Process integration and exergy analysis of the autothermal reforming of glycerol using supercritical water

Energy 42 (2012) 192e203

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Process integration and exergy analysis of the autothermal reforming of glycerol using supercritical water F.J. Gutiérrez Ortiz*, P. Ollero, A. Serrera, S. Galera Departamento de Ingeniería Química y Ambiental, Universidad de Sevilla, Camino de los Descubrimientos s/n, 41092 Sevilla, Spain

a r t i c l e i n f o

a b s t r a c t

Article history: Received 1 January 2012 Received in revised form 26 March 2012 Accepted 28 March 2012 Available online 27 April 2012

The most thermodynamically favorable operating conditions at which glycerol can be converted into hydrogen with maximum hydrogen yield by autothermal reforming using supercritical water were identified in a previous paper. As a second part of the study, a conceptual design based on energy integration and exergy analysis of the whole process has been performed. In the proposed scheme, the huge pressure energy of the gas product just at the outlet of the reforming reactor is converted into electrical power and a fraction of the expanded gas used to provide energy support for the process by burning it in a furnace, if needed. By using the optimal conditions found in the previous work, a severe deficit of energy arises in the process. Thus, both water-to-glycerol and oxygen-to-glycerol mole ratios at which thermoneutral conditions are achieved in the reformer are computed by burning all the product gas from the reformer, both for pure and pretreated crude glycerol, at different reforming and preheating temperatures. The pressure used is 240 atm. The effects of the main operating parameters are investigated by sensitivity analysis to identify optimal conditions to maximize power production under autothermal conditions, evaluating the results by energy and exergy analyses. The computations are made with the aid of AspenPlusÔ, using the predictive SoaveeRedlicheKwong equation of state as the thermodynamic method in the simulation of the supercritical region. Ó 2012 Elsevier Ltd. All rights reserved.

Keywords: Autothermal reforming Supercritical water Glycerol Hydrogen Exergy analysis Process integration

1. Introduction Glycerol is obtained as a by-product in biodiesel production using vegetable oils by a base-catalyzed transesterification reaction. The purification of this glycerol is largely based on its final purity requirements. Glycerol of high purity is an important industrial feedstock for applications in food, cosmetics, pharmaceutical and other industries; however, it is costly to refine crude glycerol, especially for medium and small-sized plants. In any case, since glycerol production and utilization have a notable impact on the economics and sustainability of biodiesel production, the development of novel processes for glycerol valorization is essential. Among all possible routes, glycerol conversion into hydrogen as an energy carrier is one of the most attractive. Different reforming processes along this line have been studied and still go on [1e6]. Likewise, glycerol bioconversion to chemicals, such as succinic acid or acetic acid, is also being studied [7,8]. This paper continues the study started in a previous paper, focused on ATR (autothermal reforming) e reforming plus partial * Corresponding author. Tel.: þ34 95 448 72 68/65/60; fax: þ34 95 446 17 75. E-mail addresses: [email protected], [email protected] (F.J. Gutiérrez Ortiz). 0360-5442/$ e see front matter Ó 2012 Elsevier Ltd. All rights reserved. doi:10.1016/j.energy.2012.03.069

oxidation e using SCW (supercritical water), which is defined as water that is heated and compressed above its critical temperature (374
C) and pressure (22.1 MPa). Supercritical water has many advantageous properties [9e11]; among others, glycerol reforming using supercritical water in a catalyst-free process arises to be a very attractive option. Autothermal reforming using SCW may be an excellent method because of its ability to achieve thermoneutral conditions by adjusting the feed ratios, which can save on energy required for heating the reactor and make it possible to avoid the use of a catalyst, as aforementioned. Studies on the ATR of glycerol can be found in the literature [3,12e15]; also on glycerol reforming in supercritical water, but non autothermal [16]. In the previous study [17], hydrogen production by autothermal reforming of pure and crude glycerol using SCW was investigated, and optimal conditions for hydrogen production from a thermodynamic point of view were obtained, but without performing an energy assessment of the process or using a process integration. In this new paper, a conceptual design based on energy integration and exergy analysis of the whole process is proposed and assessed. Although a final purified hydrogen-rich gas to use in a fuel cell may be considered as the main aim, an easier option is to convert the huge pressure energy of the reformate product at the outlet of

F.J. Gutiérrez Ortiz et al. / Energy 42 (2012) 192e203

the reforming reactor into electrical energy by means of an expander. The product gas makes it possible to sustain the energy through the overall process by further energy integration. In this way no external heat source is required and the oxygen entering the reforming reactor is controlled in order to not oxidize all the glycerol fed into the process. 2. Methodology 2.1. Exergy analysis The exergy balance is a statement of the law of degradation of energy, and an exergy analysis quantifies the efficiency lost in a process due to the loss in energy quality, indicating where the process can be improved and suggesting what areas should be given consideration. In other words, exergy analysis makes it possible to evaluate the irreversibilities of the process as the exergy flow loss or lost work, once the suitable control volume is selected. In this study, the terms energy flow and exergy flow refer to energy and exergy per unit of time (power). Likewise, the terms heat flow and work flow will be used, although power will also be used for the latter. The studied process is continuous and steady-state; it is open to interactions with the surroundings, and potential and kinetic energy are not considered. Likewise, the reactors are in equilibrium condition. Also, the product gases are at the reactor temperature. The dead-state conditions (of surroundings) are 25
C and 1 atm. All flows (mass, energy, exergy) entering the system are considered as positive, while flows exiting are indicated by a negative sign. By combining energy and entropy balances, the lost work _ can be computed as follows: (power), LW,

_ ¼ LW

X 

X  X _ _ out þ W B_ in  B_ out  W in

(1)

system

I.e., the lost work is the net sum of exergy flows of inlet and outlet streams minus the net power (work flow) produced in the system,

193

since the whole system is adiabatic. Otherwise, the heat flows ðQ_ j Þ transferred to the system (positive) and by the system (negative) affected by the Carnot factor should be taken into account:

hCarnot ¼

T 1 0 Tj

! (2)

where T0 is 298 K and Tj is the temperature of the jth heat source or sink outside the system. Additionally, the chemical energy in the reformate gas provides the energy and exergy source. Some good reviews provide further discussion about exergy terms [18]. As the final objective of reforming is to get maximum work (power) through the huge pressure energy of reformate products, the exergy efficiency of the process to be inspected is the following:

h¼

_ net W

(3)

_ net  LW _ W

The term Lost Work (flow) includes both internal and external exergy losses. By taking into account that the cooling water heated in the process could be used as a local source of heating, a combined heat-power exergy efficiency (hh-p) may be defined as follows:

hhp ¼

_ net  P DB_ W cooling cEX

_ net  LW _  W

P

cEX

water

DB_ cooling

(4) water

where

DB_ cooling

water

¼

X  B_ out  B_ in cooling

(5)

water

Finally, the thermal or energy efficiency is defined as

hT ¼

_ net W _ Gly LHVGly m

(6)

Table 1 shows the summary of the exergy analysis for each component of the process.

