Experimental Study on Natural-Gas-Hydrate-Slurry Flow - OnePetro

J158597 DOI: 10.2118/158597-PA Date: 9-April-14

Stage:

Page: 206

Total Pages: 9

Experimental Study on Natural-GasHydrate-Slurry Flow X.F. Lv, SPE, J. Gong, SPE, W.Q. Li, B.H. Shi, D. Yu, and H.H. Wu, China University of Petroleum, Beijing

Summary To better understand hydrate-slurry flow, a series of experiments was performed, including water, natural gas, and diesel oil, under 4-MPa system pressure and 1.25-m/s initial linear velocity. The experiments have been conducted in a high-pressure hydrate-flow loop newly constructed at China University of Petroleum (Beijing), and dedicated to flow-assurance studies. A focused-beam reflectance measurement (FBRM) probe is installed in this flow loop, which provides a qualitative chord length distribution (CLD) of the particles/droplets in the system. First, the influence of flow rate on the hydrate-slurry flow was discussed. Then, we studied other influencing factors—such as water cut and additive dosage—on the hydrate induction period and the CLD before/ after hydrate formation. Third, a new correlation was fitted between the dimensionless rheological index n0 and water cut as well as additive dosage, according to these experimental data. Finally, a laminar-flow model for the prediction of the pressure drop for the quasisingle-phase hydrate slurry was established, and tested by comparison with the experimental results in this paper. Introduction Nowadays, subsea multiphase pipelines are commonly used for oil/gas gathering and transportation at deepwater oil/gas fields all over the world, along with the development of offshore resources exploitation. However, the complex deepwater conditions, usually presenting high pressure and low temperature, inevitably lead to flow accidents in the pipeline system. Major accidents include solid precipitation (e.g., hydrate, asphaltene, and wax). Therefore, deepwater flow assurance has become the focus of international oil- and gasfield development (Shi 2012). Hydrate (Fig. 1) is one of the critical precipitates that have to be controlled for subsea flow assurance. The controlling methods include the traditional inhibition and risk-control technique (Sloan and Koh 2007 and Sloan et al. 2010). In the traditional way, we can either inject the thermodynamic hydrate inhibitor (i.e., methanol or monoethylene glycol) or heat (and/or insulate) pipelines, to prevent the fluid falling into the hydrate-forming region in the pipeline. With the risk-control technique, the kinetic-hydrate inhibitor can be applied in small amounts to delay the hydrate formation without changing the thermodynamics by acting as a water or surface modifier. The risk-control technique also includes the cold-flow technology (Gudmundsson 2002; Turner and Talley 2008) and the antiagglomerants (AAs) method (Frostman 2000; Sinquin et al. 2004), in which hydrate formation is permitted. These generated hydrate particles are carried by the fluid in the pipeline to form a hydrate slurry flow, avoiding the hydrate plugging. The former, traditional, method is the main hydrate-control technology at present. But because exploration and production are entering deeper water, hydrate-formation avoidance would reach its economic limits. The focus of flow-assurance strategies is thus shifting from the traditional to the risk-control technique. However, the formed hydrate slurry might change the fluid’s property fundamentally (e.g., fluid viscosity or yield stress) (Peixinho et al.

