull, 504. Rase, W-F.. 1977, "Chemical Reactor Design for Process Plants", Vol. 2 John Wiley. Reid, R-C., and Sherwood, T-K., 1966, "The Properties of GSSSS and Liquids", McGraw Hill, N.Y.

- Rostrup-Nielsen, J-R., 1984, "Catalytic- Steam Reforming", In Catalysis Sci. 6 Tech., Vol, 5, p- 1, springer Verlag, Berlin.

A2 Simulation of steam reformers for methane 1805

- Sin&h, C.P.P.; Saraf. D.N., 1979, "Simula- tion of Side Fired Steam-Hydrocarbon Reform- ers", Ind. Eng. Chem. Proc. Des. Dev.. 18, 1, l-7.

- Ku. J., and Froment, G-F., 1988, "Methane steam Reforming, Methanation and Water-Gas Shift on A Ni/MgA120,+ Spine1 Catalyst. I- Intrinsic Kinetics, II. Reactor Simulation" (to be published, 1988).

- Xu, J., 1986, "Kinetic Study of Steam Refor- ming and Methanation". Ph.D. Thesis, RIJKSUNIVERSITEIT, GENT, Belgium.

Table 1 Side fired furnace, case of Table 11. in Singh and Saraf (1979)

*Heated length of the reformer tubes = 12 meters, Inside diameter of reformer tube - 0.0935 meter, Outside diameter of reformer tube = 0.1379 meter, Number of tubes = 200. Catalyst shaped and size - Raschiy rings (0.016*0.016/0.006 meter), Bulk density of catalyst = 1262.6 kg/m3

*Process gas (natural gas) Composition (volume dry basis) - 81.5% Cl&. 7% C2hi6, 5.5% C3HB. 4.5% C&H,O, 1.5% co2

*Fuel (natural gas) flow rate 7411.5 std. &/hr, Oxygen volume percent in flue gas 3 2.5%. Number of burners = 584

.Inlet conditions to the catalyst tubes, Process gas flow rate - 13945.5 std m3/hr, Pressure = 34.4 atm. steam flowrate - 3401.515 kmole/hr. Process gas temperature 727.4 OK. Methane equivalent at inlet = = 762.84 Kmole/hr Outlet conditions

Variable Plant Calcula- ted

Steam flow rate Kmol/hr 2453.3 Methane flow rate Kmol/hr 253.9 Process gas temperature, OK 1038.55 Process gas pressure, bar 32.0 Process gas composition (X volume dry basis) 10.7

C% 10.17 co 11.4 co2 67.72 H2

2493.6 255.49

1042.3 32.02

10.75 9.95

11.45 67.85

Table 2 Top-fired reformer, Abou-Qir plant data

*Total tube length = 12 m, heated length = 10 m, Tube inner diameter = 0.1 m, tube wall thickness = 0.015, Number of tubes = 280, Catalyst used (ICI 5713 nickel oxide)

*Process gas (Natural gas) composition (volume dry basis), 95.3% CH,+, 2.8% C2H6, 0.91% CJHe, 0.07% n-butane. 0.045% I-butane, 0.63 CO2, 0.0086% I-pentane, 0.18% N2, and 0.016% hexane

oFue1 used (Natural gas) flow rate 8700 m3/h, at temperature 343-K, Pressure 4 bar (same composition as process gas), Combustion air flow rate 11.7*104 m3/hr, at pressure 40 bar. temperature 583°K

*Number of burners = 120 forced draught burners arranged in the reformer ceiling in 6.7 rows

eInlet conditions, 26000 m3/b,

Process gas flow rate - temperature 723 K. pressure 36.5

bar, Steam flow rate 5278 Kmol/hr, temperature 723 K. Pressure 36.5 bar, Methane equivalent at inlet - 1196 Kmol/hr Outlet conditions

Plant Calcula- ted

Steam flow rate Kmol/hr Methane flow rate Kmol/hr Process gas temperature, OK Process gas pressure, bar Process gas composition (X volume dry basis)

CH4 co CO2 H2 N2

4062.37 4076.72 364.96 371.322

1056 1063 33.7 33.4

8.6 8.75 8.6 8.45

11.9 11.8 70.6 70.71 0.3 0.29

1806 M. A. SOLIMAN et al. A2

+- (bae 3.8) t-)1.68 +13.9 -4.9 -6.5 -6.3 -1.68 +x7.5 -5.3 -5.8 -5.6 -1.3 +2.02 442.7 +a.7

a- - act-1.0 -39.3 -44 i6.1 l .4 +5.2 -24.0 -27.5 t5.95 +5.* 6.3 Ml.9 +39.24 T6.6 4.m

_____ ----__-_I -----_-_--- ____ _ --_-_-__-_____ --_--_-______

i Differ- l 3s ur kasis -33.7 -36.6 +*.4 +a., da.3 -26.6 -30.2 t3.3 +7.4 +,.1 +2 +9.2 -2.91 -54.9

a incr- l 9.2 *XL9 -3.3 -4.5 -4.2 +4.8 +53., -3.4 -3.95 -3.7 -1.05 +8.X! r37.9 eel.4