Table 1 Exergy analysis of the individual process units. Equipment Mixers (MIX) Splitters (SPL) and Mechanical Separators (SEP) Pumps (P) and Fans (FAN)

Exergy analysis (exergy losses and exergy efficiency) P_ _ ¼ B_ _ LW Bin lost ¼ Bout þ P _ ¼ B_ LW ¼  B_ out þ B_ lost

_ _ ðW useful Þmin ¼ W min ¼ _ _ ðW useful Þreal ¼ W real ¼ B_ out  B_ in hP ¼ ðB_ out  B_ Þ þ B_ in

lost

in

B_ out  B_ in > 0 B_ out  B_ þ B_

in lost > 0 ðB_ out  B_ in Þ ¼ _ ðB_ out  B_ Þ  LW in

P

P

 ðH_ out  H_ in Þj 1 

Multi-stage Compressor (COMP) þ intermediate cooling stages

_ ¼ B_ _ _ LW lost ¼ Bout þ Bin þ

Expanders (TURB)

_ _ _ _ ðW useful Þmax ¼ W max ¼ ðBout  Bin Þ > 0 _ _ _ out  B_ Þ  B_ ðW Þ ¼ W ¼ ð B in useful real real lost > 0 _ ðB_ out  B_ in Þ  B_ lost ðB_ out  B_ in Þ  LW hTURB ¼ ¼ ðB_ out  B_ in Þ ðB_ out  B_ in Þ P P _ _  ðB_ out  B_ in Þhot fluid  LW ðBout  B_ in Þcold fluid h EX ¼ ¼ P P _  ðBout  B_ in Þhot fluid  ðB_ out  B_ in Þhot fluid P _ P_ P _ ¼ PðB_ out  B_ Þ LW ðBout  B_ in Þcold fluid ¼ Bin  B_ out or also in hot fluid  P P P P_ _ ¼ T LW ðS_ out  S_ in Þhot fluid þ T0 ðS_ out  S_ in Þcold fluid ¼ T0 S_ out  T0 Sin 0 P_ _ _ _ _ LW ¼ B ¼ Bout þ B ; Q ¼ 0

Heat exchangers (HE)

Reforming reactor (R) FurnaceeCombustor (FeC)

Heater

lost

Q_ FeC ¼ H_ out  H_ in < 0

_  W i

i;c stage

j;c int cool

R

in

    P_ T _ ¼ B_ _ LW Bin  1  0 Q_ FeC  lost ¼ Bout þ TFeC     T0  _ _ _ _ ¼ B_  Q LW lost ¼ Bout þ Bin þ 1  TFeC  FeC  

T0 ðTout þ Tin Þ=2

 j

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2.2. Process integration and simulation

gas is expanded in a turbine, converting the pressure energy into electrical energy. The gas purification for obtaining a hydrogen-rich gas to be used in a fuel cell is not carried out. The product gas is cooled to 60
C by pumping in cooling water, and condensate water present in the product gas is removed by means of a gaseliquid separator. In order to reduce the glycerol fraction to oxidize in the reformer, a fraction of the gas stream at the separator outlet enters a furnaceecombustor along with preheated air. The furnaceecombustor is simulated as a stoichiometric reactor where the entire gas fraction that can be oxidized is combusted. The furnace is modeled as an isothermal reactor in such a way that flue gas is at 1000
C (a real value) and the heat released is sent to a heater located just at the reformer inlet. This heater would be physically integrated inside the furnace, like steam tubes in a boiler; i.e., furnace plus heater are a single unit. In the furnace, the gas is burnt to preheat the feed up to a set preheating temperature of the glycerolewater mixture before entering the reforming reactor. The obtained flue gas is used for first heating the glycerolewater mixture, and then the compressed air needed in the reformer. The scheme includes eight process-integrated heat exchangers. Air enters the process at two points. High-pressure air is needed in the reformer, using a five-stage compressor (compression ratio around 3 per stage) with intermediate cooling (to 40
C per stage) to avoid excessively high temperatures and optimize the compression from a thermodynamic point of view. Since the compressor is rated with an isentropic efficiency, it is modeled as an isentropic machine, so an isentropic power is calculated; in fact, it is an adiabatic but not reversible machine. The same is true for the expander. A fan is used to feed the air required in the furnaceecombustor, so there is an O2 content of 3%vol in the flue gas.

A conceptual design of the process is proposed by including a number of heat exchangers suitably located in the process to improve the energy use. The strategy followed is based on minimizing the lost work or exergy losses, so small temperature-driving forces must be achieved, by using countercurrent flow and small temperature approaches at the ends of the exchangers. In this ‘from inside to outside’ approach to the heat exchanger network, the internal streams close to each other at high thermal levels are the first to contact each other (the high-temperature hot streams heat the high-temperature cold streams) and the external more separated streams at low thermal levels are the last to contact each other (the low-temperature hot streams warm up the lowtemperature cold streams), by taking the reformer as a central point, where maximum temperature is required. Then, by assessing the redistribution of exchanger duties in alternative flow sheets, a final flow sheet that provides the best heat integration performance is finally selected. As a consequence, the flow sheet of the process is considerably more complex than that used in our previous work [17], as illustrated in Fig. 1. A glycerolewater mixture is pumped into the supercritical reformer after heating as much as possible by three heat exchangers in order to minimize glycerol oxidation by incoming air (using a minimum O2/Glycerol mole ratio). In the reactor, nonoxidized glycerol is reformed at a temperature and pressure previously established. Hence, the reforming reactor operates by providing the energy needed to sustain the endothermic reaction and to reach the specified temperature. The reforming reactor has been simulated as an R-Gibbs reactor, in which the gas product composition and the heat of the overall reaction are calculated under conditions minimizing the Gibbs free energy. The product

CW4 HE01

MIX P1

W 01

HE03

HE02

02

R1

HEATER

04

05

TURB 06 SG01

GLY

AIRIN

Supercritical Water Reformer

GC3 GC2 HE04 GC

Q CW3

C5 GC4

SGOUT

GC6

C4

SG02

HE06

SG04

SGIN

C3 SGIN2 C2

SPLIT

HE05

SG03

GC5 W1

AIR3IN C1

SEP

P2

FURNACE

C6 HE07

CW2

AIR3C

AIR

Fig. 1. Flow-sheet of the simulated process.

