Experimental performance of a double-lift absorption heat transformer for manufacturing-process steam generation

Energy Conversion and Management 148 (2017) 267–278

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Experimental performance of a double-lift absorption heat transformer for manufacturing-process steam generation Arnas Lubis ⇑, Niccolo Giannetti, Seiichi Yamaguchi, Kiyoshi Saito, Naoyuki Inoue Department of Applied Mechanics and Aerospace Engineering, Waseda University, Shinjuku-ku, Tokyo 169-8555, Japan

a r t i c l e

i n f o

Article history: Received 1 March 2017 Received in revised form 29 May 2017 Accepted 30 May 2017

Keywords: Absorption Heat transformer Double-lift Waste heat Steam generation

a b s t r a c t As widely known, some industrial processes produce a large amount of waste heat while others require a large amount of steam to heat the process flow. The main difference involves the temperature level of these heat quantities. Absorption heat transformers play a strategic role in waste heat recovery and heat supply to manufacturing processes due to their ability to utilize heat at a certain temperature level and release the enthalpy of mixing of the refrigerant at a different temperature level with a negligible amount of mechanical work input. However, given the lack of examples that find application as operative plants, the feasibility of the technology is questioned in academic and technical domains. In this study, the operability of a double-lift absorption heat transformer that generates pressurized steam at 170 °C is studied across a full range of operative conditions. The results demonstrate and clarify the manner in which the system can operate steadily and efficiently when driven by hot water temperature at approximately 80 °C while safely generating steam at a temperature exceeding 170 °C. The conditions yielding maximum system efficiency and capacity are identified, and the obtained experimental results are used to define an optimal control strategy. Ó 2017 Elsevier Ltd. All rights reserved.

1. Introduction In a manner similar to several other countries, Japan has developed a cautious awareness with respect to global environmental issues that has led to ecological strategies to drastically reduce CO2 and air pollutant emissions and limit the effects of global warming. Industrial emissions and residual heat release are responsible to a significant extent for the fore-mentioned environmental issues. Conversely, addressing these problems with innovative, clean, and efficient solutions can significantly contribute in the realization of an environmentally compatible, technologically advanced, and economically rewarding industry. In May 2007, as a part of the Japanese environmental strategy against global warming, a plan termed as ‘‘Beautiful star 50 (Cool Earth 50)” was stipulated by stating the long-term goal of ‘‘having the global greenhouse emissions halved by 2050 as compared to the current situation” [1]. More recently in 2013, a survey performed by the Ministry of Economy [2] pointed out that over 40% of the national energy consumption is related to industrial processes. The development of steam generation heat pumps is cited as one of the most promising elements to reach this objective since it simultaneously

⇑ Corresponding author. E-mail address: [email protected] (A. Lubis). http://dx.doi.org/10.1016/j.enconman.2017.05.074 0196-8904/Ó 2017 Elsevier Ltd. All rights reserved.

addresses waste heat recovery and clean energy provision for industry. Steam is widely used for several industrial processes with cross manufacturing procedures from heating to drying purposes. The process-steam is conventionally generated with boilers that then reject a large amount of waste heat as hot water between 40 °C and 100 °C and/or exhaust gases at 100 °C–250 °C [3–5]. In this context, heat pump technology represents a critical possibility of recovering and reutilizing this amount of heat as a source to generate steam that can be reintroduced in an industrial productive route and to further lower the release of heat and pollutants into the environment. Parallely, this technology can be directed towards the utilization of heat available from renewable sources, such as solar or geothermal energy, at equivalent temperature levels. A previous study [6] indicated an effective solution to utilize solar thermal energy with an air-conditioning system that could be used throughout the year and being installed in a tropical country. The refrigerant stability and compressor oil durability at high temperatures limits the employment of vapour compression type heat pumps for heat recovery purposes to condensation temperatures below 120 °C [7–9]. Absorption heat transformers are heat-driven heat pumps that function based on the possibility of utilizing the enthalpy of

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A. Lubis et al. / Energy Conversion and Management 148 (2017) 267–278

Nomenclature Q T COP _ m h

e

heat transfer rate (kW) temperature (°C) coefficient of performance (–) solution mass flow rate (kg/s) enthalpy (kJ/kg) temperature effectiveness

