data reconciliation and gross error detection in

Proceedings of the International Conference on Emerging Trends in Mechanical Engineering (ICETiME’15) December 16-17, 2015, CUSAT, Kochi, India

DATA RECONCILIATION AND GROSS ERROR DETECTION IN THERMAL POWER PLANTS Sushil Cherian1, Jino Pattery2, Dr. James Varghese3 1,2

Kalkitech, Bangalore, India Asst. Professor, Mechanical Engineering Division Cochin University of Science and Technology Kochi- 682 022, Kerala, India.

3

ABSTRACT Process measurements are prone to errors which can lead to significant deterioration in thermal power plant performance. An effective sensor validation method that detects and identifies faulty sensors is required before doing plant performance analysis and optimization. Gross errors in measurements can be caused by biases in sensors or gross errors in steady state conservation constraints due to unknown leaks. Data reconciliation problem is formulated as weighted least squares estimation under non linear constraints, and test statistics are derived using constraint residuals. Since the measurement variables are all related through the constraints, a gross error in one measurement may cause the test statistics of good measurements to exceed the test criterion which depends on many factors such as redundancy level, differences in standard deviations, and gross error magnitude.

INTRODUCTION Accuracy of operational data of a power plant is essential for power plant performance monitoring, optimization and fault diagnosis. However, due to inevitable occurrence of systematic and random errors in measurements, errors are propagated to the results of performance calculations. This is turn reduces the effectiveness of optimization functions, and fault diagnostics. In this paper, we propose a data reconciliation approach to detect and reduce errors in the measured data and thus enhance the measurement accuracy. The reconciled data can then be used in performance monitoring, optimization and fault diagnosis systems to 1

Corresponding author. E-mail: [email protected]

Data reconciliation and gross error detection

improve their accuracy. Data reconciliation (DR) improves the accuracy of measurements by reducing the effect of random errors in the data, by explicitly making use of process model constraints. It obtains estimates of process variables by adjusting process parameters so that the estimates satisfy the constraints. Gross error detection is a companion technique that can identify and eliminate gross errors in the measurements and process models. These two techniques are applied together to improve accuracy of measured data, exploiting spatial and temporal redundancy properties of measurements [1]. Classical DR methods evolved in the chemical engineering industry. There are certain differences in processes in thermal power plants that introduce additional complexity compared to chemical processes – primarily related to significant pressure variations and phase changes of the operating fluid at various points in the thermodynamic cycle. Previous work has focused on reconciliation of mass flow measurements [2] [3] [4] in thermal power plants, considering operating pressure to be constant. This paper extends previous work to include reconciliation of pressure and temperature measurements also. The uncertainty of key parameters in the Water/Steam system of the boilers, steam turbine stages including the extractions, condenser, feed water heater, can be reduced. A simple case study is presented to illustrate the concepts of gross error detection and data reconciliation. The study is limited to steady state operating conditions. Further the constraints are modeled as linear equations.

MODEL ERROR MODEL The measured data are subject to unavoidable measurement errors. It can be modeled as follows: y = x+ e

...(1)

where y is a vector of n measured values x is the corresponding vector of true values e is a vector of unknown random error in the measurement with mean zero They always comprise some part that is of random character and can thus be treated by statistical methods. Here, we shall restrict our attention mainly to the random errors that oscillate around zero and are characterized by standard deviation σ. As a motivation, one can perhaps consider the fact that with probability 0.95, the absolute value of the error will not exceed the value (approximately) 2σ, for Gaussian distribution of the errors. On the other hand, also in practice an error whose 2

Proceedings of ICETiME’15 December 16-17, 2015, CUSAT, Kochi, India.

probability is less than 0.05 is regarded as a gross error due to instrument miscalibration, measurement device failure, nonrandom events affecting process, such as process leaks. SYSTEM MODEL The models employed in data reconciliation represent variable relationships of the physical system of the process. The reconciled data takes information from both the measurements and the models. In reconciling steady-state measurements, the model constraints are algebraic equations. On the other hand, when dealing with dynamic processes, dynamic models that are differential equations have to be used. The constraints matrix for mass flow reconciliation of a typical plant is given Fig. 1. The constraints can be represented in general by Ax = c

...(2)

where A is a matrix of dimension m x n c is a constant m x 1 vector

Fig 1: Simple power plant cycle, and corresponding model equations

PROBLEM FORMULATION Data reconciliation and gross error detection problems have the following four components:  Detection problem: Ability to detect the presence of one or more gross  errors in the data  Identification problem: Ability to identify the type and location of the gross error 3


   

Multiple gross error identification problem: Ability to locate and identify multiple gross errors which may be present simultaneously Estimation problem: Ability to estimate the magnitude of the gross errors

Detection utilizes the fact that gross errors cause violation of model constraints. An allowance has to be made for violation of constraints due to random errors. Under assumed probability distribution for the random errors, a probabilistic approach is used for gross error detection. Normalized errors fall inside a (1-alpha) confidence interval at a chosen level of significance alpha. Any value of normalized error which falls outside the confidence region is declared a gross error. Global Test (GT) is used for detecting whether one or more gross errors are present. Generalized Likelihood Ratio (GLR) test is used for identifying the type and location of gross error. For multiple gross errors, GLR test with serial elimination strategy is used [1]. With the assumption of normally distributed measurements, a weighted least-squares objective function is conventionally formulated for the data reconciliation problem. At process steady state, the reconciled data are obtained in general by minimizing Min (y-x)T Σ-1 (y-x) x

...(3)

subject to Eq (2), where Σ is an n x n covariance matrix of the measurements.

CASE STUDY In this case study the water/steam cycle of a large conventional electricity power station with a design power of 600 MW is considered. CycleTempo software is used to simulate the plant behavior. In the water/steam cycle shown in Fig. 2, steam of 180 bars and 530 °C expands in the high pressure turbine 2 to a pressure of around 38 bars. The steam is reheated (4) to 530 °C and expanded further in two sections of the intermediate-pressure turbine and in six sections of the low-pressure turbine. The intermediate- and low-pressure turbines are represented by one turbine (5). The steam then condenses in condenser 6 at a pressure of 0.027 bar. The condensate is preheated in eight preheaters: five low-pressure and three high pressure preheaters. The last low-pressure preheater (13) is a deaerator, in fact a mixing preheater. The other low-pressure preheaters (8, 10, 11 and 15) are surface preheaters. The surface preheaters, with the exception of the first preheater, are equipped with a subcooler. The condensate of these four preheaters is collected, via

4


Fig 2: Layout of power plant under study any preceding preheaters, in the first preheater (with the lowest extraction pressure). From here the auxiliary condensate pump 19/8 pumps this condensate to the main condensate line. The three high-pressure preheaters (15, 16 and 20) are surface preheaters, each equipped with a desuperheater and a subcooler. The condensate from this is fed into the deaerator, via any preceding preheaters, with the main condensate stream. Feed pump 14 pumps the feedwater of the deaerator to boiler 1. The feed pump is driven by auxiliary turbine 18. This turbine, like the deaerator, is supplied with extraction steam from the outlet of the intermediate-pressure turbine. 5


In the auxiliary turbine the steam is expanded to a pressure of 0.027 bar . This model has 25 equipments connected by 40 pipes, which obeys 29 mass flow equations and 11 energy equations. Random errors which are less than 2 % are present in all the measurements. HEAT RATE CASE STUDY The impact of gross error in heat rate is studied by simulating a gross error of 5.09 % in main steam temperature. All other measurements are having random errors. It can be observed that the data reconciliation and gross error detection algorithm has successfully detected, identified and substituted a more accurate value for the main steam temperature. The error in turbine heat rate of 3% has been reduced to 0.04%. Any heat rate optimization functions running without data reconciliation in place is likely to be ineffective. This approach can be extended to other critical operational measurement in order to obtain more accurate reconciled data for online power plant performance monitoring, optimization and fault diagnosis. Table 1: Heat rate accuracy improvement with DR Pipe No

Mass flow (kg/s) Field

Reconciled

Temperature (C) Field

Pressure (bar)

Reconciled

Field

Reconciled

1

396.66

400.37

503.0

530.6

139.6

141.2

3

327.84

330.74

305.9

308.0

29.5

29.7

4

333.46

330.74

527.4

531.0

29.3

29.7

35

396.26

400.37

268.8

273.2

186.1

189.3

heat rate Field

Reconciled

7666.32

7904.21

CONCLUSIONS This paper explores the practical aspects of linear steady state data reconciliation problem for a typical thermal power plant. The standard deviation of measurement error plays a key role in data reconciliation and gross error detection techniques and its estimation is challenging as the true standard deviation is not known. In addition to mass flows, pressure and temperature measurements are reconciled by formulating pressure and temperature drop relations as linear constraints. A case study shows improvement that can be achieved in power plant performance calculations. An essential prerequisite for gross error detection is redundancy in measurements as gross error detection is possible only in redundant measurements. 6


Methods based on matrix approach are used for observability and redundancy classification of process variables. Further studies involving non-linear constraint equations can potentially improve the accuracy.

REFERENCES [1] Shankar Narasimhan, Cornelius Jordache (2000), “Data Reconciliation and Gross Error Detection. An Intelligent Use of Process Data”, Gulf Publishing Company, Houston, TX [2] Frantisek Madron, et al, (2007), “Process data validation in practice”, ChemPlant Technology s.r.o., Report CPT-229-07, (www.chemplant.cz) [3] Xialong Jiang, et al: “Data reconciliation and gross error detection for operational data in power plants”, Energy 75 (2014) 14 - 23 [4] Xialong Jiang, et al: “Gross error isolability for operational data in power plants”, Energy 74 (2014) 918 - 927

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Proceedings of the International Conference on Emerging Trends in Mechanical Engineering (ICETiME’15) December 16-17, 2015, CUSAT, Kochi, India, 9- 30

FLOW STRUCTURE ANALYSIS AROUND A SQUARE CYLINDER WITH AN OFF-CENTRE SPLITTER PLATE Vivek Harris1, Dr. R. Ajithkumar2, P.J. Joshy3 1

Student, 3Associate Professor Cochin University of Science and Technology Kochi- 682 022, Kerala, India. 2

Professor, Amrita University, Amritapuri Campus, Kollam, Kerala 690525

ABSTRACT In this work, flow around a stationary square cylinder of edge length 30 cm with splitter plates of six different lengths placed behind the cylinder at three different off-center positions with respect to the square cylinder centerline and at six different gaps is visualized. Flow visualization is done in a water channel, the flow pattern is video graphed and a frame-by-frame analysis of the flow pattern is carried out. Reynolds number (Re) of 1800 is maintained throughout. Various new patterns in the vortex shedding mechanisms are observed. Variations in the values of St and Cd are analyzed here. CFD analysis of the square cylinder with splitter plates is also done in ANSYS Fluent 14.5; the vorticity profile, value of Cd and St are obtained and validated with the experimental results. One of the significant findings of this study is the identification of a critical gap (gcr) of 1B, where B is the edge length of the square cylinder. A sudden change in the flow regime is observed after the Critical gap. The vortex shedding mechanism and the values of St and Cd exhibited after the critical gap is almost identical to that of a lone square cylinder and hence placing of splitter plates above gcr does not affect the flow behind a square cylinder.

1


Flow Structure Analysis around a Square Cylinder

INTRODUCTION Flow over bluff bodies has been an area of extensive research in fluid engineering, especially due to its large number of applications. Despite the numerous studies on the subject, both numerical and experimental, the subject still remains a challenge. Reduction of drag and prevention of structural damage to the body are the chief motivating factors for the studies. Unsteady wakes behind a bluff body can bring to bear great unsteady forces on the body and compromise its structural integrity. Thus, flow analysis over bluff bodies is of great significance to designers of bridges, chimneys and large building structures. At low Reynolds numbers (Re), the flow over a cylinder is smooth and it divides and reunites around the cylinder. At Re~30, the flow separates downstream and two symmetric eddies are formed. These grow in size and remain symmetrical up to Re~40. When Re approaches 47, the wake begins to shed vortices into the stream. When one vortex is shed, another one forms in its place causing an asymmetrical flow pattern around the body with consequent pressure distribution changes. This induces a lateral force on the body. If alternate shedding of vortices takes place, the body begins to vibrate and if the frequency is sympathetic to the natural frequency of the body, the body will resonate. The shedding frequency continues to increase with increase in Re and at a value of 90 the shedding pattern appears similar to footprints on a street and thus the term von Karman vortex street. The von Karman instability for a single bluff body is characterized by the vortex shedding frequency (f). This frequency can be non∗ dimensionalised to a number called the Strouhal number = , where d is the cylinder diameter and U the freestream velocity. Many studies have been conducted to control or inhibit the von Karman instability and the methods adopted can be classified as passive and active. Passive methods include placing a splitter plate in the near wake of a cylinder. VORTEX SHEDDING MECHANISMS FOR SQUARE CYLINDER Vortex shedding is an oscillating flow that occurs when a fluid flows past a bluff body at a certain velocity. Vortices are usually formed behind the body and shed periodically from both sides of the bluff body. The vortices are formed due to the inverse pressure gradient formed behind the bluff body. The low pressure vortices tend to move the bluffbody towards the low pressure region. The three prominent vortex shedding mechanisms for a square cylinder are:   

Gerrard Mechanism (GM) Passive Pushing (PP) Reattachment (RA)

A large number of flow structure studies have been conducted on various bluff bodies like circular cylinders, semicircular cylinders, cubes, elliptical bodies, square 10


cylinders, etc. It is seen that flow structure analysis of square cylinders is very limited, especially experimental studies. However, a few numerical studies have been done in the case of square cylinders. The experimental work by Roshko, A (1954) is one of the earliest to have found the effectiveness of a splitter plate as a bluff body wake control. Here he used various splitter plates of lengths varying from 1 to 5 times the diameter (D) of the cylinder. With the splitter plate of dimension 1D, he experienced no change in the vortex formation mechanisms, but observed a slight variation in the vortex shedding frequency. Upon moving the splitter plate slightly downstream, leaving a gap between the cylinder and splitter plate, he observed a decrease in vortex shedding frequency and an increase in the base pressure. He slowly increased the length of the splitter plate and observed the variations and noted that when the splitter plate dimension was changed to 3.85D, the vortex shedding frequency reaches its minimum value and the base pressure its maximum value. He was able to give a direction in the use of gaps in the application of splitter plates. He came to the conclusion that at some finite gap width, the use of a splitter plate would be effective. Bearman, P. W. (1965) conducted tests on a semi elliptical body (Re=1.5 X 105). He concluded that the vortex shedding disappears for a connected splitter plate when the length of the splitter plate approximates to twice the diameter (i.e., (Lsp=2D). He also found that for long splitter plates, St decreases with an increase in splitter plate length for a circular cylinder and a half elliptical cylinder. Apelt, C. J. et al. (1973) carried out experiments with two combinations, a circular cylinder connected with a splitter plate and a normal flat plate connected with a splitter plate to show the influence of the splitter plate length on the pressure drag and vortex shedding. Their conclusion was that short length splitter plates can impart significant change in the downstream flow from the circular cylinder and the normal flat plate. Sohankar et al. (1998) found that the flow separation points for a square cylinder are always located at the sharp corners. Flow interaction studies of a square cylinder attached with a splitter plate have not been done extensively. One of the recent numerical studies was done by Doolan, C. J., (2009) on a square cylinder with a detached splitter plate downstream of the square cylinder. This study came to the conclusion that the splitter plate contributed to a significant reduction in St and also towards the reduction of the total drag coefficient of the square cylinder. Ali et al. (2012) had numerically studied the flow over a square cylinder with a detached flat plate at low Re. Here they varied the gap between the splitter plate and the square cylinder and observed a critical gap (2.3D) which separates two flow regimes. All the above studies come to the conclusion that a splitter plate downstream of a bluff body could be used as an effective method of controlling a fully developed vortex shedding. It is seen that in all the above studies, the splitter plates were placed in a central position downstream of the bluff body. No study with splitter plates placed in an off-centered position could be identified. Hence, an opportunity presented to carry out flow study with splitter plates behind the square cylinder and 11


thus, in this study, an analysis of the flow around a square cylinder with an offcenter splitter plate is carried out.

EXPERIMENTAL INVESTIGATIONS PROBLEM DESCRIPTION Flow around a square cylinder with an off-center splitter plate downstream as shown in Fig.4, placed in a water channel, is to be analyzed. A 30 x 30-mm square cylinder, and a 0.8-mm thick splitter plate is used in the analysis. Throughout the experiment, the square cylinder is stationary; only the splitter plate length and position are varied. A constant Re of 1800 is maintained throughout the study and the effect of L/B, g/B and h/B on the near wake flow structures of a square cylinder is deduced, where:    

L = Length of Splitter Plate g = gap between the end of the square cylinder and beginning of splitter plate h = offset of the splitter plate with respect to the square cylinder edge B = Square cylinder size B

h

U∞

B

SQUARE CYLINDER

L PLATE

g

Fig.1 Square cylinder and splitter plate arrangement A total of 109 cases are studied, of which 108 were with splitter plates and one without splitter plate. The experimental analysis of the square cylinder without splitter plate is done for the validation of the results (i.e., the value of Cd and St) with the available standard results. The various positions of splitter plates are shown in the table. Table 1 Splitter plate positions at which analysis done Sl. No. 1. 2. 3. 4. 5. 6.

L/B 0.5 1 1.5 2 3 4

g/B 0 0.25 0.5 0.75 1 1.5 12

h/B 0.15 0.3 0.483


METHODS METHODOLOGY OF ANALYSIS In this work, both experimental and numerical analyses are carried out. Numerical analysis is done using ANSYS Fluent 14.5. Experimental and numerical values of Cd are obtained and compared with the standard available results. The variations in the vortex shedding with the introduction of splitter plates are mainly studied experimentally. The vortex shedding frequency and Strouhal number are calculated and compared with the standard results. CFD analysis is done for a limited number of cases only, due time consumption for the 109 experimental cases. EXPERIMENTAL SETUP The square cylinder and splitter plate are placed in a circulating water channel and the flow pattern around the square cylinder and splitter plate at various speeds of flow is visualised. Very fine aluminum powder (80 μm) is mixed in the water used and serves as tracer in the flow to aid visualization of the flow regime. Four halogen lamps, one in each corner of the water channel, are used to illuminate the tracer particles. The visualised flow patterns are captured for analysis using a high definition video camera. FLOW VISUALISATION EQUIPMENT The flow visualisation of the square cylinder and splitter plates is carried out in a setup consisting of a recirculating water channel 2.5 x 1.5 x 0.15-m, (length x breadth x depth), Fig.2. At one end of the water channel is a pair of paddle wheels (aluminium discs with vanes on one side) driven using a system of bevel gears by a variable speed electric motor. The paddle wheels cause the water to flow through the side channels and the two streams meet at the nether end and flow through the centre channel, which is the test section of the channel. A mesh is placed at the beginning of the test channel to achieve a smooth and steady flow. The AC motor is driven via a variable frequency drive (VFD).

13


Fig.2 Top view of water channel with test model - schematic TEST PIECE- SQUARE CYLINDER WITH SPLITTER PLATE A 30 x 30-mm square cylinder of 170-mm height was chosen. The 150-mm water depth in the test channel results in a 20-mm projection of the cylinder above water. The square cylinder is mounted on a machined aluminium baseplate and the splitter plates secured in position with screws on the baseplate. The plate mounting holes are grooved to facilitate positioning in any offset position. Six splitter plates of length 0.5B, 1B, 1.5B, 2B, 3B, and 4B, where B is the square cylinder size (30 mm), were used.

Fig.3 Test Piece

14


EXPERIMENTAL INVESTIGATIONS ESTIMATION OF WAKE PARAMETERS For each case, the moment at which a fully developed vortex forms is noted and a snapshot of the video at this particular moment is taken. The parameters of the vortex formation are measured using Microsoft Paint. The vortex formation length lf, Fig.4, which is the stream-wise distance from the axis of the square cylinder to the core of the fully developed vortex, is to be measured. Towards this, the vortex formation length in pixels (l) is measured. The value of l is non-dimensionalised by dividing it with the edge length of the square cylinder in pixels (w); Fig.5. This non-dimensionalised value is the vortex formation length lf. Now from lf, the base pressure coefficient, Cpb is found out. C is an empirical constant whose value varies from 1.6 to 1.8 for Re between 1000 and 9000. From Cpb, the value of K, the base pressure parameter, is obtained.

Fig.4 Vortex Formation Length

Fig.5 Edge Length of Square cylinder

CALCULATION OF Cd AND St St is found out by analysing the video. From the video, the time taken for 10 vortices to develop fully and shed is noted. From this, the time required for 1 vortex to form and shed is calculated. This exercise is carried out 8 times and the average of these times is calculated. This time, Tavg, is the vortex shedding time. From this, the vortex shedding frequency f is determined and the Strouhal number Re is calculated. From the value of Re, the Cd is obtained using Eq. (4). lf =

(1)

lf =

(2)

K=√(1 −

)

(3)

Cd*St = 0.4777 K- 0.45167

15

(4)


ANALYSIS AND DISCUSSIONS RESULTS In this section the results for the square cylinder with splitter plates are presented and discussed. The flow visualisation photographs have been given with the corresponding positions of the splitter plates (108 cases) and for the lone square cylinder case. In all photographs the flow takes place from left to right. The variations in the values of Cd and St, the vortex shedding mechanisms, and the effect of each splitter plate on the vortex shedding mechanism are discussed here. The results from the numerical analysis done in ANSYS Fluent are also given and compared with the experimental results. LONE SQUARE CYLINDER As shown in Fig.6, it can be seen that for the lone square cylinder, Gerrard Mechanism (GM) was the dominant vortex shedding mechanism. This is because there are no disturbances in the wake region and it allows the Upper Shear Layer (USL) and Lower Shear Layer (LSL) to interact freely. The values of Cd and St are found to be 2.11 and 0.1226 respectively.

Fig.6 - GM for Square Cylinder EFFECTS OF VORTEX SHEDDING MECHANISMS With the introduction of a splitter plate in the near wake region of a square cylinder, it would bring about variations in the vortex shedding mechanisms. The effects of splitter plates at various lengths (L), various off-centered positions (h) and at various gaps (g) on the vortex shedding mechanisms are discussed here. As shown in Fig.6, it can be seen that for the lone square cylinder, Gerrard mechanism was the dominant vortex shedding mechanism. The general trend observed in the vortex shedding mechanisms was that, all the three vortex shedding mechanisms were found repeating as a pattern. The main finding was that the Gerrard mechanism (GM) was the dominant mechanism for the off-centered position 16


at h=0.15B. The reattachment mechanism (RA) was dominant for h=0.3B and the passive pushing mechanism (PP) was dominant for h=0.483B, Fig.7. However, it is observed that these vortex shedding mechanisms are not seen repeating throughout the experiment. The various other observations made in the experiment are stated below. a.