AIR3

CW1

F.J. Gutiérrez Ortiz et al. / Energy 42 (2012) 192e203

Specifications of the elements used in the simulation are shown in Table 2. The equilibrium compositions in the reforming reactor have been calculated for given operating conditions. The computation has been done with the aid of AspenPlusÔ version 2006.5 [19]. The thermodynamic method used has been the predictive PSRK (SoaveeRedlicheKwong), since it provides the most accurate binary interaction parameters and gives more satisfactory results for mixtures of non-polar and polar components, as in the case of the crude/pure glycerol and water mixture [17]. Hydrogen, carbon monoxide, carbon dioxide, methane, ethane, propane, water, methanol, ethanol, glycerol and oxygen as well as pure carbon were added manually, as they were considered the possible species in autothermal reforming of glycerol using SCW. The simulation did not predict coke formation for any of the experimental conditions in this paper, and it did not add any other compound (mole fractions lower than 1012). The contents of ethane, propane and ethanol were also so low that they have not been included in the analyses. For all operating conditions simulated, glycerol and methanol conversions are always 100% at equilibrium condition. As in the previous paper [17], two feeds have been studied: pure and pretreated crude glycerol. The pretreated crude glycerol feed of glycerol of variable methanol content (10, 20 and 30 wt%), without water. Based on reforming reactions (7) and (8), two hydrogen yields are computed using Equations (9) and (10):

195

Glycerol reforming: C3H8O3 þ 3H2O ¼ 3CO2 þ 7H2

(7)

Methanol reforming: CH3OH þ H2O ¼ CO2 þ 3H2

Pure glycerol :

hH2 ¼

mol=h H2 1  mol=h C3 H5 ðOHÞ3 7

(8)

(9)

Pretreated crude glycerol :

h H2 ¼

mol=h H2   mol=h C3 H5 ðOHÞ3 7 þ mol=h CH3 OH 3

(10)

2.3. Case studies Equilibrium compositions of reforming gas obtained were determined as a function of reforming temperature (600e1000
C), preheating temperature of aqueous solution of glycerol fed to the reformer (from 500 to reforming temperature for each value of this variable), water-to-pure or crude glycerol mole ratio (W/G) from 3 to 99, and oxygen-(entering as air)-to-pure or crude glycerol mole ratio (O2/G), as well as MeOH content in the crude glycerol from 10 to 30 wt%), as a measure of the glycerol purity of the crude feed. A

Table 2 Specifications of the components for the simulation. Code

Equipment

Specifications

MIX P1

Mixer Pump

HE1

Heat exchanger

HE2

Heat exchanger

HE3

Heat exchanger

HEATER

Heat exchanger

R1

Reactor

TURB

Expander

HE4

Heat exchanger

HE5

Heat exchanger

SEP

Gaseliquid separator

FeC

FurnaceeCombustor

C1eC5.

Compressor (five stages)

Intermediate Coolers for C1eC5

Four heat exchangers

C6

Compressor

HE6

Heat exchanger

HE7

Heat exchanger

P2

Pump

SPLIT

Splitter

Pressure drop: 0 atm. Efficiency: 80% Outlet pressure: 240 atm Pressure drop: 0 atm. DT (hot fluid outlet  cold fluid outlet) ¼ 2
C Pressure drop: 0 atm. DT (hot fluid outlet  cold fluid outlet) ¼ 3
C Pressure drop: 0 atm. DT (hot fluid outlet  cold fluid outlet) ¼ 3
C Pressure drop: 0 atm. Outlet temperature ¼ Variable (500e800
C) Inlet temperature: Variable (800
C as nominal) Pressure drop: 0 atm Isentropic turbine (isentropic efficiency: 0.72) Outlet Pressure: 1.5 atm. Pressure drop: 0 atm. DT (hot fluid outlet  cold fluid outlet) ¼ 3
C Pressure drop: 0 atm. Outlet temperature of hot fluid ¼ 60
C Adiabatic Pressure drop: 0 atm. Combustion of everything able to be oxidized Operating temperature: 1000
C Operating pressure: 1 atm Surplus reaction heat / HEATER Isentropic machine Compression ratio: 2.992 per stage Final pressure at the outlet: 240 atm. Pressure drop: 0 atm. Outlet temperature in each stage ¼ 40
C Isentropic machine (isentropic efficiency: 0.72) Outlet pressure: 1.1 atm Pressure drop: 0 atm. DT (hot fluid outlet  cold fluid outlet) ¼ 3
C Pressure drop: 0 atm. DT (hot fluid outlet  cold fluid outlet) ¼ 3
C Efficiency: 80% Outlet pressure: 1.1 atm Split fraction stream TGE: variable on heat required in heater; final value to get null SGOUT Pressure drop: 0 atm

196

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sensitivity analysis has been carried out by varying these parameters. The total mass flow rate of glycerol fed to the system is always 1000 kg/h. The required W/G ratio is achieved by adding more or less water. Pressure did not significantly affect the process in the range from 200 to 300 atm and is not included in the sensitivity analysis. Thus, a fixed value of 240 atm chosen taking into account that it is convenient to have a margin above the critical pressure (218 atm), thus accounting for pressure drops in the equipment and through the pipes. Under given operating conditions, simulations run to calculate the O2 needed in the reforming reactor for achieving thermoneutral conditions, at which the process takes place without external heat to sustain the reactions and the overall heat flow in the reformer is null. Nevertheless, a fraction of the product gas energy must be fed back in order to preheat the feed and to achieve the thermoneutral conditions in some cases, as shown below. Besides, oxygen entering the reformer must be limited to avoid the complete oxidation of glycerol fed to the reformer. Thus, the O2/G ratio should be lower than 3.5; however, for reforming to take place, the O2/G ratio should be considerably lower. In the computations, this computed parameter never reaches 2.0. Accordingly, simulations have been also run to calculate the fraction of the reformate gas to be burnt in the furnaceecombustor. This fraction will be higher for high reforming temperature, low preheating temperature and high W/G mole ratio.

