**5. Oil venturi condenser with water spraying by steam jet**

To prevent the condenser cooling surface could be contaminated with carbon soot the pyrogas is proposed to be condensing for oil in the condenser of venturi type which is operating now with cooling-spraying water instead of pyrolysis oil cycling before as it referenced in [2]. Water is well corresponded to the pyrolysis process with steam so as it is spraying and evaporating for steam too while the pyrogas cooling and condensing, including the carbon soot catching with oil droplets at the same time. The new condenser is developed with a steam jet for both spraying water and ejecting pyrogas from reactor (Fig.3).

Fig. 3. The flow structure in the oil venturi condenser: pyrogas, steam, water, condensed oil.

With steam the oil condensing temp is well provided as near 100°C for the flash point to be at near 80°C. Moreover, the Venturi tube part is actively cleaning against the carbon soot by the same steam jet for spraying. There are three acting specific points with pyrogas-steam reactor flow mixing with steam-water jet and resulting in condensed oil droplets with incondensable pyrolysis off-gas and all steam residual flow in venturi tube as shown in Fig.3:

Inlet point A:

224 Material Recycling – Trends and Perspectives

pre-mixing and ignition with air in a vortex-flame tunnel just before the furnace (see Fig.3) by which way there is an area in the tunnel where the gas flame is every moment torching

> Operating process in Taiwan (1000 kg/hr)

Table 3. Modern Process Calculation in Comparison With Operating Process in Taiwan-2008

To prevent the condenser cooling surface could be contaminated with carbon soot the pyrogas is proposed to be condensing for oil in the condenser of venturi type which is operating now with cooling-spraying water instead of pyrolysis oil cycling before as it referenced in [2]. Water is well corresponded to the pyrolysis process with steam so as it is spraying and evaporating for steam too while the pyrogas cooling and condensing, including the carbon soot catching with oil droplets at the same time. The new condenser is developed with a steam jet for both

Fig. 3. The flow structure in the oil venturi condenser: pyrogas, steam, water, condensed oil.

**5. Oil venturi condenser with water spraying by steam jet** 

spraying water and ejecting pyrogas from reactor (Fig.3).

Oil combustion rate, kg/hr 20 - - Off-gas burning rate, m3/hr 100 180 180 Combusting air flow, nm3/hr 2650 5000 5000 Furnace heat capacity, МW 1.2 1.95 1.95 Furnace gas temp inlet reactor, °C 900 800 800 Furnace gas temp outlet reactor, °C 500 590 590 Furnace gas temp inlet the boiler, оC 500 590 500 Air injecting rate in boiler, nm3/hr - - 1350 Steam self-producing rate, kg/hr 350 1300 1000 Steam residual with off-gas, kg/hr 15 1300 1000

Modernized process (1000 kg/hr)

Steam producing limited with air injection by (13)

Steam selfproducing rate by (10)

and so igniting just near from the furnace.

Pyrolysis oil heat value: 42 МJ/kg Off-gas heat value : 39 MJ/m3


Mixing point B:


Condensing point C:


Proposing a simple sonic type of steam nozzle for spraying water, initially a self-cooling and wetting effect with steam jet discharged at near the sonic velocity (throttle effect) to be considered, which can be maximally estimated by [40] under the next steam min pressure and temp conditions to be: min *p* = 0.4 MPa – steam pressure (abs), *T =1* 142°C = 415 K – steam boiler temp, *T2* – steam temp after the throttle effect,

$$T\_2 = T\_1 \left(\frac{2}{k+1}\right) = 415 \cdot \left(\frac{2}{1.4+1}\right) \equiv 72 \text{ °C}.\tag{15}$$

Due to steam jet after its discharge is really cooling down only to 100°C and next a few steam is condensing at the same temp and normal pressure, it is acting for steam selfwetting as follows:

$$w\_{s(\text{max})} = \frac{c\_{p(s)}(100 - T\_2)}{h\_s} \equiv 3\,\%\,\tag{16}$$

where *p s*( ) *c* = 0.5 kcal/kg·C is steam specific heat capacity, *<sup>s</sup> h* = 540 kcal/kg is steam specific heat value at the normal (atmospheric) pressure.

#### **5.1 Spraying water specific flow rate**

With author's reference to [35-37] the tire (rubber) pyrolysis specific heat required for its thermal destruction at near to steady pyrolysis temp 400-450°C is experienced

Waste Tire Pyrolysis Recycling with Steaming:

*g* is gas flow velocity,

*<sup>s</sup>* ≅ 500 m/s (water density is

σ

wood particle combustion by its gasification too.

venturi tube as follows:

where

less ϑ ϑ

first part only:

is a few as not above

spraying way:

**5.2 Water droplets spraying and evaporating** 

Heat-Mass Balances & Engineering Solutions for By-Products Quality 227

With all reference to [39] and other fundamentals, the Nukiyma-Tanasawa equation can be applied for calculation on a liquid droplet diameter sprayed with a high-speed gas flow in

*g g*

=+ ⋅

is liquid density and

both of the flows presented in terns of the volume flow rate, the liquid flow *G* is usually much less then that of spray gas *Gg* and equation is used conveniently as simplified to the

0.585

*spr*

Proposing the steam pressure could be so much as the jet velocity to be near of sonic as not