L/B=0.5 g/B=0 h/B=0.15 (GM)

b.

L/B=0.5 g/B=0.5 h/B=0.3 (RA)

c.

L/B=1 g/B=0.5 h/B=0.483 (PP) Fig.7 Gerrard Mechanism, Reattachment and Passive Pushing

EFFECTS OF GAP (g) Vortex shedding mechanism varies for various gaps at which the splitter plates are placed. Mohamed Surki et al. has shown that after a particular gap, the vortex shedding mechanism would start to show a change. This gap was selected as the critical gap (gcr). For all positions of h and g=0.25B and some cases of g=0.5B, Fig.8, there is a common pattern of vortex shedding. The pattern followed is that through the small gap fluid is seen moving in an upward and downward motion. Fluid from the LSL is seen moving in the upward direction and the USL moving in the downward direction. The fluid moving in the upward and downward direction can be seen cutting the USL and LSL respectively.

17


a.

b.

c.

L/B=0.5 g/B=0.25 h/B=0. 3 (UF) L/B=2 g/B=0.25 h/B=0.3 (UF) L/B=4 g/B=0.5 h/B=0.3 (UF) Fig.8 Upward Flow

For all the cases where the gap is set at 0.5B and 0.75B (Fig.9) and at all offcentered positions, it can be observed that the GM gradually disappears with the increase in length of the splitter plate. This is because the presence of splitter plate in the wake region prevents the formation of a fully developed vortex, mainly in the case of the USL. The LSL gets the opportunity to develop into a fully developed vortex because the disturbance of the square cylinder is minimal due to the offcentered position of the splitter plate. For cases where g=1B we can observe that partial GM can be seen in many cases. The reason behind this could be attributed to the fact that the gap would be sufficient for the LSL and USL to interact, but the presence of the splitter plate would prevent the vortex to develop into a fully developed, Fig 10. a.

L/B=1 g/B=0.75 h/B=0.3 (no GM)

b.

L/B=1.5 g/B=0.5 h/B=0.3 (no GM)

18


Fig.9 No Gerrard Mechanism a.

L/B=3 g/B=1 h/B=0.3 (p GM)

b.

L/B=4 g/B=1 h/B=0.3 (p GM) Fig.10 Partial Gerrard Mechanism

When the gap g is increased from 1B to 1.5B Fig.11, it is observed that vortex shedding is GM in all the cases. This is because, at this huge gap the splitter plate barely stands as a disturbance for the USL to develop into a fully developed vortex. a.

L/B=0.5 g/B=1.5 h/B=0.3 (GM)

b.

c.

L/B=3 g/B=1.5 h/B=0.15 (GM) L/B=2 g/B=1.5 h/B=0.3 (GM)

Fig.11 Gerrard Mechanism for a gap-ratio of 1.5 CRITICAL GAP (gcr) As mentioned above, increasing the gap from 1B to 1.5B would allow the free interaction of the USL and LSL. The presence of the splitter plate after this gap (1B) does not have an effect on the vortex shedding mechanisms. The USL and LSL roll 19


up completely and they attain full strength to interact and shed. So this abrupt change in the vortex shedding mechanism shows a transition in the flow fied mechanism after this gap. Hence it is inferred that the value of g=1B is the critical gap. EFFECTS OF LENGTH (L) In the case of L=0.5B and g=0, Fig.7a it is observed that the vortex shedding mechanism is GM. This is because the USL moves past the splitter plate and interacts with the LSL and vortex shedding takes place. In the case of L=0.5B and L=1B, Fig.12, GM dominates when the gap is set at 0.5B for all positions of h. This is because it can seen from the vortex shedding mechanism that the vortex formed from the USL becomes huge and it attains the strength to contain the whole splitterplate inside the vortex and the LSL is rolled upward to cut the USL and thus achieve GM. a.

L/B=0.5 g/B=0.5 h/B=0.483 (GM)

b.

L/B=1 g/B=0.5 h/B=0.3 (GM)

Fig.12 - Splitter plate contained in vortex For splitter plate positioned at L/B=4 and g/B=0.25, Fig.13, at all off-centered positions, it is observed that a backward flow takes place through the bottom portion of the splitter plate. But this backward flow has very small influence on the flow regime. a. L/B=4 g/B=0.25 h/B=0.3 (BF) Fig.13 - Backward Flow 20


In the case of L=4B and g=0, fully developed vortices are seen rolling almost at the center of the splitter plate, Fig 14. a.

L/B=4 g/B=0 h/B=0.15 (Centre) Fig.14 Fully Developed Vortex formed at centre of splitter plate

EFFECTS OF OFF-CENTERED POSITION (h) For various splitter plate offset value h, a significant observation is the variotions in size of upper and lower vortices formed. Upon change in offset of splitter plate from h=0.15B to 0.483B, it is seen that the upper vortices reduced in size and the lower vortices increased in size. The reduction in size of the upper vortices upon increase in value of h, is due to the interference of the splitter plate increasing with the USL, i.e., the presence of the splitter plate disallows the full roll up of the vortex. The increase in size of the lower vortices upon increase in value of h, is because the lower vortices gets the opportunity to roll-up fully and the interaction with USL is delayed. SECONDARY VORTICES Secondary vortices are the not-so-strong vortices which are formed due to the interaction of fluid with the splitter plates. The formation of secondary vortices starts when the splitter plate is placed at g=0.75B and L=1B and h=0.15B, Fig.15 The secondary vortices are formed from the trailing and leading edge of the splitter plate. At this position the the size of the secondary vortices are small and they donot interfere in the vortex shedding mechanisms. As the gap and length of the splitter plate are increased, the the secondary vortices gain size and strength and they start to interfere in the vortex shedding mechanism slowly. Large secondary vortices are observed for L=1B at g=0.75B until L=1.5B and L=2B for all positions of h, Fig.16. In these positions, the secondary vortices are seen interfering in the vortex shedding mechanisms. Secondary vortices from the trailing edge (both large and small) interfere with the LSL and the secondary vortices from the trailing edge interfere with the USL. Large secondary vortices are usually formed from the trailing edge and these suppress the vortices formed from the LSL. The small secondary vortices formed from the leading edge of the splitter plate are seen cutting the USL. This gradual increase in size and strength of the secondary vortices is observed only for splitter plate lengths up to L=2B. For values of L>2B, gradual 21


reduction in the size of secondary vortices is observed and ultimately the secondary vortices can be seen disappearing, Fig.17. a.

L/B=1.5 g/B=0.75 h/B=0.483 (SV)

b.

L/B=2 g/B=0.75 h/B=0.3 (SV) Fig.15 Secondary Vortices formed at leading edge of splitter plate

a.

L/B=1 g/B=0.75 h/B=0.3 (SV)

b.

L/B=1 g/B=1 h/B=0.3 (SV)

c.

L/B=3 g/B=0.5 h/B=0.483 (SV)

Fig.16 Large Secondary Vortices a.

L/B=3 g/B=0.5 h/B=0.15 (No SV)

22


Fig.17 No Secondary Vortices

EFFECTS OF Cd AND St The effect of splitter plates on the values of Cd and St are discussed in this section. The results are mainly plotted as a set of two graphs [(St vs. g) and (Cd vs. g)] each for all three off-center positions. STROUHAL NUMBER VARIATION For h=0.15B 0.160 0.155 0.150 0.145 0.140 0.135 0.130 0.125 0.120 0.115 0.110 0.105 0.100

L= 4B L= 3B L= 2B L= 1.5B L= 1B L= 0.5B 0

0.25 0.5 0.75

1

1.5

Fig.18. Variation of St at h=0.15B at various positions of g. From Fig.18, it is seen that except in the case of L=4B and L=3B the St variations follow a pattern. The value of St decreases till g=0.75B. At g=1B it is seen that there is an increase in the value of St. Another common pattern for the graphs are the values of St after gcr, all the values almost coincide with the value of St of the square cylinder i.e., 0.12. This is because at g>1B the splitter plate gap becomes so high that vortex shedding is not affected by the splitter plate. For h=0.3B From Fig.19, it is seen that except in the case of L=1.5B St decreases until g=0.75B and thereafter exhibits the pattern, seen in Fig.18, of St approaching the value of the square cylinder at g=1.5B.

23


0.180 0.175 0.170 0.165 0.160 0.155 0.150 0.145 0.140 0.135 0.130 0.125 0.120 0.115 0.110 0.105 0.100 0.095 0.090

L= 4B L= 3B L= 2B L= 1.5B L= 1B L= 0.5B 0

0.25

0.5

0.75

1

1.5

Fig.19. Variation of St at h=0.3B at various positions of g For h=0.483B As seen from Fig.20, for all the cases barring L=1.5B, there is a constant reduction of St until g=0.75B and at g=1.5B the value of St apporoaches the value of a square cylinder. 0.160 0.155 0.150 0.145 0.140 0.135 0.130 0.125 0.120 0.115 0.110 0.105 0.100 0.095 0.090

L= 4B L= 3B L= 2B L= 1.5B L= 1B 0 0.25 0.5 0.75 1 1.5

L= 0.5B

Fig.20. Variation of St at h=0.483B at various positions of g

24


CO-EFFICIENT OF DRAG VARIATION For h=0.15B It is seen from Fig.21, that there is a common pattern for the variation of Cd. The value of Cd decreases with the increase in gap until g=1B. After this point it is seen that the value of Cd approaches the value of a square cylinder. The reason behind this is that at g=1.5B the splitter plate has no effect on the flow field since the gap is large. 2.200 2.100 2.000 1.900 1.800 1.700 1.600 1.500 1.400 1.300 1.200 1.100 1.000

L= 4B L= 3B L= 2B L= 1.5B L= 1B L= 0.5B 0 0.25 0.5 0.75 1

1.5

Fig 21. Variation of Cd at h=0.15B at various positions of g. For h= 0.3B It is seen from Fig.22 that there is a common pattern in the variation of the value of Cd, for all the splitter plates except for L=0.5B and L=1B. The common pattern is that the value of Cd decreases until g=1B and thereafter there is a suuden increase in the value of Cd; all the values approach the square cylinder value. 2.200 2.100 2.000 1.900 1.800 1.700 1.600 1.500 1.400 1.300 1.200 1.100 1.000

L= 4B L= 3B L= 2B L= 1.5B L= 1B 0 0.25 0.5 0.75 1 1.5

25

L= 0.5B


Fig 22. Variation of Cd at h=0.3B at various positions of g. For h= 0.483B 2.200 2.100 2.000 1.900 1.800 1.700 1.600 1.500 1.400 1.300 1.200 1.100 1.000

L= 4B L= 3B L= 2B L= 1.5B L= 1B L= 0.5B 0 0.25 0.5 0.75 1

1.5

Fig 23. Variation of Cd at h=0.483B at various positions of g. It is seen from Fig.23 that there again is a common pattern in the variation of the value of Cd. With increase in value of g, value of Cd decreases until g=1B and thereafter there is a sudden increase in the value of Cd to the square cylinder Cd range.

NUMERICAL AND EXPERIMENTAL RESULTS - COMPARISON The numerical results which were obtained from ANSYS Fluent 14.5 and the experimental results are compared in this section.

26


Numerical Results

Experimental Results

27


Table 2 Cd - Numerical vs. Experimental Cd - Numerical 2.2 1.74 1.71 1.503 1.656

Cd - Experimental 2.11 1.67 1.65 1.482 1.59

CONCLUSIONS The variations in the values of St and Cd and the variations in the vortex shedding mechanisms occurring due to the splitter plate placed in the near wake region of a square cylinder are studied here. The main finding from the analysis of vortex shedding mechanism is the identification of a critical gap 1B. It is seen that the value of Cd shows a decreasing trend with increase in the value of g until g=1B and thereafter shows a sudden increase for values of g>1B, approaching a Cd value of a lone square cylinder at g=1.5B. The St shows a common pattern only for L=4B for all h/B ratios, a diminishing trend until g=1B and a sudden spike after g>1B. The St does not demonstrate a definitive trend (barring L=4B) on increasing the value of g but does undergo a rapid change for a value of g>1B, that tends towards St value equal to that of a lone square cylinder. Thus, the value g=1B is determined to be the critical gap, gcr. The other conclusions made from the analyses are: 1.

2. 3.

4.

5. 6.

For g=0.25B there is an upward and downward flow through the small gap. This upward and downward flow contributes to the vortex shedding by cutting the USL and LSL simultaneously. This phenomenon is seen repeating for all positions of h. For small splitter plate, L=0.5B, GM dominates for most of the cases because the USL rolls past the splitter plate and interacts with the LSL. As g is increased, GM gradually disappears. Secondary vortices start forming when the g is set at 0.75B for L=1.5B from the edges of the splitter plate. Large secondary vortices are found interfering in the vortex shedding mechanism. The size of secondary vortices increase with the increase in g and L and cease to appear for large splitter plates. As the length of the splitter plate is increased to 3L, at moderate gaps a backward flow is seen through the bottom portion of the splitter plate. This backward flow has no effect on the flow mechanism. For large splitter plates, L=4B and g=0, the vortices are seen rolling almost at the center of the splitter plate. For all values of h and L, at g=1.5B, the vortex shedding mechanism is largely GM.

Vorticity profile, value of Cd and St for few cases are found out numerically and validated with the experimental results. 28


REFERENCES R. Ajith Kumar, ChangHyun Sohn, and B.H.Lakshmana Gowda (2009). “Influence of corner radius on the near wake structure of a transversely oscillating square cylinder”, Journal of Mechanical Science and Technology, 23(9):2390–2416. Y. Nakamura (1996), “Vortex shedding from bluff bodies with splitter plates”, Journal of Fluids and Structures 10, 147 – 158 Mohamed Sukri Mat Ali, Con J. Doolan, and Vincent Wheatley (2011), “Low Reynolds number flow over a square cylinder with a detached splitter plate”, International Journal of Heat and Fluid Flow, 36:133-141. E. A. Anderson and A. A. Szewczyk, (1997) “Effects of a splitter plate on the near wake of a circular cylinder in 2 and 3-dimensional flow configurations”, Experiments in Fluids 23(2):161-174. C.J. Apelt, G.S. West, A. Szewczyk (1973), “The effects of wake splitter plates on the flow past a circular cylinder in the range 104 < Re < 5104”, Journal of Fluid Mechanics, 61:187-198. P.W. Bearman (1965), “Investigation of the flow behind a two-dimensional model with a blunt trailing edge and fitted with splitter plates”, Journal of Fluid Mechanics. 21 (2) 241-255 P W Bearman (1984), “Vortex shedding from oscillating bluff bodies”, Annual Review of Fluid Mechanics, 16(1):195–222. C. J. Doolan (2009) “Flat-plate interaction with the near wake of a square cylinder”, AIAA J. 47(2): 475-479. Gerrad. J.H (1966), “The mechanics of the formation region of vortices behind bluff bodies”, Journal of Fluid Mechanics 25, 401 - 413. R.H. Hernandez, C. Baudet, and S. Fauve (2000), “Controlling the benard-von Karman instability in the wake of a cylinder by driving the pressure at the front stagnation point”, The European Physical Journal B – Condensed Matter and Complex Systems, 14(4):773–781. Kiyoung Kwon and Haecheon Choi (1996), “Control of laminar vortex shedding behind a circular cylinder using splitter plates”, Physics of Fluids (1994-present), 8(2):479–486. A. Prasad and C. H. K. Williamson (1997), “A method for the reduction of bluff body drag”, J.Wind. Eng. Ind. Aerodyn. 69–71, 155. 29


D Rockwell and E Naudascher (1979), “Self-sustained oscillations of impinging free shear layers”, Annual Review of Fluid Mechanics, 11(1):67–94. A. Roshko (1954), “On the drag and shedding frequency of two dimensional bluff bodies”, National Advisory Committee for Aeronautics, Technical Note 3169, 1954, pp. 1-29. A. Sohankar, C. Norberg, and L. Davidson (1998), “Low-Reynolds-number flow around a square cylinder at incidence: study of blockage, onset of vortex shedding and outlet boundary condition”, International Journal for Numerical Methods in Fluids, 26(1):39–56. M.M. Zdravkovich (1987), “The effects of interference between circular cylinders in cross flow”, Journal of Fluids and Structures, 1(2):239–261.

30


EXPERIMENTAL INVESTIGATION OF PERFORMANCE ON A PCM ASSISTED SOLAR WATER HEATER Vibin kumar C S 1, Swaraj kumar B 2 1

M.Tech Student,2Professor L B S College of Engineering, Kasaragod – 671 542, Kerala, India.

ABSTRACT Latent heat thermal energy storage is one of the most efficient ways to store thermal energy in solar water heaters. The present experiment focuses to convert the temperature of 60 liters of water to 70oC, from 27 oC during the day time through a flat plate collector by thermo syphon effect and to maintain 50oC temperature till morning with the assistance of PCM. The flat plate collector made of Copper and a heat storage unit of stainless steel consisting of PCM (paraffin-melting point 56oC) functioning simultaneously. The storage unit utilizes cylindrical tubes made of copper and curved cylindrical tubes made of steel filled with paraffin wax as the heat storage medium. The unit is constructed with cost effective materials such as wood, polyurethane form and expanded polystyrene for insulation. The hot water temperature can be increased to 5 o C by the application of PCM than the normal conditions.

INTRODUCTION The primary energy sources have finite supplies in the earth and are depleting day by day due to the higher consumption rates. The increasing population and the human comfort level standards speed up the crisis level and it leads to the effective utilization of secondary energy sources. The large scale application of fossil fuels causes atmospheric pollution by the evolution of harmful gasses in large scale like nitrogen oxides (NOx), sulphur oxides (SOx), carbon monoxide (CO), carbon dioxide (CO2), hydrocarbons, suspended particles and fly ash. Due to these pollutants causes ozone layer depletion, health hazards, global climate change and global warming 1

Vibin Kumar C S. E-mail: [email protected]

Experimental Investigation Of Performance On A PCM Assisted Solar …

etc. The secondary sources are mainly of solar, wind, tidal, wave, hydro and geothermal energies. In which the main energy source is of the sun and these alternative sources are the manifestation of solar energy. The solar energy is freely available, is environmentally clean and is accepted as the most promising alternative energy. By the effective utilization of solar energy can control the energy crisis and atmospheric pollution all over the world.

Fig.1. The experimental setup.

LITERATURE REVIEW Nosa Andrew Ogie et al (2013) constructed a solar flat-plate collector water heating system for domestic purposes by using locally available materials in Nigeria. Maximum fluid output temperature, the collector temperature, and insolation obtained are of 55 °C, 51 °C, and 1,480 W/m2, on a clear sunny day through thermo 32


syphon principle.B. R. Vijay Prithiv et al (2014) conducted an experimental setup at Chennai, focuses on Solar Thermal Energy storage system with the help of encapsulated PCM (Paraffin & Honey wax) by using thin copper tube of 12mm diameter. I. Al-Hinti et al (2010) conducted an experimental investigation by using Paraffin wax as the PCM which is filled in small cylindrical aluminum containers. Vikram D et al (2006) conducted an experimental work in which the PCM is encapsulated in aluminium cylinders of internal diameter 34mm. M.H. Mahfuz et al (2014) conducted an experimental study at Malaysia to analyze the performance of a vertical arrangement of shell and tube latent heat TES unit with Paraffin wax. Mohammad Ali Fazilati et al (2013) had done an experimental research on PCM in spherical capsules as storage medium at Iran. R. Meenakshi Reddy et al (2012) conducted an experimental investigation on TES system using different phase change materials (paraffin 61 °C and Stearic acid 57 °C) for various sizes of spherical capsules. The efficiency enhancement is resulted in all works using paraffin in thin copper tubes, small diameter cylinders or spherical capsules

EXPERIMENTAL SETUP The experimental setup consisting of a solar flat plate collector, an insulated storage water tank, cylindrical copper tubes and curved cylindrical steel tubes filled with phase change material (paraffin), four numbers of Thermo couples and a digital temperature indicator. The unit is associated with an auxiliary heater to maintain the temperature in the storage water tank. Table 1 Specifications of the experimental setup. Description Collector dimensions Number of glass covers Glass cover transmissivity Cover thickness Absorber plate dimensions Absorber plate, Header and Riser material Thermal conductivity of the plate material Plate thickness Absorber area Absorptivity Header tube size Riser tuber size Number of riser tubes used Insulation material &Thermal conductivity Insulation thickness The average solar insolation

Specification 1.7m x 0.9 m x 0.1 m 1 0.88 4 mm 1.50 m x 0.80 m Copper 386 W/m °C 1mm 1.2 m2 .65 Φ19.5 mm Φ9.5 mm 10 Polyurethane form,0.64W/mK Expanded polystyrene,0.034W/m k 5cm 709W/m2


EXPERIMENTAL PROCEDURE The wind speed in the location is taken as 2m/s.The storage tank is filled with water at a temperature of 27o C. The thermo syphon circulation is started accordingly with the temperature increase of flat plate collector. The ambient temperature, the collector and the storage tank temperature at two locations are taken in an interval of 30 minutes and noted. The experiments with the copper tube assembly and steel tube, filled with paraffin wax PCM were conducted. The temperature variations are noted till the next day 8am.

ANALYSIS AND DISCUSSIONS PERFORMANCE OF THERMO SYPHON SYSTEM. The collector performed well in the passive circulation mode by achieving a storage water temperature of 59oC. The maximum temperature obtained by the collector is 86oC at 12.30pm and the maximum temperature obtained for the storage water is at 3.30pm such as the thermo syphon effect is continuing till 3.30pm. After then the system is not performed well due to the inadequate solar radiation to maintain the collector temperature.