3. Results and discussion The optimal thermodynamic conditions obtained in the previous work [17] were first tested under the scheme proposed, i.e., 900
C and 240 atm for 5 wt% pure glycerol in aqueous solution (equivalent to a W/G mole ratio of 99). Tables 3 and 4 show the overall energy balance at reforming temperatures of 900 and 800
C, respectively, using an O2/G of 1. It can be observed that a heat flow of 17,289 kW is needed in the reformer, but the net power is 6588 kW when reforming at 900
C; i.e., the overall heat flow in the reformer cannot become null. Therefore, it is not possible to make the process autothermal, since the energy demanded is much higher than that provided by the system itself. In fact, there is a severe deficit of energy within the overall process, which cannot be compensated for even by converting the power obtained in the expander into thermal energy, since the reformer requires much more energy; an external energy source would be necessary and the process cannot be autothermal. By simulating at O2/G mole ratios of 2 and 3.5, the results are qualitatively similar. Thus, it is necessary to know what the maximum W/G mole ratio is.

Table 3 Overall energy balance for autothermal reforming of glycerol using SCW at 900
C, 240 atm and W/G of 99 by burning the entire product gas, using an O2/G of 1.0. Work/Heat entering the system (kW)

Enthalpy of the inlet-streams (kW)

P1 P2 (C1eC5) þ inter. cool. C6 Reformer (R1)

Glycerol Water Air Air3 CW1

220.36 1.11 72.99 14.93 17,289.35

Work leaving the system (kW) TURB

6897.39

OVERALL

10,701.35

2053.65 83,766.50 0.09 0.28 875,940.00

Enthalpy of the outlet-streams (kW) GC6 W1 CW4

7949.23 81,131.44 861,978.61 10,701.24

Table 4 Overall energy balance for autothermal reforming of glycerol using SCW at 800
C, 240 atm and W/G of 99 by burning the entire product gas, using an O2/G of 1.0. Work/Heat entering the system (kW)

Enthalpy of the inlet-streams (kW)

P1 P2 (C1eC5) þ inter. cool. C6 Reformer (R1)

Glycerol Water Air Air3 CW1

220.36 0.83 72.99 14.93 15,761.07

Work leaving the system (kW) TURB

OVERALL

6170.67

2053.65 83,766.50 0.09 0.28 825,246.39

Enthalpy of the outlet-streams (kW) GC6 W1 CW4

9899.52

7984.78 81,089.42 812,093.06 9899.65

This has been carried out for all reforming and preheating temperature ranges.

3.1. Limits of water-to-glycerol mole ratio for a thermoneutral condition Depending on the feed temperature at the reformer inlet, for a given reforming temperature, the constraint of attaining null overall heat in the reformer cannot be achieved at high water-toglycerol mole ratios:  At very high W/G, the autothermal process is not possible because an unavailable energy input is required, even though the O2/G mole ratio is increased up to what is required to burn all the glycerol fed in, due to the high water flow rate. Thus, there is an unreachable absolute value of W/G (about 23 for 600
C reforming temperature).  When the preheating temperature is very different from the reforming temperature, as W/G increases the required oxygen becomes relatively elevated since the energy input needed in the reformer is higher, especially at high reforming temperatures. This establishes a practical W/G maximum for reforming a fraction of the glycerol. Furthermore, there is also a practical minimum for the W/G mole ratio, in order to compare all the ATR cases within the same framework. For low W/G ratio values and when the preheating temperature differs significantly from the reforming temperature, only a fraction of the reformate gas is needed, and the unused fraction should be taken into account in the global assessment. This is not as simple as may be thought since the gas excess must be purified and its chemical energy converted into electrical energy; this is outside the scope of the present paper, but will be treated in the next one, although using a much more complex simulation scheme. Likewise, the cases with very low W/G ratio and high preheating temperatures (equal to or near the reforming temperature) lead to significant methane formation and low conversion of glycerol to hydrogen in equilibrium conditions. The methanation reactions of CO and CO2 (Equations (11) and (12)) are exothermic and thus, the exothermicity in the reformer makes it unnecessary to add oxygen. There is even a heat flow surplus, which would be released to the environment. Then, the process is no longer autothermal. CO þ 3H2 ¼ CH4 þ H2O e methanation of CO

(11)

CO2 þ 4H2 ¼ CH4 þ 2H2O e methanation of CO2

(12)

F.J. Gutiérrez Ortiz et al. / Energy 42 (2012) 192e203

197

Table 5 shows the range of possible W/G mole ratios for every reforming temperature and preheating temperature, as well as the O2/G mole ratio range corresponding to the W/G range. The lower limit of the W/G range matches the stoichiometric, i.e., 3. It is interesting to note that the upper limit of the W/G range corresponds to the case in which all the product gas must be burnt in the furnace. Beyond this limit, all the glycerol is oxidized in the reformer. Thus, in this study, with the aim of maximizing power production, the cases in which all the product gas is used in the process are considered for both pure and pretreated crude glycerol. This can be taken as a starting point to optimize the process performance energetically. In any case, and although all the product gas is burnt, the hydrogen yield is also taken into account in this study because its production is a focal point in the overall research, as included in the series of papers, and will be addressed in the next one, as mentioned above. 3.2. Effect of reforming temperature under thermoneutral conditions By increasing the reforming temperature, the equilibriums shift toward a higher production of H2, reducing the CH4 formation. Thus, the endothermicity is much more marked at high temperatures. As a consequence, the required O2 increases and both real limits of W/G decrease. Fig. 2 depicts the O2/G mole ratio and the fraction of product gas to be burnt when the W/G varies between the allowed limits for reforming temperatures from 700 to 1000
C when preheating the feed to the reforming temperature. At 700
C, the operation window for W/G dropped to a range between about 19 and 20, and the O2/G is quite low (from 0.012 to 0.025, respectively), burning almost all the product gas. However, at 1000
C, the W/G ratio ranges from 3 to about 15, the O2/G ratio from 0.21 to 0.46 and the fraction of product gas burnt from 0.36 to 0.98. Likewise, as intermediate conditions, when the reforming and preheating temperatures are 800
C, the lower W/G limit is about 8, the O2/G ranges from 0.012 to 0.216 (for a W/G of 18) and the fraction of product gas to burn ranges from 0.52 to 0.98, at a W/G of 8e18, respectively. Similarly (but not shown), by increasing the reforming temperature up to 1000
C and preheating at 800
C, the O2/G from 0.520 to 0.965, at a W/G of 3e14, respectively. At a W/G of 15, all the glycerol is oxidized to CO2 and H2O and the net power decreases from 1325 kW to 1093 kW. In this case, reforming does not take place. Likewise, just the opposite occurs at low reforming temperatures, and the W/G range is widened. Thus, by reforming at 600
C, if the inlet feed temperature is also 600
C, no O2 is required for a W/G between 3 and 23. Furthermore, heat is