ϑ

0.585 0.05 500 1000 *<sup>i</sup> <sup>d</sup>* ≤⋅ ≤ 10 μ<sup>m</sup>

There are many analytical and experimental data and references on that, concerning the liquid fuel combustion and specially considering the sphere droplet lifetime by its quicktransient heating and vaporizing (gasification) at the flame thermal condition, referencing that for example to [41]. And there is another approach which is all of near the same consideration but at quasi-steady thermal conditions – for example, in [42] concerning the

To estimate the time under question we use the latter as a method at quasi-steady thermal conditions, being there is not so evidence as weakness of that relatively to transient analyzing method. So way it can be analyzed by the convenient heat-mass transfer analogy, taking into account the heat for evaporating the water droplet of diameter 2 *<sup>i</sup> d r* = is provided under the heat transfer criterion condition *Nu*min( )*<sup>i</sup>* = 2 which with reference to [9] is a minimal criterion value of that for spherical particle at a zero-flow velocity condition when the heat is transferred by the gas conductivity only. Due to near the same condition is proposed for a fine water droplet injecting and moving by steam at near the same flow velocity (zero-flow velocity relatively one other), it can be differentially formulated as

bellow, beginning from the heat transfer coefficient by the criterion above and so on:

α

min( ) min( ) <sup>2</sup> *i s <sup>s</sup> <sup>i</sup>*

*Nu k k d d*

*i i*

= = . (26)

*i*

*d*

ρ

 μ

σρ

σ

ρ

*<sup>G</sup> <sup>d</sup>*

0.585 59.7 *<sup>i</sup>*

ϑ

ρ

σ

 ρ 0.45 1.5

*G*

σ

≤ 0.05 N/m), we have the next the water droplets by steam jet

≅ . (25)

= 1000 kg/m3 and water surface tension at near 100°C

, (24)

is liquid surface tension. Being

approximately as *<sup>t</sup> h* = 640 kJ/kg. Proposing the oil condensed is max 45% of tire pyrolysis mass, we have the next condenser heat capacity provided with spraying and evaporating water:

$$Q\_{oil} = 0.45 G\_t h\_t \,\text{.}\tag{17}$$

Heat with pyrogas condensing for oil and co-pyrolysis steam from reactor cooling down to Venturi operating temp 100°C is following:

$$Q\_1 = Q\_{oil} + c\_{p(s)} G\_{s1} (T\_p - 100) \, , \tag{18}$$

heat for spray water heating-evaporating is following:

$$Q\_3 = G\_w \left[ \varepsilon\_{p(w)} (100 - T\_a) + h\_s \right] \tag{19}$$

heat for steam wetness evaporating is following:

$$Q\_2 = w\_s G\_{s2} h\_s = w\_s r G\_w h\_{s\
u} \,\,\,\,\,\tag{20}$$

steam jet as ratio of spraying water rate is following:

$$\mathbf{G}\_{s2} = rW\mathbf{\hat{W}}\_s\tag{21}$$

and condenser total heat-with-mass balance is following:

$$Q\_1 = Q\_2 + Q\_3 \, . \tag{22}$$

By substitution-solution (10)-(15) we have:

$$\frac{W}{G\_t} = \frac{0.45h\_t + \frac{G\_{s1}}{G\_t}c\_{p(s)}(T\_p - 100)}{c\_{p(w)}(100 - T\_a) + h\_s(1 - w\_s r)}.\tag{23}$$

As noted here initially in 4 and with reference to [35, 36] the co-pyrolysis steam for reactor supply is self-producing after heating the latter and it is formulated, calculated and well tested relatively to tire mass in the next terms and rates:

$$\frac{G\_{s1}}{G\_t} = 25 \text{--} 30\%$$

With reference to steam common application for liquid spraying (e.g. for heavy oil fuel combustion), the factor of steam-water spraying mass ratio *r* = 1 is well enough. Taking that and all other given condition above into account, by (23) we have the next analytical solution on the water spraying for venturi condenser performance:

$$\frac{W}{G\_t} \equiv 0.2$$

#### **5.2 Water droplets spraying and evaporating**

226 Material Recycling – Trends and Perspectives

approximately as *<sup>t</sup> h* = 640 kJ/kg. Proposing the oil condensed is max 45% of tire pyrolysis mass, we have the next condenser heat capacity provided with spraying and evaporating

Heat with pyrogas condensing for oil and co-pyrolysis steam from reactor cooling down to

1 ( )

*G c T h wr*

As noted here initially in 4 and with reference to [35, 36] the co-pyrolysis steam for reactor supply is self-producing after heating the latter and it is formulated, calculated and well

*<sup>G</sup>* <sup>=</sup> 25-30%

With reference to steam common application for liquid spraying (e.g. for heavy oil fuel combustion), the factor of steam-water spraying mass ratio *r* = 1 is well enough. Taking that and all other given condition above into account, by (23) we have the next analytical

> *t W G*

≅ 0.2

*s*1 *t G*

*<sup>G</sup> h cT*

0.45 ( 100)

+ −

*<sup>s</sup> t ps p t t p w as s*

(100 ) (1 )

( )

*W G*

0 45 *Q . Gh oil* = *t t* . (17)