Fig.2. Time v/s Temperature of the thermo syphon system. PERFORMANCE OF SWH WHILE USING PCM IN CU TUBE. The copper tubes were used for the encapsulation of 55g paraffin. The solidifying temperature of paraffin helps the storage water to stay at 57o C for 75 to 80 minutes due to its higher thermal conductivity properties. The temperature obtained in the morning 8.00am is 51oc and is sufficient for the domestic applications

34


Fig.3. Time v/s Temperature plot for the Copper tube geometry. PERFORMANCE OF SWH WHILE USING PCM IN STEEL TUBE. The paraffin of 77g is encapsulated in the curved steel tubes and the solidifying temperature of wax helps to stay the storage water at 57o C for a period of 55 to 60minutes.The temperature of storage water is decreased slightly in the beginning and then it reaches uniform. This may be due to the higher thickness of tube and the low thermal conductivity property. The temperature obtained at 8am is 50o C which is sufficient for the domestic applications.

Fig.4. Time v/s Temperature for the steel tube geometry. PERFORMANCE OF SWH WITHOUT USING PCM. The performance of the system without PCM is also analyzed for comparing it with the PCM geometries. The temperature maintained is 65oC for the experiment and the temperature obtained in the morning is 46oC.


Fig.5. The Time v/s Temperature plot -Without PCM

PERFORMANCE COMPARISON. The PCM encapsulations in different geometry were studied. The paraffin wax used is of 770 g for 60 liter water. The temperature obtained for Cu tube geometry is 1oC higher than the steel tube. The PCM performance in the experiment is proved to an achievement of 4 to 5oC temperature than a system without PCM. By the application of PCM, the storage water temperature of 57oC is maintained constant for 55 to 80 Minutes. The PCM in the copper tube geometry shows more effective. The temperature obtained without the use of PCM is 46oC and by using PCM, a temperature range of 50 to 51oC is obtained.

Fig.6. The Time v/s Temperature plot –PCM in Cu, Steel and Without PCM.

36


EFFICIENCY ANALYSIS WITH TIME The efficiency v/s Time of the flat plate collector is plotted. The efficiency of the collector decreases with the increase in temperature. As the temperature increases, the losses also increases thus by the efficiency decreases.

Fig.7. The time v/s Efficiency For better melting and solidification, the cylindrical tubes or spherical capsules of diameter below 10mm and low wall thickness below 1mm should be effective to be experimentally analyzed. By introducing small metallic strips in the tube or sphere cavity can improve the thermal conductivity. The present setup is fabricated with cost effective materials for the domestic applications. By using nickel –chromium based black enamel coating; the reflection losses can be minimized. The heat loss can be further minimized by the application of glass wool insulation to the storage tank. The thermo syphon effect can further improved by increasing the head of the storage tank.

CONCLUSIONS The solar water heater with phase change material is experimentally analyzed and the various parameters were studied. 1.

2.

3.

The thermo syphon effect is a major advantage to the domestic solar water heater system; otherwise an additional electrical energy source is needed for the active circulation of the heat transfer fluid. The thermo syphon effect can further improved by increasing the head of the system. The performance depends on the collector location, collector area collector tilt, wind velocity, and the solar time. The performance of paraffin wax in different geometries is studied. The PCM can improve the performance of solar water heater. The PCM with in the Copper tube is more effective than the steel tube.


4.

The temperature variations were obtained throughout the testing period due to the variations in the incident solar insolation. The performance of the collector should be increased in a clear sunny day.

The work is done with 770 g of PCM. If the amount of PCM increases, the performance also increased. The different PCM geometries and its performance can be analyzed.The possibility of application of flat plate collector as a heat pump can be studied.

REFERENCES B. R. Vijay Prithiv, R. Prabakaran, R. Sukumar,Dr. A. Rajendra Prasad, (2014), “Heat Enhancement of Domestic Solar Water Heater using PCM Filled Thin Tubes”, ISR National Journal of Advanced Research in Mechanical and Production Engineering and Development Volume: 1 Issue: 1 . I.Al-Hinti, A. Al-Ghandoor, A. Maaly, I. Abu Naqeera, Z. Al-Khateeb, O. AlSheikh, (2010) “Experimental investigation on the use of water-phase change material storage in conventional solar water heating systems”,Energy Conversion and Management 51, 1735–1740 M.H. Mahfuz, M.R. Anisur, M.A. Kibria, R. Saidur, I.H.S.C. Metselaar, (2014) Performance investigation of thermal energy storage system with Phase Change Material (PCM) for solar water heating application, International Communications in Heat and Mass Transfer 57, 132–139 Mohammad Ali Fazilati, Ali Akbar Alemrajabi, (2013) “Phase change material for enhancing solar water heater, an experimental approach”, Energy Conversion and Management 71, 138–145. Nosa Andrew Ogie, IkponmwosaOghogho, Julius Jesumirewhe, (2013), “Design and Construction of a Solar Water Heater Based on the Thermo syphon Principle” Ashdin Publishing Journal of Fundamentals of Renewable Energy and Applications Vol. 3 Article ID 235592, 8 pages R. Meenakshi Reddy, N. Nallusamy, and K. Hemachandra Reddy, (2012), “Experimental Studies on Phase Change Material-Based Thermal Energy Storage System for Solar Water Heating Applications”, Ashdin Publishing Journal of Fundamentals of Renewable Energy and Applications, Vol. 2 , Article ID R120314. Vikram D, Kaushik S, Prashanth V, Nallusamy N, 2006 “An Improvement in the Solar Water Heating Systems using Phase Change Materials”, Proceedings of the International Conference on Renewable Energy for Developing Countries. 38


EXPERIMENTAL DETERMINATION OF WETTING FRONT SPEED DURING JET IMPINGEMENT COOLING OF A HOT SURFACE OF DIFFERENT TEMPERATURE Chitranjan Agrawal1, Mahesh Singh2, Manoj Verma2, Ankit Sukhwal 2 1

Assistant Professor,2B. Tech. Student College of Technology and Engineering, Maharana Pratap University of Agriculture and Technology, Udaipur-313001, Rajasthan, India.

ABSTRACT The quenching of a hot Stainless Steel (SS-304) horizontal surface at 450 – 620 ºC initial temperature is performed with a water jet of 33 ºC temperature and 3 mm diameter. The quenching performance is investigated with water flow rate of 1.2 and 5.10 lpm on the surface of 150 mm long, 150 mm wide and 2 mm thickness. Initially surface is heated up to desired initial temperature in a furnace and cooled by downward impinging water jet. The quenching process is recorded by a digital camera and the wetting front progression over the hot surface is determined. The wetting front speed on the hot surface is found in the range of 2 – 35 mm/s for downstream spatial locations of 10 mm – 40 mm. The wetting front speed reduces with the increase in spatial locations and surface initial temperature, however, increases with the rise in water flow rate.

INTRODUCTION The jet impingement cooling technique has been applied in several industries e.g. manufacturing, metal processing, electronics, nuclear, automobile, due to its high heat removal capability (Hatta et al 1983, Webb and Ma 1995). The jet impingement 1

1Chitranjan Agrawal. E-mail: [email protected]

Experimental Determination of wetting Front Speed during Jet Impingement…

quenching performance for a hot surface has been investigated under steady sate and transient state cooling condition many times (Agrawal et al. 2012) with water. (Agrawal et al. 2012, 2013) and other fluids (Kumar et al. 2014) as coolant. The transient cooling performance of hot surface is normally evaluated on the basis of rewetting temperature (Agrawal et al. 2012, 2014, 2015a), Wetting delay (Mozumder et al. 2005), maximum surface heat flux (Hall et al. 2001, Hammad et al. 2004) or the wetting front speed (Agrawal 2013, 2015b, Akmal et al. 2008). As the jet of sub- cooled fluid strikes onto the hot surface, the formation of vapor bubbles restricts the direct liquid surface contact. These vapor bubbles prohibits the downstream progression of the coolant wetting front over the hot surface. Soon after, as the hot surface attain the rewetting temperature, at a certain downstream spatial location, the advancement of coolant wetting font take place from that spatial location. The liner progression of the coolant wetting front per unit time, in downstream direction is termed as wetting front speed. The wetting front speed is the measure of rapidity of surface quenching as desirable in certain industries like nuclear, metal processing respectively for safety purpose under LOCA (Loss of Coolant Accident) and controlling the material property. The wetting front speed increases with the rise in jet diameter and coolant flow rate, however, reduces for the downstream locations as compared to the region near to stagnation point (Agrawal et al. 2012, 2013, 2015a). The wetting front speed is larger for the ferrous surface as compared to the non ferrous surfaces due to lower thermal diffusivity posses by the ferrous surfaces (Hammad et al. 2004). The hot surface quenching performance has been evaluated several times for rewetting temperature, wetting delay, maximum surface heat flux with different surface initial temperature, coolant flow rate and coolant temperature (Agrawal et al. 2012, Hammad et al. 2004). The determination of wetting front speed for ferrous surface has been reported with the jet of different diameter and coolant flow rate (Akmal et al. 2008, Karwa et al. 2011). However, the effect of surface initial temperature on the wetting front speed with ferrous surface has not been observed particularly with jet impingement cooling of hot flat surface for downstream locations. Therefore, an experimental investigation has been carried out to detriment the progression of wetting front for downstream spatial locations by varying the surface initial temperature.

EXPERIMENTAL INVESTIGATIONS MATERIALS A hot flat surface of stainless steel (SS-304) was cooled with a downward impinging round water jet of 33 ºC temperature and 3 mm diameter. The schematic of experimental setup is shown in Figure 2. Initially, water was stored in a reservoir (1) 40


and supplied to the nozzle (6) with the help of a pump (2). The control valve (4) were used to regulate the water flow towards the nozzle through a rota-meter (3) and bypassed back to the reservoir. The hot surface temperate was recorded by an ungrounded ‘K’ type thermocouple (8) and a temperature indictor (9). The hot test surface position underneath to the nozzle can be adjusted by a handle (11). The nozzle (6) was fixed on a base that is further attached onto the two vertical supports of the experimental set up through nut and bolt arrangement, such that nozzle can be moved in vertical and horizontal direction. The various operating parameters used for the investigations are shown in Table (1).

1. Reservoir 2. Water pump 3. Rota-meter 4. Control valve 5. Straight pipe 6. Nozzle 7. Test-surface 8. Thermocouple wire 9. Data-acquisition system 10. Base 11. Handle

Fig. 2 Schematic of experimental set up Table 1 Operating range of experimental parameters Experimental parameter

Operating range

Water flow rate, lpm

1.2, 5.1

Jet Reynolds number

11500, 48000

Water temperature, oC

33

Jet diameter, mm

3

Jet exit to surface spacing, mm

12

Test surface length and width, mm

150x150

Thickness of test-surface, mm

2

Spatial locations, mm

10 , 40

Initial surface temperature, ˚C

450 , 550, 620

41


METHODS Initially, the test surface was heated up to desired initial temperature in a furnace and then kept on the experimental set up underneath the nozzle. The nozzle exit to test surface spacing, z, was maintained at 12 mm such that dimensionless nozzle exit to

test surface spacing remained as z/d = 4. The quenching performance of the flat hot test surface was determined by analyzing the wetting front progression towards the radial downstream locations, r. The videos of quenching process were captured with the rate of 30 fps and analyzed by using Dartfish video analysis software. The progression of wetting front up to a certain downstream location per unit time is considered as wetting front speed, (u = r/t), in similar manner as reported earlier by Agrawal et al. 2013, 2015b, and Akmal et al. 2008. The experiments were performed with different surface initial temperature by varying coolant flow rate as mentioned in Table 1. The experimental uncertainty for the wetting front speed was found 10 % - 15 % for 10 mm spatial location and 2.5 % - 10 % for 40 mm spatial location. For every experiment a new test surface was used to avoid the effect of change in surface properties and oxidation due to previous experiments

ANALYSIS AND DISCUSSIONS The impingement quenching experiments were performed on the hot flat surface at 450 – 620 ºC initial temperature of 2 mm thickness. During experiments it was observed that the hot surface at the stagnation point get cooled immediately as the jet strikes to the surface. However, with the progression of wetting front towards the downstream spatial location, some amount of coolant fluid splashes obliquely away from the hot surface in the upward direction, as shown in Fig. 3.

Fig. 3 Visual observation of coolant flow over the hot surface 42


It is observed that the violent boiling of fluid takes place at the periphery of the wetting front, possibly this is the region of transition boiling. Since, initial temperature of test surface is of the order of 620 - 450 ºC, thus, the temperature of coolant at the edge of wetting front may reach to the superheated stage. With the progression of wetting front the high temperature coolant further absorbed the heat from the hot surface and leads to formation of vapor bubble. The frequent bubble formation and subsequent collapse may be the possible reason for this splashing

phenomenon of the coolant from the hot test surface. With the expense of time the intensity of splashing phenomena reduces particularly for the downstream spatial locations, perhaps due to reduction in surface temperature from its initial temperature. The wetting front speed for the investigated initial surface temperature at 10 mm and 40 mm locations with water flow rate of 1.2 lpm and 5.1 lpm is shown in Fig. 4.

Fig. 4 Effect of surface initial temperature and flow rate on wetting front speed

From Fig. 4 it has been observed that for a certain flow rate and spatial location the wetting front speed reduces with the increase in surface initial temperature. The reduction in wetting front speed exaggerated for the higher order of surface initial temperature. The wetting front speed reduces in the range of 10 - 14 % with the water flow of 1.2 lpm and in the range of 20 - 25 percent with 1.5 lpm, for the investigated spatial locations by increasing initial surface temperature from 450 ºC to 550 ºC. Whereas, with further rise in temperature up to 620 ºC, the reduction in wetting speed is observed in the range of 50 -55 % at 10 mm location and 35 – 45 % at 40 mm location for the investigated water flow rate. With the rise in surface initial temperature, the associated stored energy increase. Thus, lager amount of heat has to be absorbed for a certain coolant flow rate and hence the wetting front speed 43


reduces. The reduction in the wetting front speed is larger for the 10 mm location as compared to the 40 mm location with the increase in surface initial temperature. The tendency of coolant splashing may possible reason for this observation. The greater amount of coolant splash out away from the surface with the surface of higher

initial temperature. The higher surface temperature lead to enhanced frequency of bubble formation and collapsing, resulting greater amount of coolant splashing, away from the surface. Therefore, with reduces initial amount of coolant the reduction in the wetting front speed for 10 mm location is higher as compared to the 40 mm spatial location. In is also observed that the wetting front speed reduces for the downstream spatial locations. The rise in spent out fluid enthalpy, flow retardation for downward direction, larger peripheral surface area to be cooled with the available coolant and increase in thermal / hydraulic boundary layer thickness for downstream locations may be the possible reason for this. The spatial reduction in the wetting front speed is higher for lower temperature surfaces as compared to the higher temperature, which is further higher with higher coolant flow rate. This result reflects that with the rise in coolant flow rate, the surface near the impingement point get cooled much earlier as compared to the farthest downstream location. However with the rise in surface temperature, the initial splashing of fluid hampers the quenching performance of surface even for the location near to the impingement point.

CONCLUSIONS The jet impingement surface quenching experiments of 450 – 620 ºC temperature at 10 mm and 40 mm spatial location with different coolant flow rate reveals that. 1. The quenching performance increase with the rise in coolant flow rate however, reduces with the rise in surface initial temperature and for the downstream spatial location. 2. For a certain spatial location the reduction in the wetting front is larger for the extreme temperature surfaces due to possible increase in stored energy. 3. The spatial reduction in the wetting front is larger for the surface of lower initial temperature as compared to the higher temperature surfaces due to comparatively higher amount of jet fluid splashing as the jet strikes to the hot surface. This experimental investigation incorporated some of the parameters associated with the jet impingement hot surface quenching, however an exhaustive research can also be performed by incorporating other operating parameters. The results can be presented in the form of generalized correlation that may be helpful to the researchers and industrialist particularly for controlling specific metal properties during casting and extrusion processes. 44


REFERENCES Agrawal, C., Kumar, R., Gupta, A., Chatterjee, B. (2012), “Effect of Jet Diameter on the Rewetting of Hot Horizontal Surfaces During Quenching”, Experimental Thermal and Fluid Science, 42, 25-37.. Agrawal, C., Kumar, R., Gupta, A., Chatterjee, B. (2013), “Determination of Rewetting Velocity during Jet Impingement Cooling of a Hot Surface, ASME Thermal Science and Engineering Application, 5, 011007-1-9. Agrawal, C., Kumar, R., Gupta, A., Chatterjee, B. (2014), “Effect of Nozzle Geometry on the Rewetting of Hot Surface during Jet Impingement Cooling”, Experimental Heat Transfer, 27, 256 – 275. Agrawal, C., Kumar, R., Gupta, A., Chatterjee, B. (2015a), “Rewetting of Hot Vertical Rod during Jet Impingement Surface Cooling”, Heat and Mass Transfer, DOI:10.1007/s00231-015-1637-9. Agrawal, C., Kumar, R., Gupta, A., Chatterjee, B., (2015b) “Determination of Rewetting Velocity during Jet Impingement Cooling of Hot Vertical Rod”, Journal of Thermal Analysis and Calorimetry, DOI 10.1007/s10973-015-4905-5. Akmal, M., Omar, A.M.T., and Hammed, M.S. (2008), “Experimental Investigation of Propagation of Wetting Front on Curved Surfaces Exposed to an Impinging Water Jet”, International Journal of Microstructure and Material Properties, 3, 654681. Hall, D. E., Incropera, F.P., and Viskanta, R. (2001), “Jet Impingement Boiling From a Circular Free-Surface Jet During Quenching: Part 1—Single Phase Jet, ASME Journal of Heat Transfer, 123, 901-909. Hammad, J., Mitsutake, Y., Monde, M. (2004) Movement of maximum heat flux and wetting front during quenching of hot cylindrical block, International Journal of Thermal Sciences, 43, 743–752. Hatta N., Kokado J., Hanasaki, K. (1983), “Numerical Analysis of Cooling Characteristics for Water Bar”, Transaction of ISIJ, 23, 555-64. Mozumder, A.K., Monde, M., and Woodfield, P.L. (2005), “Delay of Wetting Propagation During Jet Impingement Quenching for a High Temperature Surface”, International Journal of Heat and Mass Transfer, 48, 5395–5407. Karwa N., Roisman, T.G., Stephan, P. and Tropea, C. (2011), “Experimental Investigation of Circular Free- Surface Jet Impingement Quenching: Transient Hydrodynamics and Heat Transfer”, Experimental Thermal and Fluid Science, 35, 1435-1443. 45


Kumar R., Jha J. M., Mohapatra S. S., Pal S. K., Chakraborty S. (2014), “Surfactant Experimental Investigation of Effect of Different Types of Surfactants and Jet Height on Cooling of a Hot Steel Plate”, ASME Journal of Heat Transfer, 136, 072102-1-10. Webb B.W., and Ma, C.F. (1995), “Single Phase Liquid Jet Impingement Heat Transfer”, Advance in Heat Transfer, 26, 105- 217.

46


EXPERIMENTAL EVALUATION OF TWIN-CHAMBER COMMUNITY SOLAR COOKER FOR EFFICIENT COOKING Mandeep Singh Sekhon1, VP Sethi*1, Amanpreet Singh Dhaliwal1 1

Department of Mechanical Engineering Punjab Agricultural University Ludhiana- 141 004, Punjab, India

ABSTRACT In this study, a new efficient design of rectangular box type Twin-Chamber Community (TCC) solar cooker is presented. The main feature of this new design is that it can be used for cooking food for at least 100 persons at one time. The major improvement of new design is conversion of single rectangular chamber into two separate square chambers of size 40”x40” each for cooking in sections with food which need lesser time and other for cooking food which needs more cooking time simultaneously without disturbing the stagnation temperature of each chamber which improves the cooking efficiency of food. The developed cooker was experimentally tested for computing first figure of merit (F1), second figure of merit(F2) and time taken to boil as per BIS standards in the month of May 2015 at Ludhiana climate (latitude of 30.910 N). It was observed that the developed cooker is A grade category cooker with F1 of 0.128, F2 of 0.41 and time taken to boil is 187 minutes. If cooking of one meal per day for 280 days per year is assumed, then the proposed community cooker is capable of saving 1100 kg of LPG which is equivalent to Rs 34,650 at subsidized rate of LPG. Thus it can recover its cost in just less than 18 months. This cooker can also earn 3.3 carbon credits for the community.

INTRODUCTION The design modifications of solar cooker have undergone a phase of rigorous development since last half a century. Domestic box type solar cookers were designed for use in single family of 4-5 members, which costs approximately Rs 3500. Thus, for a community of 25 families (100 people), the initial investment cost of domestic cookers *1


Experimental evaluation of twin-chamber…..

will approach to an unaffordable price of Rs 90000. Also, box type solar cooker has been tried in the past to cook both types of foods (one which take less time to cook, other which take more cooking time). However, single chamber design has a major drawback that when the cooker is opened for unloading of cooked one type of meal, it’s stagnation temperature lowers significantly which effects the cooking performance for other meal. Yadav and Tiwari (1987) performed the transient analysis of the box type solar cooker and expressions for the temperatures of the utensil, utensil water, glass cover, stagnant air between the cooker plate and the glass cover were developed. Mullick et al (1987) developed a thermal test procedure to test the performance of solar cooker using two figures of merit F1 and F2 and found to be 0.12 and 0.25 resp. for satisfactory performance. Jubran and Alsaad (1991) developed a mathematical model using the heat balance analysis of the various components of the cooker. Negi and Purohit (2005) presented a box type solar cooker utilizing non-tracking concentrator optics to enhance the solar energy availability in the box of the cooker. A stagnation temperature of 15–22 0 C higher than that of the conventional box type solar cooker using a booster mirror was observed. Sethi et al (2014) presented an optimally inclined box type solar cooker design with single booster mirror and improved parallelepiped cooking vessel design. The first and second figure of merit (F1) and (F2) for inclined cooker were found to be 0.16 and 0.54 as compared to 0.14 and 0.43 for horizontally placed cooker. While conducting the literature review it was observed that more research has been performed on domestic box type solar cookers. Efforts are still to be made to develop and test community scale solar cooker for lowering the cooking cost. Thus this design is a step forward in the direction for efficient cooking of both types of food simultaneously by determining the first figure of merit, second figure of merit and standard boiling time of the TCC solar cooker as per BIS standards.