Fig. 2. O2/G mole ratio and fraction of product gas to be burnt versus W/G (up to the allowed limits), for different reforming temperatures when preheating the feed to the reforming temperature.

released from the reformer to the environment. This is due to the high CH4 flow generated at such low temperatures so that hardly any H2 is produced. This is a non-true autothermal case, and has been excluded in the analyses. Fig. 3 depicts the same parameters as Fig. 2 for reforming temperatures from 600 to 1000
C but using a preheating temperature of 500
C. If the reforming temperature is 600
C, the O2/G ranges from 0.017 to 0.320, at a W/G of 10e23, respectively. When the W/G is less than 10, there is an excess of heat in the reformer, and above 24 all the glycerol fed to the reformer is burnt. For this reforming temperature, the O2/G mole ratio required is only 0.09 at a W/G mole ratio of 13, burning a fraction of product gas of 58.28%. If the reforming temperature was 800
C the O2/G ratio and the fraction of burnt product gas would be 0.986 and 72.22%, respectively. Nevertheless, by reforming at 1000
C, the O2/G mole ratio increases up to 1.70 at the same W/G, using 94.40% of the product gas. Thus, in this case, the upper W/G limit decreases as compared to the previous case owing to the high oxygen flow rate required. Since high reforming temperatures are recommended to achieve better performance in reforming and to minimize methanation, reforming and preheating temperatures of 800
C have been used as a base case for pure glycerol at 240 atm. In this case, a W/G mole ratio (maximum) of 18.54 is required and an O2/G mole ratio of 0.22, if there is no product gas leaving the system. For this base

Table 5 W/G and O2/G mole ratio ranges for different preheating and reforming temperature (pure glycerol at 240 atm); upper limits of W/G correspond to the maximum integer and O2/G is computed for these values. Reforming temperature Preheating temperature (
C)

W/G and O2/G ranges (mol/mol)

600 (
C)

700 (
C)

800 (
C)

900 (
C)

1000 (
C)

500

W/G O2/G W/G O2/G W/G O2/G W/G O2/G W/G O2/G W/G O2/G

10e23 0.017e0.320 Non an ATR case

3e20 0.032e0.797 7e20 0.013e0.390 19e20 0.012e0.025

3e17 0.319e1.20 3e17 0.136e0.843 4e18 0.039e0.544 8e18 0.013e0.216

3e15 0.624e1.478 3e15 0.422e1.163 3e16 0.257e0.905 3e16 0.104e0.606 5e17 0.035e0.317

3e13 1.028e1.699 3e13 0.837e1.428 3e14 0.674e1.217 3e14 0.520e0.965 3e14 0.368e0.711 3e15 0.212e0.459

600 700 800 900 1000

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F.J. Gutiérrez Ortiz et al. / Energy 42 (2012) 192e203

Fig. 3. O2/G mole ratio and fraction of product gas to be burnt versus W/G (up to the allowed limits), for different reforming temperatures when preheating the feed at 500
C.

case, Table 6 shows the overall energy balance; Table 7 gives the heat flows, mass flow rates and temperatures for all seven heat exchangers plus the heater located at the reformer inlet; and Table 8 shows the changes in exergy flows, work and heat flows, and exergy loss in the individual process units. The overall thermal efficiency, overall exergy efficiency, combined heat-power exergy efficiency and hydrogen yield are 31.36%, 29.46%, 32.59% and 45.03%, respectively. Fig. 4 shows how both the net power and turbine power increase when the reforming and preheating temperatures raises from 700 to 1000
C, when the exact W/G ratio required to burn all the product gas and the O2/G ratio to achieve null net heat flow in the reformer have been calculated. With a preheating temperature of 500
C, the same trends are also illustrated. Fig. 5 illustrates how the different efficiencies change when the reforming temperature increases by using a preheating temperature of 500
C, and Fig. 6 shows the trend of the same parameters when preheating at the reforming temperature. As can be observed, the three efficiencies always raise as reforming temperature increases when preheating at the reforming temperature or less (500
C). Likewise, the hydrogen yield increases as reforming temperature increases when preheating at the reforming temperature, but it reaches a maximum at about 900
C reforming temperature when preheating at 500
C. Furthermore, as shown below (Table 9), if the Table 6 Overall energy balance for autothermal reforming of glycerol using SCW for thermoneutral condition by burning the entire product gas (240 atm. 800
C reforming and preheating temperature). Work entering the system (kW) P1 P2 (C1eC5) þ inter. cool. C6

