Energy savings on heavy oil transportation through core annular flow pattern: An experimental approach

International Journal of Multiphase Flow 122 (2020) 103127

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Energy savings on heavy oil transportation through core annular flow pattern: An experimental approach Nelize Maria de Almeida Coelho a,b,˚, Maria Elena Santos Taqueda b, Nayara Mota Oliveira Souza a,b, José Luis de Paiva b, Aldo Ramos Santos a, Luis Renato Bastos Lia a, Marlene Silva de Moraes a, Deovaldo de Moraes Júnior a a

Department of Chemical Engineering, Santa Cecilia University, Rua Oswaldo Cruz, 277, 11045-907 Santos, SP, Brazil Department of Chemical Engineering, Polytechnic School, University of Sao Paulo (USP), Av. Prof. Luciano Gualberto, Travessa 3, 380, 05508-010 Sao Paulo, SP, Brazil

b

a r t i c l e

i n f o

Article history: Received 27 March 2019 Revised 28 September 2019 Accepted 29 September 2019 Available online 11 October 2019

1. Introduction The current world economic scenario has its directions strongly oriented to the availability of oil and the expectation is that this context should be maintained in the medium term (International Energy Agency, 2013). However, light oil reserves which have dominated the exploration and production scenario since the beginning of the petrochemical industry are expected to deplete in the coming decades. As a result, it will be necessary to explore heavy oil fields in an economically viable way and in sufficient volume to meet the demands of the market. Heavy oils account for about a third of the world’s hydrocarbon reserves. They are mostly located in Alberta, Canada, as well as in the Orinoco belt, Venezuela (Martinez-Palou et al., 2011). However, the production of this oil still has little impact on the world market. In terms of Brazil, heavy oils make up about 11% of current production, especially in the reserves of Rio de Janeiro, Ceara and Espirito Santo States (ANP, 2019). The production of highly viscous oil presents major technological challenges because of the great pump energy consumption, the high pressure drops and the environmental risks, especially in its transportation. One of the possible solutions transporting

˚ Corresponding author at: Department of Chemical Engineering, Santa Cecilia University, Rua Oswaldo Cruz, 277, 11045-907 Santos, SP , Brazil. E-mail addresses: [email protected] (N.M.d.A. Coelho), [email protected] (M.E.S. Taqueda), [email protected] (N.M.O. Souza), [email protected] (J.L. de Paiva), [email protected] (A.R. Santos), [email protected] (L.R.B. Lia), [email protected] (M.S. de Moraes), [email protected] (D. de Moraes Júnior).

https://doi.org/10.1016/j.ijmultiphaseflow.2019.103127 0301-9322/© 2019 Elsevier Ltd. All rights reserved.

this good at lower pumping costs is to reduce the effects of viscosity by adding solvents, heat or diluting heavy oil with light oils. Nevertheless, these methods have several operational limitations, they are very expensive and are only economically feasible for short distances (Bensakhria et al., 2004; Martinez-Palou et al., 2011; Strazza et al., 2011b). Among the potential techniques to overcome the heavy and ultra-heavy oil disadvantages, there is the core annular flow (CAF). CAF is characterized by the least amount of electric power consumption to pump oil (Bannwart, 2001). With this technique, water is carefully injected to the oil stream in order to perform an annular film along the pipe wall, encapsulating the core region where the oil flows. With this configuration, since the oil barely touches the pipe wall, the wall shear is comparable to the shear of a pure water flow under similar flow conditions (Ooms et al., 1983; Ghosh et al., 2009). Isaacs and Speed (1904) were the pioneers to discuss the use of a core annular pattern for oil transportation. However, the first scientific papers were only published in the late 1950s by researches from the University of Alberta (Charles et al., 1961; Charles and Redberger, 1962; Russell et al., 1959; Russell and Charles, 1959). Since then, several experimental works have been aimed at identifying and describing the various possible oil-water flow patterns for different oil viscosities and pipe arrangements, as well as the pressure gradient reduction concerning the CAF pattern (Al-Wahaibi et al., 2014; Bai et al., 1992; Bannwart et al., 2004; Hanafizadeh et al., 2015; Loh and Premanadhan, 2016; Sotgia et al., 2008; Yusuf et al., 2012). However, according to Table 1 there are few studies concerned to oil with viscosities as high as 2750 mPa. s.

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N.M.d.A. Coelho, M.E.S. Taqueda and N.M.O. Souza et al. / International Journal of Multiphase Flow 122 (2020) 103127 Table 1 Experimental data on oil–water flow pattern maps. Author

Charles et al., 1961 Bai et al., 1992 Trallero et al., 1997 Fairuzov et al., 2000 Bannwart et al., 2004 Rodriguez and Oliemans, 2006 Sotgia et al., 2008 Vuong et al., 2009 Strazza et al., 2011a, b Yusuf et al., 2012 Al-Wahaibi et al., 2014 Mukhaimer et al., 2015 Hanafizadeh et al., 2015 Loh and Premanadhan, 2016 Luo et al., 2017

CAF

Oil viscosity (mPa. s)