2010), resulting in transportation problems or in the alteration of the system production capacity. Consequently, it is important to investigate the hydrate-slurry-flow behavior for the slurry transportation. In this field, Andersson and Gudmundsson (1999) performed experiments on the flow behavior of oil-based hydrate slurry flow, and found that the viscosity of the hydrate slurry increases with the increasing amount of hydrate in the system at laminar flow. Boxall et al. (2008) investigated the influencing factors (i.e., water cut and flow rate) experimentally; they focused on their effects on the trend of hydrate blockage. It was then concluded that the system with relative lower water cut faces fewer transportation problems, whereas the transporting feasibility depends on the actual flow rate for a system with a relative higher water cut. Peysson (2005) described the viscosity of the hydrate slurry by using a power-law fluid model. Sjo¨blom et al. (2010) discussed the factors of hydrate blockage and recommended capillary attractive forces between hydrate particles as the key cause of agglomeration and plugging. Furthermore, Turner et al. (2005) and Turner and Talley (2008), Boxall et al. (2008), Dellecase et al. (2008), and Wang et al. (2008a and 2008b) all conducted research on the hydrate slurry in the flow loop. However, those studies on the flow characteristics of the hydrate slurry are different from each other in such respects as experimental method, laboratory equipment, and experimental medium. Thus, there is still no consensus on the slurry-flow law and how the various factors influence the hydrate blockage. In summary, it is still necessary to perform hydrateslurry experimental research. The gas-hydrate model developed for oil-dominated systems follows the conceptual process presented in Fig. 2 (Turner 2005). Initially, water droplets are emulsified and dispersed in the continuous oil phase. As the conditions necessary for hydrate formation are satisfied, natural gas dissolved in diesel would react with the water droplets, then form a thin hydrate shell on the surfaces of the droplets, yet keeping the particle size unchanged. The growth of the hydrate shell in this shrinking-core model is controlled by the intrinsic growth kinetics, as well as the mass and heat transfers at the interfaces between water and gas (Shi et al. 2011). Along with the continuous reaction, the surface shell becomes thick, and the hydrate quantity keeps growing. What is worse, this finally leads to the hydrate accumulation and agglomeration until blockage. In this paper, a high-pressure hydrate-flow loop has been newly constructed by a multiphase-flow research group at China University of Petroleum (Beijing). A series of flow experiments (introduced in the Experimental Part section) has been performed with natural gas þ diesel þ water system on this flow loop to investigate the effects of influencing factors [i.e., flow rate, water cut, and AA dosage (see these subsections: The Influence of Flow Rate on the Hydrate-Slurry Flow; The Influence of Water Cut and AA Dosage on the Hydrate-Induction Period; and The Influence of Water Cut and AA Dosage on CLD)]. Furthermore, a laminarflow model for the prediction of the pressure drop for the quasisingle-phase hydrate slurry was established in the QuasisinglePhase Laminar-Flow Pressure-Drop Model of Hydrate Slurry subsection, according to these experimental results.

C 2014 Society of Petroleum Engineers Copyright V

This paper (SPE 158597) was accepted for presentation at the SPE Annual Technical Conference and Exhibition, San Antonio, Texas, USA, 8–10 October 2012, and revised for publication. Original manuscript received for review 18 June 2012. Revised manuscript received for review 25 March 2013. Paper peer approved 7 May 2013.

Experimental Part High-Pressure Hydrate Experimental Loop. Experimental tests have been conducted in a high-pressure hydrate experimental

206

April 2014 SPE Journal ID: jaganm Time: 14:55 I Path: S:/3B2/J###/Vol00000/130056/APPFile/SA-J###130056


(a)

(b)

Fig. 1—Two examples of pipeline blockages removed from slug catchers (Boxall et al. 2011): (a) solid hydrate plug, (b) liquid hydrate sludge.

Water Entrainment

Hydrate Growth

Stage:

Page: 207

Total Pages: 9

loop devoted to flow-assurance studies (Fig. 3). The loop separately injects gas and liquid by a plunger compressor (2200 Nm3/ h) and uses a custom-made magnetic pump to maintain a flow rate of up to 12 m3/h. It should be noted here that although this pump is designed to have a minimal destructive impact on the hydrates, its influence on the hydrate size is inevitable and should be taken into consideration, as discussed in the The Influence of Flow Rate on the Hydrate-Slurry Flow subsection. Two sight glasses sit in the test sections. The gas-injection point is the test-section inlet. At the outlet of the test section, gas and liquid collect in an insulated separator and are redirected toward the test-section compressor (from the upper part) and pump (from the bottom), respectively. Several tanks allow the maintenance of loop and separator pressure as hydrate forms.