EXPERIMENTAL INVESTIGATIONS CONSTRUCTIONAL DETAILS This community solar cooker comprises of symmetrical twin-chamber design. Surface area of reflecting mirror for each chamber is 1.403 m2 (115cm x 122cm). Reflecting mirror is attached to the inner side of opening lid of cooker, which is erected using pulley arrangement, having 9 spoke pulley with diameter 15 cm. The major benefit of this arrangement is that any angle for mirror can be adjusted according to use of solar cooker at different latitudes. Glass cover area measures 1.2544 m2 (112cm x 112cm). Double glazed arrangement having air gap of 10 mm in between two glass covers is used. This will reduce the radiative heat loss from the glass plate to ambient air significantly Glass pane arrangement (acting as cover lid of chambers), is attached using shocker arrangement as shown in diagram. Each shocker is 50 cm long and it helps in easy uplifting of the glass pane arrangement despite its heavy weight (large dimensions due to community sized cooker). Each chamber is having tapered walls and depth of 11cm and absorber plate area of 1.02 m2 (101cm x 101cm). Insulation of 4 cm is provided between 48


chamber and four outer side walls and 12 cm is provided between chamber and base of the cooker. Also, insulation of 12 cm thickness is provided between both the chambers. It will reduce the conducting heat losses between two chambers. 18 vessels of which 9 of 7 kg capacity (big vessels having depth and dia. of 10 cm and 30 cm each) and other 9 of 6 kg capacity (smaller vessels having depth and dia. of 9 cm and 29 cm resp. each) are placed in this cooker. Two different sizes are selected as rice is to be cooked as major food item, thus larger vessels are required for cooking rice as compared to channa daal vessels which are of smaller size. Thus approximately food for 100 people can be cooked easily in this cooker. Moreover, twin-chamber design provides us the additional ability to cook both types of meals (one requiring less time for cooking and other requiring more cooking time) simultaneously without disturbing stagnation temperature. For example, if rice and channa daal are to be cooked simultaneously, rice should be placed in chamber-1 and channa daal in chamber-2. Sun will be on the side of chamber-2 for a longer period of time in afternoon, that’s why this placement of meals is recommended.

Chamber-2 Chamber-1

Fig.1 Twin-chamber design of community solar cooker with vessels

METHODS The designed TCC solar cooker is tested as per BIS standards to calculate first figure of merit (F1), second figure of merit (F2) and standard boiling time. To calculate first figure of merit, no load test is conducted in which cooker is placed in stationery mode facing south without using booster mirrors. Heating of cooker is done until steady state conditions are achieved. For calculation of second figure of merit, sensible heating of water is done at load of 8 kg water per m2, with each vessel filled with equal amount of 49


water. It is then boiled without using booster mirrors. Using first figure of merit and second figure of merit, standard boiling time is calculated.

ANALYSIS AND DISCUSSIONS First figure of merit (F1) is calculated using parameters such as absorber plate temperature (Tp), ambient air temperature (Ta) and incident solar radiation (G) during steady state. F1 is given as:

F1 

Tp  Ta G

… (1)

Second figure of merit (F2) is calculated using standard load of water (8 litre/m2), difference in time taken (t2 – t1) for rise of temperature of water between 60oC (tw1) to 90oC ( tw2 ) and is given as:

  Tw1  Ta  1   F 1( MC ) w   F 1G  F2  ln A(t 2  t1)   Tw 2  Ta  1   F 1G    

… (2)

Standard boiling time is calculated as:

tboil 

 F 1( MC ) w  X  ln1   60 F 2 A  F

…(3)

Where X is (100 – ta)/G. Testing of TCC solar cooker for calculation of F1 and F2 is done on 8 and 10 May 2015 resp. Average value of incident solar radiation during steady state for F1 and during temperature rise for F2 was recorded to be 743 and 747 W/m2 resp. First figure of merit (F1) and second figure of merit (F2) are 0.128 and 0.41 resp. It indicates that our designed cooker is A grade solar cooker according to BIS standards. It is thus recommended for extreme cold climatic conditions. The standard boiling time is calculated as 187 and 195 minutes for chamber 1 and chamber 2 resp. Variation of ambient and plate temperature (fig. 2) and water in vessels of each chamber (fig. 3) is plotted as:

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Temp (oC)

Proceedings of ICETiME’15 December 16-17, 2015, CUSAT, Kochi, India. 140 120 100 80 60 40 20 0

Tp1 Tp2 Ta

time

Temp (oC)

Fig.2 Variation of ambient and plate temperature on 8 May, 2015

120 100 80 60 40 20 0

Twc1 Twc2 Ta

time

Fig.3 Variation of ambient and water temperature on 10 May, 2015 where Tp1, Tp2 are absorber plate temp of chamber 1 and 2 resp, Ta is ambient air temperature, Twc1 and Twc2 are temp of water in vessel of chamber 1 and 2 resp. ENVIRONMENTAL IMPACT ANALYSIS AND FUEL SAVING In northern Indian climatic conditions, approximately 280 clear sky days are observed in a year. Thus, a total of 28,000 meals can be prepared with our designed TCC solar cooker in a year. Assuming a community of 20 families with 5 members in each family, using LPG for cooking meal 3 times a day, one cylinder lasts for approximately 22-26 days (average of 24 days) (or 360 meals) for each family. Thus, if they use community solar cooker, it will save an equivalent of 77 cylinders in a year which approximates to 1100 kg LPG. Taking the cost of LPG cylinder at subsidized rate (i.e. Rs. 450), 77 cylinders will be equivalent to Rs. 34,650. It implies that cost of our designed TCC solar cooker (Rs. 50000) will be recovered within 18 months. Moreover, energy equivalent of 14.2 kg LPG (CV is 46.4MJ/kg) cylinder is 658.88 MJ. Thus, a whopping 50,733 MJ of 51


non renewable energy can be saved by using the proposed community solar cooker. As burning of 1kg of LPG produces 3kg of CO2 during complete combustion and by saving one tonne of CO2 emission in atmosphere, 1 Carbon Credit (CC) is earned as per Kyoto protocol, therefore by saving 1100 kg of LPG per year, 3.3 tonnes of CO2 emission into atmosphere can be prevented by using the green energy. Hence, in terms of carbon crediting, use of our designed community solar cooker for 280 days in one year will earn 3.3 carbon credits for the community.

CONCLUSIONS On the basis of this study, following important conclusions ca be derived: 1. First figure of merit (F1) and second figure of merit (F2) of the developed TCC solar cooker are 0.128 and 0.41 respectively which shows that our designed cooker is ‘A’ category cooker as per BIS standards. 2. As standard boiling time of TCC solar cooker is 187 minutes which shows that this cooker can cook at least one meal per day. 3. Experimental impact analysis and fuel savings shows that TCC solar cooker’s reduction of greenhouse gas emission is 3300 kg CO2/ year on basis of just 280 days of use. This reduction in greenhouse gas emission is equivalent of 3.3 carbon credits earned for community. 4. Our designed TCC solar cooker can save 1100 kg LPG and can recover its cost in just less than 18 months.

REFERENCES Jubran B A and Alsaad MA (1991), “Parameteric study of box-type solar cooker”, Energy Conversion and Management 32, 223–34. Mullick S C, Khandpal T C and Saxena A K (1987), “Thermal test procedures for box type solar cooker”, Solar Energy, 39, 353–60. Negi B S and Purohit I (2005), “Experimental investigation of box type solar cooker employing a non-tracking concentrator”, Energy Conversion and Management, 46, 577– 604. Sethi V P, Pal D S and Sumathy K (2014), “Performance evaluation and solar radiation capture of optimally inclined box type solar cooker with parallelepiped cooking vessel design”, Energy Conversion and Management, 81, 231–41. Yadav Y S and Tiwari G N(1987), “Transient analytical study of box type solar cooker”, Energy Conversion and Management, 27, 121–25.

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NUMERICAL ANALYSIS OF GEOMETRIC EFFECTS ON ALTERNATE WEDGE STRUT INDUCED MIXING IN A SCRAMJET COMBUSTOR Aswith R. Shenoy1, Tide P. S2 1

Assistant Professor Universal College of Engineering Vallivattom, Thrissur, Kerala, India. 2 Professor Cochin University of Science and Technology Kochi- 682 022, Kerala, India.

ABSTRACT The desire for faster response and cheap access to space continues to push the envelope in terms of altitude and airspeed. The advancement in space vehicles and high speed flights critically depends on the development of engines capable of delivering thrust to attain wide range of Mach numbers. As Turbofan and Ramjet engines are unsuitable for high supersonic and hypersonic Mach numbers, the only alternative is the Scramjets in which the flow through the combustor remains supersonic. Due to the very short residence time, an efficient and rapid mixing of fuel/air is hard to achieve. In supersonic flows a rapid fuel/air mixing, additionally suffers from inherently low mixing rates due to compressibility effects at high convective Mach numbers. Struts are used in Scramjet combustors for fuel injection, fuel air mixing, flame holding and ignition. Vorticity is the main driving mechanism for rapid near-field mixing. Strength and size of the vortices depends on the strut geometry. The increase in height/width of the struts from the baseline value exhibits an increase in the total pressure loss across the combustion section. This total pressure loss is correlated with the low pressure region created by the flow displacement caused by the strut. Testing the configuration in the actual size demands huge infrastructure development. Hence a scaled down model was designed for the cold flow analysis using CFD, by varying the wedge angles. The numerical predictions clearly show that mixing is enhanced with increase in strut 1


Numerical Analysis of Geometric Effects on Alternate Wedge Strut…

wedge angle up to 14o on analyzing the velocity contours. Other mixing parameters such as momentum flux, degree of mixing and pressure drop factors were evaluated for determining the optimum operating condition.

INTRODUCTION In the past two-three decades many researchers were in the process of developing the scramjet engines for hypersonic flights, with limited implementation. The major challenge was in developing supersonic combustors successfully. The length of a typical supersonic combustor is around 1.5 m for a Mach 5 cruise vehicle, which results in a flow residence time of 1millisecond. With this short time the fuel has to be mixed thoroughly with the air and burnt. Moreover, for reducing the skin friction drag and the weight of the engine, the combustor length has to be kept small, which leads to further reduction in the flow residence time. Additionally, self-ignition and stable flame holding devices are essential. Scramjet (Supersonic Combustion Ramjet) engine is a variant of a ramjet air breathing combustion jet engine in which the combustion process takes place in supersonic airflow. As in a Ramjet, a Scramjet relies on high vehicle speed to forcefully compress and decelerate the incoming air before combustion. Airflow in a Scramjet is supersonic throughout the entire engine. This allows the Scramjet to efficiently operate at extremely high speeds. Theoretically, Scramjets provide superior specific impulse performance over a range of hypersonic Mach numbers compared to rockets as they do not carry the oxidizer internally. A Scramjet's lower operating range starts around Mach number 4 - 5. The practical upper limit expected for hydrogen-fueled Scramjets is around Mach numbers 12 - 16, and the practical upper limit expected for hydrocarbon-fueled Scramjets is around Mach numbers 9 10. The Scramjet is composed of three basic components: a converging inlet, where incoming air is compressed and decelerated; a combustor, where gaseous fuel is burned with atmospheric oxygen to produce heat; and a diverging nozzle, where the heated air is accelerated to produce thrust. The inlet section compresses the hypersonic free stream flow and lowers the Mach number by about 40%. The combustion section injects, mixes, and ignites the fuel in a supersonic flow, and the nozzle section accelerates the flow back to hypersonic speeds. Supersonic combustion is the most challenging part of successful Scramjet operation, and fuel injection strategies emerge as a primary field of research in supersonic combustion. While Scramjets are conceptually simple, actual implementation is limited by extreme technical challenges. Hypersonic flight within the atmosphere generates immense drag, and temperature shoots to nearly six-times that of the surrounding air. Maintaining combustion in the supersonic flow presents additional challenges, as the fuel must be injected, mixed, ignited, and burned within milliseconds.

54


Mixing, ignition and flame holding in the combustor of a ground test facility are the critical challenges in the development of scramjet engine. Enhancing the mixing, and thus reducing the combustor length, is an important aspect in designing Scramjet engines. The complex phenomenon of supersonic combustion involves turbulent mixing, shock interaction and heat release in supersonic flow. The flow field within the combustor of Scramjet engine is very complex and poses a considerable challenge in design and development of a supersonic combustor with an optimized geometry. Such combustor shall promote sufficient mixing of the fuel and air so that the desired chemical reaction and thus heat release can occur within the residence time of the fuel-air mixture. In order to accomplish this task, it requires a clear understanding of fuel injection processes and thorough knowledge of the processes governing supersonic mixing and combustion as well as the factors, which affects the losses within the combustor. Due to the limited mixing capabilities of parallel high speed streams, techniques for mixing enhancement are required. This can be achieved either by the use of shock waves or by creation of stream wise vorticity. Various methods for mixing enhancement are broadly classified under two categories - active and passive methods. Active method includes inducing turbulence, shock interactions, swirls etc. into the flow by active components like cavities, struts etc. Passive method includes changing the initial condition of the jet by changing nozzle geometry.

A Strut is defined as an in-stream geometric structure that spans the entire width or height of the combustor and attached to two walls. Struts are used in Scramjet combustor for fuel injection, fuel air mixing, flame holding and ignition. Strut mixing devices covers a wide range of designs and includes both normal and parallel injection methodologies. A vertical strut with a wedge leading edge is seen in most of the strut configurations. The strut is connected to both the bottom and top of the combustion section. Since it is across the whole combustion section, fuel injection occurs at several locations and allows the fuel to be added throughout the flow field. Many researchers looked at modifying the trailing edge of the vertical strut to increase mixing. In the basic strut design, the strut was connected to the top and bottom of the test section and the leading edge was a wedge. The major difference for the present configuration is the trailing edge design as seen in Fig 1. The different trailing edges, called alternating wedge designs, create either co-rotating or counter-rotating vortices that are used to enhance mixing. All these designs use parallel fuel injection at the trailing edge of the strut so that the fuel is entrained into the vortices which cause the increased mixing in the combustion section. Strut type injector used in the present analysis is based on the design concept introduced by Sunami.T et al. (2002), which is called "Alternating Wedge Strut (AW-Strut.)" Fig 1 show a schematic of Alternating Wedge Strut geometry used in Sunami.T et al. (2002). Wedges were formed by extended compression ramp from the upstream side, and wedges and expansion ramps were arranged alternately in the span width direction. Moreover, the composition of alternating-ramp-wedge formed a cavity at under the compression wedge. 55


Fig. 1 Alternating Wedge strut (Sunami.T et al. (2002)

Strut type hypermixer tested in the present experiments is based on the design concept introduced by Sunami et al. which is called "Alternating Wedge Strut (AWStrut.) Peter Gerlinger et al. (2008) studied a new concept for a three Lobed Strut injector with holes and compared with a three lobed Strut Injector with slots concept. Srikrishnan and Job Kurien (1996) performed experimental study on mixing enhancement in supersonic flow using a radially lobed (petal) nozzle. Desikan and Job Kurien (2006), studied the flow field of injectant plume produced by strut based injectors by Schlieren photography and Mie scattering technique. They found out that ramp strut performed better to plain strut to achieve better mixing due to formation of weak oblique shocks middle region which creates a subsonic region, which aids flame holding in case of combustion. Sujith et al. (2013) conducted experimental and numerical studies on supersonic mixing of air injected from struts of various trailing ramp angles with free stream air at Mach number 1.63 by providing an alternate trailing ramp on the strut. Hsu et al. (2009) analyzed effects of wedge angle, strut root length, and shape of strut base on fuel distributions. Samitha et al. (2007) studied the mixing performance in clover nozzle in which they found to have higher mixing performance when compared to conical nozzle. Also pressure drop factor is lower than petal nozzles. Momentum flux distribution, degree of mixing and pressure drop factor was taken from the mixing tube for analyzing the results. In this work, the comprehensive numerical analysis of non-reacting supersonic flow in a Scramjet with an alternate wedge strut injector was carried out.

2. DESCRIPTION OF TEST NOZZLE

56


The cold flow mixing studies was conducted on an alternating wedge type strut in a supersonic air flow. The boundary conditions and the input parameters were decided based on the available resources for experimental validation. 2.1 Primary and Secondary Flow rates: Primary Air flow rate of 1.6 kg/ sec and a secondary flow rate of 0.0263 kg/sec of GHe -Gaseous helium (which is equivalent to 0.0372 kg/sec of GH2 -Gaseous hydrogen for an equivalence ratio 0.8) was taken as input for analysis. In actual rockets the fuel used is GH2 but for the ease of conducting the cold flow test in future , GHe is used in present analysis. Air flow rate was decided based on the feasibility for experimental validation with available resources. 2.2 Supersonic Nozzle: A rectangular supersonic nozzle was designed to accelerate the air at constant pressure of 10 bar to Mach 2, considering the placement of an intrusive strut in the flow field. The throat and exit areas were calculated from the conventional gas dynamic equations and the contour is finalized using Method of Characteristics (MOC). 2.3 Strut geometry: The configuration of the strut is shown in Fig 2. The geometry of the strut up to the wedge is same for all the models of study. The wedge angle is varied for different configurations for analysis (viz. 0, 5, 7, 9, 11 and 14 degree). The secondary injection GHe is parallel to the wedge surface. There are 7 numbers of injection holes for GHe each of 1.5 mm diameter.

Fig. 2 Three dimensional view of Test Section with Strut

57


3. NUMERICAL APPROACH The numerical analysis was carried out using a commercial CFD package. The simulations were carried out on a 3D species reacting steady state RANS solver. The turbulence model used was SST k-ω (a two equation model) for its improved performance for flows involving boundary layers subjected to adverse pressure gradients. The conductivity and viscosity is defined by Sutherland’s law. The CFD code has inherent advantage of adaptive mesh refinement that allows the local modification of the mesh during the computation. This option provides a better resolution of the shock waves without significant increase in the number of cells. Flow domain including the supersonic nozzle and injector was subjected to flow analysis to validate the rectangular nozzle configuration. Struts with varying wedge angles (0, 5, 7, 9, 11 and 14 degrees) with secondary injection holes were modeled for analysis. Adiabatic, non-reactive 3D flow analysis was carried out for GHe injection using alternate wedge strut into a Supersonic Air flow at Mach 2. Ambient air at 1 MPa is supplied at the inlet of supersonic nozzle at 1.6 kg/s. GHe flow rate of 0.372 kg/sec is injected into the flow stream through 7 holes of 1.5 mm diameter (Simulation equivalence ratio  =0.8 for Air & GH2). GHe is injected at angle parallel to the strut wedge surface. The size of the grid was varied and the results were compared after CFD simulation for grid independence study. The grid independent studies were conducted on meshes with 2×105, 3×105, 6×105 and 8×105 cells. The results showed that the increase of grid size beyond 6×105 does not yield any significant variation in the results. The study also showed that 2×105 grids was quite insufficient to capture the flow in a refined manner and 3×105 mesh was having slight deviations from the predictions of 6×105 (Fig 3). Hence the grid size was finalized as 6×105. Both structured and unstructured grids were used (602228 hexahedral grids). The nozzle has about 640 and 175 cells in stream wise direction and normal direction respectively, with a maximum wall y+ value of 1.2 and average y+ value of 0.97..

58


Fig. 3 Meshed model used for analysis

4. RESULTS AND DISCUSSIONS Mach contour distribution for plane strut without and with secondary injection is shown in Fig. 4. Flow analysis result for plane strut shows Mach 1.96 at the beginning of the test section, which validates the nozzle contour design for the supersonic rectangular nozzle using MOC.

Fig. 4 Mach contour for plane (00) strut without and with secondary injection A subsonic region prevails behind the strut. Flow pattern is almost same with and without secondary injection at strut tail edge. Both primary fluid (air) and secondary fluid (GHe) flow in separate stream within the test section at an axial distance of 'H' which is equal to the height of the test section (Fig. 5). An increase in Mach number is observed at the strut tail edge, where there is a sudden increase in flow area. which is equal to the height of the test section (Fig. 5). An increase in Mach number is observed at the strut tail edge, where there is a sudden increase in flow area.

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Fig. 5 Mach contour for plane (00) strut axial normal plane at Secondary injection and at exit Mixing of primary and secondary fluids with in the axial distance of ‘H’ is not seen for 5° strut. Changes in mixing pattern are seen for 7° and 9° struts (Fig. 6 and Fig. 7).

Fig. 6 Mach contour for 7° strut (Left) axial parallel plane through side hole; (Right)Vertical axial plane through centre hole

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Fig. 7 Mach contour for 9° strut (Left) axial parallel plane through side hole; (Right)Vertical axial plane through centre hole

Fig. 8 shows enhancement in mixing with in the test section for 11° strut. Velocity vector plots show recirculation just after the strut tail where pilot flame is introduced in actual Supersonic combustors. The recirculation caused by vortices is a desirable feature for flame holding.

Fig. 8 Mach contour for 11° strut (Left) axial normal plane through side hole; (Right)Vertical axial plane through centre hole Mach number distribution for 14° strut through two vertical axial planes are shown in Fig. 9. Gross changes in mixing pattern were seen with formation of weak oblique shocks. A subsonic region exists just after the secondary injection ports, which is desirable for flame holding.

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Fig. 9 Mach contour for 14° strut (Left) axial parallel plane through side hole; (Right)Vertical axial plane through centre hole Fig. 10 shows the different normal planes where Mach number at axial distances 0.15H, 0.4H, 0.7H and H .Static pressure values are taken at these locations for evaluating the mixing effectiveness. The Mach number and static pressure values of 17 points in the cross section at different axial distances at 0.15, 0.4, 0.7, 0.8, and 1H from secondary injection port are tabulated.

Fig. 10 Mach profiles at axial distances 0.15H, 0.4H, 0.7H and H for 14° strut In order to compare the mixing performance for different strut configurations, based on a quantitative assessment of level of mixing achieved, a dimensionless parameter called uniformity factor Ф is calculated.

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4.1 Uniformity factor ф Uniformity factor Ф is defined as

Ф =1−[у μ (x)/μ av(x)] where, σ μ (x) is the standard deviation of radial distribution of momentum flux at a given axial location along the mixing tube and μav(x) is the average of momentum flux along a radial line at the location considered.