Subscript AH high temperature absorber AL absorption evaporator C condenser

condensation of a refrigerant into an appropriate absorptive solution. This enables a system to realize the operative pressure jump of the refrigerant at a liquid state with low mechanical power requirements. Several extant studies involved theoretical and experimental investigations on different configurations and applications. A previous study [10] discussed the effectiveness of these systems with under different criteria to attain meaningful comparison terms. A more recent study [11] defines a thermodynamic criterion for a preliminary design of optimized three-thermal heat transformers and a method to perform the diagnostics of existing plants to provide guidelines for their performance improvement. An extant study [12] presented a thermodynamic model for irreversible heat transformers operating between four sourcetemperatures and included the effect of heat leaks, heat resistance, and internal irreversibility given a Newtonian heat transfer law. This was followed by a more general model [13] that was developed by considering a generalized heat transfer law to study the relation between coefficient of performance (COP) and heating load. New fluid pairs [14] were proposed to alleviate some of the main technical problems of the well-established lithiumbromide/water pair and different applications for achieving efficient waste heat recovery and broadening the application field of the technology (for further details refer to Ref. [15]). The possibility of using heat sources at lower temperatures to generate hightemperature steam would significantly enhance the usefulness and broaden the implementation field of these systems. However, multiple-lift configurations are needed to achieve temperatures above 120 °C, and except for [16], research effort in this area is limited to a theoretical standpoint in which the performance of different configurations are investigated with lumped parameters models [17–23]. Thus, given the lack of examples that find application as operative plants or prototypes, the feasibility and reliability of the technology is questioned in academic and technical domains. The technical and economic feasibility of single-effect heat transformers were demonstrated by early studies. However, more advanced configurations encounter various technical challenges, and it is necessary to overcome the same before they can be operationalised as market alternatives. The present study demonstrates the real performance of a double-lift heat transformer operated with a lithium bromidewater mixture. The prototype overcomes the main technical challenges that have limited the application cases of this technology, showing that the system can operate steadily and efficiently when driven by hot water at approximately 80 °C while reaching output steam temperatures above 170 °C. The experimental performances are analysed with respect to the main operative parameters and control variables in a manner that is useful in indicating the best condition for a system recovering from heat discharged from a real plant and to clarify the functioning of the main components.

EH EL G in out s S SL SH RL RH

refrigerant separator evaporator low temperature generator inlet outlet solution steam separator low temperature solution heat exchanger high temperature solution heat exchanger low temperature refrigerant heat exchanger high temperature refrigerant heat exchanger

2. System description Vapour absorption is among methods of refrigeration that were initially used widely. Therefore, given that absorption refrigeration machines are used for a long time, this technology has a broad albeit inconclusive theoretical background [24–26]. Absorption heat transformers, or absorption second type heat pumps achieve their final objective by means of the same thermo-chemical processes as those of an absorption chiller albeit occurring with a diverse rearrangement of the same components. Given this viewpoint, single- and multiple-lift heat transformers can be studied and developed as an extension of the fore-mentioned well-rooted counterpart. A single-stage heat transformer can raise the temperature of hot water from 90 °C to 120 °C, and actual machines have tested the same [27]. 2.1. Basic cycle An initial systematic assessment of the ideal performance of these systems constitutes the starting point for a comparative evaluation of the experimental performance of an actual plant. Preliminarily, a steady endo-reversible single-stage absorption heat transformer (Fig. 1a) is considered based on a systematic approach presented in a previous study [28] wherein heat transfer through the heat exchangers occurs isothermally, with zero temperature difference and heat losses towards the external environment. The effects of potential and kinetic energies of the refrigerant are neglected, the circulating solution amount is assumed to be infinite, and the Dühring rule (Eq. (4)) is applicable yielding a linear relationship between the temperature and saturation temperature of the solution. A second type absorption heat pump (heat transformer) is driven by the intermediate temperature level heat delivered to a generator QG. Refrigerant vapour extracted is subsequently condensed at a lowest temperature source QC, whereas the rich liquid solution is pumped to a higher pressure level at low work expense. In which the refrigerant evaporated at the same pressure level (requiring QE as an input) can be absorbed by lowering the solution concentration (intended as the lithium-bromide mass fraction) and releasing heat QA at a higher equilibrium temperature TA. The efficiency of the resulting combination of endo-reversible cycles is defined as follows:

COP ¼

QA QG þ QE

ð1Þ

The algebraic form of first and second laws of thermodynamics can be expressed as Eqs. (2) and (3), respectively, as follows:

QG þ QE ¼ QA þ QC

ð2Þ

A. Lubis et al. / Energy Conversion and Management 148 (2017) 267–278

269

Fig. 1. A schematic of a single-stage heat transformer (a) and simplified Dühring and T-s diagrams (b).


n TH TC

QG QE QA QC þ ¼ þ TG TE TA TC

ð3Þ

TA ¼ TH

TE TA ¼ TC TG

ð4Þ

Finally, the general efficiency of a cycle with n effects is expressed as follows [29]:

Eq. (5) gives the efficiency of this cycle as follows:

COP ¼

TG 1 ¼ T G þ T C 1 þ TT C G

ð5Þ

Given that the Dühring rule constitutes a valid approximation on a wider equilibrium temperature range, previous studies [28–29] extended the same manner of reasoning to a multi-stage cycle as shown in Fig. 2. The concept is expounded by assuming that the heat generated by a single-stage cycle can be used to achieve a higher evaporation temperature TEH, which leads to a higher equilibrium temperature TAH of an additional absorber. Dühring rule is generalized to a generic number of effects (Eq. (6)) by assuming that the temperature difference between the generator and the evaporator is negligible (TEL = TG = TH) as follows:

Fig. 2. The simplified Dühring diagram of a double-lift heat transformer based on previous studies [28–29].