<sup>1</sup> () 1( 100) *Q Q c GT* =+ − *oil p s s p* , (18)

3 () [ (100 ) ] *Q Gc T h* = −+ *w pw a s* , (19)

*Q w G h w rG h* 2 2 = = *s s s s ws* , (20)

*G rW <sup>s</sup>*<sup>2</sup> = , (21)

*QQQ* <sup>123</sup> = + . (22)

<sup>=</sup> −+ − . (23)

water:

Venturi operating temp 100°C is following:

heat for spray water heating-evaporating is following:

heat for steam wetness evaporating is following:

steam jet as ratio of spraying water rate is following:

By substitution-solution (10)-(15) we have:

and condenser total heat-with-mass balance is following:

tested relatively to tire mass in the next terms and rates:

solution on the water spraying for venturi condenser performance:

With all reference to [39] and other fundamentals, the Nukiyma-Tanasawa equation can be applied for calculation on a liquid droplet diameter sprayed with a high-speed gas flow in venturi tube as follows:

$$d\_i = \frac{0.585}{\vartheta\_\mathcal{g}} \sqrt{\frac{\sigma}{\rho}} + 59.7 \left(\frac{\mu}{\sqrt{\sigma \rho}}\right)^{0.45} \cdot \left(\frac{G}{G\_\mathcal{g}}\right)^{1.5} \tag{24}$$

where ϑ*g* is gas flow velocity, ρ is liquid density and σ is liquid surface tension. Being both of the flows presented in terns of the volume flow rate, the liquid flow *G* is usually much less then that of spray gas *Gg* and equation is used conveniently as simplified to the first part only:

$$d\_i \equiv \frac{0.585}{\vartheta\_{spr}} \sqrt{\frac{\sigma}{\rho}} \,\,\,\,\tag{25}$$

Proposing the steam pressure could be so much as the jet velocity to be near of sonic as not less ϑ*<sup>s</sup>* ≅ 500 m/s (water density is ρ = 1000 kg/m3 and water surface tension at near 100°C is a few as not above σ ≤ 0.05 N/m), we have the next the water droplets by steam jet spraying way:

$$d\_i \le \frac{0.585}{500} \cdot \sqrt{\frac{0.05}{1000}} \le 10 \text{ \textmu m}$$

There are many analytical and experimental data and references on that, concerning the liquid fuel combustion and specially considering the sphere droplet lifetime by its quicktransient heating and vaporizing (gasification) at the flame thermal condition, referencing that for example to [41]. And there is another approach which is all of near the same consideration but at quasi-steady thermal conditions – for example, in [42] concerning the wood particle combustion by its gasification too.

To estimate the time under question we use the latter as a method at quasi-steady thermal conditions, being there is not so evidence as weakness of that relatively to transient analyzing method. So way it can be analyzed by the convenient heat-mass transfer analogy, taking into account the heat for evaporating the water droplet of diameter 2 *<sup>i</sup> d r* = is provided under the heat transfer criterion condition *Nu*min( )*<sup>i</sup>* = 2 which with reference to [9] is a minimal criterion value of that for spherical particle at a zero-flow velocity condition when the heat is transferred by the gas conductivity only. Due to near the same condition is proposed for a fine water droplet injecting and moving by steam at near the same flow velocity (zero-flow velocity relatively one other), it can be differentially formulated as bellow, beginning from the heat transfer coefficient by the criterion above and so on:

$$
\alpha\_{\rm min(i)} = \frac{N u\_{\rm min(i)} k\_s}{d\_i} = 2 \frac{k\_s}{d\_i} \,. \tag{26}
$$

Waste Tire Pyrolysis Recycling with Steaming:

taken into account:

Heat-Mass Balances & Engineering Solutions for By-Products Quality 229

the droplets save the initial flow velocity as 50 m/s above, it is min 0.02 sec which is

The steam counter-feeding effect for the carbon black purification at the end of processing just inside the reactor of screw type is illustrated in Fig.4 where with author's reference to [38] the longitudinal diagram of scrap tire pyrolysis is presented and where a multi-tube reactor is simplified as one line. Steam is well penetrating into the moving-mixing bed of scrap tire simply by its diffusion, as well as into the every porous fragment or particle of that too, acting so for cleaning them off the volatile residue matters at the end of processing, even there is not convective steam flow inside the bed and most of steam is flowing above that. Along with such cleaning there is evidently some of inner heating the scrap tire by steam diffusion into the bed which question is a quite easy for estimation by value of inner specific surface per 1 m3 bulk of scrap tire minimally as *<sup>i</sup> f* = 20 m2/m3, including the

obtained simply by the conic length divided by the venturi velocity above.

**6. Steam: Inner heating, carbon black cleaning & air sealing lock** 

standard mean-average thickness of chips *<sup>i</sup> d* = 10 mms, bulk density of tire scrap

kg/m3 and Nusselt number for particle heating without convection is *Nu*min( )*<sup>i</sup>* = 2 to be

Fig. 4. The longitudinal and linearized diagram of scrap tire pyrolysis recycling in reactor of

screw-tubular type with steam (the helix of the screw inside reactor is not shown): 1 – geared motor of screw, 2 – reactor tube shell, 3 – reactor heating box, 4 – scrap tire, 5 – release of tire volatile matters, 6 – carbon black, 7 – cleaning steam flow, 8 – resulting steam flow with pyrolysis gases (pyrogas), 9 – some of possible and allowed air inflow at reactor loading side, 10 – steam counter-flow pulse impact toward the air inflow as a steam

seal-lock at the reactor unloading side.