4.1.1 Momentum Flux Primary air and secondary fuel enters the chamber at different momentum and stagnation pressures. Momentum flux distribution at the exit of supersonic combustor in radial direction is the measure of bulk mixing. Momentum Flux µ is calculated as

µ = p (1 + γ M2) where p is the static pressure and M is the Mach number. Mach number can be calculated from the measured values of stagnation pressure. The momentum flux at which uniformity is attained indicates the axial distance where mixing is complete Uniformity factor Ф is a measure of the uniformity of the momentum flux distribution in the radial direction at a given location. For a perfectly mixed flow, the distribution has to be uniform across the section. The Uniformity Factor ф of 00, 70, 90, 110 and 140 struts are tabulated in Table 1. Table 1: Uniformity Factor

Uniformity factor ф

Position of axial Plane

140

110

90

70

0.15H

0.725

0.197

0.105

0.088

0.4H

0.734

0.775

0.802

0.829

0.7H

0.745

0.687

0.628

0.569

H

0.753

0.718

0.668

0.618

From the Table 1 it is clear that 140 wedge strut has got the better momentum flux distribution as the value uniformity factor is more or less uniform. For 14° strut, a Uniformity factor of 0.753 is obtained at H and 0.725 at 0.15 H. Corresponding

63


values for 11° struts at H and 0.15 H are 0.718 and 0.0197 respectively. The weak oblique shock formed in the case of 14° is the cause for higher uniformity factor.

CONCLUSIONS A parametric study of influencing parameters of alternate wedge strut for supersonic combustor applications has been carried out. GHe at a flow rate 0.0263 kg/s is injected through a strut into stream of air at Mach 2 and flow rate of 1.6 kg/s. A quantitative assessment of mixing effectiveness is obtained by calculating uniformity factor and it is observed that mixing of air with fuel using alternate wedge strut improves with the increase in strut angle. Weak oblique shocks are formed for 14° strut angle. It may be inferred that the formation of these shocks gives better mixing compared to other lower angles which aid the combustion process

REFERENCES A.R. Sreekrishnan and Job Kurien, (1996) "Experimental Study on Mixing Enhancement by Petal Nozzle in Supersonic Flow", 1996, Journal of Propulsion and Power, Vol. 12, No. 1, pp. 165-169. Kuang-Yu Hsu, Campbell Carter, Mark Gruber, Chung-Jen.Tam, (2009) ”Mixing Study of Strut Injectors in Supersonic Flows” 45th AIAA/ASME/SAE/ASEE Joint propulsion conference exhibit AIAA-2009-5226. Peter Gerlinger, Peter Stoll, Markus Kindler, Fernando Schneider and Manfred Aigner, (2008) "Numerical Investigation of mixing and Comnustion Enhancement in Supersonic Combustors by Strut Induced Streamwise Vorticity", Journal of Aerospace Science and Technology, Vol. 12, Issue 2, pp. 159–168. S.L.N. Desikan and Job Kurien, (2006), “Strut based Gaseous injection into a Supersonic stream”, Journal of Propulsion and Power, Vol. 22, No. 2, pp. 474-477. S. Sujith, T. M. Muruganandam, and Job Kurian, (2013) "Effect of Trailing Ramp Angles in Strut-Based”. Journal of Propulsion and Power, Vol. 29, No. 1, pp. 6678. Sunami. T, Atsu Murakami, Kenji Kudo, Kodera M, Michio Nishioka, (2002)“Mixing and Combustion Control Strategies for Efficient Scramjet Operation in Wide Range of Flight Mach Number”. AIAA/AAAF 11TH International Space Planes and Hypersonic Systems And Technologies Conference AIAA.20025116 64

Proceedings of ICETiME’15 December 16-17, 2015, CUSAT, Kochi, India. Z A Samitha, B Swaraj Kumar, P Balachandran, (2007) “Experimental Study on supersonic mixing using clover nozzle” 45th AIAA Aerospace science meeting and exhibit AIAA-2007839.

Kodur, V.K.R.(2003), “Fire Resistance Research Needs for High Performing Materials”, NRCC-45405, Making the Nation Safe from Fire- A Path forward in Fire Research, NRC, National Academy of Sciences, Washington D.C., 1-7. Sanjayan, G. (1993), “Spalling of High Strength Silica Fume Concrete in Fire”, ACI Materials Journal, 90(2), 170-173.

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66


MODELLING OF INDIRECT NATURAL CONVECTION ON A TWO DIMENSIONAL DOMAIN Nidhin Roy V1, Joshy P J2 1

Assistant Professor,2Associate Professor K V M College of Engineering and IT Cherthala 2 Cochin University of Science and Technology Kerala, India.

1

ABSTRACT Two dimensional convection on a fluid with Boussinesq apprroximation is an impotant area of study with lots of practical applications. In this study we are trying to create a numerical model for the flow. Indirect natural convection over a horizontal surface is modelled with appropriate boundary conditions. We are focussing on the nearwall region. The governing equations for a two dimensional Rayleihgh-Benard convection is solved in a 2D domain. The results obtained from the model are useful for studying about boundary layer region as well as plume region in a natural convection phenomena. Instabilities of boundary layer region and formation of plume can be visualized. Variations in the flow with respect to changes in properties such as Rayleigh number and Prandtl number can be visualized. For each time step development of boundary layers can be observed from velocity contour and temperature contour obtained from the numerical code. Keywords: Rayleigh – Benard convection, Boundary layer, Near wall region, Plumes .

INTRODUCTION Thermal energy from a hot surface is transferred through an initially inert fluid in two ways conduction and convection. Conduction heat transfer occurs near to the wall and convection is within the bulk fluid. When the heat flux exceeds a threshold value boundary layer instability is created and plumes are generated. 1


Modeling of indirect natural convection……

Indirect natural convection comes in an indirect manner by way of an induced pressure gradient. In a fluid whose density decreases with increasing temperature, a boundary layer flow forms at horizontal hot plate. The temperature is everywhere T∞, so that, as in the static field, there is a pressure distribution with the gradient (∂p/∂y) = ρgg. The temperature in the boundary layer region is larger than T∞ and so the density is lower than ρ∞. The decreased pressure gradient leads to a reduced pressure in the boundary layer region. Hence there is a pressure drop in the horizontal direction. This pressure gradient is the origin of the flow parallel to the plate. This paper mainly deals with the near wall dynamics of the flow. The region near to the wall has great importance in the heat transfer. The sheet plumes formed near the heated wall will transfer a large amount of heat from the surface as per the findings of Theerthan (1998).Natural convection in three dimensional rectangular enclosures had been analyzed numerically by Sik Lee et al. (1989) the effect of the Rayleigh number was mainly investigated. Sabeur et al. (2007) performed a numerical investigation of the influence of the hot surface geometry on a laminar natural convection in a cubical cavity filled with air differentially heated. The results obtained showed that the hot wall geometry affects the flow and the heat transfer rate in the cavity. Fusegi et al. (1991) studied numerically a transient three dimensional natural convection in a differentially heated cubical enclosure at Rayleigh number of 106. Bocu et al. (2011) studied numerically laminar natural convection heat transfer in 3D rectangular air filled enclosures, with pins attached to the active wall. The Rayleigh numbers considered in their study ranges from 105 to 107.The results presented revealed a good agreement not only in 2D also in 3D and for a wide range of Ra numbers. Numerical model for this phenomenon is cheaper than an experimental model. To get reasonably accurate model suitable computational methods has to be used. We are trying to write a program in MATLAB to solve the governing equation.

GOVERNING EQUATIONS Two dimensional governing equations for the flow are continuity equation, momentum equation and energy equation. Flow is assumed to be steady, viscous and incompressible with Boussinesq approximation, which implies that except for the density in the gravitational term all other properties in the governing equations are kept constant and there is no source or sink in the system. Continuity Equation

+

=0

……… (1) 68


Momentum equations +

+ Energy equation +

+

= −

+

= −

+

=

+

+

+

+

………. (2) +

+

…… (3)

…….. (4)

Here ‘Ra’ represents Rayleigh number; Peclet number is represented by ‘Pe’ and ‘Pr’ represents the prandtl number. These are the three numbers governing the flow in which Ra and Pe are flow properties and Pr is a fluid property. = = Pr =

ℎ

Boundary conditions for the rectangular 2D domain are, U=0, V=0, T = 1 U=0,

= 0, T = 0

U=0, V=0,

= 0,

for bottom isothermal wall, for top isothermal wall, for adiabatic side walls.

NUMERICAL METHOD AND MODEL VALIDATION

69


Fig 1: Staggered control volume A 2D staggered control volume is used in the computational domain. Scalar variables like pressure temperature are kept at the centre of the control volume and velocity vectors are placed at the faces of the control volume. The indices ‘i’ and ‘j’ are used to denote the nodal points in x and y directions respectively. An enlarged view of the control volume for pressure and temperature is given in the fig.2 and the control volume for velocity terms are staggered from the given figure.

Fig 2: Enlarged View of the control volume Control volumes for scalars and vectors are different because of the pressure gradient terms in the momentum equations. If we consider scalars and vectors at the same nodal points it will produce no source of momentum even when we consider checker board pressure distribution. This is against the physics of actual problem. In order to avoid this problem we are considering a staggered grid arrangement. Since velocity terms are arranged at the faces of control volume, there is no need to conduct an interpolation for calculating the flow variables at the control volume surfaces. So staggered grid arrangement offers one more advantage. The general form of a convection diffusion equation can be represented as, (

)

+

(

)

=

+ 70

………….. (4)


Where φ is the general variable, ϒ is the diffusion co-efficient and S is the source term. All governing equations continuity, momentum and energy equations come under this category. After discretization the general form of a 2D convection diffusion equation can be written as per the theory of Patankar (1980), = Where,

+

+

(| |) + ⟦− , 0⟧

=

(| |) + ⟦

=

=

∆ ∆

=

+

+

…. (5)

, 0⟧

(| |) + ⟦− , 0⟧

=

=

+

+

(| |) + ⟦ , 0⟧ +

+

−

∆ ∆y

Values at previous time step. = ∆ ∆ + aE, aW, aN, aS etc represents the neighbouring co-efficient and the corresponding φ values represents the nodal values. D is the conductance term and F is the flow rate. ∆ = ( ) This is for N and S nodes. =

(

∆

)

∆

This is for E and W nodes

F = ρu∆y For E and W nodes F = ρu∆x For N and S nodes. A (P) = 1 – 0.5 |P| P is Peclet number. Second order finite difference method is used for discretization. Implicit time march is used for both convective and viscous terms. Pressure Poisson equation is solved in a projection method to enforce incompressibility using sparse matrices. The domain size is fixed as L=15 and H=5. Domain is suitably selected to capture the various flow structures. A suitable grid sizing of 300x200 is used for calculations. Initially all matrices are equated to zero. Water is assumed to be the fluid in the domain so ‘Pr’ number is set to 7. For a set of Ra numbers ranging from 10-2000, variations in the flow can be visualized. We can observe the flow variations in the domain with respect to the change in time. Time step Δt is suitably selected as 0.01. LU decomposition method is used for the solution of each matrix instead of conventional TDMA method. Results are obtained in the form of Temperature and Velocity matrices and contour plots are obtained from these matrices.

RESULTS AND DISCUSSIONS 71


Results obtained from the code are classified into two categories. First one is the variations in the flow with time for a constant Rayleigh number and the other one is for different Rayleigh numbers at constant time. We need to analyse the development in flow with respect to the change in time. Effects of different Rayleigh numbers on the boundary layer instability can be found out with the help of temperature contours. FLOW VARIATION WITH TIME

Following plots are obtained for a Ra number equal to 2000 in different time steps. For each time step changes in boundary layer region can be clearly visualized.

(A)

(B)

(C)

(D)

(E) (F) FIG 3. Flow development with change in time. A,B,C,D ,E and F represents contour plots at time steps 0.5,1,2,5,10 and 15 respectively.

Rayleigh-Benard convection is an unsteady phenomenon but for specified combinations of suitable flow parameters and domain sizes steady convergence is possible. From this numerical model we can observe the development of convective flow over the horizontal surface. During the initial stage of convection a narrow boundary layer is formed near the bottom wall. At time t= 0.5 we could observe some instabilities in the boundary layer. Then the boundary layer breaks and vertical column of fluid flows through the surrounding fluid. This is known as thermal plumes. As time increases plumes become erect and steady. As the value of Ra increases flow instability is large. For higher Ra numbers boundary layer breaks at the earlier stages of convection as well as time required for steadiness is also large. 72


For steady flow cases plumes are evenly spaced with mean plume spacing. Pressure drop in the boundary layer region is the reason for horizontal flow. Flow become steady after t=10, and the temperature contour shows negligible variation with time. Within the plume region horizontal component of velocity is zero. Flow domain can be visualized as three distinct regions such as boundary layer zone, plume region and the surrounding fluid.

FLOW DOMAIN FOR DIFFERENT RAYLEIGH NUMBERS Rayleigh number is an important flow property which governs the flow. It is the ratio between buoyancy forces and viscous forces. Change in Ra number will affect the initialization convection and the breaking instance of boundary layers.

Ra = 10

Ra = 100

Ra = 500

Ra = 1000

Ra = 2000

FIG 5.Flow modelling for different Ra numbers Steady flow indirect natural convection is a pressure driven flow unlike the direct effect of buoyancy in direct natural convection. Ra number is the governing parameter of the flow. As the value of Ra increases buoyancy effects dominates the viscous effects. So that convection becomes dominant for high Ra numbers. For a fixed time step we are extracting the temperature contours of the two dimensional domain. For Ra = 10 and 100, only conduction layer is developed. 73


When Ra = 500 instabilities are formed in the boundary layer region and it breaks and for further increase in Rayleigh number plume development is more distinct. At Ra = 1000 steady two dimensional sheet plumes are obtained. When Ra number increases from 1000 to 2000 number of plumes increases from two three. So the numerical model can easily capture the significance of Ra number variation in natural convection process. As Ra number increases spots of instability also increases . One of the interesting results is that, as Ra increases thickness of boundary layer decreases. Due to an increase in Ra number buoyancy effects dominates and boundary layer breaks earlier to initiate the vertical flow. This reduces the thickness of boundary layer. So we can say that Ra number and boundary layer thickness is inversely proportional. When Ra number exceeds a certain limit flow becomes highly unstable and steadiness cannot be achieved.

CONCLUSION Indirect natural convection is a pressure driven flow process having several practical importance. A numerical model for the 2D domain is proposed which solves the viscous incompressible governing equations with appropriate boundary conditions. Numerical code is useful for,  Modelling the flow for different combinations of Ra and Pr numbers.  Studying the effects on boundary layer with change in time.  Boundary layer instability and plumes can be visualized. The proposed numerical model and computational procedure is helpful for getting reasonably good results about indirect natural convection. Modifications in the numerical code can give better results and can be used for versatile studies about this phenomenon. Obtained results are closely related to the previous experimental results in this field. Proposed code has to be improved for predicting the heat transfer during convection. So that suitable model can be correlated with actual cases. This model can be extended for the studies of cooling of electronic components, Room temperature control and ventilation etc. with reasonable accuracy.

REFERENCES Ananda Theerthan and Jaywant H Arakeri.(1998), “A model for near-wall dynamics in turbulent Rayleigh Benard convection”, J. Fluid Mech, vol. 373, pp. 221-254. A.Sabeur-Bendehina, O.imine, L.Adjlout, B.Imine.(2007), Effect of the hot surface geometry on laminar Natural Convection in Cubical Air Filled Enclosures, Third International Conference on Thermal Engineering: Theory and Applications. 74


Suhas V Patankar, (1980) “Numerical Heat Transfer and Fluid Flow”, McGrawHill Book Company. T. Fusegi, J. M. Hyun, K. Kuwahara, (1991), “Transient three-dimensional natural convection in a differentially heated cubical enclosure”, Int. J. Heat Mass Transfer. Vol. 34. No. 6, pp. 1559-1564. T.Sik Lee, G. Hun Son, J.Sik Lee, (1989), “Numerical study on natural convection in three-dimensional rectangular enclosures”, KSME Journal, Vol. 3, No. 1, pp. 5055. Z.Bocu, Z.Altac, (2011), “Laminar natural convection heat transfer and air flow in three-dimensional rectangular enclosures with pin arrays attached to hot wall”, Applied Thermal Engineering vol. 31, pp 3189-3195.

75



76


NUMERICAL INVESTIGATION ON PULSATING JET IMPINGEMENT COOLING Sabu Kurian 1, Vipin R 2, Tide P. S 3 & Biju N 4 1

Research Scholar, 3Professor, 4Associate Professor, Mechanical Engineering Division, School of Engineering, Cochin University of Science and Technology, Cochin, Kerala, India 2 Post-Graduate Student, M. A. College of Engineering, Kothamangalam

ABSTRACT The aim of this work is to present the results of a numerical investigation of the effect of pulsating frequencies on the local and average heat transfer characteristic of an impinging air jet. The calculations were done using a commercial software. Temperature plots obtained from the numerical analysis was used to calculate the average and local Nusselt Number for different pulsating frequencies at a Reynolds number of 5000. The pulsating frequencies were between 5Hz - 150 Hz. Results obtained show that the local Nusselt number calculated were higher at all radial position away from the stagnation point. The pulsed jet Nusselt number was higher than the average steady jet Nusselt number for all values of frequencies due to the higher localized heat transfer.

INTRODUCTION Impingement heat transfer is considered as a promising heat transfer enhancement technique. Among all convection heat transfer enhancement 1


Numerical investigation on pulsating jet impingement cooling

methods, it provides significantly high local heat transfer coefficient. Jet impingement produces a rapid cooling or heating on the surface where it impinges. Study on the effect of pulsating frequencies on the impingement heat transfer has been focused by many researchers in the past. Pulsating flow is widely believed to increase the heat transfer rate. Zumbrunnen and Aziz (1993) studied the importance of pulsation frequency and amplitude. They investigated convective heat transfer with a planar impinging water jet on a surface subjected to a constant heat flux. A twofold enhancement of heat transfer was reported. Sheriff and Zumbrunnen (1998) expanded the study to pulsating array of jets. An array of nine convergent jets in a square matrix was used to cover a Reynolds number range of 2500–10000. Their study focused on jet height to diameter ratio of 2–6, pulse frequency up to 65 Hz, Strouhal number of 0.028 and flow pulsation magnitude up to 60%. It was observed that the increased jet interactions at large magnitude of flow pulsations reduced the jet potential core length by 20%. Also the turbulence intensities were higher by 7–15% than in steady jets. Fisher (2001) conducted a numerical study to investigate the heat transfer to an axisymmetric circular impinging air jet using the k-ϵ turbulence model. Of particular interest is the effect of jet Reynolds number on convective heat transfer at fixed nozzle-plate spacing. The paper showed that the k-ϵ model over predicts the turbulent kinetic energy in the stagnation region and also the maximum away from the stagnation region. The study reveals that the k-ϵ model works best in simple shear flows. Zuckerman and Lior (2006) focused on the applications, physics of the flow, heat transfer phenomena, available empirical correlations and their values. They used numerical simulation techniques and predicted heat transfer for impinging jet devices. Xu et al. (2010) conducted a numerical study for pulsating turbulent slot impinging jet. The jet velocity was varied in an intermittent (on-off) fashion. The effects of the time-mean jet Reynolds number, temperature difference between the jet flow with respect to the impinging surface and nozzle-to-target distance were examined. Parametric studies show that increase of the mean jet Reynolds number, frequency of pulsation and reduced temperature difference enhance the time-averaged local Nusselt number. The on-to-off jet time ratio and nozzleto-plate distance show significant effects on the heat transfer rates, which can be properly adjusted to achieve optimized performance. Rasool et al. (2014) studied the flow structure and heat transfer of air jet normally impinging on a flat plate using numerical and experimental methods.

The purpose of this study is to investigate steady and pulsating single circular jet heat transfer characteristics. The focus of the study is given on

78


the effect of flow pulsation frequencies on the average and stagnation Nusselt number. The study is also trying to find out the possibility of controlling the flow structure in pulsating air jet which leads to enhancement in the heat transfer characteristics. Comparisons between steady and pulsed jet heat transfer were discussed in details together with available data in the literature. In this paper the stagnation point Nusselt number of a pulse and steady jet means the time average value at the impingement point of the jet axis. The local Nusselt number of a pulse jet is the time average at a point on the impingement surface. The local Nusselt number is assumed to be radially symmetrical about the stagnation point. The average Nusselt number of a pulse jet is both a time average and an area average over the impingement surface. The total heat flux is proportional to the average Nusselt number. Interaction of flow structures can be influenced by the mixing within the boundary layer and a marked increase in turbulence intensities has been noted with pulse flows. Recent findings on the enhancement of heat transfer due to pulse air jets have encouraged new research in this subject. Comprehensive data showing the effect of pulse frequency on local and average heat transfer profile are still limited and there is need for further investigation.

NUMERICAL PROCEDURE The pulsated impinging jet heat transfer problem is numerically computed with the commercial finite-volume code using the time-averaged NavierStokes and energy equations with the standard k-ɛ turbulence model. The kɛ model is chosen due to its simplicity, computational economy and wide acceptability. The circular air jet is assumed to have constant thermophysical properties such as density, specific heat and thermal conductivity. Hence, the geometric boundaries and physical conditions are symmetric about the axis of the jet; a 2D axisymmetric model is constructed. It neglects gravitational effect during the impinging jet. The finite -volume code ANSYS 14.0 is used to solve the thermal and flow fields using t he standard turbulence model. Diffusion terms of all the governing equations are discretized using the central difference scheme. Convective terms of the momentum and energy equations are discretized using the third order QUICK interpolation scheme and convective terms of the turbulent kinetic energy and turbulent dissipation rate equations are discretized using a second-order upwind differencing scheme. Pressure-velocity coupling is handled using the SIMPLEC algorithm. The axi-symmetric domain is shown in Fig. 1.

79


Fig.1 Computational Domain Successful computation of the turbulent model requires some consideration during the mesh generation. Since turbulence plays a dominant role in the solution of transport equations, it must be ensured that turbulence quantities are properly resolved. It is therefore proposed to use fine meshes as shown in Fig. 2 to resolve the near-wall region adequately. Computational domain contains around 40000 elements. Edge sizing for jet axis and wall region was set at an appropriate value for relevance. Relevance centre was set as fine and smoothing is high.

. Fig. 2 Domain with mesh

80


Grid independent study The average Nusselt number shows much variation from the accurate value with the grid size. The accuracy of the result increases with the element size and it reaches an optimum value, above which the variation is insignificant (Fig. 3).

Nuavg vs Mesh size

20

Nuavg

15 10 5 0 0

10000

20000

30000

40000

50000

60000

Mesh size

Fig.3 Variation of average Nusselt number with number of elements The average Nusselt number shows much variation from the accurate value with the grid size. The accuracy of the result increases with the element size and it reaches an optimum value, above which the variation is insignificant.