COP ¼ 1þ

ð6Þ

1 n  i X T

ð7Þ

C

TH i¼1

The thermodynamic endo-reversible approach indicates the upperbound of cycle efficiency and maximum temperature lift, and these values are further reduced by irreversible transfer processes and real characteristics of the absorptive medium.

2.2. Experimental system Sterilization processes [30–32] require saturated steam at temperatures at approximately or above 150 °C. The plant described and tested in [16] is suitable for this application case, being able to generate a maximum steam temperature of 150 °C when operating with inlet cooling water at 23 °C and inlet hot water at 89 °C. However, in manufacturing processes, steam at a saturation pressure of 0.8 MPa (170.4 °C) is needed in order to be reintroduced in the production line. These requirements can be achieved by operating sufficiently beyond the crystallization limit and with acceptable safety factors by means of the double-lift cycle target of this analysis. As shown in Fig. 3, the refrigerant is evaporated inside the absorber and sent directly to the high temperature absorber (AH) once it is separated from the liquid in the refrigerant separator (EH). The selected system configuration delivers the concentrated solution from the generator to the high-temperature absorber (AH in Fig. 3). This enables the system to maximize the temperature of the output heat at AH [16] as a result of releasing the heat of absorption of the vapour from the high temperature evaporator (EH). The solution then enters the first absorber (AL) at an intermediate concentration. The hot water passes through the generator (G), and subsequently the first evaporator (EL), while the cooling water extracts heat at the condenser (C). The design conditions of the experimental prototype (Fig. 4) are specified in Table 1.

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Refrigerant vapour

Separator (EH)

Solution Heat Exchanger (SH)

Refrigerant vapour

Evaporator (EL) Refrigerant liquid

QE

QC

Refrigerant vapour

QA

Absorber (AH)

Absorber (AL)

Refrigerant liquid Solution Heat Exchanger (SL) Refrigerant vapour

Condenser (C)

Generator (G)

QG

Fig. 3. Conceptual configuration of the double-lift heat transformer.

According to the previous idealized calculation, a system that operates between 25 °C at the condenser and 88 °C at the generator is bounded by an ideal COP corresponding to 0.399. A small-scale test apparatus (Fig. 4) was built based on the forementioned specifications to clarify the characteristics and verify the potential of the cycle as well as to eventually optimize the operability of an industrial scale plant. A 200 kW system was then constructed as a first practical application case [33]. The geometri-

Table 1 Design condition of the prototype. Steam temperature Heat output Hot water inlet temperature Cooling water inlet temperature Solution charge Refrigerant Absorbent Additive

170 14 88 25 30.0 Water Aqueous LiBr 2-Ethylhexanol

°C kW °C °C kg

cal features involved in the small scale manufacturing of the prototype are reported in Table 2. Falling film heat exchangers are employed to realize a high transfer coefficient with negligible pressure losses. The generator and the evaporator consist of enhanced tubes with end-cross outer surfaces [34]. Surfactant additives are used to promote transfer rates by causing surface gradient-induced convection and mixing (also referred to as ‘‘Marangoni convection” [35]). The liquid distributer is appositely designed to uniformly drop a fluid on the transfer surface with minimum pressure losses in a wide flow rate range. 3. Instrumentation

Fig. 4. A double-lift absorption heat transformer.

The system was instrumented to collect the information required for a comprehensive characterization of the system performance in a wide range of the main operative parameters and to understand the system response to different boundary conditions. The specifications of the measurement equipment and the accuracy of the sensors are reported in Table 3. The system flow diagram is shown in Fig. 5. The vapour desorbed in the generator G is condensed in C. It passes through the low-temperature refrigerant heat exchanger (RL) and the hightemperature refrigerant heat exchanger (RH) and is pumped to the low temperature absorber where it gains the heat of absorption released on the solution side. The high-temperature evaporator is driven by the output heat of the low stage absorber AL. The liquid phase of the refrigerant in EH is delivered to EL and passes through RH. At this point, part of the refrigerant is evaporated and absorbed by the solution into AL. The remaining liquid refrigerant reaches the condenser after passing through RL. From the viewpoint of the solution cycle, the concentrated solution from G is preheated in the low-temperature (SL) and high-temperature (SH) solution heat exchangers prior to reaching the high temperature absorber (AH) where it absorbs the vapour from EH, is diluted, and produces

271

A. Lubis et al. / Energy Conversion and Management 148 (2017) 267–278 Table 2 Features of the prototype heat exchangers. Component