Enthalpy of the inlet-streams (kW) 48.69 0.17 15.78 19.62

Work leaving the system (kW) TURB

1478.14

OVERALL

1393.89

Glycerol Water Air Air3 CW1

2053.65 15,987.49 0.02 0.10 155,340.25

Enthalpy of the outlet-streams (kW) GC6 W1 CW4

7428.99 14,482.31 152,864.33 1394.13

preheating temperature is was 600e900
C, the H2 yield when reforming at 900
C would be always higher than that obtained by reforming at 1000
C, when preheating at the same temperature. This unexpected result cannot be explained merely by the higher O2/G value (and, hence, higher glycerol oxidation), since when reforming at 800
C and preheating at 500
C, the O2/G is clearly lower and the hydrogen yield is also lower than at 900
C. This is part of the explanation because, obviously, by adding more O2, more glycerol is consumed and less remains to be reformed. The CO2 molar flow rate resulting from the oxidation increases as O2/G increases (and as the reforming temperature also rises). It is verified that the H2-to-available-glycerol mole ratio (once the oxidized glycerol is discounted) is actually higher at 1000
C (4.85) than at 900
C (4.78), and at 900
C is higher than at 800
C (4.04); i.e., by taking the available glycerol, H2 yield does not reach a maximum, but continuously raises as the reforming temperature does. At 1000
C, H2 comes mainly from glycerol decomposition to CO and H2 (only 4 mol of H2 produced for each mole of glycerol) rather than from glycerol reforming (7 mol of H2 obtained for each mole of glycerol). Thus, the sum of CO2 and CO coming from decomposition and reforming of glycerol reaches a maximum at 900
C, after discounting the CO2 coming from glycerol oxidation and the CH4 formed by reactions (11) and (12) that remove CO and CO2 and also prevents more H2 production. The reason is because, at a lower temperature, a significant amount of CH4 is produced from CO and CO2 and at a higher temperature, more CO is produced because the water-gas shift reaction hardly takes place. Hence, maximum H2 yield is achieved at 900
C. On the other hand, when reforming and preheating at 1000
C, the H2O flow rate is higher and the O2 needed is considerably lower than when preheating at 500
C, since the process heat required is only for the reforming reaction and not for heating the feed up to 1000
C. Thus, the H2 production significantly increases with respect to the case of preheating at a lower temperature (even 900
C); less glycerol is oxidized, less CH4 is formed and, although CO2 from reforming is somewhat lower than at 900
C reforming and preheating temperatures, much more CO is produced, so the sum of CO2 and CO is higher, and hence the higher H2 yield. 3.3. Effect of preheating temperature under thermoneutral conditions At a given reforming temperature, and for a water-to-glycerol mole ratio higher than the minimum at which no O2 is needed in the reformer, the oxygen required in the reforming reactor increases when the aqueous solution of glycerol is fed at a temperature very different from the reforming temperature. Fig. 7 shows the O2/G mole ratio and the fraction of product gas to be burnt versus the W/G mole ratio, which varies within a range of lower and upper limits, when the reforming temperature is 800
C for inlet feed temperatures from 500 to 800
C. In this Figure, it can be observed that when the inlet feed temperature is 500
C, W/G mole ratio ranges from 3 to 17 and the O2/G mole ratio required for those W/G limits are 0.319 and 1.198, respectively. If the inlet feed temperature is 600
C, the limits for W/G and the corresponding O2/G ratio are from about 3 to 17 and 0.136 to 0.843, respectively. When preheating the feed up to 700
C, the maximum W/G ratio is about 18 and the O2/G needed is 0.544, which is lower than those obtained when preheating at 500 or 600
C. Fig. 8 illustrates how the turbine power and net power (after discounting the power consumption in pumps, compressors and fans) increase when the preheating temperature decreases for a given reforming temperature of 800, 900 or 1000
C, respectively. Although one might think that this result is unexpected, that is not so, since for a low preheating temperature more oxygen is needed

F.J. Gutiérrez Ortiz et al. / Energy 42 (2012) 192e203

199

Table 7 Heat flows, mass flow rates and temperatures for all the seven plus one heat exchangers for thermoneutral condition by burning the entire product gas (240 atm. 800
C reforming and preheating temperature). Heat exchanger

Q_ (kW)

(Tin)cool

(Tout)

HE01 HE02 HE03 HE04 HE05 HE06 HE07 HEATER

288.65 525.04 1716.73 19.36 2764.42 38.02 149.94 2113.71

28.67 82.99 183.34 194.39 25.00 60.00 36.40 394.81

82.99 183.34 394.81 387.50 92.03 107.69 123.09 802.39

(Tin)

cool

(Tout)

hot

92.03 390.49 1000.00 397.81 286.11 126.09 186.33 e

hot

85.00 186.33 397.81 390.49 60.00 110.69 126.09 e

_ cool (kg/h) m

_ hot (kg/h) m

4626.30 4626.30 4626.30 322.42 35,234.42 1632.58 6123.66 4626.30

35,234.42 7756.24 7756.24 7756.24 4948.71 7756.24 7756.24 e

Temperatures in
C.

Table 8 Changes in exergy flows, work and heat flows, and exergy loss in the individual process units for thermoneutral condition by burning the entire product gas (240 atm, 800
C reforming and preheating temperature). Equipment

DB_ (kW)

_ (kW) W in

MIX1 P1 HE01 HE02 HE03 R1 TURB HE05 SEP HE6 C6 HE7 Furnace HE4 P2 COMP þ cool-inter. Heater Total exergy losses

25.80 38.98

48.69

474.90 1862.70 0.00 1.73E-04 14.30

_ out (kW) W

1478.14

Q_ out (kW)

DS_ cool (kW/K)

DS_ hot (kW/K)

288.65 525.04 1716.73

288.65 525.04 1716.73

0.88 1.30 2.96

0.80 0.94 1.81

0.00 2764.42

0.00 2764.42

8.37

7.31

38.02

38.02

0.11

0.10

149.94 0.00 19.36 0.00 0.00 2113.71

149.94 2113.71 19.36 0.00 57.34 0.00

0.43

0.35

0.03

0.03

19.62

2910.75 0.13 46.53 1329.02

Q_ in (kW)

0.17 73.12 0.00

B_ lost (kW) 25.80 9.71 24.33 105.71 342.91 474.90 384.56 314.95 0.00 2.84 5.32 23.03 1292.10 1.69 0.03 40.04 289.70 3337.62

and more air must enter the reformer. Thus, the product gas mass flow rate will rise and more power will be obtained in the expander. Once most of the water present in the product gas is condensed, this is completely consumed in the furnace. Although the compression power also increases, the net power increases when the inlet feed temperature is reduced. Table 9 summarizes the exergy analysis of the pure glycerol simulations, showing the overall thermal efficiency, the overall exergetic efficiency, the combined heat-power exergy efficiency and the hydrogen yield. It can be observed that the energy, exergy and combined efficiencies increase as the reforming temperature rises when the feed is preheated to the reforming

temperature. Likewise, these efficiencies increase when the preheating temperature decreases at a given reforming temperature, but conversely, the hydrogen yield drops when the preheating temperature is low.

