Recent developments in thermally-driven seawater ...

DES-12312; No of Pages 16 Desalination xxx (2014) xxx–xxx

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Recent developments in thermally-driven seawater desalination: Energy efficiency improvement by hybridization of the MED and AD cycles Kim Choon Ng a,b,⁎, Kyaw Thu a, Seung Jin Oh a, Li Ang a, Muhammad Wakil Shahzad b, Azhar Bin Ismail b a b

Department of Mechanical Engineering, National University of Singapore, Singapore King Abdullah University of Science & Technology, Water Desalination & Reuse Center (WDRC), Thuwal, Saudi Arabia

H I G H L I G H T S • Recent development in adsorption desalination • A new isotherm model based on energy distribution function (EDF) for all types • An exergy-based method for fuel cost apportionment in cogeneration plants

a r t i c l e

i n f o

Article history: Received 8 September 2014 Received in revised form 14 October 2014 Accepted 16 October 2014 Available online xxxx Keywords: Seawater desalination Adsorption desalination Multi-effect distillation MEDAD cycles

a b s t r a c t The energy, water and environment nexus is a crucial factor when considering the future development of desalination plants or industry in the water-stressed economies. New generation of desalination processes or plants has to meet the stringent environment discharge requirements and yet the industry remains highly energy efficient and sustainable when producing good potable water. Water sources, either brackish or seawater, have become more contaminated as feed while the demand for desalination capacities increase around the world. One immediate solution for energy efficiency improvement comes from the hybridization of the proven desalination processes to the newer processes of desalination: For example, the integration of the available thermally-driven to adsorption desalination (AD) cycles where significant thermodynamic synergy can be attained when cycles are combined. For these hybrid cycles, a quantum improvement in energy efficiency as well as in increase in water production can be expected. The advent of MED with AD cycles, or simply called the MEDAD cycles, is one such example where seawater desalination can be pursued and operated in cogeneration with the electricity production plants: The hybrid desalination cycles utilize only the low exergy bled-steam at low temperatures, complemented with waste exhaust or renewable solar thermal heat at temperatures between 60 and 80 °C. In this paper, the authors have reported their pioneered research on aspects of AD and related hybrid MEDAD cycles, both at theoretical models and experimental pilots. Using the cogeneration of electricity and desalination concept, the authors examined the cost apportionment of fuel cost by the quality or exergy of working steam for such cogeneration configurations. © 2014 Elsevier B.V. All rights reserved.

Contents 1. 2.

3. 4.

5.

Introduction . . . . . . . . . . . . . . . . . . . . . . . . . . Basic adsorption phenomena . . . . . . . . . . . . . . . . . . 2.1. Sorption thermodynamics . . . . . . . . . . . . . . . . 2.2. Universal site-energy probability distribution function (EDF) . The design of AD batch-operated cycle . . . . . . . . . . . . . . The MED-AD hybrid cycle . . . . . . . . . . . . . . . . . . . . 4.1. MED-AD simulation . . . . . . . . . . . . . . . . . . . 4.2. MED-AD experimentation. . . . . . . . . . . . . . . . . Exergy analysis for operational cost apportionment . . . . . . . .

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⁎ Corresponding author.

http://dx.doi.org/10.1016/j.desal.2014.10.025 0011-9164/© 2014 Elsevier B.V. All rights reserved.

Please cite this article as: K.C. Ng, et al., Recent developments in thermally-driven seawater desalination: Energy efficiency improvement by hybridization of the MED and AD cy..., Desalination (2014), http://dx.doi.org/10.1016/j.desal.2014.10.025

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6. Conclusions. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Acknowledgment . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . References. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .

1. Introduction Fresh water is a key resource that is necessary for the economic development of every country of the world, and it is consumed by all economic sectors such as agriculture, industrial, commercial and dwellings. Annually, the growth of water demand in the world is estimated to be around 3–4% [1–12], and two factors contributed to this growth, namely, (i) the continued economic growth of all countries and (ii) the exponential increase in the world's population, have imposed increasing demand for potable or drinkable water. Despite the fact that more than 70% of the earth is covered by water, much of it is not directly available for consumption or usage in all sectors due to the high water salinity, either in the form of brackish, waste or seawater, as shown in Table 1 [13–18]. In many water-stressed countries of the world, the annual water availability is less than 250 m3 per capita per year [19], risking the health of population due to poor water quality and sub-standard sanitation. Also, there is a clear disparity of water consumption levels amongst the economic sectors of developed and developing countries, as shown in Fig. 1(a): The water consumption is most dominant in the agriculture sector of developing countries while the developed countries would have higher water consumption in the industrial sectors [20]. It is predicted that future potable-water demand of the world will be led by the East Asia and Asian developing countries, contributing to more than 50% of the world's water requirement and the remaining water demand trend would come from the America, Europe etc. [21–25]. With the increasing trend of water consumption, it is foreseeable that the future water demand will exceed the existing level of sustainable water supply: In 2030, the estimated demand for fresh water is 6900 billion cubic meters (Bm3), however, total available water from the natural water cycle remains only at 4500 Bm3 [26–29], as shown in Fig. 1(b). Hence, the sustainable natural water cycle is unable to meet future water demand of the world. To bridge this water demand–availability gap in the years to come, new solutions to water supply must be sought. Improvements in desalination efficiency of processes alone can only alleviate partially the projected water shortage gap, leaving a large portion of water deficit to be met by further increase in desalination capacity offered by engineering solutions [30]. From 2010 to 2016, the desalinated water production capacities of many water stressed countries have been collated by the International Desalination Association (IDA), as shown in Fig. 2. Over the past 5 years, the increase in the installed capacities has been rising at 9 to 10% per annum [31], with the total global installed desalination capacity increasing from 44 million cubic meters per day (Mm3/day) in 2006 [32] to 75 Mm3/day in 2010. This increasing trend is expected to double by 2015 [33,34]. Owing to the dry climate of GCC (Gulf Cooperation Council) countries, it is understandable that the growth rate for