Superficial velocity (m/s) Oil

Water

Yes Yes No No

65.0 600 28.8 5.071

0.02–1.21 0.0031–0.33 0.01–1.0 0.3–1.1

0.03–1.07 0.0036–0.40 0.01–1.6 0.01–0.4

Yes

488

0.007–2.5

0.04–0.5

No

7.5

0.02–3.0

0.02–2.55

Yes No Yes

919 1070 900

0.19–0.97 0.1–1.0 0.03–0.7

0.006–5.5 0.1–1.0 0.2–3.0

Yes Yes

12 12

0.10–2.0 0.10–2.0

0.10–2.6 0.10–2.6

No

1.85

0.05–3.0

0.05–3.0

Yes

4.5

0.05–0.65

0.1–1.0

No

30 and 300

0.05–2.0

0.05–2.0

Yes

1025.3–11232.8

0.08–0.10

0.015–0.90

The flow patterns are the result of the interaction between the gravitational, inertial and interfacial tension present in the flow. Thus, the spatial arrangement of the phases in the flow is dependent on the physicochemical properties of the fluids and the characteristics of the installation. In the first case, is reported the superficial velocity of the phases, the volumetric fraction of the components, the difference in density and viscosity of the fluids and the wettability characteristics of the duct walls. In the second case, they refer to the position, the material (roughness), the diameter and the geometry of the pipe, besides the presence of hydraulic fittings (Angeli and Hewitt, 1998; Bannwart et al., 20 04; Brauner, 20 03; Hanafizadeh et al., 2015; Loh and Premanadhan, 2016). In view of this, flow pattern maps are specific to the operating conditions employed in each study. For instance, if the density difference between the fluids is too large, the CAF pattern is not observed. Therefore, the introduction of different observations and new experimental data to the literature is fundamental for the development of mathematical models capable of accurately predicting in real time flow patterns and pressure drops of the biphasic oil-water flow as well as estimating the reduction of the pump energy consumption in an extended scenario of practical situations. Previous studies available in the literature describe theoretically and experimentally the benefits of transporting heavy oil with water, especially in CAF patterns (Arney et al., 1993; Grassi et al., 2008; Joseph et al., 1997; Rodriguez et al., 2009; Rovinsky et al., 1997). The energy savings provided by this technique are usually presented in terms of the pressure-gradient reduction factor (fP ), which corresponds to the single-phase oil over the oil–water two phase (same oil flow rate) pressure gradient in a long straight test section. Furthermore, some works estimate the pump power reduction factor ( f ) using this basis (Table 2). However, the power reduction factor estimated through pressure drop refers only to friction issues and is usually based on straight section measurements. Hence, it does not consider the energy loss due to the pipe fittings, which is not within the same order of magnitude as that determined on straight sections. To the best knowledge of the authors there is no experimental data concerning the overall power reduction factor of a CAF pumping facility, considering the impact of several hydraulic fittings, as

valves, elbows, long bends, couplings and directional flow changes. According to Prada (1999) and Joseph (1997) oil fouling is more severe near these line irregularities and pumping stations, therefore, it would not be expected similar yield to the straight section. Although it represents more conservative numbers, they would be more likely to the ones that can be obtained in conventional industrial plants. This paper presents a detailed follow-up of the flow patterns evolution in three sequential sections of tests (two horizontal sections separated by an intermediate vertical section). In addition, the influence of hydraulic fittings present in the experimental unit will be discussed, with emphasis on the 90° long bend. The energy coefficients of the heavy oil flow will also be reported using traditional monophasic oil pumping and biphasic oil–water pumping, especially in the CAF pattern, allowing quantitative evaluation in the attainable energy savings provided by the CAF technique. 2. Materials and methods 2.1. Test facility The experiments reported in this paper were performed at the Unit Operations Laboratory of Santa Cecilia University (UNISANTA) located in Santos, Brazil, where the test facility was designed, built and installed. The oil-water flow bench (Figs. 1 and 2) was fully constructed using transparent acrylic tanks and transparent clear PVC pipes for full visual monitoring. It is composed of one separation tank, one oil accumulation tank, and two cargo tanks, one for oil and one for water. A regenerative turbine water pump (D) and a gear oil pump (B) transport the fluids from their respective tanks to the test rig. Both pump rotational speeds are controlled by individual frequency inverters which are also responsible for flow rate variations in the experiments. Oil flow rate is measured using a calibrated pump rotation (maximum experimental uncertainties of 1.5%), while water flow rate is provided by a rotameter (maximum experimental uncertainties of 5%). Oil and water are introduced into the pipeline concentrically in an oil core and water annulus flow by means of an injection nozzle (A) fitted centrally in the pipe. This injection device is very similar to the ones used by Bensakhria et al. (2004) and

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Table 2 Pressure gradient and power reduction factors from previous studies. Author Research category Pipe internal diameter -ID

Pipe position Flow type

μO

Oil cut (%)

f P

f a

pmPa. sq

ρO

pkg{m3 q

Russell and Charles (1959) Theoretical Rovinsky et al. (1997) Theoretical

f P “

Horizontal CAF fully eccentric

–

–

f P “ 40

–

80.0b 92.2c

f Ppmaxq “ 2053b f Ppmaxq “ 894c f P “ 15μμOW

f  “ 981.2d f  “ 19.μ4 OμW

95.0

–

f  pmaxq “ 225.3

90.0 45.0

f Ppmaxq “ 95 f Ppminq “ 30

–

96.0 96.5

f Ppmaxq “ 60 f Ppmaxq “ 80

–

Vara (2001) Experimental ID 28.4 mm (Glass)

Horizontal CAF eccentric

Sotgia et al. (2008) Experimental ID 26.0 mm (Plexiglas)

Horizontal CAF eccentric

Strazza et al. (2011b) Experimental ID 21.0 mm (Plexiglas) ID 22.0 mm (Glass)

Horizontal CAF eccentric

c d

μO

–

Vertical CAF concentric

b

f “

–

Prada (1999) Experimental ID 27.6 mm (Steel)

a

μO

Horizontal CAF concentric

64, 660 972 1193 946 919 889 900 886

2 μW

2.78 μW

f  pmaxq “ 1694

Indirectly estimated by pressure drop measurements. Best scenario. Worst scenario for fP(max) , which depends on the superficial velocity of the fluids. f  is the average f (max) .

Strazza et al. (2011). The 27-mm-ID pipe loop for the biphasic flow is, approximately, 8 m long with two horizontal sections (H1 and H2), one upward vertical flow section (V) and several pipe fittings as valves, unions, elbows, 90° long bends, nipples and couplings. The facility allows experiments to be conducted for a range of superficial oil velocity from 0.07 to 0.40 m/s and a range of superficial water velocity from 0.15 to 0.46 m/s with variations in the input water cut from 32 to 87%. Cleaning procedures for oil fouling removal are also possible using a pig launcher (C) placed next to the injection nozzle that introduces into the pipeline a flexible foam pig, which is driven towards the loop end by clean water. Those are suggested to establish a similar initial pipe condition for all running tests. Temperature is measured by a thermometer placed in the oil cargo tank. Once the 200-L-available oil volume is used, 48 h rest is needed in order to separate the oil and water to be able to start a new test. All the experimental runs were performed under temperature conditions close to 23.5 °C.