Agglomeration

Plugging

Oil

Water

Fig. 2—A schematic for hydrate formation and transportation (Turner 2005).

T07 P05

Separator

FBRM probe Ft02 T08

NDRI P01

Magnetic Pump

T06 T01

Pressure reducing valve

T02

Mixer

P02 Ft01

Ft03

DP4 DP01

Jacket

P07

Sight glass

Gas supply line

T03

P04

P03

DP02

NDR2

T05

P04

Buffer tank

DP03

Chiller T04

Gas supply Compressor

Gas charge line Liquid charge line

Instrument codes

Test section

P: Pressure Transducer; DP: Differential pressure; T: Temperature Transducer NDR: Nuclear Densitometer; Ft: Mass flow meter Fig. 3—A schematic of the high-pressure hydrate experimental loop. April 2014 SPE Journal

207 ID: jaganm Time: 14:55 I Path: S:/3B2/J###/Vol00000/130056/APPFile/SA-J###130056


Stage:

Page: 208

Total Pages: 9

Image illustrating the view from the FBRM Probe Window

Cutaway view of FBRM In-process Probe

Outgoing Light Return Light Optics

Rotating Optics FBRM Probe Tube Sapphire Window

Focused beam

Probe installed in process stream R

Fig. 4—A principal diagram of the FBRM particle-size analyzer (from LasentecV 2010).

TABLE 1—THE COMPOSITION OF GAS SAMPLES (MOL%)

TABLE 2—COMPOSITION OF 220# DIESEL OIL (MOL%)

Composition

Mol%

Composition

Mol%

Composition

Mol%

Composition

Mol%

N2 CO CO2 C1 C2

1.53 2.05 0.89 89.02 3.07

C3 iC4 iC5 nC6þ —

3.06 0.33 0.04 0.01 —

C11 C12 C13 C14 C15

0.89 3.36 5.38 6.2 6.78

C16 C17 C18 C19 C20þ

6.83 7.99 7.46 6.38 48.73

The 30-m stainless-steel test section consisted of two rectilinear horizontal lengths joined together to form a pipe with a 2.54cm (1-in.) internal diameter and a 5.08-cm (2-in.) -diameter jacket circulating a water/glycol blend that surrounded the test section. Process temperature control ranged from 20 to 80 C. The Hydrate Experimental-Loop Instrumentation. This experimental flow loop is equipped with several sensors. Thermocouples lie along the pipe, inside the separator, inside the water/ glycol system, and on the different gas utilities. A Coriolis flowmeter measured the liquid-mixture density and flow rate. Two FM1000 gamma ray densitometers were also available to measure the multiphase fluid’s mean density. Differential-pressure sensors along the loop followed the evolution of the linear pressure drop. Rapid data acquisition permitted detection of quickly occurring events. A gas-flow metering instrument in the gas-supply system recorded the gas consumption. It should be noted here that we have not yet obtained a specific apparatus to measure the viscosity 25

Pressure (MPa)

20

of the hydrate slurry, which was instead back calculated from the slurry’s pressure drop. An additional consideration with viscosity is that it is nonuniform along the pipe in the experimental process; thus, the back-calculated result could present a more average value for the whole loop. An FBRM probe allowed monitoring the evolution of objects—droplets, bubbles, solid particles—carried inside the flow. The FBRM was composed of a low-intensity rotating laser beam reflected when intercepting a particle (Fig. 4). Measuring the reflection time allows the deduction of a chord length. Assignment of a CLD and a mean chord length followed every measurement equal to 10 seconds. The CLD gives an idea of the particlesize distribution (PSD) of objects carried by the flow. The lower and upper limits of the FBRM CLD are 0.5 and 1000 lm, respectively. A representative sampling of the PSD recommended the installation of the particle-size analyzer on the straight vertical pipe ahead of the experimental-loop inlet. The analyzer’s probe window cuts the streamlines at a 45 angle, beginning at the center of the pipe. The FBRM probe estimated the initial water-droplet size inside the fluid and followed the hydrate-particles’ agglomeration with time (Boxall 2009). The mean-square-weighted chord length gives more weight to the longer chord length and is particularly well adapted to agglomeration phenomena (Darbouret et al. 2008).