ρ

*t b*( ) = 500

**6.1 Inner heating and increasing the tire pyrolysis rate with steam** 

Evaporating surface of the water spherical droplet is following:

$$f\_i = \pi d\_i^{\text{\tiny \tag{27}}},\tag{27}$$

droplet mass evaporated with the surface layer *dr* is following:

$$dm\_i = \rho\_i f\_i dr \,, \tag{28}$$

heat for the layer *dr* above to be evaporated is following:

$$dQ\_i = dm\_i h\_{s\\_\prime} \tag{29}$$

the same heat to be transferred during in *d*τis following:

$$dQ\_i = \alpha\_i f\_i(T\_p - 100)d\tau \,. \tag{30}$$

By (26)-(30) substitution and integration we have the next solution on the question in title:

$$d\pi = \frac{\rho\_i h\_s}{k\_s(T\_p - 100)} r dr \,\,\,\,\,\tag{31}$$

$$\tau = \frac{d\_i^2 \rho\_i h\_s}{8k\_s \left(T\_p - T\_{oil}\right)} \ln \frac{T\_p - 100}{T\_{oil} - 100} \text{ }, \tag{32}$$

where *<sup>s</sup> k* = 0.02 kcal/m·hr is steam thermal conductivity at 100 C,*Toil* = 105-110°C is pyrolysis oil condensing temp to be proposed, and where the current arithmetic temp difference ( 100) *Tp* − by its integrating in (30) is logically resulted in a mean-logarithmic temp difference Δ*T* as following:

$$
\Delta T = \frac{(T\_p - 100) - (T\_{oil} - 100)}{1 \text{n} \frac{T\_p - 100}{T\_{oil} - 100}} = \frac{T\_p - T\_{oil}}{1 \text{n} \frac{T\_p - 100}{T\_{oil} - 100}} \,\tag{33}
$$

So way we have the next numerical estimation on the droplet 10 μm evaporating time :

$$\pi = 3600 \cdot \frac{10^{-12} \cdot 1000 \cdot 540}{8 \cdot 0.02 \cdot (400 - 105)} \ln{\frac{400 - 100}{105 - 100}} \cong 0.015 \text{ sec}$$

With venturi condenser or scrubber under consideration, it is about 50 m/s of gas flow velocity as a minimal value of that to be in the narrow part of the venturi tube for its effective performance. Next in the conic part of the tube the flow is extending with a spherical angle about 10° so as the flow velocity dropping down about by one order as for an ordinary gas pipe to be. In particularly, considering the venturi condenser above, the narrow tube for 1 t tire pyrolysis per hour is proposed to be about 4 inches in diameter, proposing so the length of the conic tube to be not less then *L2* = 1 m (see Fig.3). The minimal exposition time for water droplets evaporation in the venturi tube, even proposing

*i ii dm f dr* = ρ

τ

By (26)-(30) substitution and integration we have the next solution on the question in title:

( )

( 100) ( 100)

*<sup>T</sup> T T*

− −− − Δ = <sup>=</sup> − −

So way we have the next numerical estimation on the droplet 10 μm evaporating time :

−

where *<sup>s</sup> k* = 0.02 kcal/m·hr is steam thermal conductivity at 100 C,*Toil* = 105-110°C is pyrolysis oil condensing temp to be proposed, and where the current arithmetic temp difference ( 100) *Tp* − by its integrating in (30) is logically resulted in a mean-logarithmic

ln ln

*p oil p oil p p oil oil*

*T T*

( ) <sup>12</sup> 10 1000 540 400 100 <sup>3600</sup> ln 8 0.02 400 105 105 100

With venturi condenser or scrubber under consideration, it is about 50 m/s of gas flow velocity as a minimal value of that to be in the narrow part of the venturi tube for its effective performance. Next in the conic part of the tube the flow is extending with a spherical angle about 10° so as the flow velocity dropping down about by one order as for an ordinary gas pipe to be. In particularly, considering the venturi condenser above, the narrow tube for 1 t tire pyrolysis per hour is proposed to be about 4 inches in diameter, proposing so the length of the conic tube to be not less then *L2* = 1 m (see Fig.3). The minimal exposition time for water droplets evaporation in the venturi tube, even proposing

*T T TT*

( 100) *i s s p <sup>h</sup> <sup>d</sup> rdr k T* ρ

<sup>2</sup> 100 ln 8 100 *p i is s p oil oil d h T kT T T* ρ

100 100

100 100

− −

<sup>=</sup> ⋅⋅ − ⋅ ≅ ⋅⋅ − − 0.015 sec

α

τ

τ

2 *i i f* = π

is following:

 τ

*d* , (27)

, (28)

*i is dQ dm h* = , (29)

*f T d* . (30)

. (33)

<sup>=</sup> <sup>−</sup> , (31)

<sup>−</sup> <sup>=</sup> <sup>−</sup> <sup>−</sup> , (32)

Evaporating surface of the water spherical droplet is following:

droplet mass evaporated with the surface layer *dr* is following:

heat for the layer *dr* above to be evaporated is following:

( 100) *i ii p dQ* = −

the same heat to be transferred during in *d*

temp difference Δ*T* as following:

τ

the droplets save the initial flow velocity as 50 m/s above, it is min 0.02 sec which is obtained simply by the conic length divided by the venturi velocity above.