0 0 0

desalination capacity is higher in this region than the rest part of the world. It is noted that the GCC countries suffer a rapid depletion of ground water due to high extraction rates from agriculture while the population growth rates of these countries are increasing exponentially [35–41]. Despite the higher installed desalination capacities in recent years, the annual water availability per capita in GCC countries is deemed to remain at an acute level as they are having less than 300 m3/year per capita. In order to meet the future water demand, GCC countries have to pursue energy efficient desalination methods. Owing to the ease of fuel oil supply, higher salinity of seawater feed and the frequent occurrences of the harmful algae blooms, the majority of desalination methods found in the region is designed to have the thermally-activated type [42–46] to be collocated together with power (electricity) plants: Electricity is produced by high exergy steam that emanates from the boilers at high pressures and temperatures while the thermally-activated desalination processes form the bottoming cycles, powered by low temperature and pressures bled-steam of low exergy, such as the multi-stage flash (MSF) or the multi-effect distillation (MED). The energetic analysis of thermal system is presented in literature [47–55] in detail and operating fuel cost of MSF or MED cycles, hitherto, has been computed based only on the energetic analysis [55–64], omiting the role of the quality or exergy of expanding steam. Thermal system exergy analysis is conducted by many researchers [65–80] for overall system performance investigation.

a

8%

11%

22%

8% 10%

59% 70% 82% 30% World Developed countries Agricultural

Developing countries Industrial Domestic

b

Table 1 Distribution of earth surface water. Source

Oceans Ice sheets, glaciers Ground water Surface water Atmosphere Total Grand total

Volume (cubic kilometer) Fresh water

Salt water

0 24,364,000 10,530,000 122,210 12,900 35,029,110 1,386,000,000 (estimated)

1,338,000,000 0 12,870,000 85,400 0 1,350,955,400

Fig. 1. (a): Share of different sectors (by percentage) in water consumption in different parts of the world and total Global water consumption [20]. (b): Fresh water supply demand gap: current and future estimates.



Fig. 2. Trend in installed desalination capacities in the past and the future for all countries in the world [31].

Some scientists performed exergy of single system component efficiency improvement [81–89]. Exergy destruction across MED and MSF components is investigated by many researchers and presented in the literature [90–99]. Despite a large number of publications reported on the aspect of exergy analysis of the cogeneration system, the low exergy bled-steam from turbines (with high latent heat contribution) has insignificant work contribution to the power plants should it were to continue its expansion through the turbines. The authors consider the exergy effect of steam for the fuel cost apportionment between the electricity and the MSF/MED cycles. Otherwise, the conventional energetic methodology may give an over-estimate of the fuel cost in comparison with electricity production. In this review, the impact of exergetic analysis is demonstrated in terms of the unit cost of water production. Although energy efficiency consideration for desalination processes is important, today's goal for desalination industry embraces a more holistic approach where the planners of desalination plants consider other factors such as the environmental impact from brine and chemical discharge of chemical for long-term sustainability of the desalination industry. The intricate nexus between water, energy and environment in many water stressed economies has encouraged scientists, planners and engineers to look for innovative desalination processes. The thermodynamic limit for desalination is dependent of the salinity and the temperature of seawater and the accepted specific energy consumption of seawater, with total dissolved solids (tds) of 35,000 to 45,000 ppm, ranges from 0.78 to about 1.2 kWh/m3 [100]. However, practical desalination processes, irrespective of the methods employed, are operated at acceptable water production rates consuming their specific energy efficiencies at several folds higher: The dissipative losses incurred in practical desalination processes are dictated by the Second Law of thermodynamics. Recent improvements in the membrane-based RO processes are known to have set an efficiency benchmark of 3 to 3.5 kWh/m3, yet they draw electrical energy input to overcome the osmotic pressures of RO cycles. Table 2 depicts the various energy requirements for the conventional desalination methods that are found in the industry [101–103]. With the

3

exception of hydro-electric plants, every kWh of electricity production is accompanied by much emissions of carbon dioxide and sulfur dioxide gases, as well as the rejection of more than a kWh of waste heat to the environment. The energetic efficiency of conventional fossil fuels based power plants is less than 45%. On the other hand, there is an excellent opportunity for energy efficiency improvement from the thermally-driven cycles of desalination: Firstly, the synergetic effects from temperature cascaded evaporation– condensation processes can be exploited where the re-utilization of latent energy could be configured in multiple times as the working steam expands. Recent integration of the low temperature or waste heatdriven adsorption cycle with the multi-stage flash (MSF) or the multieffect distillation (MED) [104–128] has, in principle, increase the top-brine-temperature (TBT) to lower-brine-temperature (LBT) range from conventional 70 °C–40 °C to 70 °C–10 °C, increasing the potential number of MED or MSF theoretical stages from the conventional 15 to 60 stages with the temperature drop per stage varying from 2 to 2.5 K. Hence, the kWhthermal/m3 of the thermally-driven cycles reduces from the conventionally known values of about 32 to 11 kWhthermal/m3, energetically. The higher TBT–LBT range is due to the vapor-uptake by the AD cycle that lowers the designed LBT of combined cycle from 35 °C to 10 °C [104–107]. Thus, the key advantage of the hybridized cycles is that the ability to exploit the low temperature heat source of cogeneration plants. If such waste heat is unused otherwise, it would have been purged into the ambient. Thus, hybridization of the MED to AD cycle or simply called MEDAD cycle increases the overall efficiency of the desalination cycles. Secondly, the latent energy input at the TBT stage has much lower exergy or availability to perform mechanical work, due primarily to the low pressure of the bled-steam. Most engineers or operators of cogeneration plants have erroneously equated the exergetic or quality of bledsteam supplied to MED or MSF processes to have the same quality as the steam entering the high pressure turbines. Although the energetic contributions in two mentioned situations are similar, it could have led to an unfair fuel-cost apportionment for the thermally-activated desalination processes. In this paper, the authors have distinguished the exergetic to the energetic contributions where the former can accurately capture the inferior quality of steam supplied for desalination purposes in a cogeneration plant. In large seawater basin with only a single inlet and narrow channel or straights, the quality of seawater feed of desalination plants may suffer from high salinity, typically as high as 45,000 ppm as compared with less than 32,000 ppm of open seas or oceans. When such basins have river's estuaries discharging un-treated river discharges, such as the Gulf of Arabian peninsula, the seawater tend to have high silt suspension and frequent harmful algae blooms (HABs) occur in the coastal waters of the basin where feed intake ducts of desalination plants are located. These pollutants cause high fluctuations in seawater feed, making it less suitable for membrane-based RO plants. However, these mentioned pollutants are less susceptible to the thermally-driven MSF or MED processes due to evaporative effect. Consequently, up to 80% of the large desalination plants along the GCC countries comprise mainly the MSF or the MED plants [129,130]. In addition to the suspended pollutants, the main cause of concern emanates from the release of toxins by algae microbs in the seawater feed, such as neuro-, paralytic-, and