15

10

5

0

0

2

4

6

8 10 12 14 Temperature (°C)

16

18

20

Fig. 5—Hydrate-formation curve of the testing natural gas.

Fluids. The testing used deionized water, civil natural gas, and 20# diesel (Tables 1 and 2). The AAs used in this work are recompounded ones containing Span20; thus “AAs” is adopted throughout this paper. An electronic balance weighed the quality of AAs, with a measuring error of 60.01 and a high-pressure piston-pump-adjusted concentration of AAs in the water phase to 0 wt%, 1 wt%, 2 wt%, and 3 wt%. The preliminary Chen-Guo model (Chen and Guo 1996, 1998) determined the curve of hydrate formation (Fig. 5) for the defined natural-gas composition. Test Protocol. Cooling experiments occurred under constant pressure (4 MPa) and initial linear flow velocity (1.25 m/s) in

208


Stage 1

3.10

Stage 2

5.92 6.67 Stage 3

12

2200

Stage 1

Total Pages: 9

3.10

Stage 2

2000

18 10

12

6 hydrate formation equilibrium temperature at 4 MPa

10

4

1800

8

1600 1400 1200 1000

Stage 4

Hydrate volume fraction

1.5

800

2 0

1

2

4 5 3 Relative time (h)

6

50

40 1.0

35 30 25

0.5

2.4

20

1.8

1.2

0.6

15 0.0

7

3.0

45

600

6

60 55

Velocity

Velocity (m/s)

8

Temperature Flow rate Pressure drop

14

5.92 6.67 Stage 3

Viscosity

Flow rate (kg/h)

16

Pressure drop (kPa)

Temperature (°C)

1.89

2.0

Stage 4

Page: 209

Viscosity (cP)

1.89

20

Stage:

Hydrate volume fraction (%)


0

1

2

3 4 5 Relative time (h)

6

10 7

0.0

Fig. 6—Temperature, flow rate, pressure drop as functions of time during natural hydrate formation (15 vol% water, 1.25-m/s flow velocity, 4-MPa pressure, 1 wt% AA).

Fig. 7—Viscosity, velocity, and hydrate fraction as functions of time during natural hydrate formation (15 vol% water, 1.25-m/s flow velocity, 4-MPa pressure, 1 wt% AA).

accordance with actual operating conditions as much as possible. The test protocol is briefly described as follows: The loop is loaded with diesel and water (100 vol% liquid loading) by considering the defined water-volume fraction (water cut) for each test. The gas-supply unit begins to inject gas into the separator at the room temperature (18 C) until achieving the aimed experimental pressure (4 MPa in this paper). The water/oil mixture is circulated at a constant flow rate to form a homogeneous and stable emulsion with the set AA dosage for each test. Here, the stable emulsion means the size of the droplets detected by FBRM fluctuates little in the system (the boundaries of FBRM fluctuations are 60.2 lm). Under the maintained pressure and initial flow velocity, the temperature decreases from 18 C to 4 C. The gas consumption is compensated by gas injection from gas tanks into the separator. During the hydrate-formation process, the data-acquisition system continuously collects data about the temperature, pressure, pressure drop, flow rate, density, and the chord length. At the end of the formation process, when all measured data are stable, the feasible conditions for hydrate-slurry flow are explored by increasing or decreasing the flow rate step by step until blockage. The dissociation of hydrate is conducted by increasing the temperature to 30 C. The system is kept at these conditions for 12 hours. A high flow rate is then applied to create a fine emulsion again.