#### **6. Steam: Inner heating, carbon black cleaning & air sealing lock**

#### **6.1 Inner heating and increasing the tire pyrolysis rate with steam**

The steam counter-feeding effect for the carbon black purification at the end of processing just inside the reactor of screw type is illustrated in Fig.4 where with author's reference to [38] the longitudinal diagram of scrap tire pyrolysis is presented and where a multi-tube reactor is simplified as one line. Steam is well penetrating into the moving-mixing bed of scrap tire simply by its diffusion, as well as into the every porous fragment or particle of that too, acting so for cleaning them off the volatile residue matters at the end of processing, even there is not convective steam flow inside the bed and most of steam is flowing above that. Along with such cleaning there is evidently some of inner heating the scrap tire by steam diffusion into the bed which question is a quite easy for estimation by value of inner specific surface per 1 m3 bulk of scrap tire minimally as *<sup>i</sup> f* = 20 m2/m3, including the standard mean-average thickness of chips *<sup>i</sup> d* = 10 mms, bulk density of tire scrap ρ*t b*( ) = 500 kg/m3 and Nusselt number for particle heating without convection is *Nu*min( )*<sup>i</sup>* = 2 to be taken into account:

Fig. 4. The longitudinal and linearized diagram of scrap tire pyrolysis recycling in reactor of screw-tubular type with steam (the helix of the screw inside reactor is not shown): 1 – geared motor of screw, 2 – reactor tube shell, 3 – reactor heating box, 4 – scrap tire, 5 – release of tire volatile matters, 6 – carbon black, 7 – cleaning steam flow, 8 – resulting steam flow with pyrolysis gases (pyrogas), 9 – some of possible and allowed air inflow at reactor loading side, 10 – steam counter-flow pulse impact toward the air inflow as a steam seal-lock at the reactor unloading side.

Waste Tire Pyrolysis Recycling with Steaming:

**6.2 Steam feeding rate for reactor air-lock sealing** 

Heat-Mass Balances & Engineering Solutions for By-Products Quality 231

At last, with reference to [38] there is other effect with steam feeding into reactor of screw type which really acts as a hydrodynamic seal-lock preventing any possible air inflow through the reactor unloading system which is usually proposed to be seal but should be taken in mind as possible to be otherwise too. In any case the steam feed forms a local hydrodynamic counter-pressure pulse (steam seal-lock) which precisely keeps air from entering the reactor. With purpose of the uniform steam inlet and sealing impact all over the reactor cross-section in-side, it is feeding into there via the multi-jet deflector as shown in Fig.5. The same is shown by dashed arrows as a steam counter-flow acting for sealing toward the possible air inflow in the apposite direction in Fig.4. The steam pulse above is

> 2 <sup>1</sup> 2 *ss ss p* ϑΔ =

where the velocity is to be calculated by half of reactor cross-section square whose second

<sup>2</sup> <sup>2</sup> *<sup>s</sup> <sup>s</sup> ss r ss ss G g*

ρ

Proposing the reactor unloading system would be not sealed with a double-gate or doubleflap valve etc., it would be a chimney draft effect acting as a static low-pressure by the temp difference between the reactor inside and outside which is additionally depended on the

*S*

Fig. 5. Reactor unloading system with a water-cooling screw and double-gates as for consideration on the steam sealing effect against the possible air inflow from outside.

ρ

 ρ

, (42)

= = . (43)

formulated usually as its dynamic pressure depended on the flow velocity:

half is filled with scrap tire initially and carbon black finally (see Fig.5):

height of the reactor installation as shown in Fig.5:

ϑ

$$\sum F\_t = f\_t \frac{G\_t}{\rho\_{t(b)}},\tag{34}$$

$$\alpha\_i = \frac{\mathbf{Nu}\_{\min(i)} \, k\_{ss}}{d\_i} \, \tag{35}$$

$$
\Delta T = \frac{T\_p - T\_a}{\ln \frac{T\_{ss} - T\_a}{T\_{ss} - T\_p}} \text{ \textsuperscript{\textsuperscript{\textsuperscript{\textsuperscript{\textsuperscript{\textsuperscript{\textsuperscript{\textsuperscript{\textsuperscript{\textsuperscript{\textsuperscript{\boxscript{\boxscript{\boxscript{\boxscript{\boxscript{\boxscript{\boxscript{\boxscript{\boxscript{\boxscript{\boxscript{\boxscript{\boxscript{\boxscript{\boxdot{\phantom{\boxdot{\phantom{\boxdot{\phantom{\boxdot{\phantom{\boxdot{\mu}}{\boxdot{\phantom{\frac{\boxdot{\mu}}{\cdots}}}}}}}}}}}}{\over }}}} \text{\overleftarrow{\phantom{\frac{\boxdot{\mu}}{\cdot}}}}\text{\textsuperscript{\boxminus{\frac{\boxdot{\mu}}{\cdot}}}}\text{\textsuperscript{\boxminus{\frac{\boxdot{\mu}}{\cdot}}}}\text{\textsuperscript{\boxminus{\frac{\boxdot{\mu}}{\cdot}}}}\text{\textsuperscript{\frac{\boxdot{\mu}}{\cdot}}}\text{\textsuperscript{\frac{\boxdot{\mu}}{\cdot}}}\text{\}}\text{\}}\tag{36}
$$