Table 2 Primary energy requirements for proven desalination methods (without exergy consideration). Method of desalination

Thermal energy (kWhthermal/m3) ηboiler = 0.95 (A)

Electricity consumption (kWhe/m3) ηth = 0.45 (B)

Total primary energy consumption kWhPe/m3 (C) = (A) / ηboiler + (B) / ηth

Multi-stage flash (MSF) Multi-effect distillation (MED) Vapor compression (VC) Reverse osmosis (RO) — single pass Reverse osmosis (RO) — double pass

19.4 16.4 – – –

5.2 3.8 11.1 5.5 7.5

31.97 25.7 24.67 12.34 16.67


4


(a) SEM picture of mesoporous (b) SEM picture of Zeolite FAM silica gel. The insert has a magnification of Z01(alumina-phosphate). The insert has a 80,000. magnification of 160,000.

30 C

0.50

35 C

40 C

45 C 30 C

Adsorbate Uptake, q* [kg/kg of Adsorbent]

Adsorbate Uptake, q* [kg/kg of Adsorbent]

0.40 0.35 0.30 0.25 0.20 0.15 0.10 0.05 0.00

40 C

45 C

60 C

0.25

0.45

Pressure [kPa]

0

2

4

6

8

0.20 0.15 0.10 0.05 Pressure [kPa]

0.00 10

(c) The isotherms of silica gel at 30 C to 45C at low water vapor (single component) pressures. The isotherms are of Type 1 of the IUPAC classification and their characteristics make them suitable for AD cycles in the tropics, ambient temperatures less than 35o C.

0

2

4

6

8

10

12

(d) The isotherms of aluminaphosphate (Zeolite Z01) at 30 C to 45C at low water vapor (single component) pressures. The Type 5 isotherms of Zeolite (Z01) has provided a unique ability for the AD cycle to operate even at high ambient temperatures in hot summers of desert countries, with ambient dry bulb exceeding 45oC.

Fig. 3. (a) SEM picture of mesoporous silica gel. The insert has a magnification of 80,000. (b) SEM picture of Zeolite FAM Z01 (alumina-phosphate). The insert has a magnification of 160,000. (c) The isotherms of silica gel at 30 °C to 45 °C at low water vapor (single component) pressures. The isotherms are of Type 1 of the IUPAC classification and their characteristics make them suitable for AD cycles in the tropics, ambient temperatures less than 35 °C. (d) The isotherms of alumina-phosphate (Zeolite Z01) at 30 °C to 45 °C at low water vapor (single component) pressures. The Type 5 isotherms of Zeolite (Z01) has provided a unique ability for the AD cycle to operate even at high ambient temperatures in hot summers of desert countries, with ambient dry bulb exceeding 45 °C.

diarrheic-toxins. Such microbial-based toxins are of the same molecular sizes and they posed health hazards to humans when traces of toxins are ingested. When water vapor is evaporated by thermally-driven processes, the toxins are separated from the distillate product of MED or MSF plants [131–133].

The commercial grade SiO 2 ·nH 2 O has a pore surface area of 700–800 m 2 /g, and it is available in abundance as well as cost effective when compared to Zeolite (Z01). Silica gel has isotherms that are suitable for ambient below 32 °C while the Zeolite's sorption behavior can be tuned to suit a higher cooling temperature, up to 45 °C and the latter adsorbent is suitable for the summer conditions of the GCC regions. Table 3 depicts the common physical

2. Basic adsorption phenomena For a thorough introduction, we present the adsorption desalination phenomena first and then followed by description of the batchoperated adsorption cycle configuration, as well as its integration of the AD cycle to a conventional thermally-driven cycle such as the multi-effect distillation (MED). For desalination application, the adsorbate would be water vapor while the adsorbent is usually the inexpensive micro- and meso-porous materials, prepared either in the granular of the powder form, as shown in Fig. 3(a) and (b) [134]. Their respective vapor uptake at different temperatures is depicted accordingly in Fig. 3(c) and (d).

Table 3 Thermo-physical properties of the silica gel type 3A and Zeolite FAM Z01. Properties 2

BET surface area [m /g] Porous volume [ml/g] Apparent density [kg/m3] Thermal conductivity [W/m·K] Heat of adsorption (H2O) [kJ/kg of H2O] Specific heat capacity [kJ/kg·K]

Silica gel Type 3A

Zeolite FAM Z01

680 0.47 770 0.174 2800 0.921

147.3 0.071 600–700 0.113 (30 °C) 3110 (25 °C) 0.805 (30 °C)



5

each containing an average local adsorption energy level but when assembled, they give the overall energy distribution surface. Mathematically, the adsorption isotherms across a heterogeneous surface is given by the integral of the product between the energy distribution and the local surface coverage or uptake, i.e.,

properties of both adsorbents that are suitable for desalination cycles [135–137]. Superficially, the Zeolite may appear to be inferior, but its isotherms is of Type 5 category with tunable uptake slopes for the isotherms. 2.1. Sorption thermodynamics

X

θ¼

In designing an adsorption based system, the basic equilibria-vapor uptake information at assorted pressures and temperatures of the vapor must be available. These are commonly represented in the form of isotherms of an adsorbent–adsorbate pair which, hitherto, have been captured by many empirical and semi-empirical models. Under the IUPAC classification, sorption phenomena comprised six categories in terms of their uptake shapes and behavior but there is yet a universal theory to unify them: Each empirical or semi-empirical model can handle only a limited uptake shape of the adsorbent–adsorbate pair. For example, the Langmuir, Langmuir–Freundlich, Dubinin–Astakhov, Dubinin– Raduskevich and Tóth isotherms are some of the widely used models for assorted types of isotherms and they have different adsorption site-energy distributions. Thus, there is motivation for scientists and engineers to develop a theoretical framework that can describe not only the isotherms accurately but it should span across all sorption phenomena with known energy distributions. The simplest adsorption isotherm model is the classical Langmuir model [138] where it assumes a homogeneous surface with a monolayer vapor uptake with all surfaces containing a uniform charged energy (a flat distribution). Each pore vacant site is filled by a single vapor molecule to form a single sorption event. Invoking the fundamental rate of gas molecules filling the adsorption sites, as given by Ward and coworkers [139–142],