RESULTS AND DISCUSSIONS Numerical results are presented for the heat transfer characteristics of steady and pulsed jet. The local Nusselt number corresponding to the stagnation point and the average Nusselt number are considered for the evaluation of the performance of the jet impingement. The jet impingement heat transfer rates of circular (D=5 mm) un-confined jets were evaluated for Reynolds number of 5000. Effect of pulse jet frequency Variation of average Nusselt number corresponding to different frequencies is shown in Fig. 4.

81


Nuavg Vs f

50

NU avg

40 30

41.32 30.99

34.59

42.063

41.32

35.1659

32.197

20

Nusselt no

10 0 steady

f=5

f=12.5

f=20

f(Hz)

f=50

f=100

f=150

Fig.4 Variation of average Nusselt number for steady and pulsed jet with different frequencies at Re=5000 and Z/D=3

Fig. 4 shows the variation of the average heat transfer coefficients along the impinging surface for different pulse jet frequencies. Here, a jet with square wave form is used as the impinging jet, with different frequencies, for the analysis of heat transfer rate from the impinging plate. The analysis was studied under steady state and then compared with jets of different frequencies. The pulse jet shows the higher heat transfer rate than steady jet at different frequencies. The average heat transfer coefficient increases for the frequency range of 5Hz to 50 Hz. However, the average heat transfer coefficient value is less for 100Hz and 150 Hz when compared with low frequency values such as 5 Hz, 12.5 Hz, 20 Hz and 50 Hz. An enhanced heat transfer coefficient is observed when compared with steady jet. The effect of pulse jet under different frequencies on the stagnation region is shown in Fig. 5. For higher frequency; the Nusselt number in the stagnation region of the pulse jet is even larger than that in the steady case. Stagnation point Nusselt number increases with increase in frequency for the pulsed jet, reaches an optimum value and then decreases. At very high frequencies the pulse jet behaves like as steady jet and thus a reduction in Nusselt number occurs in the stagnation region.

82


NU stag

Nu stag Vs f 90 80 70 60 50 40 30 20 10 0

80.86

75.91

84.8556 84.556

80.86

67.62

59.02

Nusselt no

steady

f=5

f=12.5

f=20

f=50

f=100

f=150

f(Hz)

Fig. 5 Variation of stagnation Nusselt number for steady and pulsed jet with frequency at Re=5000 and Z/D=3 At higher Z/D ratios, the jet becomes shorter and broadens its velocity profile. This reduces the length for which the potential core persists and degrades the mean velocity of the jet at the core. However, the distance is not large enough to allow the increased mixing provided by the pulsating jet to impact the potential core of the jet. Pulsed jets generate large-scale eddy patterns around the exit nozzle, resulting in unsteady boundary layers (thermal& hydrodynamic) on the target that may produce higher heat transfer rate with frequency. The unsteady disturbances of the boundary layer can cause the increased rate of heat transfer rate with frequency. As with a pulsed jet, the variation in local fluid velocity over the target prevents the development of a steady boundary layer. This effect was counteracted by an increased tendency of the precessing jet to mix with the surrounding fluid at higher Z/D ratios, loose energy and reach the target at lower velocities than would be found with a stationary jet.

The effect of pulse jet frequency on the local Nusselt numbers is shown in the Fig. 6. The enhancement of heat transfer by intermittent pulsation in the wall jet region is clear from the figure. Compared with the steady impinging jet of the same mean Reynolds number, the decrease of local Nusselt number along distance from the stagnation region slows down in the wall jet region for the pulsed case. The relatively strong vortex is believed to increase flow entrainment and mixing, and contribute to the predicted heat transfer.

83


Nu local Vs r/D 100 steady

Nu local

80

f=5Hz

60

f=12.5Hz

40

f=20Hz

20

f=50Hz

0 0

1

2

3

r/d

4

5

f=100Hz f=150Hz

Fig. 6 Variation of local Nusselt number for steady and pulsed jet with different frequencies

CONCLUSIONS The heat transfer enhancement in pulsed jet impingement cooling at different frequencies was studied and the following conclusions made. For all frequencies under study (5 Hz, 12.5 Hz, 20 Hz, 50 Hz.100 Hz, 150 Hz) for pulsed jet it is observed that there is a considerable increase in average Nusselt number. Significant enhancement of heat transfer at the target surface by the intermittent pulsation in a turbulent impinging jet is observed. At 20Hz and 50Hz the stagnation and local nusselt number of the pulsed impinging jet is larger than that in the steady jet. For a frequency of 100Hz, there is a significant enhancement of local Nusselt number, where as Nusselt number at stagnation region remains same. For a Reynolds value of 5000 and a nozzle to target plate distance (Z/D) of 3, the optimum pulsed frequency is observed to be 50Hz. The stagnation and local Nusselt number at this frequency is higher, that ensures a higher value of heat transfer coefficent over the target plate.

84


REFERENCES D. A. Zumbrunnen and M. Aziz, 1993, “Convective heat transfer enhancement due to intermittency in an impinging jet,” Trans. ASME, J. Heat Transf., Vol. 115, pp. 91–98. H. S. Sheriff and D. A. Zumbrunnen, 1998, “Means to improve the heat transfer performance of air jet arrays where supply pressure is limiting,” Trans. ASME, J. Heat Transf., Vol. 120, pp. 787–789. Lance Fisher, 2001, A Numerical Investigation of Jet Impingement on a Heated Flat Plate, ME 513, pp. 3-20. N. Zuckerman, N. Lior, 2006, Jet Impingement Heat Transfer: Physics, Correlations, and Numerical Modelling, Advances in Heat Transfer, Vol-39, pp. 565-631. Peng Xu, Boming Yu, Shuxia Qju, Hee Joo Poh, Arun S. Mujumdar, 2010, Turbulent impinging jet heat transfer enhancement due to intermittent pulsation, International of Thermal Sciences, Vol.49, pp. 1247-1252. Abdul Rasool A. A, Jirunthanin V, Faik Hamad, 2014, Numerical and Experimental Study of Flow Structure and Cooling Behaviour of Air Impingement on a Target Plate, Int. J. of Thermal and Environmental Engineering, Vol. 8, No.1, pp. 33-43.

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86


NUMERICAL STUDY ON LINEAR AEROSPIKE NOZZLE AND EXPERIMENTAL VALIDATION OF SPIKE PRESSURE DISTRIBUTION Amal Dev NV1, Mathew George2, Khalid Rashid3, JC Pisharady4, Tide PS4 1

Assistant professor, Department Of Mechanical Engineering, MVJ College of Engineering, Whitefield, banglore-560067, 2, 3, 4 Liquid Propulsion systems center, ISRO, Trivandrum, Kerala-695547 5 Professor, Cochin University of Science and Technology Kochi- 682 022, Kerala, India.

ABSTRACT Linear aerospike nozzles are studied here, as a possible device to improve launcher engine performance. The main property of these nozzles is the possibility of good integration with vehicle. To improve the knowledge of flow-field and spike pressure distribution a 3-D aerospike nozzle is studied numerically and experimentally at atmospheric conditions and by varying chamber conditions. The spike plate pressure variation is measured by attaching pressure sensors at four different points of the plate and is compared with CFD results. Keywords: Linear Aerospike, Spike pressure. NOMENCLATURE pa pc M PR

Ambient Pressure Chamber Pressure Mach number Pressure Ratio

1


Numerical study on linear Aerospike….

INTRODUCTION Development of a new generation of space launchers requires the preliminary design of propulsion systems that can provide good performance in a wide range of altitudes. Among the advanced engine configurations a great interest has been devoted to those based on the aerospike nozzle concept. This external expansion nozzle permits an intrinsic adaptation of the exhaust jet to the varying ambient pressure and, at the same time, high design expansion ratios. Moreover, its compact shape provides an easy integration with the vehicle body. For these reasons it was selected as the propulsion system during the studies for the X-33 experimental vehicle. In that case, as in the general case of lifting body vehicles, the best structural integration with the vehicle was achieved by linear aerospike nozzles, which necessarily features a three dimensional flow about the plug due to the side truncation. The study of the three-dimensional flow generated by a linear aerospike is mandatory for a correct prediction of loads and performance during the vehicle ascent phase. In this framework it has to be emphasized that, because of the external expansion, the effect of the vehicle velocity on the flow-field around the aerospike nozzle cannot be neglected during the flight within the atmosphere. Moreover, the role played by ambient pressure on side expansion, and therefore on performance, has to be adequately addressed. Aim of the present paper is to analyze the flow-field details of a linear aerospike nozzle whose shape is designed by the two-dimensional approach and to measure the spike plate pressure distribution computationally and experimentally.

LITERATURE REVIEW Analytical and Numerical Studies Klaus Gross [1] conducted performance analysis of aerospike rocket engines. During this time period, it was believed that the thrust had two components, one being the effect from the wall surfaces guiding the flow, the other from the base area. Based on the chamber to ambient pressure ratio, two nozzle operation modes were possible. For low pressure ratios, the base flow separates and is dependent on the flow conditions through the trailing wake corridor. At higher pressure ratios, the flow field is unaffected by the atmospheric conditions. Ruf and McConnegay [2] presented a paper to provide an understanding of aerospike plume physics responsible for aerospike performance characteristics. Aerospike plume physics responsible for thrust generation and altitude 88


compensation were explained. The interaction of free stream and aerospike plume i.e. slipstream effect was also explained, based on his study. Ito et al. [3] performed a CFD study that focused on the flow field of an axisymmetric plug nozzle. A plug nozzle design was then designed using method of characteristics. Different lengths were designed, starting with the full nozzle and truncating at 20, 30, 40, 50, and 100% of that length. One important detail was that the area ratio was the same for all plug nozzles designed (1.7 for the inner nozzle and 6.5 for the whole nozzle). It was found that the optimum expansion occurred when the pressure ratio was 71. The Mach number was tested between values of 0.0 and 3.4. The flow field and thrust were studied at pressure ratios between 5 and 1000. Marco et al. [4] presented a paper which describes the flow field of aerospike nozzle in 3D model. To improve the knowledge of the flow field and performance of aerospike nozzles, they were studied numerically, with particular attention to the differences between the basic two -dimensional nozzle, usually considered in the design phase, and the more realistic three- dimensional nozzle. The study focused on different plug lengths and ambient pressures to assess the role of the linear plug side truncation on the base pressure behavior. Numerical tests were carried out at supersonic flight Mach numbers. Seiji Tsutsumi et al. [5] reported investigation on flow field of clustered linear aerospike nozzle numerically to reveal the correlation between flow structure and performance loss due to the clustering. The numerical results revealed stream wise vortices on the aerospike nozzle wall, leading to separations and separation shock waves. These shock structures produce three high pressure regions bounded by the separation lines over the aerospike nozzle surface. The flow structure over the aerospike nozzle surface were compared with supersonic base flow. However, this study revealed that the real flow field is more complicated, and that the separations and shocks determine pressure field and the thrust characteristics. Contribution of each high pressure regions to the thrust performance was evaluated, and mechanism of the loss caused by the clustering was clarified. Andrea et al. [6] reported a multi-chamber 1,300 lbf thrust LOX/Ethanol aerospike engine integrated into an instrumented test vehicle designed to reach supersonic conditions at approximately 15,000 ft and burnout shortly thereafter. The paper focused on the CFD analysis of this engine under various slipstream and overexpanded conditions. Slipstream effects were significant at sea-level conditions but rapidly disappeared with decrease in the ambient pressure. For example, an 8% drop in efficiency was predicted at sea-level between Mach number of 0.2 and 1.2. However, at 15,000 ft, this drop reduced to 3% and the phenomenon totally disappeared by 40,000 ft. These results suggest that slipstream effects in practical launch vehicle applications are likely to be minimal compared to the gains in thrust coefficients provided by the aerospike nozzle. Results also show that 2 to 5% in 89


nozzle efficiency was lost due to thruster interactions and inefficiencies which could be addressed independently of the slipstream effects. EXPERIMENTAL STUDIES Tomita et al. [7] reported the effect of base bleed on thrust performance of linear aerospike nozzle. The variation of thrust was studied with the help of side plate and found that with the absence of side plate thrust decreased and ideal base bleed percentage is 2%. Tomita et al. [8] investigated the effect of side wall on flow fields and thrust coefficients. Cold flow tests were conducted with and without side wall for 80% spike length and shock formations were captured with the help of shadowgraph technique. Zhang et al. [9] investigated a simplified design and optimization method of aerospike nozzle contour and the results of tests and numerical simulation of aerospike nozzles were presented. The primary nozzle contour was approximated by two circular arcs and a parabola: the plug contour was approximated by a parabola and a third-order polynomial. The maximum total impulse from sea level to design altitude was adopted as objective to optimize the aerospike nozzle contour. Experimental studies were performed on a 6-cell tile-shaped aerospike nozzle, a 1cell linear aerospike nozzle and a 3-cell aerospike nozzle with round-to-rectangle (RTR) primary nozzles designed by method proposed in the present paper. Three aerospike nozzles achieved good altitude compensation capacities in the tests and still had better performance at off-design altitudes compared with that of the bellshaped nozzles. The performance was obtained under two nozzle pressure ratios (NPR) lower than design altitude. Efficiency reached 92.0-93.5 % and 95.0-96.0 % respectively. Pressure distribution along plug ramp was measured and the effects of variation in the amount of base bleed on performance were also examined in the tests. Flight Tests Besnard et al. [10] performed first flight validation of aerospike plug nozzle by integrating it to prospector-2 vehicle. After a smooth countdown and nominal engine ignition, the thirteen-foot long P-2 quickly accelerated up a 60-ft launch rail and entered stable flight. Soon afterwards, the thrust decreased and, then an off-axis component arose which ultimately caused the vehicle to enter unstable flight. Post flight engine inspection showed that a significant amount of flow escaped around the outer surface of the graphite convergence ring and burnt the back of the chamber, leading to several secondary plumes. These secondary plumes eventually led to asymmetric thrust and sent the vehicle out of control until it transitioned into a unpowered ballistic terminal descent.

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Trong et al. [11] conducted a flight research on an aerospike rocket nozzle using high power solid rockets. Flight test results were compared with ground test results. Flight data showed that all the rockets successfully reached supersonic speeds with a maximum Mach number of 1.6 and a peak pressure altitude of nearly 30,000 ft. The aerospike nozzle efficiency was determined to be 0.96 from computational fuid dynamics (CFD) analysis. NUMERICAL SIMULATION Numerical simulations were carried out using a commercial CFD code, ANSYS FLUENT 15. It is based on a second order finite volume discretization and the SIMPLE pressure correction technique for enforcing the divergence-free condition of the velocity field; the time integration is three-level fully implicit. The time step size, the number of iterations per time step, the total number of time steps, and the convergence limit for each time step must be specified if the unsteady solver is used. The total number of iterations and the convergence limit must be specified if the steady solver is used. CFD simulation consists of three main parts:  Pre-processor  Solver  Post processor The Pre-processing module is used to model the computational domain and appropriate boundary conditions. During Pre-processing computational domain with boundary is selected. Grid generation is also a part of Pre-processing. Fluid properties are also selected before simulation. Solver used is ANSYS FLUENT 15. Here the mesh file is imported from a modeling and meshing software GAMBIT. Selection of solver, turbulence model, specification of material properties and boundary conditions has been made during pre-processing. The convergence criteria and the required number of iterations are approximately calculated. Once the solution is converged the Post processing is done in ANSYS CFD POST and plots generated in Techplot 360. The velocity contours, Mach Contours and pressure distribution plots are compared during Post processing.

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COMPUTATIONAL DOMAIN

Fig 1: CFD DOMAIN OF LINEAR AEROSPIKE NOZZLE The main objective of this CFD study is to find the pressure distribution along the symmetric line on spike plate. So major fine grids are made on plate surface. A fine boundary layer has been made on spike surface with y+ value 0.01mm in order to capture the boundary layer separation and shock formations.So the domain is made to be 10D width and 25D height. GRID INDEPENDENCE STUDY A grid independence study has been conducted for 3d linear aerospike nozzle for different grids of size, such as 2.5 lakhs, 5 lakhs, 6 lakhs, 8 lakhs, 9 lakhs and 10 lakhs. The inlet conditions used for grid independence study are Pressure: 8 bar Temperature: 300 K Turbulence model used: SST k-ɷ The outlet conditions are standard atmospheric conditions. Pressure: 1 bar Temperature: 300 K 92


Table 1: Results Of Grid Independence Study Sl.No 1 2 3 4 5 6

Grid size (in lakhs) 2.5 5 6 8 9 10

Mach No. Along Axis 3.68 3.56 3.54 3.51 3.50 3.50

The values of Mach number did not show much deviation (less than 2%) beyond a grid size of 9 lakhs and hence this grid was selected for the remaining calculations.

Fig 2: GRID OF 9 LAKH CELLS EXPERIMENTAL MODEL A cold flow experiment has been done on a linear aerospike nozzle to get an insight on the pressure distribution along the spike wall surface of the nozzle. Pressure at four different points on the spike surface is determined by attaching pressure sensors to the spike as shown in Fig 4. The maximum chamber pressure obtained during the test was 16 bar. The ultimate aim of this experimental study is to obtain the pressure values obtained from the CFD simulation of 3D linear aerospike nozzle on spike wall and compare it with the present experimental values.

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FIGURE 1: EXPERIMENTAL MODEL WITHOUT SENSOR POINTS

FIGURE 2: SENSOR POINTS S1 S2 S3 S4 (FROM LEFT TO RIGHT). The Figure 3 shows experimental aerospike nozzle without sensor points and Figure 4 shows the enlarged view with sensor points. In Fig. 4, the four sensor ports on spike plate are labeled as S1, S2, S3, S4 for computational and calculation purposes. The sensor used for pressure measurement is Keller pickup whose value is calibrated between 0 bar to 3000 milli bar (3bar). The value is calibrated in milli range in order to capture a slight deviation in the spike pressure. The chamber sensor is calibrated in the range of 0-20 bar. RESULTS AND DISCUSSIONS 1.

Numerical Results

A comprehensive numerical study was carried out on the designed linear aerospike nozzle using ANSYS 15 CFD package. Pressure based solver was initially used for the simulation as it has higher accuracy for high Mach number flows and convergence is faster than density based solver. The SST k-ɷ two equation model was used for turbulence modeling because of its higher accuracy for high speed separated flows. Moreover, the predictions of the above model are better for nozzle flows.

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The pressure variation along the spike surface for various inlet pressures ranging from 8 bar to 15 bar are plotted and compared.

Figure 3: Mach Number Contour At 8 Bar Inlet Pressure (Symmetry)

Figure 4: Mach Number Contour At 12 Bar Inlet Pressure (Symmetry)

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Figure 5: Mach Number Contour At 15 Bar Inlet Pressure (Symmetry) The above figures clearly depicts the increase in exit Mach number with increase in inlet pressure and maximum Mach number after expansion in the ambient is 4.3 at 15 bar inlet pressure. As the shock diamond becomes stronger, the pressure acting on the plate shows an increasing trend. 1.1 Spike Plate Pressure Distribution The following graphs show the variation of pressure at symmetric plane on spike surface for different inlet pressures.

Fig 8: Variation In Pressure With Axial Distance From Nozzle Exit (Numerical) 96


2.

EXPERIMENTAL RESULTS

The following graphs show the comparison of pressure distribution along the symmetric line of spike plate of 3 D aerospike nozzle obtained from CFD results with the experimental values. The experimental data were obtained from the four sensors that are attached to four different points at spike plate on the experimental model.

FIGURE 9: COMPARISON OF VARIATION IN PRESSURE WITH AXIAL DISTANCE FROM NOZZLE EXIT 97


The above graphs show better comparison for low values of inlet pressure (6 and 10 bars). However, slight deviation is observed for the inlet pressures of 14 and 16 bars. CONCLUSIONS The flow field and pressure distribution along the linear aerospike nozzle wall was studied numerically and validated with the data obtained from experiments. In CFD plot the peak pressure region is formed due to the recirculation of air on the plate surface, which forms a high pressure region. The pressure in this region increases as the chamber pressure increases. This region is just below the first shock diamond as visible in Mach contour. The separation of flow also takes place in this region but immediately joins the plate surface after this recirculation zone. From the Mach contours it is obvious that the recirculation zone moves towards the trailing edge of the nozzle with increase in inlet pressure. Both CFD predictions and experiment data closely matches for low values of inlet pressures.

REFERENCES [1] Gross, Klaus W., “Performance Analysis of Aerospike Rocket Engines.” 1972. [2] J. Ruf, P. McConaughey. The plume physics behind aerospike nozzle altitude compensation and slipstream effect, 33rd Joint Propulsion Conference and Exhibit, 1997-3218,1997. [3] Takashi Ito, Kozo Fujii and A. Koich Hayashi. Computations of Axisymmetric Plug-Nozzle Flowfields: Flow Structures and Thrust Performance, Journal of Propulsion and Power, Vol. 18, No. 2 (2002), pp. 254-260. [4] Marco Geron, Renato Paciorri, Francesco Nasuti, Filippo Sabetta, Emanuele Martelli. Transition Between Open and Closed Wake in 3D Linear Aerospike Nozzles, 41st AIAA/ ASME/ SAE/ ASEE Joint Propulsion Conference Exhibit, 2005-4308, 2008. [5] Seiji Tsutsumi, Susumu Teramoto, Toshio Nagashima. Flow Structure and Performance Analysis of Clustered Linear Aerospike Nozzle, 41st AIAA/ ASME/SAE/ ASEE Joint Propulsion Conference Exhibit, 2005-4307, 2005. [6] Andrea Eric Besnard and John Garvey. CFD Performance Analysis of a MultiChamber Aerospike Engine in Over-Expanded, Slipstream Conditions, 45th AIAA/ASME/SAE/ASEE, 2009-5486, 2009. [7] Takeo Tomita, Mamoru Takahashi, Takuo Onodera, Hiroshi Tamura. Thrust loss due to design of linear aerospike nozzles, 36th AIAA/ ASME/ SAE/ ASEE Joint Propulsion Conference and Exhibit, 2000-3290, 2000. [8] Takeo Tomita, Mamoru Takahashi, Takuo Onodera, Hiroshi Tamura. A simple performance prediction model of clustered linear aerospike nozzles, 37th Joint Propulsion Conference and Exhibit, 2001-3560, 2001. 98

Proceedings of ICETiME’15 December 16-17, 2015, CUSAT, Kochi, India. [9] Wuye Dai, Yu Liu, Zhengke Zhang, Lizi Qin, Yibai Wang. Numerical

investigation on linear aerospike nozzles, 37th Joint Propulsion Conference and Exhibit, 2001-3568, 2001. [10] Eric Besnard. Development and Flight-Testing of Liquid Propellant Aerospike Engines, 40th AIAA/ ASME /SAE /ASEE, 2004-3354, 2004. [11] Trong Bui, James Murray, Charles Rogers, Scott Bartel, Anthony Cesaroni, Mike Dennett. Flight Research of an Aerospike Nozzle Using High Power Solid Rockets, 41st AIAA/ ASME/ SAE/ ASEE Joint Propulsion Conference Exhibit, 2005-3797, 2005. [12] Eric Besnard. Design, Manufacturing and Test of a Plug Nozzle Rocket Engine, 38th AIAA/ ASME/ SAE /ASEE, 2002-4038, 2002. [13] Andrea Eric Besnard and John Garvey. CFD Performance Analysis of a MultiChamber Aerospike Engine in Over-Expanded, Slipstream Conditions, 45th AIAA/ASME/SAE/ASEE, 2009-5486, 2009. [14] G. Angelino. Approximate method for plug nozzle design, AIAA Journal, 1964, Vol.2: 1834-1835.