$$
\Delta Q\_t = \alpha\_i \sum F\_t \Delta T \tag{37}
$$

where the inner tire heating by steam Δ*Qt* to be provided by steam superheating as Δ*Qss* :

$$
\Delta Q\_t = \Delta Q\_{ss} \tag{38}
$$

$$
\Delta Q\_{ss} = c\_{p(s)} G\_s (T\_{ss} - T\_s) \tag{39}
$$

and the heat for tire pyrolysis is formulated as before too:

$$Q\_t = G\_t \left( c\_{p(t)} (T\_p - T\_a) + h\_t \right) \,. \tag{40}$$

With reference to [38] the numerical data on that are presented in Table 4 where it is calculated by (34)-(40) per 1 t/hr tire rate at the minimal thermal pyrolysis condition as follows: ambient air temperature is *Ta* = 20оС, boiler steam temperature is *Ts* = 100оС, tire pyrolysis temperature is *Tp* = 350оС, steam super-heating temperature is *Tss* = 400оС, steam feeding rate is *Gs* = 300 kg/hr, tire pyrolysis heat is *<sup>t</sup> h* = 630 kJ/kg and others by the nomenclature:


Table 4. Numerical Data (34)-(40) on Inner Heating with Steam per 1 t/hr Tire Pyrolysis.

The additional tire inner heating (37)-(39) could be objectively corresponded to increasing the pyrolysis rate, being that compared to the similar tire processing without steam and being all other conditions equaled. Taking into account the calculating accuracy of that as 2-3% the obtained result can be concluded as 10-15% which is a theoretical limit for increasing the tire pyrolysis rate with steam at the given conditions:

$$\frac{\Delta Q\_{ss}}{Q\_t + \Delta Q\_t} = \frac{52}{350 + 52} = 13\,\%. \tag{41}$$

#### **6.2 Steam feeding rate for reactor air-lock sealing**

230 Material Recycling – Trends and Perspectives

*<sup>t</sup> t t t(b)*

Numin( )*i ss*

*d*

*k*

*i*

*p a ss a ss p*

−

*T T <sup>T</sup> T T ln T T*

<sup>−</sup> Δ = <sup>−</sup>

Δ= Δ *Q FT tit* α

where the inner tire heating by steam Δ*Qt* to be provided by steam superheating as Δ*Qss* :

With reference to [38] the numerical data on that are presented in Table 4 where it is calculated by (34)-(40) per 1 t/hr tire rate at the minimal thermal pyrolysis condition as follows: ambient air temperature is *Ta* = 20оС, boiler steam temperature is *Ts* = 100оС, tire pyrolysis temperature is *Tp* = 350оС, steam super-heating temperature is *Tss* = 400оС, steam feeding rate is *Gs* = 300 kg/hr, tire pyrolysis heat is *<sup>t</sup> h* = 630 kJ/kg and others by the

40 8 162 52 52 350

The additional tire inner heating (37)-(39) could be objectively corresponded to increasing the pyrolysis rate, being that compared to the similar tire processing without steam and being all other conditions equaled. Taking into account the calculating accuracy of that as 2-3% the obtained result can be concluded as 10-15% which is a theoretical limit for

> 52 350 52

Table 4. Numerical Data (34)-(40) on Inner Heating with Steam per 1 t/hr Tire Pyrolysis.

increasing the tire pyrolysis rate with steam at the given conditions:

*ss t t Q Q Q*

*<sup>i</sup>* , W/(m2·оС) Δ*T* , оС Δ*Qt* , kW Δ*Qss* , kW *Qt* , kW

*i*

α

and the heat for tire pyrolysis is formulated as before too:

nomenclature:

*Ft* , m2

α

*<sup>ρ</sup>* , (34)

= , (35)

, (36)

, (37)

Δ =Δ *Q Q t ss* , (38)

(s) ( ) Δ= − *Q c GT T ss <sup>p</sup> s ss s* , (39)

( ) (t)( ) *Q Gc T T h t t* = −+ *p p a t* . (40)

<sup>Δ</sup> = = +Δ + 13%. (41)