Flow characterization was accomplished through multiple photographs and recordings from each combination of oil and water flow rates. Those were obtained with a high-definition digital camera model Cannon EOS-500D equipped with a 50 mm focal length lens, 1/40 0 0 s shutter speed and f/1.8 aperture operational conditions. This set up generated a freeze aspect of the flow and the details that could not be seen with plain eyesight were revealed. To improve the quality of the photographs, a glass pipe immersed in a water-filled-Plexiglass-box surrounded by an external illumination chamber was introduced in the test section replacing part of the original pipe. Thus, optical distortions from the pipe wall curvature were minimized and sharper images were obtained as the glass is brighter and smoother than the transparent PVC.

2.2. Working fluids

 “ Q.ρ .g.H

The viscous fluid selected for the tests was LUBRAX GEAR 680 Newtonian lubricating oil with viscosity (μo ) of 2750 mPa s and density (ρ o ) of 945 kg/m3 (at the temperature of the experiments). The aqueous phase was distilled water with viscosity (μw ) of 1 mPa s and density (ρ w ) of 10 0 0 kg/m3 . Weekly maintenances were performed to ensure clean water in the system and to restrain microorganism’s proliferation.

where Q is the volumetric flow rate of the fluid (m3 /s), ρ is the density of the fluid (kg/m3 ), g is the acceleration of gravity (9.81 m/s2 ) and H is the head produced by the pump (m). One can obtain the liquid horsepower (WHP) by Eq. (2) which is dependent of Q and P (differential pressure across the pump – Pa) and involves parameters that can be directly measured in the experimental unit:

2.3. Flow pattern identification The biphasic oil–water flow patterns were evaluated in three different locations, the first one was in the horizontal H1 section, approximately 65 internal pipe diameters away from the injection nozzle; the second one was in the upward vertical section V, about 40 internal pipe diameters far from the first 90° long bend; and the third one was in the horizontal returning line H2, also distant around 40 internal pipe diameters from the second 90° long bend.

2.4. Power reduction evaluation The hydraulic power ( ), or energy imparted on the fluid being pumped, can be calculated by Eq. (1):

W HP “ Q.P

(1)

(2)

Considering the pressure gauge upstream the pump negligible, Eq. (2) turns into Eq. (3):

 “ Q.P

(3)

Pressure measurements downstream of each pump were carried out in parallel when the pipe loop was completely full of fluids and under steady state operation. Static pressures (difference between total and dynamic pressures) were taken in pressure taps

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Fig. 1. Schematic of the experimental setup and a zoomed photography of the injection nozzle.

with perpendicular flow holes with a maximum of 1/10 of the internal diameter of the pipe and distant from hydraulic fittings of at least 10 internal diameters of the pipe. In the water system (PW ), it was used an analog Glycerin-filled pressure gauge, 0–2 kgf/cm2 scale range and accuracy of 4.4%; in the oil system (PO ), four digital pressure gauges, two 0–10 kgf/cm2 scale range and accuracy of 0.15% and two 0–2 kgf/cm2 scale range and accuracy of 0.5% were used for every data acquisition. The hydraulic power calculation for the single oil phase flow considered only the oil pump data, while the hydraulic power of the two-phase fluid flow considered the sum of both pump data. The evaluation of the energy savings on pumping viscous oil using the core annular flow technique was carried out by comparing the total hydraulic power required to pump a certain amount of oil using the traditional single-phase oil flow ( S ) to that required to pump the same quantity of oil through two-phase oil-water fluid flow ( B ), Eq. (4):

f “

S S ; f “ B  W ` O

(4)

Thus, the higher this power consumption reduction factor p f  q, the better the energy efficiency of the two-fluid flow process; also,

power reduction factors greater than 1 indicate that the two-phase system is energetically more advantageous than the single-phase oil system. It is important, however, to emphasize that the total energy consumption concerning the oil transportation with water goes beyond hydraulic costs. In this case, it is necessary to count additional costs related to dewatering processes, water treatment and disposal in downstream facilities. Also, the optimum process condition in terms of energy consumption must take into account not only greater f attainable in the unit, but also the thickness of the water ring, since this is the safety factor that guarantees minimum contact between the oil and the pipe wall and, consequently, the efficiency of the transport. 2.5. Governing parameters The superficial velocity Ji (m/s) is the average velocity of a given fluid i (W for water and O for oil) as if each fluid is flowing in a completely filled pipeline as a single-phase (Eq. (5)):

JW “

4 QW 4 QO ;J “ π D2 O π D2

(5)

N.M.d.A. Coelho, M.E.S. Taqueda and N.M.O. Souza et al. / International Journal of Multiphase Flow 122 (2020) 103127

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Fig. 2. Overview of the experimental unit.

where Qi represents the volumetric flow rate of the fluid i (m3 /s) and D is the pipe internal diameter (m). The input phase fraction or fluid cut α i (%) is defined as the ratio of an individual flow rate to the total flow rate (Eq. (6)):

αw “

QW QW ` QO

. 100;

αO “

QO QW ` QO

. 100

(6)

The absolute pump energy consumption on transporting one volumetric unit of heavy oil (J/m3 ) by means of a single-phase conventional transport (ES ) and of a biphasic oil-water transport (EB ) is defined in Eqs. (7) and (8):

ES “ EB “

S QO

;

(7)

W ` O QO

;

(8)

The indirect calculation of the power reduction factor is defined by Eq. (9) and depends on the pressure gradient measured in two pressure taps (1/10 maximum inner pipe diameter), within H1 test section, located approximately 27 and 70 internal pipe diameters downstream the injection nozzle and with a total length of 1.20 m. For this purpose, a pressurized piezometer composed by two vertical transparent tubes connected in the top and pressurized with air (maximum experimental uncertainties of 4.4%) was employed. This apparatus has the advantage of high measuring accuracy and low cost assembly. | f  “ f P .