Z

∞ 0

i

χ ðεÞ e θðεÞdε

ð3Þ

where, e θðεÞ represents the local surface coverage and χ(ε) corresponds to the energy distribution function (EDF) describing the energetically heterogeneous surface over all sites. The total EDF when integrated across all energy levels would sum to unity, i.e., Z

∞ 0

χ ðεÞ dε ¼ 1:

ð4Þ

Invoking the earlier published concept of condensation approximation principles [145,146] for moderate temperatures, the local surface coverage can be assumed to have a step-like behavior, spreading with a “dirac-delta” function, then, the local adsorption uptake function or isotherm (θ) have two regions: One region of zero uptake when the energy level is below εc, and the other region is above the threshold energy, i.e., Z θ¼

Z

¼ μ g −μ α μ α −μ g dθ 0 ¼ K exp −exp dt RT RT

χ ðεi Þ e θðεi Þ ¼

εc 0 ∞ εc

Z 0 χ ðεÞdε þ

∞ εc

1 χ ðεÞdε ð5Þ

χ ðεÞdε:

ð1Þ We reckon that it is possible to have a universal energy distribution function (EDF) to approximate the site distribution function χ(ε) to represent all six types of isotherms of IUPAC. This depends on the adsorbent surface characteristics during adsorption interaction, in which either a symmetrical or an asymmetrical Gaussian function could be assumed. In this paper, we present the energy distribution function known individually as shown in Table 4, which summarizes the functions of χ(ε) for the classic Langmuir–Freundlich [147], Dubinin–Astakhov [148], Dubinin–Raduskevich [149] and Tóth isotherm models. By performing integration of these distribution functions over all site energy, the exact form of these isotherm models would be obtained.

where μ is the chemical potential of the gaseous (“g”) and adsorbed (“a”) phase [143,144] the expression of the Langmuir isotherm model can be derived as follows; ε P RT θ¼ ε : 1 þ K exp P RT K exp

ð2Þ

where ε is the activation energy, and K is the pre-exponential constant. When the geometrical roughness and heterogeneity of the adsorbent (inherently formed during production) are considered, the gas molecules depositing on the pore surfaces may have a preferential energy potential distribution or shape. Assuming that the overall energy distribution surface can be sub-divided into many sub-regions,

2.2. Universal site-energy probability distribution function (EDF) Guided by the available site-energy probability distribution function, χ(ε), for the four empirical models, the authors have proposed, for the

Table 4 Forms of adsorption site energy distribution functions for the classical Langmuir–Freundlich, Dubinin–Astakhov, Dubinin–Raduskevich and Tóth isotherm models. Model

Adsorption site energy distribution χ(ε)

Isotherm equation εc

∞

∞

c

θ ¼ ∫0 0 χ ðε Þdε þ ∫ε 1 χ ðε Þdε ¼ ∫ε χ ðε Þdε c

Langmuir–Freundlich Dubinin–Astakhov Dubinin–Radushkevich Tóth

ε−ε

χ ðεÞ ¼

1 exp ð c 0Þ c 2 ε−ε ½1þexp ð c 0 Þ

χ ðεÞ ¼

rðε−ε 1 Þr−1 Er

1Þ χ ðεÞ ¼ 2ðε−ε E2

χ ðεÞ ¼

1 RT

θ¼

h r i 1 exp − ε−ε E h 2 i 2 exp − ε−ε E

½ exp ð

1þ½ exp

ð

ε−ε3 RT

ε−ε 3 RT

t

Þ tþ1 t t

Þ

RT

ε

½KP exp ðRT0 Þ c 1þ½KP exp

ε

RT

ðRT0 Þ c h θ ¼ exp − RT E ln RT θ ¼ exp − E ln

θ¼

KP exp 1þ½KP exp

P0 P P0 P

r i 2

ε

ðRT3 Þ ε

t

ðRT3 Þ

1 t


6


first time, an energy distribution function (EDF) that can capture six types of IUPAC adsorption isotherms [150]. This universal EDF is presented in the form of Eq. (6), and it has only four regression parameters, namely, ϕ, C, t and Ec. We believe that using the measured data of simple isotherms of an adsorbent–adsorbate pair, these parameters can be readily regressed while the more complex isotherms may require a superposition of two EDF functions to fully capture the isotherm patterns.

P P P Aϕ exp β þC q Ps Ps Ps ot ¼ n q0 1 þ ϕ exp β PP PP s

ð6Þ

s

and, β ¼ exp

Ec RT

ð7Þ

Fig. 4. (Left) Isotherms of water vapor adsorption on assorted adsorbent with regressed isotherm from Eq. (7) (solid black line) alongside experimental data with 5% error bars [Table 5] (Right) The corresponding asymmetrical Gaussian representing the EDF from the Eq. (10) for Type I to V adsorption isotherm characteristics.



7

Table 5 List of adsorbent + adsorbate pair tested. Type

Symbols used in Fig. 4

Adsorbate

Adsorbent

Temperature (°C)

References

Type I

◊

Water

Type RD, silica Gel

43

Type I

Δ

Water

Type A, silica gel

37

Type II

◊

Water

Pseudoboehmite

22

Type II

Δ

Water

Poly(vinyl pyrrolidone)