*<sup>G</sup> F=f*

At last, with reference to [38] there is other effect with steam feeding into reactor of screw type which really acts as a hydrodynamic seal-lock preventing any possible air inflow through the reactor unloading system which is usually proposed to be seal but should be taken in mind as possible to be otherwise too. In any case the steam feed forms a local hydrodynamic counter-pressure pulse (steam seal-lock) which precisely keeps air from entering the reactor. With purpose of the uniform steam inlet and sealing impact all over the reactor cross-section in-side, it is feeding into there via the multi-jet deflector as shown in Fig.5. The same is shown by dashed arrows as a steam counter-flow acting for sealing toward the possible air inflow in the apposite direction in Fig.4. The steam pulse above is formulated usually as its dynamic pressure depended on the flow velocity:

$$
\Delta p\_1 = \frac{\vartheta\_{ss}^2}{2} \rho\_{ss'} \tag{42}
$$

where the velocity is to be calculated by half of reactor cross-section square whose second half is filled with scrap tire initially and carbon black finally (see Fig.5):

$$
\sigma\_{ss} = \frac{2G\_s}{S\_r \rho\_{ss}} = \frac{2g\_s}{\rho\_{ss}}.\tag{43}
$$

Proposing the reactor unloading system would be not sealed with a double-gate or doubleflap valve etc., it would be a chimney draft effect acting as a static low-pressure by the temp difference between the reactor inside and outside which is additionally depended on the height of the reactor installation as shown in Fig.5:

Fig. 5. Reactor unloading system with a water-cooling screw and double-gates as for consideration on the steam sealing effect against the possible air inflow from outside.

Waste Tire Pyrolysis Recycling with Steaming:

Thermal efficiency of the boiler: 600 200

1050 x (770–570) x 0.35 x 1.16 = 0.085 MW

\$0.06 x 85 = \$5.1

**8. Acknowledgement** 

**9. Nomenclature** 

ρ

μ

*w* – wetness of components, %

– density of components, kg/m3;

 – molecular weight of components; *m* – mass of steam or gas components, kg; *p* – pressure of steam-gas components, bar; *V* – volume of steam-gas components, m3; *<sup>p</sup> c* – specific heat of components, J/kg·°C;

*T* – temperature of components and others, °C; *Tg*1 – furnace gas temp for reactor heating inlet, °C; *Tg*2 – furnace gas temp for reactor heating outlet, °C;

Heat for steam generation: 0.7x1.075 MW = 0.7525 MW; Steam rate (by index 1400 kg/hr per 1 MW) = 1050 kg/hr; Thermal efficiency of small steam turbine (max) = 35%;

Carbon black (per 1000 kg/hr pyrolysis) = 350 kg/hr Carbon black rate additionally: 0.1 x 350 = 35 kg/hr Economy for sale the carbon black: \$0.3 x 35 = \$10.5

Horng Jiang (Rhine) for assistance and good co-operation in Taiwan.

Heat-Mass Balances & Engineering Solutions for By-Products Quality 233

Pyrolysis off-gas combustion heat value (pure-dry gas as without steam) = 39 MJ/m3; Heat capacity by max 18% off-gas after-burning (relatively to tire mass) = 1.7 MW; Heat emission to outside with both of furnace and reactor by 5% totally = 0.1 MW; Heat consumption for tire heating and pyrolysis (process without steam) = 0.35 MW Heat capacity for 1000 kg/hr steam super-heating to 400°C (for turbine) = 0.175 MW Heat residual and available for steam generation after all of these above = 1.075 MW; Exhaust gas flow temperatures inlet/outlet the steam second-heat boiler = 600/200°C

> 600 20 *<sup>E</sup>* <sup>−</sup> = ≅ <sup>−</sup> 70%;

By the steam enthalpy operating range in turbine as max h1 = 770 kcal/kg (4.0 MPa, 400°C) and h2 = 570 kcal/kg (0.05 MPa, 40°C), it is the next power per 1000 kg/hr tire processing:

Proposing as max \$0.06 price per 1 kW-hr power, we have the next economy if to sale that:

By min 10% waste tire pyrolysis rate to be more by steaming way and with reference to [32] as the carbon black price is min \$0.3 per 1 kg, we have the next economy if to sale that more:

The author wishes sincerely and friendly to thank Mr. Wu Chun Yao (Morgan) and Mr.

$$
\Delta p\_2 = \rho\_a gH \left( 1 - \frac{T\_a}{T} \right),
\tag{44}
$$

which is to be equal to the dynamic pressure (42) with steam feeding above and by which substitution the steam feeding rate under question is following:

$$\mathbf{g}\_{s1} = \frac{\mathbf{G}\_{s1}}{S\_r} = \sqrt{\frac{\mathbf{g}\_a \rho\_{ss} H}{2} \left(1 - \frac{T\_a}{T}\right)}\,\mathrm{}.\tag{45}$$

With reference to [38], as well as simply and evidently by equation (45), the steam feedingsealing effect is to be objectively enhanced with cooling of the carbon discharge from reactor, being the opposite draft effect for air inflow is depended on the cooling temp just by another way. The numerical data on that are presented in Table 5 and theoretically it even would be nothing of sealing required if the carbon discharge temp could be entire equal to outside so as nothing of chimney effect appeared. At the same time in Table 6 there are numerical data on the steam feed required as obtained by (45) at the different ambient temperature which is other factor of the chimney draft effect simply by outside condition of the pyrolysis plant location in different region. With the same reference to [38] it was well experienced practically and particularly in Taiwan at about 30оC where it was no evidence of the air penetration inside with the steam feeding rate appropriated minimally as 200–250 kg/hr for the reactor diameter 0.6 m, or near 1 t/hr per 1 m2 of cross-section square of that in specific terms.