QO pQW ` QO q

; f P “

PS PB

(9)

2.6. Experimental procedures The following procedure was used during the experimental campaigns: (1) At the beginning, water was introduced into the pipeline at the largest possible flow rate until the pipe loop was fully filled and no air bubbles were present; (2) Oil was then introduced to the flow at the selected flow rate QO ; (3) Water flow was adjusted to promote the desired flow rate QW ; (4) Visual monitoring was performed for approximately 1 min, when the pressure readings on the downstream of the oil pump appeared stable; at this time, photographs and videos were taken and pressure measurements were executed; (5) Water flow rate was decreased, keeping constant the oil flow rate, to cover different water contents in the twophase oil–water flow; (6) Step 5 was repeated until the lowest water flow rate, at a fixed QO , was reached or until the pipeline showed signs of oil fouling. In this case, pigging cycles were performed for pipe reconditioning and the new run starts over step 1; (7) Water was set again to the largest possible flow rate and a new QO value was set;

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N.M.d.A. Coelho, M.E.S. Taqueda and N.M.O. Souza et al. / International Journal of Multiphase Flow 122 (2020) 103127

Fig. 4. Photographs of the flow patterns observed in the horizontal flow sections. Fig. 3. Photographs of the flow patterns observed in the upward vertical flow section.

(8) Steps 3–7 were repeated until all oil-water combinations were explored. The experiments were performed in duplicate (power consumption tests) and in quintuplicate (pressure drop tests) for each flow combination and the average pressure values were used for the results analysis. The flow patterns on the test sites were evaluated individually by means of image generation. It was conducted a minimum of 357 tests. 3. Results and discussion 3.1. Flow patterns The nomenclature of the flow patterns, as well as the transitional boundaries are not a consensus among the various published works. In addition to the particularities regarding fluids and installation, the individual and subjective perception of the researcher is added. In the present work, the nomenclatures adopted were strongly inspired by the publication of Bai et al. (1992) and the descriptions of flows, although subjective, they are in agreement with the reports published by Bannwart et al. (2004), Charles et al. (1961), Sotgia et al. (2008), Trallero et al. (1997), Vuong et al. (2009) and Yusuf et al. (2012). 3.1.1. Upward vertical heavy crude oil–water flow Fig. 3 shows photographs of the five flow patterns identified for the vertical section, operating with upward flow: DO/W – oil in water dispersion, B – oil bubbles in water, S – slugs of oil in water, CAFS – core annular with swirling motion and waves at the interface and CAFB – core annular with bamboo waves at the interface. Oil water dispersion (DO/W ): presence of almost spherical oil droplets dispersed randomly in the continuous phase. This occurrence was associated with high water/oil flow ratio. The turbulence of the aqueous phase caused instabilities to appear at the interface, disrupting the continuity of the oil phase and producing oil droplets with flow velocity close to the free water flow velocity. Oil bubbles in water (B): characterized by the agglutination of the oil droplets from DO/W pattern and the formation of larger

fragments (bubbles) which flow in the continuous medium of water. It occurred when the water flow was reduced, or the oil flow was increased in relation to the DO/W flow. In this configuration, it was already possible to observe that the larger oil segments tended to be positioned in the center of the duct. Slugs of oil in water (S) or intermittent flow: coalescence of the oil bubbles and formation of elongated and discontinuous oil blocks. Long slugs of oil in water resembled the CAFS pattern, exhibiting helical waves due to hydrodynamic torques. This event occurred with the increase of the superficial oil velocity or the reduction of the superficial water velocity in comparison to the B flow pattern. Core annular with swirling motion and waves at the interface (CAFS): merge of the oil slugs in the central portion of the pipe and appearance of a continuous oil stream surrounded by water. Presence of a distorted interface, helical movements in the direction of flow and bulky intermittent oil regions. Appearance of crests and valleys as in a transverse wave, however, without symmetry. With high water flow rates, dispersed oil drops flow freely in the water phase near the interface. As the oil flow rate increase, these drops are absorbed by the core. It occupied a large area in the flow map, comprising flows containing between 37 and 55%vol. of oil. Core annular with bamboo waves at the interface (CAFB): occurrence with high oil flow rates, above 60%. The oil core phase occupies almost the entire transversal section of the duct, leaving only a small annular space, through which the water flows. The thin film of water limited the helical movements of the oil core and the fluid flow trajectory became practically linear. In this case, buoyancy forces accelerates the oil core and it moves faster than the water annulus, which is subordinated to the wall effects (Bai et al., 1992). 3.1.2. Horizontal heavy crude oil–water flow The flow patterns observed for the horizontal sections are reproduced in Fig. 4. The basic flow patterns were classified as: DO/W – oil in water dispersion, B – oil bubbles in water, S – slugs of oil in water, CAFS – core annular with swirling motion and waves at the interface, CAFW – core annular with waves at the interface and S – stratified. Even though it was promoted concentrically from the injector, the horizontal biphasic flow presented eccentric-

N.M.d.A. Coelho, M.E.S. Taqueda and N.M.O. Souza et al. / International Journal of Multiphase Flow 122 (2020) 103127

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Fig. 5. Flow pattern maps of heavy crude oil-water transportation in a 27-mm-ID clear PVC pipe pμo “ 2750 mPa s, ρo “ 945 kg{m 3 q for (A) Upward vertical flow; (B) Horizontal H1 flow; (C) Horizontal H2 flow. The numbers above each marker represent the input oil volume fraction (%).

ity in all combinations tested. Thus, the oil tended to flow into the upper quadrants of the pipe. The horizontal flow has as characteristic to be very susceptible to the effect of gravity, since the kinetic component and the gravitational component are in perpendicular axes. Oil water dispersion (DO/W ): occurrence similar to DO/W of the vertical flow, with dispersed droplets in the continuous phase and high water/oil flow ratios. The presence of the dispersed phase (oil) was higher in the upper part of the duct. Oil bubbles in water (B): analogous to B pattern of the vertical flow and originated from the coalescence of the oil drops present

in the upper part of the pipe, forming small oil blocks that moved over a free layer of water. Flows with maximum 16% of oil cuts exhibited this pattern. Slugs of oil in water (S) or intermittent flow: discontinuous blocks of oil in the upper region of the conduit and occurrence of some oil-pipe wall contact points. In this flow pattern, an increase in the superficial oil velocity led to the CAFS pattern, while a decrease in the superficial water velocity established the stratified pattern. Therefore, this configuration was associated with a transition zone.