30

Type III

◊

Water

Activated carbon

20

Type III

Δ

Water vapor

Single walled carbon nanotubes

20

Type IV

◊

Water vapor

Polymer carbon

35

Type IV

Δ

Water vapor

Polymer carbon

35

Type V

◊

Water vapor

Molecular sieve

28

Type V

Δ

Water vapor

Molecular sieve

28

JIAYOU, QIU. “Characterization of silica gel–water vapor adsorption and its measurement facility.” MEng. diss., 2004. JIAYOU, QIU. “Characterization of silica gel–water vapor adsorption and its measurement facility.” MEng. diss., 2004. Wang, Shan-Li, Cliff T. Johnston, David L. Bish, Joe L. White, and Stanley L. Hem. “Water-vapor adsorption and surface area measurement of poorly crystalline boehmite.” Journal of colloid and interface science 260, no. 1 (2003): 26–35. Zhang, Jiang, and George Zografi. “The relationship between “BET” and “free volume”‐derived parameters for water vapor absorption into amorphous solids.” Journal of pharmaceutical sciences 89, no. 8 (2000): 1063–1072. Kim, Pyoungchung, and Sandeep Agnihotri. “Application of water-activated carbon isotherm models to water adsorption isotherms of single-walled carbon nanotubes.” Journal of colloid and interface science 325, no. 1 (2008): 64–73. Kim, Pyoungchung, and Sandeep Agnihotri. “Application of water-activated carbon isotherm models to water adsorption isotherms of single-walled carbon nanotubes.” Journal of colloid and interface science 325, no. 1 (2008): 64–73. Bansal, R. C., and T. L. Dhami. “Surface characteristics and surface behaviour of polymer carbons—II: Adsorption of water vapor.” Carbon 16, no. 5 (1978): 389–395. Bansal, R. C., and T. L. Dhami. “Surface characteristics and surface behaviour of polymer carbons—II: Adsorption of water vapor.” Carbon 16, no. 5 (1978): 389–395. Lagorsse, S., M. C. Campo, F. D. Magalhaes, and A. Mendes. “Water adsorption on carbon molecular sieve membranes: experimental data and isotherm model.” Carbon 43, no. 13 (2005): 2769–2779. Lagorsse, S., M. C. Campo, F. D. Magalhaes, and A. Mendes. “Water adsorption on carbon molecular sieve membranes: experimental data and isotherm model.” Carbon 43, no. 13 (2005): 2769–2779.

Fig. 5. Detailed schematics of an MEDAD cycle. The circle (filled) dots are the vapor valves and the pair of triangles refers to the liquid valves. The range of vapor pressures in evaporator and condenser are 1–2 kPa and 5 to 7 kPa.


8


Fig. 6. Detailed P& ID diagram for an experimental MED facility of nominal heating capacity of 12 kW. The first stage comprises a water-fired steam generator while the last stage (3rd stage) is open to a water-cooled cooling tower.

A¼

½1 þ ϕ exp ðβÞt −C ϕ exp ðβÞ

ð8Þ

where, the alphabet ϕ, and C, are constants; t is the surface heterogeneity factor and Ec denotes the characteristic energy of the

adsorbent–adsorbate pair. These four parameters can be mathematically tuned to get a suitable form to match the isotherm data from experiments. Utilizing this unified adsorption isotherm framework, an EDF may thus be expressed as single peak asymmetrical Gaussian function for

Table 6 MEDAD modeling equations. Equation

Remarks

Modeling equations for steam generator • • ½ðM hw :Cphw Þ dTdthw ¼ m hw; h f ;Thw;in − m hw; h f ;Thw;out −hin;o :Ain;i T hw −T tube;i dT tube;i ¼ hin;i :Ain;i T hw:i −T tube;i −hout;i :Aout;i T tube;i −T v;i M HX;i :Cphx;i dt •

dM b;i dt

•

•

¼ m f ;i − mb;i − mv;i • • • i ¼ m f ;i h f ;T f − m b;i h f ;Tb − m v;i hg;Tv þ Q in;i M b;i :Cpb;Tb þ M HX;i :CpHX dT dt Qin = ho,iAi(Tt − Tv,i) • • • dX M b;i dtb;i ¼ m f ;i X f ;i þ m b;i X b;i − m v;i X v;i Nu ¼

hin;i din;i K tube;i

¼ 0:023Re

0:80 l

Pr

0:40 l

Energy balance for metal tubes. Mass balance for the seawater inventory in the evaporator side of the brine heater. Energy balance for the evaporator side of the steam generator. Material/concentration balance Convective heat transfer coefficient equation Tube wall resistance

dout;i din;i

ln

Energy balance for the hot water flowing inside the tubes of brine heater.

Rwall;i ¼ 2πK tube;i Ltube;i 2 1=3 0:4 ¼ 0:0007 Re0:2 Pr 0:65 q″ ho ν3

Falling film evaporation heat transfer coefficient, Han and Fletcher's correlation

k g

U i Ai ¼

Overall heat transfer coefficient

1

1 þRwall;i þh 1A hin;i Ain;i out;i out;i

Modeling equations for intermediate stages dT cond;iþ1 •

¼ mv;hfg;Tv i −½hin :Ain ðT cond −T tube Þiþ1 M l;iþ1 :Cpl;Tcond dt dT tube;iþ1 ¼ hin;iþ1 :Ain;iþ1 T cond;iþ1 −T tube;iþ1 − M HX;iþ1 :Cphx;iþ1 dt hout;iþ1 :Aout;iþ1 T tube;iþ1 −T v;iþ1 •

•

•

Energy balance for the condenser side of the ith effect. Energy balance for tube metal

¼ m f ;iþ1− mb;iþ1− mv;iþ1 :Cpb þ M HX;iþ1 :CpHX;iþ1 dTdtiþ1 ¼ M • b;iþ1 • • m f ;iþ! h f ;T f − m b;iþ1 h f ;Tb − m v;iþ1 hg;Tv þ Q in;iþ1 Q in;iþ1 ¼ hout;iþ1 Aiþ1 T t;iþ1 −T v;iþ1 • • • dX b;iþ1 M b;iþ1 dt ¼ m f ;iþ1 X f ;iþ! − m b;iþ1 X b;iþ1 − m v;iþ! X v;iþ! 1=4 gh ρl;Tcond ðρl −ρv ÞTcond K 3l;Tcond h Liþ1 ¼ 0:728 fg;Tcond Nu ¼ Kin;iþ1 μ l;Tcond di ðT v;iþ1 −T tube;iþ1 Þ tube;iþ1

Brine inventory balance

U i Ai ¼

Overall heat transfer coefficient equation.

dM b;iþ1 dt

1

1 þRwall;i þh 1A hin;i Ain;i out;i out;i

MED last stage connected with AD beds • • M b;n Cpb þ M HX;n CpHX;n dTdtn ¼ m f ;n h f ;T f − m b;n h f ;Tb − M sg hg;Tv dqdtads þ Q in;n Q in;n ¼ hout;n An T t;n −T v;n

Energy balance for evaporator side

Material/ concentration balance Nusselt film condensation correlation for the calculation of the heat transfer coefficient inside the condenser tubes

Energy balance for the condenser side of the ith effect.