C

GL

EL

AL

AH

m

Condensation Cu baretubes 3.00

Falling film Cu end-cross tubes 3.30

Falling film Cu end-cross tubes 2.80

Falling film Cu baretubes 3.00

Falling film SUS baretubes 2.00

m2

2.77

3.05

2.46

2.77

1.75

Type Outer transfer area Inner transfer area

2

Table 3 Specifications and accuracy of measurement instruments. Item Mass flow rate Volumetric flow rate

Pressure

Value Solution Refrigerant Water Steam Steam HT Absorber (6) Absorbing evaporator (8) LT Generator (2)

Temperature Density of solution Height level

±0.1% ±1.0% ±0.5% ±1.0% ±5 kPa ±1.25 kPa ±0.125 kPa ±0.25 kPa ±0.5 K ±0.001 kg/L ±0.5%

the heat output to the steam generation circuit. The steam generation circuit is constructed in a manner that allows the operator to circulate the fluid either by natural or forced circulation. The water is preliminarily preheated with the hot water in the heat exchanger BHX. The acquisition system is connected to thermo-resistors at the inlet and outlet ports of each component and flow meters to evaluate local and global transfer performance. These in conjunction with pressure gauges and level sensors can be used to directly monitor and promptly control the system. The system responses to different hot water and cooling water temperatures are investigated to analyse different boundary conditions with respect to the outer environment. Additionally, the solution mass flow rate that, which strongly affects transfer coefficients, size of the solution pump, and required transfer surfaces of the heat exchangers, is changed to identify maximum efficiency or maximum capacity conditions. After the liquid level, temperature, and pressure at each component are stabilized, stationary data are continuously acquired for approximately 15 min and reduced to their time averaged values. The PV (steam pressure control-valve) opening is regulated by a PID controller to target a given value of the steam pressure. The boiler feed water flow rate is controlled by adjusting the BP frequency to maintain a constant liquid level at S. The frequency of the SP is controlled by the PID controller to maintain a constant liquid level at AH. The level at AL is managed by using the PID controller to control the opening of the SV. The solution mass flow rate is indirectly adjusted by acting on the needle valve Snv at the outlet port of the AH. Similarly, the frequency of the RP is controlled to maintain a constant liquid level at EL, and the level at EH is managed to control the opening of the RV. Table 4 shows the measured values at the design condition. 4. Thermodynamic consideration As previously mentioned, the target application of the system is defined in terms of achievable steam saturation pressure (0.8 MPa or about 170 °C) as a result of the release of the absorption heat of the refrigerant in the high temperature absorber. This can be adjusted around the design point based on the operative and boundary conditions. The temperature at the evaporator

(EL) is mainly related to the temperature of the driving heat source and the circulation path of the hot water circuit. The temperatures of low temperature absorbers (AL) and high temperature absorber (AH) are related to the solution concentration, saturation temperature (or pressure) of the refrigerant at EL and EH. Boundary conditions (with respect to the temperatures of the generator and condenser) can vary from case to case based on the plant’s characteristics or weather conditions. Conversely, functioning around certain temperatures of the heat sources and sinks can be controlled through the solution mass flow rate and finally set as an overall management strategy. This latter parameter provides an additional degree of freedom to the system and significantly influences the transfer phenomena of the fundamental processes constituting the system thermodynamic cycle, and consequently affects the effectiveness of the transfer surfaces of each component and the size of the whole machine with the exception of the work required for the solution pump. Accordingly, experiments were conducted to cover the range of parameters as listed in Table 5. The results are reported and discussed in the following paragraph. 5. Experimental results Experimental results collected by operating the small scale prototype confirmed the feasibility and efficiency of the system. Specifically, the operability of the high temperature absorber and the system as a whole could be safely controlled by avoiding the occurrence of corrosion, crystallization, and/or air infiltration throughout the whole experimental campaign. These technical challenges were previously recognized as responsible for low durability and limited controllability of heat transformers with higher gross temperature lifts. The technical measures adopted in the present study as a response to durability issues include corrosion durability tests on the components material that comply with the highest safety standards for a pressurized vessel. Additionally, preliminary tests were conducted to define a ladder sequence for a safe start-up as well as steady functioning and shut-down of the system to be integrated with an automated algorithm to ensure stable controllability. Furthermore, as presented in a previous study [33], safe and sealed operability was verified on a practical scale plant with a heat capacity of 200 kW at corresponding operative conditions. The system operability (Figs. 6–8) is investigated through the experimental results obtained using a small-sized prototype. A parametric study was conducted with respect to the solution mass flow rate, the inlet cooling water, and the inlet hot water temperatures. The COP and heat transfer rate of the experimental results are calculated by using Eqs. (8) and (9). The thermo-physical properties of water (enthalpy and density) are obtained through the polynomial functions of ASHRAE handbook [36].