Fig. 4. Turbine work and net work flows versus the reforming temperature when preheating at reforming temperature and preheating at 500
C (pure glycerol, thermoneutral condition and 240 atm).

Fig. 5. Energy, exergy and combined efficiencies and hydrogen yield for different reforming temperatures by using a preheating temperature of 500
C (pure glycerol, thermoneutral condition).

3.4. Effect of the methanol concentration in the crude glycerol under thermoneutral conditions When pretreated crude glycerol is fed into the system, the molar flow rate of the feed is increased with respect to the pure glycerol feed as MeOH concentration increases, since the total mass flow

200

F.J. Gutiérrez Ortiz et al. / Energy 42 (2012) 192e203

Fig. 6. Energy, exergy and combined efficiencies and hydrogen yield for different reforming temperatures when preheating at the reforming temperature (pure glycerol, thermoneutral condition).

rate is the same but the molar weight of methanol is lower than that of glycerol. As a result, 24.2 mol% is MeOH when a crude glycerol with MeOH 10 wt% is fed into the system. Thus, both the W/G and the O2/G mole ratios will be lower for pretreated crude glycerol in the same operating conditions when the purity of glycerol decreases (MeOH concentration increases), although the difference in the O2/G mole ratio becomes insignificant when the preheating and the reforming temperatures are the same, since the absolute value of O2/G is very low in these cases. Conversely, the flow rates of both water and air (and hence, oxygen) fed to the reformer will rise with respect to those required in the pure glycerol case, thus increasing the total flow rate entering and leaving the reformer. Although this increase in total mass flow rate is not large, the machines’ power (turbine, compressors and pumps) increase in such a way that the total net power is slightly higher than with pure glycerol when operating in the same conditions. Likewise, the exergy losses are slightly higher but the computation of exergy and combined efficiencies give somewhat higher values

Fig. 7. O2/G mole ratio and fraction of product gas to be burnt versus W/G (up to the allowed limits), when the reforming temperature is 800
C, for inlet feed temperatures from 500 to 800
C.

for crude glycerol (between 0.2 and 0.5%, for a MeOH content of 10 wt%). These differences are more marked when the methanol concentration in the feed is increased. When the preheating temperature is much lower than the reforming temperature (200
C or more), the exergy losses are lower with crude glycerol; then the exergy and combined efficiencies increases between 1.0 and 1.2%, for 800 and 1000
C reforming temperatures, when the MeOH content is 10 wt%. If this concentration increases to 30 wt%, those efficiencies raise to 1.2 and 1.4%, respectively, for the reforming temperature range. Fig. 9 shows the trends of net power and turbine power as a function of methanol content in the crude glycerol feed at different reforming temperatures from 700 to 1000
C, when the preheating and reforming temperatures are the same. If MeOH content is 10 wt% these powers are 2.4% higher than when using pure glycerol; by using 20 and 30 wt%, the increases in power are 4.7 and 7.2% higher than for pure glycerol. Fig. 10 illustrates the same but using a preheating temperature (500
C) different from the reforming temperature. The increases in power

Table 9 Water to glycerol mole ratio, oxygen to glycerol mole ratio, exergy losses, turbine and net power; energy, exergy and combined efficiencies and hydrogen yield for different reforming and preheating temperatures. B_ lost (kW)

_ out (kW) W

Wnet (kW)

hThermal (%)

hExergy (%)

hcombined (%)

hH2 (%)

23.15 0.32 Non autothermal case

3526.63

1333.41

1231.10

27.70

25.88

29.49

19.65

20.11 20.50 20.83

0.80 0.39 0.04

3547.21 3458.65 3381.85

1502.09 1427.49 1361.91

1397.80 1327.26 1284.67

30.95 29.86 28.91

27.94 27.73 27.53

31.17 31.07 30.96

30.38 31.79 32.88

17.52 17.85 18.20 18.54

1.20 0.88 0.54 0.22

3552.25 3482.62 3408.75 3337.62

1658.47 1601.23 1539.21 1478.14

1510.72 1473.91 1433.73 1393.89

33.99 33.16 32.26 31.36

29.84 29.74 29.61 29.46

32.77 32.72 32.66 32.59

38.00 40.65 43.04 45.03

15.58 16.01 16.40 16.78 17.12 temperature of 1000
C 13.63 14.02 14.38 14.72 15.03 15.37

1.48 1.21 0.90 0.60 0.32

3547.11 3480.22 3416.17 3353.72 3297.03

1777.29 1721.25 1666.20 1611.00 1559.48

1615.02 1578.79 1542.65 1506.09 1471.60

36.34 35.52 34.71 33.89 33.11

31.29 31.21 31.11 30.99 30.86

34.03 34.00 33.94 33.87 33.78

39.50 43.87 48.36 52.32 55.75

1.70 1.43 1.22 0.96 0.73 0.46

3527.73 3467.54 3414.66 3360.78 3307.82 3255.71

1886.96 1837.87 1793.91 1749.09 1704.21 1659.28

1714.19 1682.50 1654.07 1624.86 1595.52 1565.74

38.57 37.86 37.22 36.56 35.90 35.23

32.70 32.67 32.63 32.59 32.54 32.47

35.30 35.29 35.29 35.27 35.24 35.19

35.65 41.45 45.83 51.01 55.68 60.75

Preheating temperature (
C) Reforming 500 600 Reforming 500 600 700 Reforming 500 600 700 800 Reforming 500 600 700 800 900 Reforming 500 600 700 800 900 1000

W/G ratio (mol/mol)

O2/G ratio (mol/mol)

temperature of 600
C

temperature of 700
C

temperature of 800
C

temperature of 900
C

F.J. Gutiérrez Ortiz et al. / Energy 42 (2012) 192e203

201

Fig. 8. Turbine work and net work flows versus the preheating temperature when reforming at 800, 900 and 1000
C (pure glycerol, thermoneutral condition and 240 atm).