Results and Discussion The Influence of Flow Rate on the Hydrate-Slurry Flow. In this natural gas þ diesel þ water system, the hydrate-slurry formation-process schematic is shown as Fig. 2. It can be found from Figs. 6 through 8 that the experimental process went through four stages (Table 3): the stable stage (Stage 1), the rapidhydrate-formation stage (Stage 2), the metastable stage (Stage 3), and the contrived-slow-down stage (Stage 4). Stage 1 comprises the time periods before the hydrate formation (0 to 1.30 hours) as well as the hydrates beginning to form on the surfaces of the droplets in the water/oil emulsion (1.30 to 1.89 hours), which is evident from the hydrate-volume fraction points in Fig. 7. We used the gas-consumption data during the experiments to estimate the hydrate-volume fraction in the system. It can be found from Fig. 8 that when the hydrates began to form (1.30 to 1.89 hours), the mean chord length of droplets/particles in the system is broadly unchanged. This could provide some indication of the hydrate formation along the water/oil interface. As shown in Figs. 6 through 8, the flow rate, pressure drop, viscosity, and the water-droplet mean chord length in the fluid leveled off in Stage 1. Stage 2 begins when a considerable amount of hydrates formed and agglomerated in a relative rapid way (compared with hydrate formation in Stage 1). In Stage 2, the viscosity, the size of particles/droplets, and the amount of formed hydrates increased notably, and the flow rate plummeted in the system. Then the flow rate entered the metastable stage (Stage 3). During this stage, the hydrate kept forming, as the points of the hydrate-volume fraction indicate in Fig. 7. Both the viscosity and the frictionalpressure-drop curves increased first and then decreased, whereas the flow rate experienced the opposite trace (decreased first and then increased) and finally approached a steady value. When the flow rate at the metastable stage tended to a certain value and lasted for a while, we began to slow down the pump speed; thus, the process entered the contrived-slow-down stage. The target of this stage is to investigate the feasible conditions for hydrate-slurry flow. It can be noted from Figs. 6 and 7 that the

13 Mean Chord Length (microns)

15% water-cut

12 11 10 9

TABLE 3—TIME BOUNDARIES OF THE FOUR STAGES DURING THE EXPERIMENT

8 7 0

1

2

3 4 5 Relative Time (h)

6

7

Fig. 8—Chord length as a function of time during natural hydrate formation (15 vol% water, 1.25-m/s flow velocity, 4-MPa pressure, 1 wt% AA).

Stage Definition

Time Boundary (hours)

1. The stable stage 2. The rapid-hydrate-formation stage 3. The metastable stage 4. Contrived-slow-down stage

0–1.89 1.89–3.10 3.10–5.92 5.92–6.67

April 2014 SPE Journal

209 ID: jaganm Time: 14:55 I Path: S:/3B2/J###/Vol00000/130056/APPFile/SA-J###130056


4 0 wt% AAs 1 wt% AAs 2 wt% AAs 3 wt% AAs

Induction Time (h)

3

2

1

0 15

25 20 Water Cut (vol%)

30

Fig. 9—The induction period under various water cuts and AA dosage (1.25-m/s flow velocity, 4-MPa pressure).