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Core annular with swirling motion and waves at the interface (CAFS): appearance of a continuous eccentric line of oil flowing completely enveloped by water, where the presence of several oil drops was noted. The oil core was characterized by the absence of symmetry and displaced a helical trajectory with nonuniform periods in the flow direction. Due to the density difference of the fluids, the water layer thickness present in the upper quadrants of the pipe was thinner than that of the lower sections. Core annular with waves at the interface (CAFW): attenuation of the core helical movements and the beginning of a linear trajectory. The upper interface gained ripples of almost periodic occurrence similar to ridges and valleys of a transverse wave. On the other hand, the lower interface developed slightly wavy, and in some cases, it was practically flat. Stratified (STR): presence of two continuous layers flowing in parallel, oil in the upper part and water in the lower part of the pipe. This configuration was characteristic of operations with low water and oil flow rates, where the flow was dominated by the effects of gravity. The interface was shaped according to the combination of the superficial velocities of the fluids and presented either smoothly or sometimes with perturbations/undulations that resembled small crests of a transverse wave. 3.2. Flow pattern maps The experimental maps of the flow patterns observed in the three test sections (Fig. 5) were constructed with reference to the photographic images obtained from the flow and had as objective to promote the identification of the occurrence window of each pattern and to verify the particularities of the flow after crossing severe conditions of perturbation and trajectory. Individual images of the two-phase oil–water flow in the three test sections for all tested combinations of superficial velocities can be seen in the supplementary material to this text. 3.2.1. Upward vertical heavy crude oil–water flow Fig. 5(A) shows the flow pattern map for the two-phase oilwater flow in section V. Through this map it was evidenced that the input oil volume fraction was the determining factor in the type of flow pattern acquired by the fluids inside the pipe. Flows with low oil cuts, less than 15%, presented DO/W configuration, combinations between 15 and 25% of oil cuts showed B pattern. Slugs of oil-in-water flow pattern (S) were identified for a range between 25 and 37%vol. of oil. Above 37%vol. of oil cuts, only CAF pattern was observed, with helically movements until about 55% (CAFS) and with linear trajectory from this point (CAFB). According to the flow images (best seen in the supplementary material), in all flow configurations the oil flowed mainly in the center of the pipe, demonstrating that in vertical installations the oil phase tends to flow preferentially through the central part of the tube. This observation indicated agreement with other studies in the literature (Renardy, 1997). In the literature, vertical biphasic flows containing up to 96% of oil were reported by Oliemans (1986) and Prada (1999). However, it was not convenient to operate with such fractions of oil in the present work due to the presence of hydraulic fittings, specially the 90° long bend between H1 e V sections, that caused severe perturbations to the flow trajectory. As previously mentioned, line irregularities strongly favor oil fouling, promoting oil accumulation zones downstream of the pipe fittings that evolve to dramatic pressure increase over time. Fig. 6 presents a photograph of the 90° long bend region after continuous operation of the experimental unit with two-phase oil-water flow containing maximum 70%vol. of oil. From Fig. 6 one may note the existence of an oil trail in the pipe wall of approximately 5-pipe ID in length, which agrees with the

Fig. 6. Oil fouling downstream the first 90° long bend (between H1 and V section) after oil-water biphasic flow operation.

observations of the literature but did not compromise the occurrence of CAF patterns during operation. On the other hand, stable operation conditions with CAF pattern at oil cuts above 70% were not feasible. 3.2.2. Horizontal heavy crude oil-water flow Fig. 5(B) shows the flow pattern map of the biphasic flow in section H1. Differently from the vertical flow, a direct and absolute relation of the input flow ratio with the observed flow pattern was not verified. As an example, a biphasic flow combination with 46.4% of oil cut has shown either CAFW or STR flow patterns. Likewise, a flow with as low as 28%vol. of oil exhibited a CAFW pattern, while another one with 56.5%vol. of oil has shown a STR pattern. Therefore, as previously reported by Oliemans (1986), there is a strong dependence between the superficial velocities of the fluids and the observed flow patterns, especially in horizontal flows with larges density differences between the fluids. Individual images of the flow allowed the evaluation of the effects of the superficial velocity of the fluids on the observed flow pattern (supplementary material). The progressive increase of JW had as consequences: intensification of the waviness at the interface; segregation of oil blocks and oil droplets to the water phase; discrete increase in the water ring thickness at the upper quadrants; and incorporation of helical motions in the core. On the other hand, the increase in JO resulted in oil particles agglutination and formation of oil blocks. Also, higher JO slight shifted the oil core towards the center of the pipe (reduction of the eccentricity) and the core diameter had growth. From the flow maps presented in Fig. 5(A) (V section) and Fig. 5(B) (H1 section) it was interesting to note that while H1 sec-

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9

Fig. 7. Comparison of pump energy consumption on heavy oil transportation by conventional method (pure oil flow) and oil–water biphasic flow. (A) General overview; (B) Unfold view of all biphasic flow combinations and their flow patterns in H1 section; (C) Surface plot for the biphasic flow.