Fig. 10. Turbine work and net work flows versus the reforming temperature when preheating at 500
C (pretreated crude glycerol feed with a methanol content of 10, 20 and 30 wt%, thermoneutral condition and 240 atm).

are 2.2, 4.7 and 7.0% for 10, 20 and 30 wt% crude glycerol, respectively, with respect to pure glycerol. Fig. 11 depicts the turbine power and net power versus the methanol concentration in the crude glycerol feed at different preheating temperatures for a given reforming temperature of 800
C. The above increases, as compared to those in Fig. 8, are 2.2e2.4% for a preheating temperature range of 500e800
C when MeOH content is 10 wt%, and 4.8e4.9% or 7.0e7.2% if the MeOH concentration is 20 or 30 wt%, respectively. Regarding the energy and thermal efficiency, the differences between pure and crude glycerol are still less, and are about 0.2e0.3% higher for a MeOH content of 10 wt% as compared with pure glycerol for any preheating temperature at a given reforming temperature. The increases in these efficiencies are 0.3e0.5% or 0.4e0.6% when using 20 or 30 wt% MeOH. The increase in H2 yield observed relative to that of pure glycerol (about 0.15, 0.50 and 1.00% for 10, 20 and 30 wt% MeOH content) was expected since methanol reforms better than glycerol under the same conditions. Likewise, since the total molar flow rate is higher than when using pure glycerol, the molar flow rate of H2 increases when the methanol concentration increases. This makes more chemical energy available to the product gas and the heat released from the furnace to the heater at the inlet of the reformer is much higher. As shown, the above differences are accentuated when the MeOH concentration in the feed is increased. Fig. 12 illustrates the different efficiencies as the MeOH content varies versus the reforming temperature using a preheating temperature of 500
C, and Fig. 13 shows the same but when preheating at the reforming

temperature. Increases in efficiencies are very small again. These two Figures show trends qualitatively similar to those found and described for pure glycerol (Figs. 5 and 6).

Although thermodynamically it would be convenient to operate at high W/G mole ratios to maximize the hydrogen yield [17], when integrating the process to approach actual conditions in a plant, it is seen that this mode of operation is not viable. A subsequent aim is to maximize the power production but minimizing the process integration complexity. For this purpose, the methanol content in crude glycerol is beneficial for the process since MeOH reforms better than glycerol and exergy and combined efficiencies are higher. However, the differences with respect to the results obtained for pure glycerol are not as relevant as using methanol in reforming instead of biodiesel production. Thus, as much MeOH as possible should be returned to the biodiesel production process. Operation at high reforming temperatures and low preheating is the more convenient mode since it allows to obtain maximum net power and energy and exergy efficiencies. However, operation at excessively high temperatures (900 and 1000
C) involves using exotic materials to build the devices, which are difficult to find and very expensive. Therefore, it seems to be recommendable to operate at temperatures no higher than 800
C.

Fig. 9. Turbine work and net work flows versus the reforming temperature when preheating at reforming temperature (pretreated crude glycerol feed with a methanol content of 10, 20 and 30 wt%, thermoneutral condition and 240 atm).

Fig. 11. Turbine work and net work flows versus the preheating temperature when reforming at 800, 900 and 1000
C (pretreated crude glycerol feed concentration for a methanol content of 10, 20 and 30 wt%, thermoneutral condition and 240 atm).

3.5. Optimal conditions for maximum work under thermoneutral conditions

202

F.J. Gutiérrez Ortiz et al. / Energy 42 (2012) 192e203

Fig. 12. Energy, exergy and combined efficiencies and hydrogen yield versus reforming temperature when preheating at 500
C (pretreated crude glycerol feed with a methanol content of 10, 20 and 30 wt%, thermoneutral condition).

study [17] were tested first, but the resulting process is not feasible from the energy point of view since there is a severe energy deficit. Another procedure has been then tried, in which the water and the oxygen entering the reformer are calculated to achieve thermoneutral conditions aided by the energy coming from the combustion of all the product gas in a furnace. The constraint imposed is null net heat within the reformer. The preheating temperature of the inlet feed (water plus glycerol) and the reforming temperature were taken as the main parameters to change. Energy and exergy have been assessed for numerous cases, illustrated in different Figures and Tables, both for pure and pretreated crude glycerol. One unpredicted result comes from the preheating temperature, which is that better exergy efficiency and higher net power are achieved when using a low preheating temperature. The next paper of this series will be aimed at maximizing hydrogen production after purifying the product gas not burnt in the furnace in order to use it in a fuel cell to obtain supplementary electrical power apart from that obtained from the expander. As described in the previous study mentioned above, the studied ATR process using SCW may be realistic and experimental investigation is required to better understand this process. Acknowledgment This research is supported by the Science and Technology Ministry of Spain under the research project ENE2009-13755, as a Project of Fundamental Research inside the framework of the National Plan of Scientific Research, Development and Technological Innovation 2008e2011. Nomenclature B_ H_ LHV _ LW _ m Q_ S_

Fig. 13. Energy, exergy and combined efficiencies and hydrogen yield versus reforming temperature when preheating at the reforming temperature (pretreated crude glycerol feed with a methanol content of 10, 20 and 30 wt%, thermoneutral condition).

T _ W Greek

h 4. Conclusions Taking the previous work [17] as a starting point, a conceptual design of the autothermal reforming of glycerol using supercritical water has been established. The selected scheme makes it possible to compute and use the overall product gas to revert its chemical energy in the process itself. In this way, the energy that can be taken out of the process comes from the expansion of the product gas in an expander; i.e., pressure energy converts into electrical power. With this aim, a sensitivity analysis has been carried out for four possible glycerol feeds to the reformer: pure glycerol and pretreated crude glycerol with a MeOH content of 10, 20 and 30 wt%. The mass flow rate of the glycerol feed has been always 1000 kg/h. The operating conditions of the process have been assessed by a complete analysis of energy and exergy, carried out for every component of the process. AspenPlusÔ has been used for computation, using the predictive SoaveeRedlicheKwong equation of state as the thermodynamic method in the simulation of the supercritical region. The optimal conditions found in a previous

Exergy flow (kW) Enthalpy flow (kW) Lower Heating Value (kJ/kg) Lost work flow (power) or exergy flow loss (kW) Mass flow rate (kg/s) Heat flow (kW) Entropy flow (kJ/(s K)) Temperature (K) Work flow or Power (kW)

Efficiency (%)

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