COP ¼

Q AH Q G þ Q EL

ð8Þ

A. Lubis et al. / Energy Conversion and Management 148 (2017) 267–278

Refrigerant separator

Safety and relief valve

HT Absorber

5 6

S

PV

B4

B3

EH

Steam pressure control valve

Steam separator

To the tank

AH

H5

R5

7'

R7

Hot water RL

Refrigerant low heat exchanger

H2

9

R8

Cooling water

LT Generator

R9

1

2

G

CP B2

C3

Hot water

H2'

C

CS1

Pressure sensor Volumetric flow meter Level meter Coriolis mass flow and density meter

C1

3

RP Refrigerant pump

BP

Temperature sensor

Hot water

R3

H4 CS2

B1 VH2'

H1 VH

Rnv

B5

Feed water

VH2

4

RV

AL

EL

Vacuum Condenser valve

VR1

C2

8

H3

VS1

BHX heat exchanger

Absorbing evaporator Snv

Evaporator

VR2

R4

Solution high SH heat exchanger

7

SL Solution low heat exchanger

RH

Rnv2

Refrigerant high heat exchanger

R6

BHX2 heat exchanger

272

SP Solution pump

Refrigerant temperature

7

6 5

9

8

4 1 2 3

Solution concentration Fig. 5. System flow diagram.

_ in  hout Þ Q ¼ mðh

ð9Þ

The effect of each parameter on the Dühring diagrams of the cycle is also illustrated to better understand the steady characteristics of this system. Fig. 6 shows the effects of the hot water inlet temperature on the components and the performance of the global system. Specifically, Fig. 6a highlights the capability of the system to reach the target temperature over the entire range of hot water inlet temperatures while the corresponding operative pressure level required at the high temperature absorber is maintained (Fig. 6c). Operability at higher inlet generator temperatures also shifts the saturation pressure level of the low temperature evaporator (EL) and absorber (AL) to higher values and increases the absorptive capability for a given cooling temperature value and the system capacity. Fig. 6b illustrates the evidence for higher absorption rates (i.e. larger solution concentration differences) at higher hot water temperatures, augmenting the amount of refrigerant circulating in the system and correspondingly leading to higher heat transfer rates

(Fig. 6f). Strong, middle and weak solution concentration refer to the lithium bromide mass fraction measured respectively at C1, C2, and C3 in Fig. 5. Finally, the saturation pressure of the condenser rises correspondingly. Additionally, the solution- and refrigerant-internal heat exchangers improve their thermal effectiveness (Eqs. (10) and (11), respectively) with increases in hot water temperature. Particularly, the parameter mainly affects the refrigerant heat exchanger (Fig. 6h) and exerts a modest influence on the solution heat exchanger (Fig. 6d).

eSL ¼

T4  T3 T9  T3

eRL ¼

T R8  T R9 T R8  T R3

eSH ¼

T5  T4 T7  T4

eRH ¼

T R6  T R7 T R6  T R4

ð10Þ

ð11Þ

Fig. 6 also describes the ultimate objective of the system in terms of heat transfer rate in the high temperature absorber, steam generation rate (Fig. 6g), and efficiency of the whole cycle (Fig. 6e). Given

273

A. Lubis et al. / Energy Conversion and Management 148 (2017) 267–278 Table 4 Measurement points at design condition (Table 1). Measurement point

Value

Unit

Temperature

Measurement point

Value

Unit

Measurement point

H5

90.0 ± 0.9

°C

1 3 4 5 7

87.8 ± 0.9 80.7 ± 0.9 119.8 ± 0.9 168.4 ± 0.9 172.2 ± 0.9

°C °C °C °C °C

S1 S2 S3 S4 S5

25.9 ± 0.9 87.9 ± 0.9 164.1 ± 0.9 171.0 ± 0.9 171.1 ± 0.9

°C °C °C °C °C

Volumetric flow rate VR1 VR2 VH VH2 VH20

70 9 R3 R4 R5

126.3 ± 0.9 123.1 ± 0.9 31.5 ± 0.9 66.6 ± 0.9 107.1 ± 0.9

°C °C °C °C °C

Pressure S 6 8 2

826.3 ± 8.7 218.6 ± 2.2 45.7 ± 0.4 4.4 ± 0.2

kPa kPa kPa kPa

R6

122.9 ± 0.9

°C

R7 R8 R9 H1 H2 H20 H3 H4

68.6 ± 0.9 78.4 ± 0.9 35.5 ± 0.9 88.0 ± 0.9 82.0 ± 0.9 78.8 ± 0.9 80.2 ± 0.9 79.7 ± 0.9

°C °C °C °C °C °C °C °C

Mass flow rate C1 C2 C3 Density C1 C2 C3

15.786 ± 0.027 15.390 ± 0.027 14.950 ± 0.026

kg/min kg/min kg/min

1.709 ± 0.002 1.642 ± 0.002 1.634 ± 0.002

kg/l kg/l kg/l

Value

Units

Steam pressure Hot water inlet temperature Cooling water inlet temperature Hot water volume flow rate Cooling water volume flow rate Solution mass flow rate