tion had shown already CAF pattern for JO as low as 0.13 m/s and JW of 0.34 m/s, V section still exhibited slugs of oil-in-water profile for the same water-oil ratio. The first CAF pattern on V section, considering the same JW , was observed only when operating with JO of 0.20 m/s. In this case, the presence of a 90° long bend between H1 and V sections plus the occurrence of buoyancy forces that accelerate upward vertical flows led to the increase of local fluid velocity in the vertical region. Thus, to demonstrate a continuous-oil-core line in the presence of higher kinetic energy conditions, CAF pattern on V section required higher oil cuts. Fig. 5(C) shows the flow pattern map of the biphasic flow in section H2. The main differences observed in this map was the increase of the window of occurrence of B and CAFS, besides the absence of STR flow. The appearance of the oil core in the H2 CAF

flows was revealed twisted and asymmetric for a large range of combinations and it was a result of the fluid velocity increase from their passage through the second 90° long bend. In this context, the higher local superficial velocity stimulated by the second 90° long bend and the favorable tendency of oil to occupy the center of the pipeline in vertical flows led the H2 section to a more advantageous operating conditions compared to H1 section even under similar initial conditions. The absence of a STR flow was very convenient since this biphasic flow configuration has low energy efficiency. Accordingly, it was possible to verify that, in general, the CAF pattern was favored in the horizontal H2 test section due to the provision of a vertical section in the system. Therefore, a vertical section installed between long horizontal segments could be

10

N.M.d.A. Coelho, M.E.S. Taqueda and N.M.O. Souza et al. / International Journal of Multiphase Flow 122 (2020) 103127

Fig. 8. Absolute energy consumption on transporting oil through oil-water biphasic flow for various combinations of superficial velocity of the fluids in a 27-mm ID PVC pipe.

used as a flow rectifier in which the curves, even being subject to a probable deposition of oil, would provide a local velocity increase beneficial to the occurrence of the CAF pattern. Thus, the optimal vertical/horizontal length ratio to obtain such positive results deserves further investigation as vertical sections also have the drawback of imposing high hydrostatic heads. 3.3. Pump energy consumption As already mentioned, the friction reduction of the viscous fluid with the pipe provided by the technique of transporting heavy oil with water promotes great savings on pumping energy. In this context, biphasic flows containing very low water cuts would be those commercially more interesting. They would minimize the energy consumption by volume of transported oil as well as the costs associated with the separation of fluids. Technically, however, an oil flow containing a very thin annulus of water would dramatically expand the probability of oil touching the pipe wall, reducing the energy efficiency of the transport, especially if one considers a real industrial setup that contains many hydraulic fittings and directional flow changes. Those line irregularities cause disruptions in the flow and, in the case of CAF pattern, its orderly configuration would only last if more conservative water flow rates were employed, which means, greater water ring thickness. Accordingly, the total energy expended by the pumps to promote flow is dependent, among other factors, on the presence of hydraulic fittings in the pipeline. In other words, to do a certain work, a facility that contains hydraulic fittings will require more energy than a similarly sized one consisting exclusively of straight sections. Fig. 7 presents the experimental data regarding the energy consumption when pumping heavy oil alone (monophasic flow) and with addition of water (biphasic flow) for various combinations of superficial velocities. In this figure the flow patterns for the biphasic flow in the horizontal test section H1 (more severe runoff condition) were also highlighted. In the single-phase flow, a linear ratio between the energy consumption and the superficial oil velocity was verified; however, for biphasic flow this relation was a 2nd order polynomial. In addition, it was observed that higher superficial water and oil velocities led the energy curves to converge to a similar region, indicating that under these conditions the biphasic flows did not present significant differences among them from the energetic point of view.

Fig. 9. Estimated pump power reduction factor in a 27-mm PVC pipe for increasing (A) superficial oil velocity and (B) input oil volume fraction. The red line represents the maximum power reduction factor for the experimental bench working with LUBRAX GEAR 680 oil. (For interpretation of the references to color in this figure legend, the reader is referred to the web version of this article.)

Space projections of the energy curves are also presented in Fig. 7 and one can observe that the least absolute energy consumption of the biphasic flows was registered mainly in the region where the first formations of the CAF profile were verified. The exception was in the low superficial velocities of the fluids (JW “ 0.15 m/s and JO less than 0.20 m/s), where the STR pattern dominated section H1, but had S pattern prevailing in sections V and H2. Thus, economically advantageous biphasic flows started at superficial oil velocities close to 0.15 m/s. Fig. 8 presents the experimental data of energy consumption from the input oil volume fraction point of view. It was observed that the region of minimum energy consumption occurred close to oil cuts between about 30 and 35%, that is, the region of onset of the CAF profile as shown in Fig. 7(B). However, it is wrong to assume that this point of minimum energy consumption is the optimum operation point of the biphasic flow, since this is an absolute energy evaluation of the flow. This minimum point marks the approximate fraction of oil from which the biphasic flow becomes competitive. Thus, the optimum operation condition should

N.M.d.A. Coelho, M.E.S. Taqueda and N.M.O. Souza et al. / International Journal of Multiphase Flow 122 (2020) 103127