9

condensation approximation of energy distribution, i.e., the summation of all EDF energy levels from εc to infinity, must sum to unity, i.e., 1st stage heat source, tube surface & evaporator temperature

Z

∞ εc

χ ðεÞdε ¼ 1:

ð11Þ

The boundary condition for the correct isotherm is guided by having ∞ θ ¼ ∫ε χ ðεÞdε, and θ = 1 only at P = Ps and hence, one obtains that parc ticularly for the proposed isotherm model of the present work, ε c ¼ Ec :

ð12Þ

The detail of adsorbent + adsorbate pair tested by different researchers is listed in Table 5. 3. The design of AD batch-operated cycle Fig. 7. MEDAD cycle component temporal temperature profiles.

Types I, II, III, IV and V as per the classical models discussed earlier (see Fig. 4). This function is described as Eq. (9) given by 8 9 ðβ þ 1Þ C > > > > > > Aϕ exp β þ > > > β β > > > > −t−1 > = < β β β ðt−1Þ 1−ϕ exp β ð9Þ ϕ exp β þ1 χ ðεÞ ¼ > > β RT β > > > > 2 > > > > β > > > > ; : −Ctϕ exp β β where the variable β* is a function of the adsorption site energy ε and expressed as,

β ¼ exp

2Ec −ε : RT

ð10Þ

Utilizing the same procedure, it is believed that two-peak asymmetrical Gaussian energy distribution functions can be derived to provide a theoretical EDF function for the Type VI behavior. In all cases, the

There are five main components of AD system namely: i) evaporator, ii) adsorption and desorption reactor beds, iii) condenser, iv) pumps and v) pre-treatment facility. The detailed process diagram of MEDAD is shown in Fig. 5. For the batch-operated AD cycle, it involves two main processes. Adsorption-assisted-evaporation In which the vapors generated in AD evaporator are adsorbed on the pore surface area of adsorbent. The heat source is circulated through the tubes of evaporator and seawater is sprayed on the tube bundle. The heat transfer through tube surfaces from heat source to seawater causes evaporation of feed film. It is also observed that the high affinity of water vapor of adsorbent also contributes in evaporation. The AD evaporator operation temperature can be controlled by

Fig. 8. MEDAD hybrid cycle performance (concentration, GOR, PR and WPR).


10


Fig. 11. MEDAD component temperature profiles at 38 °C heat source temperature.

Fig. 9. Percentage improvement in water production rate by the MEDAD cycle.

heat source temperature that is normally circulated in terms of chilled water. The AD evaporator can operate at a wide range of chilled water temperature that varies from 5 °C to 30 °C to produce the cooling effect as well at low temperature operation. The vapor adsorption process continues until the adsorbent bed reach to a saturation state. Desorption-activated-condensation In which saturated adsorbent is regenerated using the low grade industrial waste heat or renewable energy (desorption temperature varies from 55 °C to 85 °C) and desorbed vapors are condensed in a water cooled condenser and collected as a distillate water [151–154]. It can be seen that two useful effects produced by AD cycle are the cooling effect by the first process “adsorption-assisted-evaporation” and fresh water by converting the seawater by second process “desorption-activated-condensation”. Useful effects which are cooling and water production can be produced simultaneously by introducing the multi-bed technique [155–159]. In multi-bed AD system, the operation and switching technique are used. During operation, one or a pair of adsorbent reactor beds undergoes the adsorption proces and at the same time one or a pair of

Fig. 10. A pictorial view of MEDAD pilot as installed in NUS.

adsorbent reactor beds executes the desorption process. The time for adsorbent reactor bed operation either adsorption or desorption depends on the heat source temperature and silica gel quantity packed in a bed [159]. Prior to changing the reactor duties, there is a short time interval called switching in which the adsorber bed pre-heated while the desorber bed pre-cooled to enhance the performance of cycle. In AD cycles, the operation (adsorption and desorption) and switching processes are controlled by automated control scheme that can open and close the respective bed hot/cold water valves. During switching operation, all vapor valves are closed so there is no adsorption or desorption taking place. 4. The MED-AD hybrid cycle MEDAD is a hybrid of two thermal systems namely; multi-effect desalination system and adsorption cycle. The main components of this novel thermal hybrid system are: 1) multi-effect distillation (MED) system, 2) adsorption desalination (AD) cycle and 3) auxiliary equipments. In this hybridization system, the last stage of the MED is connected to adsorption beds of AD cycle for the direct vapor communication to adsorption beds. Fig. 6 shows P&ID of MED plant combined with AD system. Adsorption based desalination is investigated by many researchers and reported that optimal specific daily water production (SDWP) for four bed scheme is about 4.7 kg/kg silica gel [160,161]. Adsorption desalination based on low grade waste heat utilization is patented by Ng et al. They installed the first adsorption desalination plant in the National University of Singapore (NUS) which consists of four silica gel beds. They investigated the processes using chilled water at 12.5 °C and demonstrated that the specific water production of 4.7 kg/kg of silica gel is possible [161]. Ng et al. introduced and patented a novel hybrid desalination method “MEDAD cycle” that is a combination of conventional MED and AD cycle. This novel desalination cycle can mitigate the limitations of conventional MED system to increase the system performance.

Fig. 12. Conventional MED and hybrid MEDAD cycle water production profiles at 38 °C heat source temperature.



11

Table 7 A comparison of conventional MED and hybrid MEDAD systems' temperatures at different heating temperatures. Heat source temperature (°C)

MED only

MEDAD hybrid

MED only

MEDAD hybrid

MED only

MEDAD hybrid

SG

SG

Stage-2

Stage-2

Stage-3

Stage-3

70 65 60 55 50 45 40 38

63.9 59.6 55.4 51.5 47.4 43.9 38.7 37.4

54.4 50.7 49.2 48.5 44.6 41.1 36.6 35.4

62.8 58.7 54.4 50.5 46.7 43.4 38.4 37.2

50.8 47.0 45.7 44.8 41.6 38.6 34.8 33.5

62.3 58.1 53.8 49.9 46.2 43.0 38.1 37.0 – – – – –

Operating limit of Conventional MED. The lower operational points are from MEDAD Hybrids 35 – 30.9 – 27.9 30 – 26.0 – 23.7 25 – 22.1 – 19.9 20 – 18.1 – 14.2 15 – 13.2 – 11.5