flow rate reduced with the pump speed progressively, while the pressure drop and the viscosity of hydrate slurry presented as a Vshaped curve (i.e., initially went downhill then uphill). By analyzing the four stages, these determinations were concluded: At the rapid-hydrate-formation stage, the dramatic change/ decline of flow rate stemmed from the collision and coalescence of the hydrate particles formed in considerable amounts, which also increases the viscosity and frictional pressure drop of the hydrate slurry. Then, the flow rate climbed because the formed larger-size aggregates were crushed into smaller pieces (as the CLD shows in Fig. 8). The phenomenon that saw larger-sized aggregates broken into smaller ones might mainly result from the following two reasons. First, the viscous force between hydrate particles decreased along with the continuous hydrate generation and accumulation. (The hydrate formation might consume the water on the particle’s surface, change the wettability of the surface, and then reduce the interparticle viscous force.) Second, the hydrate aggregates in the fluid were constantly under shear stress from both the pipeline flowing and the pump. (The crushed aggregates were the cumulative results of these two shear forces.) The experimental data shown in Fig. 8 could support the two preceding reasons to some extent. The mean chord length of the particles/droplets was approximately 9 lm at Stage 1, and it increased because of the collision and the coalescence of the hydrate particles formed in considerable amounts in Stage 2. These larger aggregates are certainly under constant shear stress from both the pipeline flowing and the pump. Instead of an immediate drop, the mean-chord-length curve of the aggregates oscillated basically in a certain range and lasted for approximately 1 hour. Then, the aggregates were crushed into smaller ones suddenly, leading to a declined mean-chord-length curve. In other words, if the shear stress from the pump dominated the particle breakage in the fluid, the aggregates would have been crushed right after the slurry passed the pump. The 1-hour duration of the aggregates from the experiments contradicts this. However, the physical interaction of the magnetic pump with the hydrate slurry still cannot be quantitatively characterized because of its complication at present experimental conditions. That is, we did not figure out the weighting coefficients for different influencing factors of the hydrate particles and/or aggregates; thus the breakdown phenomenon was considered as a result of the combined effect of the reduced viscous force and the shear forces of both pipe flowing and the pump in this work. At the metastable stage, the slurry-flow rate fluctuated slightly, but the pressure drop and viscosity varied notably. The hydrate slurry showed the shear-thinning behavior of a non-Newtonian fluid because the flow rate and the fluid’s viscosity present opposite trends (Figs. 6 and 7).

Stage:

Page: 210

Total Pages: 9

At the contrived-slow-down stage, the pressure drop initially decreased then increased with the slowing flow rate. The main reason is that the pressure varied with the pump speed at first, but the shear stress in the loop reduced with the flow rate, too. In other words, the weak shear stress cannot inhibit the agglomeration between hydrate particles, thus increasing the slurry viscosity and pressure drop. Moreover, the small flow rate intensified the heat exchange between the temperature-control jacket and the loop, leading to the decline of the fluid temperature in the loop (Fig. 6). Therefore, the lower temperature benefitted the hydrate formation by reinforcing the driving force. In summary, the combined effect of slowdown as well as hydrate formation increased the risk of hydrate blockage in the pipeline. Consequently, a critical minimal flow rate ensuring safe hydrate-slurry flow is defined here as a hypothesis: When the actual flow rate exceeds the minimum, the safe slurry flow could be ensured; on the other side, there might be a hydrate-blockage accident. This phenomenon also can be explained by the formation of a hydrate bed (Hernandez 2006). When the flow rate is too low to hold the hydrate particles in the slurry, the particles would deposit to form a hydrate bed and then a hydrate plug. The critical minimal flow rate is thus consistent with the hydrate-bed explanation in a sense. The Influence of Water Cut and AA Dosage on the HydrateInduction Period. In most research, the studies of the hydrateinduction period typically performed with autoclave cells are either under static or stirring conditions (Bansal 1994; Nerheim et al. 1994; Cingotti et al. 1999; Kelland et al. 2000); relatively speaking, there are fewer research projects carried out in axialflow systems. Turner (2005) and Li (2012) observed the induction time under different experimental conditions in flow loops, respectively; Sun (2001) fitted an empirical correlation of the induction time for the methane hydrate in an axial flowing system, and he found that the hydrate induction time in a flowing system was related not only to pressure, but also to the flow velocity of the fluid. The definition of hydrate-induction time was also diverse from every researcher (Kashchiev 1991 and 2000). Therefore, this paper has defined a clear, easy, and general method to measure the hydrate-induction period on the basis of slurry experiments. The essence of this method is to determine a fixed starting point as well as an easily observable endpoint, which are typical and universal for a certain system. Accordingly, the hydrate-induction period in this paper starts when the temperature reduced to the thermodynamic equilibrium curve of hydrate formation, and ends at the time point when the temperature began to rise since the rapid hydrate formation. Thus, we have studied the influence of factors such as water cut and AA dosage on the hydrate-induction period, as shown in Fig. 9. It is evident from Fig. 9 that the experimental result in the loop of the induction time at 0 wt% AA presented a disparate tendency compared with the ones containing AAs. The 0 wt% AA experiment here is used for reference, that is, as a contrast (or “blank”) experiment to show induction periods without any AA in the system. Its disparity probably came from the difficulty for the diesel and water to form a stable emulsion without the help of the AA. We had tested the stability of the oil/water emulsions with a stirrer with a relative high shearing rate before the loop experiments, and found that this “blank” mixture would broadly stay in a layering state even during the stirring, and would revert to a clear layering state within 3 minutes after the external stirring. On the contrary, the mixtures with AAs could form emulsions as soon as the stirring began; the 1 wt% AAs emulsion therein could last for at least 30 minutes after the external stirring. The experimental result without AAs was indeed too erratic and counterintuitive for analysis; consequently, only the results with AAs are discussed in the following. The induction period was compacted at first then prolonged slowly with the increase of water cut under the same initial flow velocity (1.25 m/s) and system pressure (4 MPa). The induction