be determined from the relative energy consumption, which is expressed as the power reduction factor ( f ). The power reduction factors for the biphasic oil-water flows are shown in Fig. 9(A). It can be seen that, in general, flows with low water cuts obtained larger reduction factors; however, with the increase of the superficial oil velocity, the difference in energy efficiency decreased so that a hypothetical maximum reduction factor close to 2.2 could be established. Similarly, Fig. 9(B) shows the relation of the power reduction factor with the input oil volume fraction, to which again the 2.2 factor was manifested, evidencing the existence of a maximum finite limit of energy reduction. Moreover, it was possible to observe that, above input oil fractions of 0.5, regardless the superficial velocity of the fluids, the power reduction factor was very close to the maximum factor. This observation implies that, for the operation conditions of this work, continuous increments of oil above 50% vol. would increase oil fouling probability and would not accomplish a compensatory energy saving. Table 2 shows the restricted number of studies that proposed correlations or experimental data concerning the pump power reduction factor for the biphasic oil-water transport method. The first theoretical model presented by Russell and Charles (1959) would predict a power reduction factor close to 1100 times for the experimental conditions of this study. Clearly, this value is very different from the 2.2 obtained experimentally, however, when proposing this model, the authors reported that, in practice, it was found a maximum reduction factor of 12 times. Another mathematical model presented by Brauner (1991) found that the maximum power reduction factor occurs when JO equals to 2 times JW . This proposition was quite consistent with the experimental data for low water flow rates in the biphasic flow (Fig. 9(A)). For instance, for JW of 0.15 and 0.21 m/s, f (about 2.2) was found very close to JO of 0.3 and 0.4 m/s. On the other hand, for high water flow rates this model was not equally representative. The only experimental studies found in the literature regarding the power reduction factor were performed by Prada (1999) and Vara (2001). However, these studies used the indirect method to determine this parameter (Eq. (9)), which differs from the direct method by not considering the presence of hydraulic fittings in estimating the overall power consumption. Thus, with the objective of comparing their experimental data and the present study, it was measured in test section H1 the pressure drop. Then, the power reduction factor was estimated by the indirect method (Fig. 10). It was observed that, through this method, the maximum power reduction factor greatly increased to approximately 175. The power reduction factors presented by Prada (1999) were very high, easily reaching values as 10 0 0 and 20 0 0 times for an experimental unit with vertical pipes and working with a 10260mPa s-viscosity oil. As the power reduction factor is proportional to the oil viscosity, it was not surprising that the power reduction factors of Prada were superior. But, in this case, the vertical positioning of the pipe favored the CAF configuration as the buoyancy forces tend to accelerate the oil core, generating an upward movement that consumes lower energy of the pumps, unlike the flow in horizontal sections. Nevertheless, in the work of Vara (2001) that comprises horizontal pipes and an oil with viscosity of 1193 mPa s, the maximum power reduction factor was 225 for a volumetric combination of 95% oil and 5% water. Considering that the characteristics of Vara’s work and those of the present study were not significantly discordant, the lower maximum reduction factor presented here was associated to experimental lower oil cuts employed. As one can observe in Fig. 10, the indirect method gives the impression that it is possible to increase the input oil flow rate to obtain greater power reduction factors, as it does not clearly exhibit a maximum limit. But as previously discussed, this analysis

11

Fig. 10. Pump power reduction factor estimated by pressure drop measurements in a straight pipe for LUBRAX GEAR 680 oil being transported with water in a 27-mm PVC pipe.

Fig. 11. Pump power reduction factor estimated by pressure drop measurements in a straight pipe in terms of oil volume fraction for LUBRAX GEAR 680 oil being transported with water in a 27-mm PVC pipe.

does not represent the overall energy savings observations. Fig. 11 reinforces this inaccurate impression. It is important to note, however, that the maximum power reduction factor of the CAF technique is specific to the facility and fluid evaluated and is related to the amount of hydraulic fittings, length of straight pipe sections, pipe diameter, efficiency of the equipment involved, wetting behavior, viscosity of the oil, among others; being possible, therefore, to verify higher factors than those obtained in this study in more simplified systems or in different operational assemblies. 4. Conclusions The experimental studies have shown that the presence of hydraulic fittings in the test unit severely affects the energy gains attainable in transporting oil with water, especially in the CAF pattern which has the lowest pumping cost. When monitoring the flow, it was evident the appearance of oil fouling zones near the hydraulic accessories, what reveals the need of a higher amount of

12

N.M.d.A. Coelho, M.E.S. Taqueda and N.M.O. Souza et al. / International Journal of Multiphase Flow 122 (2020) 103127

water to keep CAF structure. With this, flows with very high oil cuts (above 70%) were unfeasible. The differences between the three pattern maps demonstrated the greatest difficulty in maintaining the CAF pattern in horizontal sections, in which the occurrence of the stratified pattern is frequent and should be controlled in order to guarantee high efficiency of flow lubrication. Since the CAF profile is favored in vertical sections, the presence of the vertical section V between the horizontal H1 and H2 sections revealed to be very convenient as it avoided the appearance of the stratified pattern in section H2. Therefore, this vertical section worked as a flow rectifier and the curves, even being subject to a probable deposition of oil, provided a beneficial increase in the local velocity of the fluids and a greater CAF window occurrence in the H2 test section. The power reduction factors obtained for the biphasic oil-water flow were found to be different to the ones previously presented in the literature. Most of them, however, were obtained by the indirect method, which reflects the gains related to the frictional loss reduction in a straight pipe. Nevertheless, under the presence of various hydraulic fittings, directional flow changes and equipment with different operating efficiency, this estimation may not represent, in quantitative terms, the actual power reduction that would be achieved in the installation. In this scenario, lower reduction factors are obtained, as was the case of the present experimental work that reached a maximum power reduction factor of 2.2. Therefore, the direct determination of the power reduction factor is the most recommended for calculating projects. Considering, also, the large volume of heavy oil transported every day around the world, this method of transporting oil, which allows a saving of 2.2 times, represents a great economic impact within the company, even though there is a need to install an oil-water separation system to complete the process.

PO were estimated through 35 nested (hierarchical) design (corresponding to 35 flow combinations), where the components of the variance (process and test variation) for each nested design were determined. Then, the pool variance for the entire experimental arrangement was calculated using a weighted average based on the degrees of freedom of each individual variance calculated in the nested designs (Box et al., 2005). It is important to point out that the analysis of the components of the variance for PO indicated that 90% of PO uncertainty was due to the process and only 10% due to test variation. In this case, the absence of a 3D symmetry in the biphasic flow naturally produces pressure readings downstream of the oil pump that are difficult to accurately reproduce. The standard errors were estimated by taking the square root of the variances. The uncertainties in all independent variables are summarized in Table 3. Dependent variables: the uncertainties associated to the dependent variables pJo, Jw , αo, αw , o, w , s , f , | f , EB , Es q were estimated by the uncertainty propagation law. Considering a generic variable Y which is not measured directly, but determined as a function of n independent variables Y “ f px1 , x2 , . . . , xn q, and their individual uncertainties σ1 , σ2 , . . . , σn , the uncertainty in Y, expressed as σ Y , can be obtained by Eq. (A-1) (Taylor and Kuyatt, 1994): d ˆ ˙2 ˆ ˙2 ˙2 ˆ BY BY BY σY “ σ1 ` σ2 ` . . . ` σn (A-1) B x1 B x2 B xn

Declaration of Competing Interest

B JO “ B QO

We wish to confirm that there are no known conflicts of interest associated with this publication and there has been no significant financial support for this work that could have influenced its outcome. Acknowledgment The authors thank Santa Cecilia University and University of Sao Paulo, especially their respective Unit Operations Laboratories for supporting this work. Sincere thanks are extended to the workshop technicians for all interesting discussions and vital help during the construction and assembly of the experimental bench. This research did not receive any specific grant from funding agencies in the public, commercial, or not-for-profit sectors. Supplementary materials Supplementary material associated with this article can be found, in the online version, at doi:10.1016/j.ijmultiphaseflow.2019. 103127.