800 80–90 20–30 99 99 6–22

kPa (A) °C °C L/min L/min kg/min

Unit

4.373 ± 0.076 1.756 ± 0.030 97.574 ± 1.690 23.862 ± 0.413 3.372 ± 0.058

L/min L/min L/min L/min L/min

VS1 CS1 CS2 Height level G

3.582 ± 0.062 0.407 ± 0.007 0.424 ± 0.007

L/min L/min L/min

34.3 ± 0.3

%

C

29.9 ± 0.3

%

EL AL AH EH S

38.8 ± 0.3 29.8 ± 0.3 30.1 ± 0.3 51.4 ± 0.4 78.7 ± 0.7

% % % % %

Table 5 Operative range of the experiments. Measurement

Value

the fixed cooling water temperature in the condenser TC, a higher generator temperature augments both the input heat transfer rate and heat transfer rate delivered as an output at the absorber. Overall, global cycle efficiency is improved and reaches a COP that exceeds 0.3 at 84 °C although the improvement is modest. Higher input source temperatures are responsible for a relentless increase in output heat transfer rates. However, an excessive increase without properly redesigning the transfer surfaces of each component can eventually lead to greater system irreversibility and lower first and second principle efficiencies [37–38]. Throughout the experimental campaign, the system is controlled to generate steam at 170 °C while varying the operative parameter under analysis. For a certain solution mass flow rate, this target is associated with a defined AH outlet solution temperature. Therefore, as the hot water temperature is increased, the capacities of the absorbers and generator increase by widening the solution concentration differences between inlet and outlet (Fig 6i). The steam generation rate increases and the capacity of the EL grows larger. Specifically, with reference to the flow diagram in Fig 5, an increase in the hot water inlet temperature increases the temperature of the EL, hence increasing the pressure of the AL. Consequently, the solution temperature rises and boosts the pressure of the EH. Since the AH outlet temperature of the solution is kept nearly constant, the middle solution concentration (at the outlet of the AH) decreases. Moreover, since the widening of the solution concentration difference within the generator is larger than the drop in the middle solution concentration, the strong solution concentration increases the mass fraction of Lithium-Bromide

Calculated parameters Temperature effectiveness SH SL RH RL Solution concentration Strong Middle Weak COP (-) Heat transfer rate of AH Steam generation rate

Value

Unit

0.93 ± 0.04 0.92 ± 0.05 0.96 ± 0.01 0.91 ± 0.01

-

61.24 ± 0.09 58.98 ± 0.10 57.43 ± 0.10 0.31721 ± 4.0E-05 16.82134 ±1.5E-04

% % %

25.46 ± 0.87

kg/h

kW

(becomes slightly stronger) while the weak solution concentration is further lowered. The expanded uncertainty of the data plotted in Figs. 6–8 is calculated by using 3 as the coverage factor, providing a level confidence of approximately 99.7%. All the expanded uncertainty of the results are: temperature uncertainty ±0.90 °C; solution concentration uncertainty ±0.10%; EH pressure uncertainty ±2.2 kPa; EL pressure uncertainty ±0.4 kPa; C pressure uncertainty ±0.2 kPa; effectiveness temperature uncertainty ±5.0  102; COP uncertainty ±4.0  104; heat transfer rate of AH uncertainty ±1.5  104 kW; steam generation rate uncertainty ±0.87 kg/h. Fig. 7 illustrates the influence of the cooling water inlet temperature on the performance of single components and on global performance. Specifically, Fig. 7a confirms the capability of the system to reach the target temperature in the high temperature absorber AH with water at 88 °C in the generator inlet when the cooling water temperature ranges between 20 °C and 30 °C. The inlet hot water temperature is fixed, and thus the saturation pressure level of the evaporator EL and low temperature absorber (AL) (as denoted by the filled red markers in Fig. 7c) remains constant over the entire range of inlet cooling water temperatures. In contrast, the pressure at the condenser increases at correspondingly higher cooling water temperatures. This reduces the vapour generation and absorption capacity of the thermodynamic cycle. Fig. 7b shows the evidence for lower absorption rates (i.e. lower solution concentration differences) at higher cooling water temperatures, which reduces the amount of refrigerant circulating in the system and leads to lower heat transfer rates (Fig. 7f). Furthermore, the thermal effectiveness of the solution and refrigerant internal heat exchangers slightly worsens when TC increases. Subsequently, Fig. 7 describes the effect of the cooling water temperature on the ultimate objective of the system, namely heat transfer rate at the high temperature absorber (Fig. 7f), steam generation rate (Fig. 6g), and thermal efficiency of the entire cycle (Fig. 7e). With respect to a given hot water temperature at the generator TG, a lower cooling water temperature augments both input and output heat transfer rates. Globally, the cycle efficiency slightly improves at lower cooling temperatures and maintains a COP that

A. Lubis et al. / Energy Conversion and Management 148 (2017) 267–278

185

0.4

180

0.35

175

0.3

COP

AH temperature, oC

274

170

0.25 0.2

165 160

Solution (6)