This combination allows the MED last stage to operate below ambient temperature typically at 5 °C as compared to traditional MED at 40 °C. This not only reduces the corrosion chances but also increases the distillate production to almost 2 folds as compared to traditional MED systems. 4.1. MED-AD simulation MEDAD cycle theoretical model has been developed [106,107] and presented in Table 6. The simulation is based on a fully transient model and the predicted results are compared with conventional MED system. It is observed that at the same input parameters such as a topbrine temperature (TBT), water production can achieve up to two-fold increase when the hybridized MEDAD is compared with the MED system alone. Fig. 7 shows the transient temperature profiles of a MEDAD cycle. It can be seen that last stages of MED are operating below ambient temperature due to hybridization. It can also be noticed that MED last stages profiles are affected by cyclic AD operation. Fig. 8 mapped the performance parameters of hybrid MEDAD cycle (concentration, GOR, PR and WPR). For each stage, three lines show heat source, tube surface and evaporator temperatures. It can be observed that the batchedmode water vapor uptake by the coupled AD cycle dominates the performance of the cycle. The performance of the MEDAD cycle with different additional effects with the conventional 8 effect MED cycle as baseline cycle is studied in terms of water production rate. Fig. 9 shows the improvement in the water production of the proposed MEDAD cycle with different numbers of MED stages (8–12) at different heat source temperatures. At each temperature, the percentage of water production by hybridization can be observed clearly for the MED stages. Another aspect of the hybridization is that the desorbed water vapor from the AD cycle can be recovered back into the MED system for further energy recovery.

MED condenser

AD evaporator

46.7 42.5 41.3 40.8 37.9 34.9 31.6 30.3

56.6 53.9 49.4 45.8 42.7 40.3 36.1 35.3

26.1 25.8 25.5 25.1 23.2 22.4 21.0 18.2

23.5 19.1 16.3 11.3 9.4

– – – – –

16.1 12.2 10.1 7.1 5.6

that this drop in pressure and temperature of the last stage also affected the operational parameters of the few proceeding stages. Experiments are conducted in two steps. In the first part, the system is operated as a conventional MED at assorted heat source temperature ranges from 38 °C to 70 °C. In the second part, experiments are conducted as a hybrid MEDAD system at assorted heat source temperature ranges from 15 °C to 70 °C and results are compared with conventional MED system. Fig. 11 shows the instantaneous temperatures of MEDAD components at heat source temperature of 38 °C. It can be seen that steady state events (minimum temperature fluctuations) occur after 1 h from start-up and experiments for distillate collection are continued for 4 to 5 h. It is noticed that the inter-stage temperature difference (ΔT) is more than twice per stage as compared to the conventional MED stages: This is attributed to the vapor uptake by the adsorbent of AD cycle, resulting in the increase of vapor production. The MEDAD cycle yields a stage ΔT from 3 °C to 4 °C as compared to 1 °C or less in the case of MED alone. Table 7 shows the comparison of MEDAD and MED components' steady state temperature values. Fig. 12 shows the distillate production trace at heat source 38 °C from MED stages, AD condenser and combined. The batch operated AD production can be seen clearly. At the start of desorption the production is higher and it drops with time to zero during the switching period while MED stage production is quite stable. Small fluctuations in MED water production may be due to the fluctuations in the spray of the feed that affect the condensation rate. Water production profiles are similar as explained in simulated results. Fig. 13 shows the comparison of water production of MED and hybrid MEDAD cycles at assorted heat source temperatures. Quantum increase in water production (2–3 folds) can be observed at all heat source temperatures. These results have a good agreement with

4.2. MED-AD experimentation A 3-stage MED system is fabricated and installed in NUS as shown in Fig. 10. In MED stages, vapor emanation from feed seawater is achieved by falling film-evaporation process. Evaporation energy is recovered by a series of re-utilization of vapor condensing energy in successive stages, those are produced in preceding stages. Process of vapor production and energy recovery by condensation continues until the last stage of MED. The vapors from the last stage are then directed towards AD beds where they adsorbed on the adsorbent surface. Adsorbent high affinity for water vapor drops the pressure and hence the saturation temperature of last stages below ambient typically up to 5 °C. It is observed

Fig. 13. MED and MEDAD steady state water production at different heat source temperatures.


12


Table 8 Detailed exergy modeling of the power plant combined with the MED desalination stages. State points

Exergy calculation

1−a (HP-T) b−2 (LP-T)

ΔE1−a ¼ m 1 ½ðh1 −ha Þ−T o ðS1 −Sa Þ

Total exergy across PP Extracted steam exergy Exergytotal, cogen % exe across turbine % exe for desalination

•

• • • • ΔEb−2 ¼ m 1 hb −m 2 h2 −T o m 1 Sb −m 2 S2 • • • m 2 ¼ m 1 −m extracted EPP = EHP − T + ELP − T • ΔEextracted ¼ m extracted h4 −h f ðT O Þ −T o S4 −S f ðT O Þ ΔETotal = ΔEPP + ΔEextracted

Comments

38,396.61

Total available work from HP-T

115,530.09

Available work from LP-T from c-4

153,926.70 7027.39

Total exergy across power plant Available work from extracted steam

160,954.09

ΔE

95.6%

Total exergy emanated from boiler (PP + steam extracted for Desal) (PP + steam extracted for Desal) Percentage of exergy utilized by power plant

ΔEExtracted ΔETotal;cogen

4.4%

Percentage of exergy utilized by desalination

pp %EPP ¼ ΔETotal;cogen

%EDesal ¼

Exergy value at 20% extraction

simulation results. It can also be seen that in conventional MED system last stage temperature is limited to 38 °C due to condenser operating with cooling water from cooling tower. While in the case of hybrid MEDAD, the last stage temperature can be as low as 5 °C because there is no condenser and last stage is connected to AD beds for vapor adsorption. This higher overall operational gap in proposed hybrid MEDAD cycle helps to insert more number of stages (up to 19 stages) as compared to conventional MED system (about 4–6 stages). More number of stages increase the vapor condensation heat recoveries and hence the water production at the same top brine temperatures.

5. Exergy analysis for operational cost apportionment An exergetic model is developed for a dual purpose plant (power and desalination) fuel cost apportionment, as shown in Fig. 13. The detailed mathematical equations are provided in Table 8. By implementing an improved method, as shown in Fig. 14, the fuel cost apportionment of a dual-purposed power and desalination plant and total water production cost are calculated as shown in Fig. 15. Traditionally an energy apportionment method was developed that is not a true representative for dual-purposed plant cost apportionment.