210



Stage:

Mean Chord Length (microns)

before hydrate nucleation during hydrate nucleation after hydrate nucleation

3 Probability (%)

Total Pages: 9

12.0

4

2

1

11.5 11.0 10.5

1

10 Chord Length (microns)

9.5 9.0 8.5

100

Fig. 10—CLD (25 vol% water, 1.25-m/s flow velocity, 4-MPa pressure, 1 wt% AA).

period with a higher water cut (20 to 30 vol%) was significantly less than that with a 15 vol% water cut. It is known that the mass transfer is a function of the interfacial area and the driving force. And the increased water cut could lead to either a larger contact area between water oil phases or a decreased gas solubility per unit volume of oil/water emulsion. In detail, the more interfacial area or nucleation sites facilitated the hydrate formation on the water-droplet surface, which could shorten the hydrate-induction period; whereas the less gas dissolved in the unit liquid confines the mass transfer. In general, the water cut influences the hydrateinduction time in both two ways. Nonetheless, along with the further larger water cut, the shear stress that flow rate and the pump exerted on the dispersed water phase got stronger, thereby hindering the critical nucleation on the water surface, then prolonging the induction period. As a result, the induction period exhibited a slowly rising trend. It also can be seen from Fig. 9 that the hydrate-induction period ascended with the AAs dosage. The probable reason is that the AAs molecule adsorbed on the water surface could inhibit the hydrate formation. In addition, the randomness of the induction period and the memory effect of hydrate particles might also affect the preceding phenomena. The Influence of Water Cut and AA Dosage on CLD. According to Turner (2005) and Taylor et al. (2007), the CLD of hydrate particles resembles that of dispersed droplets. The water droplets converted into hydrate particles directly without changing their distribution at the hydrate formation. There are many influ12.0 1 %Wt AAs 2 %Wt AAs 3 %Wt AAs

11.5

11.0

10.5

10.0

0

2

4


12

14

Fig. 12—The influence of AA doses on CLD (25 vol% water, 1.25-m/s flow velocity, 4-MPa pressure).

15 vol% water-cut 20 vol% water-cut 25 vol% water-cut

10.0

8.0

0

Mean Chord Length (microns)

Page: 211

0

2

4


12

14

Fig. 11—The influence of water cut on CLD (1.25-m/s flow velocity, 4-MPa pressure,1 wt% AA).