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Proceedings of the International Conference on Emerging Trends in Mechanical Engineering (ICETiME’15) December 16-17, 2015, CUSAT, Kochi, India, 101-106

ONLINE DETECTION OF FAULTS IN INTERNAL COMBUSTION ENGINES USING ACOUSTIC EMISSION SIGNALS Sreedhar Puliyakote1, Krishnan Balasubramaniam2 1

Scholar, 2Professor Centre for Non-Destructive Evaluation, Indian Institute of Technology Madras, Chennai 600036.

ABSTRACT Most commonly occurring failures in an Internal Combustion (IC) engine generate a characteristic indication before failure. This indication is mostly acoustic in nature, sometimes in the ultrasonic frequencies. Identification of such indicators in a timely fashion allow in taking steps to reduce further damage to the system, saving the operators precious time and money. Detecting these faults in the early stages of initiation, if possible in a nonintrusive way, is crucial and pose a major challenge for the operator. The work involves the development an efficient and cost effective acoustic emission system with custom signal conditioning unit which can be used to continuously monitor ‘online’ the AE generated during the piston movement with no modification to the engine. The signals obtained are analyzed for indication of faults using a few popular signal processing techniques. .

INTRODUCTION Internal combustion engines have become integral part of human life with its involvement in the field of transportation and power generation. With engines being developed to generate more power per unit capacity while curbing the fuel consumption and adhering to the stringent environmental norms, their running condition has been pushed to the limit. Moreover, due to economic reasons, operators tend to run the engines at the most cost effective condition. One such method is running the engine at lubricant oil levels less than the recommended amount. This strategy makes sense from an economic point of view in case of large engines like the ones used in marine and power generation applications where the lubricant oil consumption itself is in terms of several liters per hour. But such low 1


Online Detection of Faults in.......

supply of lubricant oil leads to severe reduction in the oil film thickness on the cylinder liner. This can, as seen in many cases, lead to severe wear of the cylinder liner and piston rings. The reduction in lubricant oil inside the cylinder can lead to a number of other issues also. The higher friction between piston and cylinder liner generated by lack of oil film effectively creates a drop in the power output and the brake mean effective pressure of the engine which affects the engine performance. Moreover increase in friction leads to increase in fuel consumption too. So the monitoring of the oil level becomes critical so as to maintain the efficiency of the engine while keeping the economic aspect in mind. One technique that has found a growing acceptability in the field of condition monitoring is Acoustic Emission. Acoustic Emission (AE) is the propagation of transient elastic waves that are generated within or on the surface of a material subjected to force. AE, in machinery, is generated as a result of deformation and micro fractures brought about by friction and wear which are inevitable in any kind of machinery involving moving components. AE offers the advantage of earlier failure detection due its inherent higher sensitivity as compared to the low frequency vibration signals and other such techniques (Neill et.al., 1998; Gill et.al., 1998, 2000). Over years past, many people have worked on using a number of such systems to monitor the various parameters associated with engine operation. These include cylinder temperature, pressure, oil film thickness, and oil analysis. Most of these approaches were found to have shortcomings. Over the years, Acoustic Emission (AE) has emerged as a relatively robust tool for engine condition monitoring as well as for monitoring other machinery with moving parts. A lot of work has already been done in the field of condition monitoring of engines using AE. It has been used successfully to detect exhaust valve leakage (Fog et.al., 1998), fuel injection behavior (Gill et.al., 2000) and various aspects of combustion processes (Gill et.al., 1998; El-Ghamryet.al., 2003). The aim of this paper is to identify indicators of possible fault conditions in a 4 stroke petrol engine, mainly inside the cylinder. The AE signal generated by the friction between the piston rings and the cylinder liner is studied as an indicator of possible lack of lubrication. Study was conducted to understand the relationship between the lubricating condition and the acoustic emission. Then this knowledge was used to utilize information obtained from the AE signal to provide an early warning of a possible reduction in oil which can help in the prevention of scuffing.

EXPERIMENTAL INVESTIGATIONS Experimental Setup An AE system developed in-house was used for the study. This system comprised of a 3 components. First one was a custom designed piezoelectric sensor that can 102


detect the AE off the outer surface of the cylinder liner. A broad band Piezo-electric crystal resonant at 250 kHz was used as the active element. The sensor casing was made out of aluminum. A photograph of the sensor and its layout is shown in Fig. 1.

Fig.1 Custom made sensor used to pick-up the AE (left) and its layout (right) The AE from the sensor is then fed to the next part of the system which is the Signal conditioning Unit (SCU) where the signals are amplified and filtered. The SCU had 2 input and 2 output channels and each channel is capable of 80 dB amplification. The SCU also contains appropriate filters to process the signals as these signals contain a large amount of noise. Then the signal is saved to a Personal Computer using an NI 5132 data acquisition card for further analysis. The data acquisition and analysis is done using custom-developed software based on LabVIEW. A schematic of the experimental setup is shown in fig.2. The capabilities of the sensor and the system has already been studied by the authors of this paper and can be found in Sreedhar et.al. (2010) Engine Details Table 1 Specifications of the IC Engine used in this study Operation

Four stroke Petrol

Number of cylinders

1

Speed

5000 rpm

Bore

53.5 mm

Stroke

48.8 mm

Bhp/ cylinder

8.2

Power output

6.1 kW @ 7500rpm 103


The Experiments were conducted on a 125 cc single cylinder 4 Stroke petrol engine. The sensor was mounted on the cylinder surface in such a way that it was properly coupled to the outer surface of the cylinder head and thus picks up most of the AE passing through the cylinder liner into the cylinder block. The engine was run at 3 different lubricant oil conditions-full or recommended level, half the recommended level and with no oil supply to the cylinder. AE signals are recorded for several seconds for each of the conditions. Tests were also conducted with bad piston and worn out piston rings to study the impact of these faults on AE generated. Table 1 gives the specifications of the engine used in this study.

Fig.2 Schematic Diagram of the experimental setup

ANALYSIS AND DISCUSSIONS

Fig.3 AE signal obtained from one cycle of engine operation The Fig. 3 shows a typical signal from one cycle of the engine operation. Signals were obtained at three lubricant oil levels-normal (recommended) half and no oil conditions. With the help of the output from the ignition coil, the TDC and BDC were identified and the time was converted to crank angle thus making it easy to identify the strokes which are shown in Fig 4.

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Fig.4 AE signal generated during one cycle of engine 3 operation with the corresponding stroke during engine operation Fig.5 shows Short Time Fourier Transform (STFT) plots of the signals obtained under the three lubricating conditions, severity increasing from top to bottom. Similar to the previous cases, AE activity was found to be increasing with reduction in lubrication. Besides, a couple of new features emerged, including a rise in amplitude in the higher (200-250 kHz) frequency range.

Fig.5 Comparison of STFT plots for the 3 lubricant oil conditions, 100%, 50% and 0% respectively from the top. (Feature indicating the fault marked)

CONCLUSIONS Since the engine was compact, it was easier and safer to simulate lubrication supply faults and conduct AE tests in faulty conditions. Thus, a much more controlled 105


experiment was conducted on the high speed engine. The AE generated due to the lack of lubrication was clearly identified in the STFT. Processing of the raw AE signals was found to be very effective in enhancing the required feature that arise due to the reduction in the lubrication oil while suppressing the other acoustic emission components that are primarily due to the normal activities in the engine. The results from the test indicated that there is also additional potential to AE from the cylinder of an engine can be used to diagnose the engine for other faults also, like bad piston and worn out piston rings. This method can be employed as a economically viable online monitoring system for IC Engines.

REFERENCES El-Ghamry, M.H., R.L. Reuben and J.A. Steel (2003)The development of automated pattern recognition and statistical feature isolation techniques for the diagnosis of reciprocating machinery faults using acoustic emission, Mechanical Systems and Signal Processing, 17, 805–823. Fog, T.L., E.R. Brown, H.S. Hansen, L.B. Madsen, P.S. Rensen, J.A. Steel, R.L. Reuben and P.S. Pedersen (1998) Exhaust valve leakage detection in large diesel engines, Condition Monitoring and Diagnostic Engineering Management Proceedings of 11th International Conference on Condition Monitoring and Diagnostic Engineering Management COMADAM, Clayton, Australia, 269–278. Fog, T.L. Condition monitoring and fault diagnosis in marine diesel engines. PhD Thesis, Danish Technical University, 1998. Gill, J. D., E.R. Brown, M. Twite,G. Horner, R.L. Reuben and J.A. Steel (1998) Monitoring of a large reciprocating compressor. Proceedings of 11th International Conference on Condition Monitoring and Diagnostic Engineering Management COMADAM, Clayton, Australia, December, 317-326. Gill, J.D., R.L. Reuben, J.A. Steel, M.W. Scaife and J.A. Asquith (2000) Study of small HSDI diesel engine fuel injection equipment faults using acoustic emission. Journal of Acoustic Emission, 18, 96-101. . Neill, G.D., S. Benzie, J.D. Gill, P.M. Sandford, E.R. Brown, J.A. Steel and R.L. Reuben (1998)The relative merits of acoustic emission and acceleration monitoring for detection of bearing faults, Proceedings of 11th International Conference on Condition Monitoring and Diagnostic Engineering Management COMADEM, Clayton, Australia. Sreedhar P, Janardhan Padiyar M., Maharajan R., Krishnan Balasubramaniam (2010) Development of an Acoustic Emission system for IC Engine Health Monitoring, NDE2010, Kolkata, India, December 2010. 106


HEAT TRANSFER ANALYSIS IN ELLIPTICAL TUBE HEAT EXCHANGERS Viswajith M V, Gireeshkumaran Thampi 1

Assistant Professor Universal Engineering College Thrissur 680123, Kerala, India 2 Associate Professor Cochin University of Science and Technology Kochi- 682 022, Kerala, India.

ABSTRACT The purpose of this numerical study is to explore the fundamental mechanism between the local flow structure in heat transfer enhancement and the pressure drop in elliptical tube heat exchangers. The heat transfer enhancement potential of an elliptical tube heat exchangers in which elliptical tubes are placed at an angle to the flow direction were numerically studied. The investigation was done at relatively lower Reynolds number ranging from 250-1000. It was observed that the flow disturbances generated at the downstream, when flow over inclined elliptical tube occurs, were the main reason for the heat transfer augmentation. The result showed that these elliptical tube heat exchangers without VGs can augment the heat transfer rate better than that of the circular tube heat exchangers with VGs and with a moderate pressure drop penalty.

INTRODUCTION High heat exchanger performance is very important in meeting efficiency standards with low cost and environmental impact. In practical applications, the air side convection resistance is usually dominant due to the thermo physical property of air. Efforts have been devoted to the research and development of the

Viswajith M V E-mail: [email protected]

Numerical Investigation of heat transfer enhancement in elliptical tube heat exchangers

enhancement of heat transfer surfaces, especially on the gas side, where a high thermal resistance exists. Many efforts have been made to enhance air-side heat transfer performance and variations in fin patterns like wave, louver and slit fin have been adopted (Huisseune, 2013). However, with significant heat transfer enhancement, the associated penalty of pressure drop is also tremendous for those conventional heat transfer enhancement methods. In recent years, a very promising strategy of enhancing air-side heat transfer performance is using flow manipulators, known as vortex generators (VGs). When the fluid flows through a vortex generator, stream wise vortices are generated in the flow field due to flow separation on the leading edge of the VG. These vortices causes bulk flow mixing, boundary-layer modification, and flow destabilization; heat transfer is enhanced due to these vortices (Lin, 2013). Biswas et al. [1996], in one of his studies identified that, the vortex strength and thus the heat transfer performance improved with increasing the angle of attack of the delta-winglets and the Reynolds number. He et al. (2012) numerically studied the heat transfer enhancement by punched winglet type vortex generator arrays in a fin and tube heat exchanger. This study inferred that for the punched VGs, the main vortex, located directly downstream of the delta winglet, is formed by flow separation at the leading edge of the winglet. The ‘corner vortex’ which is also formed along with the main vortex, plays a major role in the fluid flow behavior and heat transfer characteristic in the channel of a fin and tube heat exchanger than that of the main vortex [He, 2012]. Experimental investigations by Valencia et al. [1992] showed that locating the winglets symmetrically downstream of the tube is the optimal position for the generators with respect to the tube. This location significantly enhances the heat transfer in the wake of the tube, where a poor heat transfer coefficient is noticed. Torii et al. (2002) experimentally observed an increase in heat transfer coefficient using delta-winglets in common-flow-up and common flow- down configurations in a plate-fin-and-tube heat exchanger at a relatively low Reynolds number. The winglets used were not completely located on the upstream or downstream region of the tubes but were placed more on the sides of the tubes. The majority of the researchers focused on experimenting various patterns of VGs and positioning VGs for creating the vortices and hence the heat transfer enhancement. Less preference was given for introducing an alternate technique to create the vortices. It was understood that, the flow that inhibits the heat transfer enhancement could also be produced by replacing the circular tubes by elliptical tubes and placing them at an angle to the flow direction, without VG. The inclined elliptical tubes work on the same principle as VGs. They not only improve the heat transfer performance, but also relatively reveal the pressure drop. This paper establishes the physics of the flow structure and heat transfer enhancement in practical heat exchangers. 108


MODEL DESCRIPTION Physical Model In this study, a fin-and-tube heat exchanger with longitudinal vortex generators is investigated. The schematic diagram of the heat exchanger is shown in Fig.1. In the present study, we adopt the rectangular winglet pair as the vortex generator based on the literature. A pair of rectangular winglets is symmetrically mounted on the fin surface, adjacent to heat transfer tube.

Fig 1: Schematic diagram of core region of finned tube heat exchanger with vg [He, 2013]

The height of the winglets is equal to 80% of the fin spacing, Fig.2 (b). Fig.2 (a) shows the dimensions of rectangular winglets and their location with respect to the circular tube. The rectangular generators are placed in flow-up orientation. In Fig. 1, the computational domain and the coordinate system are presented, where X is the stream wise direction, Y is the span wise direction and Z direction shows the fin pitch direction. Fin spacing or the distance between the fins is set as H = 4 mm, width B = 45 mm, and length L = 100mm. The first tube of diameter D = 10 mm, is located at X = 15 mm from the inlet of the flow channel. The longitudinal tube pitch Pl = 16 mm and the transverse tube pitch Ps = 13 mm. The tube rows are arranged staggered. Each tube in the second row is placed in the center of the adjacent tubes in the first row. The fin material is aluminum and fin thickness Ft=0.16 mm. Since the geometry of the fin-and-tube heat exchanger is symmetric, the whole region of two rows of tubes shown in Fig. 3 is selected as the computational domain for both circular and elliptical tube heat exchangers. Due to the high heat transfer coefficient inside the tube and the high thermal conductivity of the tube wall, the tube temperature is set as constant. However, the temperature distribution on the fin surface is unknown and will be determined during the computational iteration process. In order to solve this problem, the computational domain should contain the whole fin surface during the numerical simulation. 109


Fig 2: (a) VG placement and its position with respective to the tube in circular finned tube heat exchangers (b)side view showing height of the VG between the fins

In this study, the fluid is considered as incompressible with constant properties. The generation of longitudinal vortices is a quasi-steady phenomenon. Consequently, due to the low inlet velocity and the small fin pitch, the flow in the channel of the compact heat exchanger is assumed to be laminar and steady. Fin thickness, heat conduction through the fins and vortex generators are taken into account for the analysis. The temperature distribution for the fins can be determined by solving the conjugate heat transfer problem in the computational domain.

Fig 3: Elliptical tube dimensions and its position with respect to its flow direction

The governing equations in Cartesian coordinates can be expressed as follows Continuity Equation: General form is given by

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(1) For a steady and incompressible flow, the change in density w.r.t. time is zero. In tensor form, it is given as (2) Similarly, for steady and incompressible flow Momentum Equation: (3) Energy Equation: (4)

Fig 4: Co-ordinate system and computational domain of circular tube heat exchangers with vg and elliptical tube heat exchangers

The boundary conditions for all surfaces are described as follows: 

At the inlet boundary U = Uin = constant, v = w = 0, T = Tin= constant

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

At the top and bottom boundaries Periodic boundary conditions are given: , w=0, At Extended region u=v=w=0,



=0

at fin region

At the front and back surface of wind tunnel region (x-z plane), symmetry boundary condition are given: , v=0,



At the outlet, the outflow boundary condition is given



Tube region : u = v = w = 0; T =Tw = constant

The governing equations were solved using numerical methods for corresponding boundary conditions.

NUMERICAL METHODS The geometry for the three-dimensional vortex-enhanced multi-row fin-andtube heat exchanger is complex and it can be expected that the velocity and temperature fields are complicated in the computational domain. In order to capture important scales and resolve the near-wall gradients appropriately, the grid must be generated with great care and effort. The computational meshes were generated by using ANSYS. Because of the complexity of the computational domain, it is difficult to use a single structured quadrilateral mesh in the whole flow passage. In order to improve the quality of the grid system, mapped face method and inflation is adopted to generate the mesh. Different strategies are employed for each subdomain to generate the mesh. For the blank zone, a structured mesh is employed because of its simplicity. So does it for the tube zone. The VG zone is complex and hence an unstructured tetrahedral mesh is employed. Generally, the meshes are generated much finer in the regions adjacent to tubes and winglets. Finally, mesh independent study is conducted to ensure the quality of mesh. The Navier-Stokes and energy equations with the boundary condition equations are solved by using a computational fluid dynamics code (Fluent). The convective terms in governing equations for momentum and energy are discretized with the second-order upwind scheme. The coupling between velocity and pressure is performed with SIMPLE algorithm. The convergence criterion for the velocities is 112


that the maximum mass residual of the cells divided by the maximum residual of the first 5 iterations is less than 1.0 x 10-5, and the convergence criterion for the energy is that the maximum temperature residual of the cells divided by the maximum residual of the first 5 iterations is less than 1.0 x 10-6.

RESULTS AND DISCUSSION In our present study, the VGs are arranged in “up-flow” orientation. In this confined passage, the fluid is accelerated. As a consequence, the boundary layer separation is suppressed and the tube wake region is narrowed. Furthermore, the fluid accelerated in the constricted passage will impinge directly on the downstream tube resulting in a local heat transfer enhancement. The combined effect of separation delay, narrowing of tube wake and impingement can significantly augment the heat transfer in the flow channel [Valencia, 1992].

Fig 5: Flow in the computational domain - elliptical tube heat exchangers for Reynolds number 500.

Basically there are three kinds of longitudinal vortices in the channel, main vortex, corner vortex and induced vortex [Toa, 2007]. The main vortex is formed due to the flow separation and friction on the leading edge of VGs. However, the corner vortex is formed by the deformation of the near-wall streamlines. The induced vortex which is formed by the interaction between the main vortex and the fin surface rotates opposite to the main vortices. The combined effect of these vortices distorts the temperature field in the channel and serves ultimately to bring about a significant augmentation of heat transfer between the fluid and its neighboring surfaces. It is clear from the figure 5 that the streamlines are stretched and bended toward the wake region behind the tube. The converged streamlines are generated due to the formation of longitudinal vortices behind the VGs. The longitudinal vortices bring high-momentum of fluid into the wake region [Biswas, 1996]. Then the wake region 113


is compressed and the size of this region is reduced. It is interesting to note that with the increase of the angle of attack, the streamlines are bended progressively closer to the center of wake region in the rear of the first tube and the size of the wake region reduces with the increasing attack angle. This phenomenon can be explained by the vortex strength of the longitudinal vortices. With the increase of the angle of attack, the vortex strength is becoming larger. The larger vortex strength will bring highermomentum of fluid into wake region. Then the accelerated flows compress the fluid in the wake region and delay the flow separation on the tube and result in a reduced wake size [He, 2012]. Influence of Elliptical tubes Viswajith et al. [2015] established that the vortex generators can significantly improve the heat transfer performance at a modest pressure loss. By changing the arrangement of vortex generators array from inline array to staggered array, the performance of heat transfer almost remains in the same level but the pressure loss is reduced by 7%. This implies that the optimization of vortex generators arrangement can result in a better pressure distribution and a reduced pressure loss without reducing the heat transfer enhancement. Similarly by changing the dimensions of elliptical tubes and keeping them at an angle to the flow direction, without employing a VG, will generate same flow disturbances at the vicinity of tubes and will perform similar to a circular tube with VG. Therefore a similar heat transfer enhancement as in a circular tube with VG can be obtained in elliptical tube heat exchangers without a VG and at lower pressure drop penalty. For the comparative investigation, the elliptical tube was fixed at 230 to the flow direction. The Reynolds number is varied from 250 to 1500 in steps of 250, by varying the inlet velocities from 1.16 to 4.66 ms-1. The figure 6 and figure 7 shows the heat transfer enhancement and pressure drop in both the cases respectively.

Fig 6: Heat transfer coefficient variation in circular tube with vg and elliptical tube without vg for Reynolds number ranging from 250-1000

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For the Reynolds number from 250 to 1000, the heat transfer coefficient and pressure drop variation is plotted and it is observed that, the heat transfer enhancement is from 16-56% for reduction in pressure drop by 9-25%. In fact, the difference of heat transfer coefficient is very small for the two cases and it only presents the changing trend of different configurations, however slightly heat transfer enhancement for slightly lesser pressure drop penalty makes its value.