0.15

(a)

Steam

Heat transfer rate of AH, kW

Solution concentration, %

64 62 60 58 56

(b)

Weak

Middle

Strong

54

Pressure, kPa

C

EL

Steam generation rate, kg/h

1000 EH

100

10

(c)

Temperature effectiveness of refrigerant heat exchanger

1

Temperature effectiveness of solution heat exchanger

(e)

0.1

155

1 SH

SL

0.95 0.9 0.85

(d) 0.8 78

80

82

84

86

88

90

92

20 16 12 8 4

(f)

0

28 24 20 16 12 8 4

(g)

0

1 0.95 0.9 0.85

(h)

RH

RL

0.8 78

Hot water inlet temperature,oC

80

82

84

86

88

90

92

Hot water inlet temperature, oC

140 89.8oC 80.0oC

Refrigerant temperature oC

120

100

80

60

40

20

(i) 0 0

20

40

60

80 100 120 Solution temperature oC

140

160

180

200

Fig. 6. Influence of the hot water inlet temperature (a) high temperature absorber, (b) solution concentration, (c) pressure inside the system, (d) solution heat exchanger effectiveness, (e) COP, (f) output heat transfer rate, (c) steam flow rate, (h) refrigerant heat exchanger effectiveness, and (i). Dühring diagram at steam pressure of 0.8 MPa, Lithium-Bromide/Water solution mass flow rate of 16 kg/min, hot water volume flow rate of 99 L/min, cooling water volume flow rate of 99 L/min, and cooling water inlet temperature of 25 °C.

275

185

0.4

180

0.35

175

0.3

COP

AH temperature, oC

A. Lubis et al. / Energy Conversion and Management 148 (2017) 267–278

170

0.25

165

0.2

160

0.15

Solution (6)

(a)

Steam

Heat transfer rate of AH, kW

Solution concentration, %

64 Weak

Middle

Strong

62 60 58 56

(b)

C

EL

EH

100

10

(c) 1

SH

Temperature effectiveness of refrigerant heat exchanger

1 SL

0.95 0.9 0.85

(d) 0.8 18

20

22

24

26

28

30

Steam generation rate, kg/h

1000

Pressure, kPa

24 20 16 12 8 4

(f)

0

54

Temperature effectiveness of solution heat exchanger

(e)

0.1

155

32

28 24 20 16 12 8 4

(g)

0 1 RH

RL

0.95 0.9 0.85

(h) 0.8 18

Cooling water inlet temperature, oC

20

22

24

26

28

30

32

Cooling water inlet temperature, oC

140 29.9oC 20.0oC

Refrigerant temperature oC

120 100 80 60 40 20

(i) 0 0

20

40

60

80 100 120 Solution temperature oC

140

160

180

200

Fig. 7. Influence of the cooling water inlet temperature (a) high temperature absorber, (b) solution concentration, (c) pressure inside the system, (d) solution heat exchanger effectiveness, (e) COP, (f) output heat transfer rate, (g) steam flow rate, (h) refrigerant heat exchanger effectiveness and (i) Dühring diagram at conditions of 0.8 MPa steam pressure, 16 kg/min Lithium-Bromide/Water solution mass flow rate, hot water volume flow rate of 99 L/min, cooling water volume flow rate of 99 L/min, and 88 °C hot water inlet temperature.

A. Lubis et al. / Energy Conversion and Management 148 (2017) 267–278

185

0.4

180

0.35

175

0.3

COP

AH temperature, oC

276

170

0.25

165

0.2

160

0.15

Solution (6)

(a)

Steam

Heat transfer rate of AH, kW

Weak

Middle

Strong

62 60 58 56

(b)

1000

Pressure, kPa

C

EL

EH

100

10

(c) 1 1 SH

SL

0.95 0.9 0.85

(d) 0.8 3

8

13

18

23

28

Steam generation rate, kg/h

64

54

Temperature effectiveness of solution heat exchanger

(e)

0.1

Temperature effectiveness of refrigerant heat exchanger

Solution concentration, %

155

20 16 12 8 4

(f) 0

28 24 20 16 12 8 4

(g)

0

1 0.95 0.9 0.85

(h)

RH

RL

0.8 3

Solution mass low rate, kg/min

8

13

18

23

28

Solution mass low rate, kg/min

140 24.87 kg/min 5.95 kg/min

Refrigerant temperature oC

120 100 80 60 40 20

(i) 0 0

20

40

60

80 100 120 Solution temperature oC

140

160

180

200

Fig. 8. Influence of the solution mass flow rate with respect to (a) high temperature absorber, (b) solution concentration, (c) pressure inside the system, (d) solution heat exchanger effectiveness, (e) COP, (f) output heat transfer rate, (g) Steam flow rate, (h) refrigerant heat exchanger effectiveness and (i) Dühring diagram at conditions of 0.8 MPa steam pressure, 88 °C hot water inlet temperature, hot water volume flow rate of 99 L/min, cooling water volume flow rate of 99 L/min, and 25 °C cooling water inlet temperature.