From the above, the uncertainties associated to all dependent variables were estimated by Eqs. (A-2–A-38) and the results are presented in Table 4. d ˆ ˙2 ˆ ˙2 B JO B JO σJO “ σQ O ` σD (A-2) B QO BD

(A-3)

B JO 8 QO “´ BD π D3 d

σJW

ˆ

“

B JW “ B QW

(A-4)

B JW σQ B QW W

˙2

ˆ `

B JW σD BD

˙2

(A-5)

4

(A-6)

π D2

B JW 8 QW “´ BD π D3

σαO

d ˆ “

B αO σQ B QO O

(A-7) ˙2

ˆ `

B αO σQ B QW W

˙2

(A-8)

B αO 100 QW “ B QO pQO ` QW q2

(A-9)

B αO 100 QO “´ B QW pQO ` QW q2

Appendix A. Uncertainty analysis In this section, the uncertainty analysis associated with the experimental measurements is presented. Independent variables: the uncertainties related to A and Qw were considered as the uncertainty of the measuring instrument (caliper and rotameter) and the ones associated to PW , PS , PS and PB were evaluated as the standard deviation of the average from independent observations. The uncertainties in QO and

4

π D2

σαW

d ˆ “

B αW σQ O B QO

(A-10)

˙2

ˆ `

B αW σQ B QW W

˙2

(A-11)

B αW 100 QW “´ B QO pQO ` QW q2

(A-12)

B αW 100 QO “ B QW pQO ` QW q2

(A-13)

N.M.d.A. Coelho, M.E.S. Taqueda and N.M.O. Souza et al. / International Journal of Multiphase Flow 122 (2020) 103127

13

Table 3 Measuring range and uncertainty for independent variables. Variable

Range/Value

Unit

Standard error

Variation coefficient (%)

QW QO PW PO PS P S P B D

8.75 x 10´5 ´ 2.63 x 10´4 3.89 x 10´5 ´ 2.28 x 10´4 27468 ´ 62784 54661 ´ 326673 120555 ´ 706320 12632 ´ 61053 72 ´ 468 0.027

m3 /s m3 /s Pa Pa Pa Pa Pa m

8.33 x 10´7 1.92 x 10´6 1865 22031 21260 324 9.7 5 x 10´5

5.0 1.5 4.4 14.2 5.8 0.9 4.4 0.2

Table 4 Measuring range and uncertainty for dependent variables. Variable

Range/Value

Unit

Standard error

Variation coefficient (%)

JO JW

0.07 ´ 0.40 0.15 ´ 0.46 0.13 ´ 0.68 0.32 ´ 0.87 2.29 ´ 74.15 4.93 ´ 159.22 2.36 ´ 16.48 0.27 ´ 2.40 6.12 ´ 175.00 140.20 ´ 479.71 130.25 ´ 701.42

m/s m/s % % W W W – – kJ/m3 kJ/m3

0.0035 0.02 1.23 1.23 2.93 0.61 2.90 0.17 3.53 22.97 23.12

1.5 6.0 3.0 2.0 10.5 7.2 4.4 11.1 5.3 8.7 5.7

αO αW O W S f | f



EB ES

d

σO

ˆ

“

B O σQ O B QO

˙2

ˆ `

B O σP B PO O

˙2

(A-14)

B O “ PO B QO

(A-15)

B O “ QO B PO

(A-16)

σW

“

B W σQW B QW

˙2

ˆ `

B W B PW

σPW

˙2

(A-18)

B W “ QW B PW

(A-19)

σS

“

˙2

ˆ `

B S σP B PS S

QO

(A-28)

PB pQW ` QO q

Bf PB . QO “´ B PB PB 2 pQW ` QO q !

Bf “ B QO

d

σE S (A-21)

B S “ QO B PS

(A-22)

(A-23)

“

(A-29)

Ps . QW

(A-30)

PB pQW ` QO q2

!

(A-20)

B S “ PS B QO

σ f

Bf “ B PS

Bf “´ B QW

˙2

g fˆ ˙2 ˆ ˙2 ˆ ˙2 f Bf B f B f e  “ σS ` σW ` σO B S B W B O

¸2

!

(A-17)

B W “ PW B QW

B S σQ B QO O

σQW

(A-27) !

d ˆ

d ˆ

g f˜ ¸2 ˜ ¸2 ˜ ¸2 ˜ f B| f f f f B| B| B| e | σ f “ σPS ` σPB ` σQ O ` BPS BPB B QO B QW

ˆ

Ps . QO

(A-31)

PB pQW ` QO q2 B ES σO B O

˙2

ˆ `

B ES σQ B QO O

˙2

B ES 1 “ B O QO B ES O “´ 2 B QO QO d ˆ ˙2 ˆ ˙2 ˆ ˙2 B EB B EB B ES σE B “ σO ` σW ` σQ O B O B W B QO

(A-32) (A-33) (A-34)

(A-35)

B f 1 “ pW ` O q B S

(A-24)

(A-36)

B f S “ B W pW ` O q2

B EB 1 “ B O QO

(A-25)

B EB 1 “ B W QO

(A-37)

B f S “´ B O pW ` O q2

(A-26)

pW ` O q B EB “ B QO QO 2

(A-38)

14

N.M.d.A. Coelho, M.E.S. Taqueda and N.M.O. Souza et al. / International Journal of Multiphase Flow 122 (2020) 103127

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