Fig. 14. Simplified model of dual purpose plant (PP + MED).



13

Fig. 15. A comparison of life-cycle unit water cost for various desalination methods [164–166].

For dual purposed plants, the consideration of quality of working fluid utilized is very important. Even though many researchers presented exergy analysis, there is no single model available for cost apportionment on the basis of exergy of working fluid utilized by the processes [162,163]. We proposed an exergy based analysis to apportion the primary fuel cost for dual purpose (power + desalination) plants. Generally, the bled-steam is nominally at 20% steam of the total steam flow and it is usually extracted from the last stages of the low pressure turbines for powering the thermally-driven desalination system. From the calculations for a 20% bled-steam extraction from the turbines (LP-T), the exergy and energetic ratios of power-to-water for the cycle described earlier are 95.7%:4.3% (exergy) and 72.2%:27.8% (energetic), respectively. Fig. 15 shows the operational cost of desalination systems by using exergy factor when it is combined with power plant. It can be seen from the life cycle costs (LCC) of assorted desalination methods, the unit water production cost is highest for PP + MSF which amounts to US$1.201/m3, while the lowest unit cost is the method of PP + MEDAD which is only at $0.485/m3. This unit cost is even lower than the LCC of the reverse osmosis (RO) plants. 6. Conclusions The trends in cogeneration of electricity and desalination plants have improved both the energetic and exergetic utilization of the thermally-activated processes. The authors have demonstrated the correct use of exergetic approach to fuel-cost apportionment rather than the conventional energetic analysis despite the high latent heat of condensation in the working steam. The lower exergy of bled-steam consumed by the MED desalination processes should cost significantly lesser by comparison to the high exergy steam which is needed for electricity production of turbines. The recent hybridization of the thermally-driven cycles, such as the AD to the MED cycles, have demonstrated their excellent thermodynamic synergy between the thermally-driven processes, improving the specific water production yields by more than two-fold at almost

similar heat input temperatures or TBTs. The quantum improvements in water production is due mainly to the operation of the vapor suction by the AD cycles, resulting in the higher temperature differences between stages as well as the lowering of the bottom-brine temperatures to below ambient. Experiments from laboratory-scale pilots have validated the water production increase by more than 2 folds. In the authors' opinion, the hybrid desalination cycles provide the key methodology in the crucial nexus between energy, water and environment that is needed for a more efficient desalination industry.

Abbreviation GDP gross domestic product Bm3 billion cubic meters Mm3 million cubic meters GCC Gulf Cooperation Council countries. MED multi-effect desalination AD adsorption desalination RO reverse osmosis HAB harmful algae blooms EDF energy distribution function CA condensation approximation SWDP specific daily water production GOR gain output ratio PR performance ratio TBT top brine temperature LBT lower brine temperature PDF probability distribution function EDF energy distribution function

Symbols C pressure constant for Eq. (6), [−] c Langmuir–Freundlich power factor, [−] E, Ec characteristic energy, [kJ/kmol]


14

K′ K, ϕ P P0, Ps R T t


Statistical rate constant, [−] pre-exponent constant for adsorption, [kPa−1] pressure, [kPa] saturation pressure, [kPa] universal gas constant, [kJ/kmol·K] temperature, [K] surface heterogeneity, [−]

Greek symbols γ Dubinin–Astakhow power factor, [−] ε adsorption site energy, [kJ/kmol, or kJ/kg] εc equilibrium adsorption site energy, [kJ/kg] ε1 to ε3 reference energy, [kJ/kmol] θ adsorption uptake, [kg/kg of adsorbent] e θ local adsorption uptake, [kg/kg of adsorbent] μ chemical potential, [kJ/kmol] χ(ε) energy distribution function, [−]

Subscript α g l/f g/v Pe e/elec th Desal Cogen PP

adsorbed phase gaseous phase liquid gas primary electrical thermal desalination cogeneration power plant

Acknowledgment The authors wish to thank National Research Foundation (NRF) Singapore (grant WBS no. R-265-000-399-281) and King Abdullah University of Science & Technology (KAUST) (project no. 7000000411) for the financial support for MED plant at the National University of Singapore. References [1] Kyaw Thu, Anutosh Chakraborty, Young-Deuk Kim, Aung Myat, Bidyut Baran Saha, Kim Choon Ng, Numerical simulation and performance investigation of an advanced adsorption desalination cycle, Desalination 308 (2013) 209–218. [2] Kim Choon Ng, Kyaw Thu, Young-Deuk Kim, Anutosh Chakraborty, Grey Amy, Adsorption desalination: an emerging low-cost thermal desalination method, Desalination 308 (2013) 161–179. [3] http://www.rubiconwater.com/news/869/usa-why-we-need-to-improve-wateruse-productivity (DoA: August 2014). [4] http://www.un.org/waterforlifedecade/scarcity.shtml (DoA: August 2014). [5] Don Hinrichsen, Henrylito Tacio, The Coming Freshwater Crisis is Already Here, http://www.wilsoncenter.org/sites/default/files/popwawa2.pdf (DoA: August 2014). [6] www.2030waterresourcesgroup.com/water_full/ (DoA: August 2014). [7] Jim S. Wallace, Peter J. Gregory, Water resources and their use in food production systems, Aquat. Sci. 64 (2002) 363–375. [8] P. Pinstrup-Andersen, R. Pandya-Lorch, M.W. Rosegrant, The World Food Situation: Recent Developments, Emerging Issues, and Long-term Prospects, IFPRI, Washington, D.C., 1997 [9] V. Frenkel, Desalination methods, technology and economics, Desalination Conference Santa Barbara, CA, USA2004. [10] http://webworld.unesco.org/water/ihp/db/shiklomanov/summary/html/summary. html#5.%20Water (DoA: August 2014). [11] Semih Otles, Serkan Otles, Desalination techniques, Electron. J. Environ. Agric. Food Chem. 4 (2004) 963–969. [12] Paul Alois, Global Water Crisis Overview, http://www.arlingtoninstitute.org/wbp/ global-water-crisis/441#2007 (DoA: August 2014). [13] Table of World Water Distribution, http://serc.carleton.edu/details/images/12447. htmlJuly 20 2008 (DoA: August 2014). [14] Earth's Water Supply, http://gen.uga.edu/documents/water/Earth.pdf (DoA: August 2014).

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