A. Lubis et al. / Energy Conversion and Management 148 (2017) 267–278

exceeds 0.3 for a wide operative range of the parameter. The proper design and high effectiveness of the internal heat exchangers and especially of the solution heat exchangers will evidently rule the capacity and thermal efficiency of the system. Hot water and cooling water temperatures generally constitute design parameters that are mainly ruled by the system boundary conditions and external sources availability. Conversely, the absorptive solution mass flow rate although influent with respect to the dimensions of the main heat exchangers, represents a parameter that can vary based on the transient operative conditions of the system, and in turn controls the instantaneous capacity and efficiency of the system. As a rule, a lower solution mass flow rate is associated with higher transfer rates per unit mass flow rate of the solution, and consequently with higher solution concentration differences (Fig. 8b). This significantly affects the heat transfer characteristics of a system with given transfer areas and fixed heat source/sink temperatures. After establishing the system boundary conditions (in terms of temperatures at the condenser, generator and high temperature absorber), the operative saturation pressures of the refrigerant inside the system are consistently fixed (Fig. 8c). Additionally, a higher transfer rate (associated with the increased absorption/release of the heat of the absorption characteristic of the Lithium-Bromide/water pair) brings the solution closer to the equilibrium solution concentration at the specific saturation pressure. The temperature effectiveness of the solution heat exchangers (Fig. 8d) increases but the refrigerant heat exchangers (Fig. 8h) is weakly effected. As the concentrated solution mass flow rate decreases and the pressure at EH is nearly constant, the solution concentration difference increases. The area of the plate heat exchanger is constant, and as the flow rate decreases, the NTU increases, thus increasing the temperature effectiveness of the solution heat exchanger. However, when the solution mass flow rate is excessively lowered, the partial wetting of the exchange surface reduces the active transfer area and eventually reduces the system capacity (shown in Fig. 8f and Fig. 8g) and thermal efficiency (shown in Fig. 8e). Since the heat transfer rates at the AH, the EL and the generator do not drastically change (except at the lowest solution mass flow rate, where partial wetting of the exchange surface starts occurring), the pressure of the EL and the middle solution concentration remain approximately constant. As the solution mass flow rate is increased, to achieve a constant solution temperature at the outlet of the AH, the solution concentration difference (the strong and weak solution concentrations) at the generator decreases (Fig. 8b). Fig. 8 demonstrates the evidence for the possibility of maximising both heat transfer (corresponding to higher steam generation rates) and efficiency of the system by properly setting the solution mass flow rate. There are differences in the manner in which input power required by the generator and the system output in the high temperature absorber are affected by the solution temperature and concentration changes due to the differences in the flow rates, the flow rate maximizing the heat transfer rate, and the flow rate that results in the highest system efficiency corresponding to slightly different values of this parameter. Generally, a significantly higher overall performance of the system can be observed when compared to the experimental results of a corresponding system configuration in a previous study [33].

6. Conclusions Despite the industrial and academic interest in multiple-lift heat transformers, previous research efforts have been limited to mathematical modelling and performance analysis, while lacking of experimental performance assessments and operative cases.

277

The study reports on the actual performance of a double-lift absorption heat transformer. The results demonstrated and analysed the manner in which the system could operate steadily and efficiently when driven by hot water temperature at approximately 80 °C while reaching temperatures above 170 °C in the high temperature absorber. The experimental results collected by operating a small scale prototype confirmed the feasibility and efficiency of the system. Specifically, the operability of the high temperature absorber and the system as a whole could be safely controlled by avoiding the occurrence of corrosion, crystallization, and/or air infiltration throughout the entire experimental campaign. Furthermore, the real performances were analysed with respect to the main operative parameters and control variables in a manner that is useful in pointing out the best condition of a system that recovers from the heat discharged from a real plant. Generally, the following conclusions can be stated: - When the inlet cooling water temperature and steam temperature are fixed the steam generation rate increases at higher inlet hot water temperatures; - When the inlet hot water temperature and steam temperature are fixed the steam generation rate increases at lower inlet cooling water temperatures; - The global cycle efficiency is improved and reaches a COP that exceeds 0.3 at 84 °C; - Both heat transfer (corresponding to higher steam generation rates) and efficiency of the system can be optimized by properly setting the solution mass flow rate; - The highest COP and the maximum steam generation rate are obtained when the solution mass flow rate gives a solution concentration difference of 5–6% between strong and weak solutions;

Acknowledgments This research work was supported by New Energy and Industrial Technology Development Organization (NEDO), Ebara Refrigeration Equipment & System Co.Ltd, and Research Institute for Science and Engineering (RISE) Waseda University as part of the research project of ‘‘Next generation heat pump technology”.

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