encing factors of CLD in the water/oil emulsion—for example, the fluid viscosity, pressure, flow rate, water cut, and AA dosage. This paper focuses on the water cut as well as AA dosage to investigate the CLD of water droplets, and thus the CLD of hydrate particles. The results are shown in Figs. 10 through 12. Fig. 10 demonstrates that the size of hydrate particles or water droplets changed little before and during the hydrate nucleation process, which is similar to the conclusions made by Turner (2005) and Taylor et al. (2007). Then, the abundance of hydrate particles gave rise to collision, coalescence, and the larger-size hydrate aggregate, as shown by the blue-triangle points in Fig. 10; the CLD moved to the larger particle direction. From Figs. 11 and 12, it can be seen that the size of hydrate particles or water droplets increased with increasing water cut but decreased with increasing AA concentration. This resembles the research findings of Lachance et al. (2012). Therefore, both water cut and AA doses indeed influence the CLD of the hydrate particles or water droplets. Gainville and Sinquin (2011) have chosen the mean diameter as 5 lm in their viscosity model of hydrate slurry; however, it is shown in our study that the hydrate mean diameter varies with different water cut and AA dosage. In summary, the appropriate mean diameter under diverse conditions is significant to promote the accuracy of related predicting models. Thus, in determining how to correlate the mean diameter with the viscosity model, the slurry pressure-drop calculation can be considered in further researches, but is outside the scope of this paper. Quasisingle-Phase Laminar-Flow Pressure-Drop Model of Hydrate Slurry. Andersson and Gudmundsson (1999 and 2000) and Peysson et al. (2003) studied hydrate-slurry rheology with the light-hydrocarbon/brine/oil system, in an experimental loop and an autoclave cell, respectively. They concluded that the slurry showed the characteristics of non-Newtonian fluid even under a very low hydrate concentration (808 >808 343.8 672.8 616.4 752.6 435.2 415.4 369.6

laminar flow laminar flow laminar flow turbulent flow turbulent flow laminar flow laminar flow laminar flow laminar flow laminar flow laminar flow laminar flow

20

25

30

where k is the consistency coefficient (Pasn) and n is the dimensionless index of the flow behavior. The shearing rate c_ is a function of the shearing stress s; then, the velocity distribution is ðr u ¼ f ðsÞdr þ C . . . . . . . . . . . . . . . . . . . . . . . . . ð2Þ The combining of Eq. 2 with the boundary conditions then gives the flow-rate expression of the power-law fluid flowing through the pipe: 1 1þ3n DP n n . . . . . . . . . . . . . . . . . . . . ð3Þ R n Q¼p 2kL 1 þ 3n Moreover, the following equations can be established for the power-law fluid: n ¼ n0 ¼

dlnsw . . . . . . . . . . . . . . . . . . . . . . . . . ð4Þ 8uav dln D

DPD . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . ð5Þ 4L 4Lk 1 þ 3n n 8uav n . . . . . . . . . . . . . . . . . . ð6Þ DP ¼ D 4n D sw ¼

where u is the velocity, m/s; Q is the volume flow rate, m3/s; R is the pipe radius, m; L is the pipe length, m; DP is the frictional

Stage:

Page: 212

Total Pages: 9

pressure drop, Pa; D is the pipe diameter, m; sw is the shearing stress at the pipe wall, N/m2; and the uav is the average velocity, m/s. And the dimensionless rheological index n0 was fitted by 90 sets of experimental data obtained at the end of the metastable stage when the hydrate particles’ chord length kept stable in the slurry. The correlation is expressed as n0 ¼ 11:557ax þ 0:26895x 0:49838a þ 1:08906 ð7Þ where a is the water’s volume fraction and x is the mass fraction of AA. With regard to the categorization of the flow regime of the non-Newtonian power-law fluids, a critical stability parameter, Z, was adopted to distinguish the laminar flow (Z 808) (Ryan and Johnson 1959). Z can be calculated by Z¼

nþ2 nþ1 qu2av 3n þ 1 2 1 . . . . . . . . . . . . . . . ð8Þ n DPR n nþ2 2L

Thus, the flow behaviors of hydrate slurry at various water cut and AA dosage in this work are shown in Table 4. Subsequently, the comparison between experimental data and predictive values calculated by the laminar-flow pressure-drop model is shown in Table 5. It can be concluded from Table 5 that the results calculated by the pressure-drop model agree well with the experimental results in this paper, with the maximal and the minimal relative deviation being 4.6% (