Table 5. Steam Feeding Rate per 1 m2 Reactor Cross-Section Square for Air-Lock Seal With Carbon Product Cooling Temp for Discharge.


Table 6. Steam Feeding Rate per 1 m2 Reactor Cross-Section Square for Air-Lock Seal With Air Ambient Temperature Outside Reactor.

#### **7. Conclusion: Brief engineering-economy analysis on steam use way**

In conclusion, in connection with possibility for steam self-producing along with tire pyrolysis recycling, it is a reason to analyze numerically and economically what is more effective way for steam use: power generation by turbine machine or carbon black production could be more as min 10% of tire rate additionally by inner heating with steam feeding into reactor as it is considered above in 6.1? By first way we have near to the next thermal data for steam generation after heating the pyrolysis reactor per 1000 kg/hr tire processing:

Pyrolysis off-gas combustion heat value (pure-dry gas as without steam) = 39 MJ/m3; Heat capacity by max 18% off-gas after-burning (relatively to tire mass) = 1.7 MW; Heat emission to outside with both of furnace and reactor by 5% totally = 0.1 MW; Heat consumption for tire heating and pyrolysis (process without steam) = 0.35 MW Heat capacity for 1000 kg/hr steam super-heating to 400°C (for turbine) = 0.175 MW Heat residual and available for steam generation after all of these above = 1.075 MW; Exhaust gas flow temperatures inlet/outlet the steam second-heat boiler = 600/200°C

Thermal efficiency of the boiler: 600 200 600 20 *<sup>E</sup>* <sup>−</sup> = ≅ <sup>−</sup> 70%;

Heat for steam generation: 0.7x1.075 MW = 0.7525 MW; Steam rate (by index 1400 kg/hr per 1 MW) = 1050 kg/hr; Thermal efficiency of small steam turbine (max) = 35%;

By the steam enthalpy operating range in turbine as max h1 = 770 kcal/kg (4.0 MPa, 400°C) and h2 = 570 kcal/kg (0.05 MPa, 40°C), it is the next power per 1000 kg/hr tire processing:

1050 x (770–570) x 0.35 x 1.16 = 0.085 MW

Proposing as max \$0.06 price per 1 kW-hr power, we have the next economy if to sale that:

\$0.06 x 85 = \$5.1

232 Material Recycling – Trends and Perspectives

<sup>2</sup> 1 *<sup>a</sup> <sup>a</sup> <sup>T</sup> p gH <sup>T</sup>* ρΔ= −

which is to be equal to the dynamic pressure (42) with steam feeding above and by which

<sup>1</sup> 1 2 *s a a ss <sup>s</sup>*

*G T g H <sup>g</sup> S T* ρ ρ== −

With reference to [38], as well as simply and evidently by equation (45), the steam feedingsealing effect is to be objectively enhanced with cooling of the carbon discharge from reactor, being the opposite draft effect for air inflow is depended on the cooling temp just by another way. The numerical data on that are presented in Table 5 and theoretically it even would be nothing of sealing required if the carbon discharge temp could be entire equal to outside so as nothing of chimney effect appeared. At the same time in Table 6 there are numerical data on the steam feed required as obtained by (45) at the different ambient temperature which is other factor of the chimney draft effect simply by outside condition of the pyrolysis plant location in different region. With the same reference to [38] it was well experienced practically and particularly in Taiwan at about 30оC where it was no evidence of the air penetration inside with the steam feeding rate appropriated minimally as 200–250 kg/hr for the reactor diameter

carbon cooling temp *Т*, <sup>о</sup> С 400 300 200 100 50

Table 5. Steam Feeding Rate per 1 m2 Reactor Cross-Section Square for Air-Lock Seal With

Table 6. Steam Feeding Rate per 1 m2 Reactor Cross-Section Square for Air-Lock Seal With

In conclusion, in connection with possibility for steam self-producing along with tire pyrolysis recycling, it is a reason to analyze numerically and economically what is more effective way for steam use: power generation by turbine machine or carbon black production could be more as min 10% of tire rate additionally by inner heating with steam feeding into reactor as it is considered above in 6.1? By first way we have near to the next thermal data for steam generation after heating the pyrolysis reactor per 1000 kg/hr tire

**7. Conclusion: Brief engineering-economy analysis on steam use way** 

ambient air temp *Ta* ,<sup>о</sup> С 30 20 10 0 -10

by (45), t/(m2·hr) 1.07 1.55 1.74 1.94 2.12

by (45), t/(m2·hr) 3.7 3.45 3.0 2.3 1.55

substitution the steam feeding rate under question is following:

1

*r*

0.6 m, or near 1 t/hr per 1 m2 of cross-section square of that in specific terms.

steam feeding rate

steam feeding rate

processing:

Carbon Product Cooling Temp for Discharge.

Air Ambient Temperature Outside Reactor.

, (44)

. (45)

By min 10% waste tire pyrolysis rate to be more by steaming way and with reference to [32] as the carbon black price is min \$0.3 per 1 kg, we have the next economy if to sale that more:

Carbon black (per 1000 kg/hr pyrolysis) = 350 kg/hr Carbon black rate additionally: 0.1 x 350 = 35 kg/hr Economy for sale the carbon black: \$0.3 x 35 = \$10.5