Fig 7: Pressure drop variation in circular tube with vg and elliptical tube without vg for Reynolds number ranging from 250-1000

CONCLUSION In this study, three-dimensional numerical simulations are employed to investigate the heat transfer characteristics and flow structure in full-scale fin-andtube heat exchangers with VGs for circular and without VGs for elliptical tube heat exchangers respectively. The same effect of thermal mixing of the fluid, delay the boundary layer separation can be produced by replacing the circular tubes with elliptical tubes and placing them at an angle to the flow direction, thereby eliminating the use of VG. The longitudinal vortices generated by inclined elliptical tubes rearrange the temperature distribution and the flow field, and as a consequence significantly enhance the heat transfer performance of the fin-and-tube heat exchanger. Comparing these elliptical tube heat exchangers without VG with the circular tube with VG, the heat transfer coefficient of the fin-and-tube heat exchanger using elliptical tubes is improved by 16%, 32%, 47% and 56% for Reynolds number 250, 500, 750 and 1000 respectively. The corresponding pressure drop is also decreased by 9%, 13%, 19% and 25% respectively. 115


REFERENCES Huisseune. H, Joen. C. T, Jaeger. P. D, Ameel. B, Schampheleire. S. D, Paepe. M. D, (2013) “Performance enhancement of a louvered fin heat exchanger by using delta winglet vortex generators”, International Journal of Heat and Mass Transfer, 56, 475–487 Lin. W. C, Ferng. Y. M, Chieng. C. C (2013). “Numerical computations on flow and heat transfer characteristics of a helically coiled heat exchanger using different turbulence models”, Nuclear engineering and design 263, 77-86. Biswas. G, Torii. K, Fujii. D, Nishino. K (1996) “Numerical and experimental determination of flow structure and heat transfer effects of longitudinal vortices in a channel flow”, International Journal of Heat and Mass Transfer 39 (16), 3441– 3451. He. Y. L, Han. H, Tao. W.Q, Zhang Y.W (2012) “Numerical study of heat transfer enhancement by punched winglet type vortex generator arrays in fin and tube heat exchangers”, International Journal of Heat and Mass Transfer 55, 5449-5458. Valencia. A, M. Fiebig. M, Mitra. N. K, Leiner. W (1992) “Heat transfer and flow loss in a fin tube heat exchanger element with wing-type vortex generators”, Institution of Chemical Engineering Symposium Series 129 (1), 327–333. Torii. K, Kwak. K. M, Nishino. K, (2002) “Heat transfer enhancement accompanying pressure-loss reduction with winglet-type vortex generators for fintube heat exchangers”, International Journal of Heat and Mass Transfer 45, 3795– 3801. Viswajith. M.V, Thampi. G.K, Varghese. J, (2015) “Heat transfer enhancement with pressure loss reduction in compact heat exchangers using vortex generators”, Journal of chemical and Pharmaceuticals Sciences, Special Issue 6, 60-64. He. Y.L, Chu. P, Tao. W. Q, Zhang Y.W, Xie. T, (2013) “Analysis of heat transfer and pressure drop for fin and tube heat exchangers with rectangular winglet type vortex generators”, Applied thermal Engineering 61, 770-783. Tao. Y.B, He. Y. L, Huang. J, Wu. Z.G, Tao. W.Q, (2007) “Numerical study of local heat transfer coefficient and fin efficiency of wavy fin and tube heat exchangers”, Internal journal of Thermal sciences 46, 768-778.

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EXPERIMENTAL INVESTIGATION OF A SOLID / MOIST AIR THERMO CHEMICAL STORAGE PROCESS FOR WASTE HEAT RECOVERY Jeffy Scaria1, Abhijith N2, Swaraj Kumar B3 1

M.Tech student, 2 B.Tech student, 3assistant professor (mechanical dept)

LBS college of engineering,Kasaragod, Kannur university contact mail id : [email protected] ABSTRACT Thermal energy storage (TES) is regarded as an enabling technology with a variety of applications, especially regarding energy efficiency and usage of renewable and waste heat. TES materials, in this aspect, provide much higher storage capacities per mass or volume compared to sensible or latent heat storage, often by a factor of 10 or more compared to water storage, which is the most often used storage type . Moreover, thermo chemical storage materials can store the heat for infinite time without insulation and are regarded as a key technology for heat transport and long term storage. TES is based on two reactions, one endothermic and one exothermic which are reversible. In this work our aim was to design a thermo chemical heat storage system that can store waste heat from a stationary IC engine exhaust for long durations and which should be able to heat water for household purposes when required. In this study we reviewed various thermo chemical materials based on its characteristics and our requirements, MgCl2·6H2O or magnesium chloride hex hydrate was chosen. Dehydration of MgCl2·6H2O at about 120°C into MgCl2·H2O stored heat energy (charging), and hydration of this mono hydrate released the stored heat (discharging). A reactor for the reversible reaction and heat transfers from the chemical to water and from exhaust to chemical was designed and fabricated. In this work the TES capability of the reactor is investigated. Making of an efficient, compact and practical reactor was the core idea of the work. Promising results were obtained from the magnesium chloride hexahydrate filled thermo chemical reactor. Upto 65.35% of the stored thermal energy were recovered from the reactor after one day of storing. And 37.72% of the stored thermal energy were recovered after one month of storage. 1


Experimental Investigation of a solid…

INTRODUCTION The main principle of thermo chemical TES is based on a reaction that can be reversed: Let A,B and C be some chemical compounds. C + heat A + B In this reaction, a thermo chemical material (C) absorbs energy and is converted chemically into two components (A and B), which can be stored separately. The reverse reaction occurs when materials A and B are combined together and C is formed. Energy is released during this reaction and constitutes the recovered thermal energy from the TES. The storage capacity of this system is the heat of reaction when material C is formed. When energy is released, the reaction is exothermic.

ADVANTAGES OF TES Thermo chemical TES systems have several advantages over other types of TES: • Components (A and B) can usually be stored separately at ambient temperature, after cooling to ambient conditions subsequent to their formation. Therefore there is little or no heat loss during the storing period and, as a consequence, insulation is not needed. • As a result of the low heat losses, thermo chemical TES systems are especially suitable for long-term energy storage (e.g., seasonal storage). • Thermo chemical materials have higher energy densities relative to PCMs and sensible storage media. Because of higher energy density, thermo chemical TES systems can provide more compact energy storage relative to latent and sensible TES. This attribute is particularly beneficial where space for the TES is limited or valuable.

OBJECTIVES OF THE PRESENT WORK The main objective of this paper is to study the feasibility of thermo chemical energy storage system. Compared to sensible and latent thermal energy storage systems, chemical thermal energy storage system has very high energy storage density. Also in the process, thermal energy is stored as chemical energy. Hence thermal energy can be stored in room temperatures, which means there will be no loss due to convection or radiation.

In this thesis, first objective is to select the source of heat used for storing. Even though the first idea was to use solar heat, due to difficulties in obtaining high temperatures and maintaining that temperature it was discarded. Then the waste heat 118

Proceedings of ICETiME’15 December 16-17, 2015, CUSAT, Kochi, India from the exhaust of diesel engine was selected. The next main objective is to select a suitable thermo chemical compound. Chemical compound is selected according to the temperature availability from the source. Chemical compound selection is through a wide literature survey. After selecting the thermo chemical compound, its reactions and enthalpy changes during the reactions are to be studied. Then a reactor for the charging and discharging reactions is to be designed and fabricated. After fabrication, the reactor is to be tested for its thermal storage capabilities under various conditions.

GENERAL CONCLUSIONS FROM THE LITERATURE SURVEY Properties of various thermo chemical compounds were reviewed on various research papers and review papers (Ali H. Abedin and Marc A. Rosen,2011). .MgCl2 6H2O and CaCl2 6 H2O had temperature levels useful for space and water heating and Mg(OH) 2 and Ca(OH) 2 were good for higher temperature levels (> 500 K ). While Ca(OH) 2 showed good reactivity and overall behaviour, but only relevant for high-temperature heat storage. Mg(OH) 2 and the hydrated sulphates showed less sufficient reactivity in the hydration process. The calcium sulphate can be used only for low temperature applications (N.B. Singh,B. Middendorf,2007). The chloride salt hydrates on the other hand showed good reactivity under all circumstances. But they are known to be quite corrosive and show a certain tendency to over-hydrate, developing a gel-like consistency which significantly hinders the ability to store and release heat. Moreover, when using MgCl2 6 H2O, care is to be taken, as the material shows thermal decomposition releasing HCl above certain temperatures. Whereas reviewed papers state that thermal decomposition of MgCl2 6 H2O starts at temperatures above 100-130°C (H.A. Zondag et al , 2010), in their experiments at low pressures. Hence comparing all factors observed from above papers , it is concluded that magnesium chloride is the best material to be used for thermo chemical application in the charging temperature range of 100°C to 130°C.

CHEMICAL REACTIONS INVOLVED Enthalpy data for the reactions were obtained from national institute of standards and technology (NIST), chemistry web book. Charging reactions 1) MgCl2.6H2O => MgCl2.4H2O + 2H2O reaction starts at 70°C , ΔH = +133.94 kJ mol-1 2) MgCl2.4H2O => MgCl2.2H2O + 2H2O 119

Experimental Investigation of a solid… reaction starts at 105°C , ΔH = +30.36 kJ mol-1 Above two are the requird reactions . further dehydration of MgCl2.2H2O is not possible by heating. Further heating starts following reaction, which is irreversible and must be avoided. 3) MgCl2.2H2O => Mg.O.HCl + HCl (g) reaction starts at 130°C Discharging reactions 1.

MgCl2·2H2O + 2H2O => MgCl2·4H2O ΔH = -30.36 kJ mol-1

2.

MgCl2·4H2O + 2H2O => MgCl2·6H2O ΔH = -133.94 kJ mol-1 (NIST chemistry web book)

Where , ΔH is the enthalpy change during reaction. Molar mass of MgCl2·6H2O is 203.3 g/mol Molar mass of MgCl2·2H2O is 131.1 g/mol Total enthalpy change -133.94 kJ mol-1 + -30.36 kJ mol-1 = -164.24 kJ mol-1 Mass of MgCl2·2H2O to test in the reactor is 600g. which is 4.57 moles. Hence total hydration enthalpy is 4.57 mol x -164.24 kJ mol-1 = 751.6 kJ. Hence 600g of MgCl2·2H2O can store 751.6 kJ of thermal energy.

EXPERIMENT SETUP The thermo chemical reaction requires sufficient quantity of moisture to come in contact with the chemical compound. The heat released during the reaction should be transferred to the flowing water. Both these requirements are made possible in thermo chemical reactor system. The reactor fabricated for the experiment is shown in Fig 1. The thermo chemical reactor was made with stainless steel material. Water inlet is supplied through the centre pipe, from which water travels through 6 branches in radial direction, which individually branches into double pipes. The double pipes are enclosed in rectangular trays ,which carries the thermo chemical compound. The branched out water joins into the outer circular pipe, and hot water is taken out from a single outlet. The moisture needed for the thermo chemical reaction is supplied from the top directly to the trays. 120

Proceedings of ICETiME’15 December 16-17, 2015, CUSAT, Kochi, India

Fig 1. Thermo chemical reactor

Fig 2. Tray half filled with MgCl2.6H2O, double pipe passing through tray visible 121

Experimental Investigation of a solid… In first test reactor trays were filled with MgCl2.6H2O. Engine exhaust was connected to the reactor inlet. Engine was started and temperature started building up. In 5 minutes the double pipe wall temperature reached 118°C and dehydration started. After 20 minutes chemical compound around the double pipe were completely dehydrated , but heat was not reaching the bulk in whole tray. The engine was turned off. After a cooling off time of 20 minutes, the discharging performance was tested ,Water inlet at 2 lpm was supplied at inlet and water sprayed on top of dome to start the reaction. hot water temperature was measured at outlet and results were plotted. In the second test , MgCl2.6H2O was dehydrated by heating it on a gas burner. So that the whole compound could be dehydrated evenly to test the compounds storing capability. In this test1 kg of MgCl2.6H2O was heated to obtain 600 g of MgCl2.2H2O. Each tray filled with 100 g of MgCl2.2H2O. Water inlet at 2 lpm was supplied at inlet and water sprayed on top of dome to start the reaction . hot water temperature was measured at outlet. In third test, the dehydrated MgCl2.2H2O was stored for 1 month in an air tight jar, so as to test the storing potential of the compound. The performance was tested for same quantity and method as in second test. And the results were noted down and plotted. By the inferences from second and third tests , fourth test was conducted with an additive mixed with the thermo chemical compound. Mg3Si4O10(OH)2 , hydrated magnesium silicate or also called as talc was added as anti caking agent to improve the heat discharge during hydration. Same quantity of 600 g and same water flow rate of 2 lpm were used in the experiment to compare it with previous performances.

RESULTS AND DISCUSSIONS In first test, MgCl2.6H2O in the tray was heated with the exhaust heat from the engine for 20 minutes. But the whole compound in the tray could not be dehydrated, as the contact area was not sufficient. The compound which was in contact with the double pipe inside the tray was dehydrated to MgCl2.2H2O. Total energy required to dehydrate 930.43g of MgCl2.6H2O into 600g of MgCl2.2H2O is 751.6 kJ. As we could not obtain required dehydration, the percentage of recovered energy was not calculated.

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Fig 3 Temperature of chemical compound inside reactor tray, test 1

Fig 4 Temperature of water outlet from the reactor , test 1 In test 2, MgCl2.6H2O was heated on a gas burner to completely dehydrate it. Dehydration started at 100°C. Above 120°C pungent smell was observed, which was emission of HCl gas. So temperature was maintained between 100°C and 120°C. After dehydration the compound was stored in an air tight jar. After 24 hrs the stored compound was taken out and 600 g of MgCl2.2H2O was filled into 6 trays 123

Experimental Investigation of a solid… 100g each. Water at initial temperature 28°C was supplied to the inlet at 2 lpm. Moist air for reaction was supplied from top of dome with water sprayer.


Fig 6 Temperature of water outlet from the reactor , test 2 A maximum temperature of 80°C was observed inside the tray. For water outlet maximum temperature rise of 7°C was observed after 120s. temperature difference alculate the heat recovered following method of integration was used Total heat recovered , Q = ρ .V .Cp 124

Proceedings of ICETiME’15 December 16-17, 2015, CUSAT, Kochi, India The function of temperature difference to integrate for this test estimated from MS EXCEL is t = 0.2829+ 0.103T –5.73x10-4T2+1.24x10-6T3 –1.23x10-9T4 +4.53x10-13 T5 Q = 0.1385 = 0.1385 x 2728.29 = 377.93 kJ The total heat of hydration being 751.6 kJ, 50.28 % was recovered from the thermo chemical compound In test 3, the long term thermal storage capability of the thermo chemical compound was tested. The dehydrated MgCl2.2H2O 600 g was stored in a closed air tight glass jar for 1 month. Being rainy season, atmosphere relative humidity was 90 – 95 %. After 1 month the compound seemed same but was clogged into small lumps. The compound was filled in reactor for test. Water initial temperature was 29 °C, and flow rate was 2 lpm.

Fig 7 Temperature of water outlet from the reactor , test 3 The maximum tray temperature was only 57 °C, and maximum temperature difference obtained for water flow was 3 °C. 14 minutes ( 840 s) after start became 0. On stirring the chemical compound inside the tray, small lumps of compound were crushed. Again a rise in chemical temperature was observed. There was a corresponding rise in water outlet temperature. After 32 minutes (1920s) from the start of the test water again was 0. This time hydration of the chemical seemed complete. The function of temperature difference to integrate for this test estimated from MS EXCEL is 125

Experimental Investigation of a solid… t = 0.933+ 0.023T – 08.84x10-5T2 +1.12x10-7T3 –5.94x10-11T4 +1.115x10-14 T5 Q = 0.1385 = 0.1385 x 2047.37 = 283.56 kJ The total heat of hydration being 751.6 kJ, 37.72 % was recovered from the thermo chemical compound even after 1 month storage . From test 3 it was observed that , MgCl2.2H2O attracts moisture from atmosphere very fast . It tends to form lumps of crystals. Which during hydration process, prevents the MgCl2.2H2O in the inner part of the lump to be in contact with moisture. Hence prevents or slows down heat release from the chemical compound. Use of anti caking agents can prevent lump formation in powdered substances. One of the most common anti caking agent is hydrated magnesium silicate also called as talc ( Mg3Si4O10(OH)2 ). In test 4 after dehydrating MgCl2.6H2O on gas burner, 600g of MgCl2.2H2O was mixed with 50g of talc and was stored in a glass jar. After 24 hrs the mixture was taken for the test. 108 g on each tray, water initial temperature was 28°C. water flow rate was kept same as other tests, 2 lpm.


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Fig 9 Temperature of water outlet from the reactor , test 4 The function of temperature difference to integrate for this test estimated from MS EXCEL is t = 0.92 + 0.143T –8.25x10-4T2 + 1.84x10-6T3–1.83x10-9T4 + 6.79x10-13T5 Q = 0.1385 = 0.1385 x 3546.4 = 491.17 kJ The total heat of hydration being 751.6 kJ, 65.35 % was recovered from the thermo chemical compound. So addition of talc improved the heat recovery efficiency by avoiding lump formation.

CONCLUSIONS Thermo chemical heat storage materials (TCM) have been subject of research for quite a long time. From the data obtained from various research papers and considering the heat source and requirements, a thermo chemical heat storage system was designed and fabricated. Magnesium chloride gave a storage density of 0.717 GJ/m3. Which is very high compared to ordinary sensible and latent heat storage systems. In first test it was found that the heat exchange from the exhaust to chemical was not efficient, because of smaller contact area. But for the given engines back pressure limitation, adding the number of pipes was not safe. Tests were conducted in the period of April to June of the year 2015 Relative humidity during tests were in the range 90%-94%. Still in long storage period test 37.72% heat was recovered after 1 month storage. As the wall temperature were highly fluctuating and 127

Experimental Investigation of a solid… partially filled pipe flows, conduction/convection heat losses were not calculated separately. Efficiency of the system in various tests were calculated by comparing the heat recovered from system to the total enthalpy of the reaction, that is the total energy stored in the chemical. Heat recovered from system was calculated by integration method, as the water outlet temperatures were varying continuously with time. Test 2 gave an efficiency of 50.28%. From test 2 and 3 ,the problem of lump formation was observed. Hence talc was mixed as anti caking agent. It reduced lump formation, and improved the efficiency drastically. Test 4 with talc additive gave an efficiency of 65.35%.

REFERENCES [1] Ali H. Abedin and Marc A. Rosen “A Critical Review of Thermochemical Energy Storage Systems” Faculty of Engineering and Applied Science, University of Ontario Institute of Technology, Oshawa, Ontario, 2011. [2] H.A. Zondag,V.M. van Essen,L.P.J. Bleijendaal, B.W.J. Kikkert, M. Bakker “Application of MgCl2·6H2O for thermochemical seasonal solar heat storage” 5th International Renewable Energy Storage Conference IRES November 2010, Berlin, Germany [3] a) N.B. Singh, b) B. Middendorf. “Calcium sulphate hemihydrate hydration leading to gypsum crystallization” a) DDU Gorakhpur University, Chemistry Department, Gorakhpur, India ,b) University of Dortmund, Department of Building Materials, Dortmund, Germany.2007 [4] National Institute of Standards and Technology (NIST), chemistry web book

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A NOVEL METHOD TO PARAMETER ESTIMATION IN LAMINAR NATURAL CONVECTION HEAT TRANSFER PROBLEM IN TWO PARALLEL VERTICAL PLATE Anu Nair P1, Aneesh M S2, Premjith S3 and Anoop G Das4 1,2,3,4

Assistant Professor Department of Mechanical Engineering, Gurudeva Institute of Science and Technology, Kottaym-686516,Kerala, India

ABSTRACT Heat transfer using natural convection and fluid flow is an important phenomenon in daily life and engineering applications such as HVAC systems of building, harvesting energy using non-conventional energy sources, extraction of heat from electronic equipment and cooling technologies integrated in computer systems. The present study investigates about the two experiments: (1) Heat transfer taking place between two parallel-Vertical Plates by natural convection, in which two walls are adiabatic and other two ends are open to the ambient air and (2) The estimation of constant C in the expression of Nusselt Number from the data of temperature distribution obtained by performing experiments using a parallel Vertical-Plate by maintaining steady state conditions. Two parallel vertical plates made of Aluminium are allowed to cool in air, to ensure that Lumped capacitance formulation is valid. In the first case the Nusselt and Rayleigh numbers were calculated experimentally and compared with the value available from the correlation which is found to be in good agreement. In the second case, the value of C in the expression of Nusselt Number was calculated using Least Square Residual Method and validated this result by introducing this value into the correlation available. The experimental setup which is required has been designed and fabricated. 1


A Novel Method to Parameter Estimation In Laminar Natural Convection Heat Transfer

INTRODUCTION Natural convection heat transfer occurs in many engineering applications such as extracting heat generated during the working of electronic equipment, cooling of electric transformer , chimneys and furnaces, cooling the reactor core in nuclear power generation to dissipate the heat generated by nuclear fission, cooling of solar collectors and geophysical flows. In these equipment’s, the source of heating, in general is either due to volumetric heat generation or due to surface heat fluxes. For instance, electronic equipment generates heat, which can be expressed in terms of volumetric heat generation. Volumetric heat generation in nuclear fuel rods is due to nuclear reaction. The performance of devices involved in thermal energy conversion depends on the energy exchange that takes place through various heat and fluid flow processes prevailing in these devices. Therefore, research on flow and heat transfer through plate’s demands significant attention. The amount of heat energy transported by the working fluid in a plate is dependent on the geometry of the plate, nature of the flow, and the thermal boundary conditions of the plate. Electronic equipment’s and devices have become essential ingredients in our day-to-day life. Among them the most widely used of these is the electronic computer, ranging in size from the hand-held personal digital assistant to large scale mainframes or servers. In numerous occurrences a computer is imbedded or integrated inside some other electronic devices and is not by any means conspicuous. The use of computers is increased in prominent spheres such as financial, defense, science and technology, banking, health sectors etc. As automation is proliferating into vital infrastructures, a computer failure can bring about a catastrophic interruption of basic administrations and can even have life threatening outcomes. As a result, efforts to improve the reliability of electronic computers are as important as efforts to improve their speed and storage capacity. Various techniques and designs for cooling, that were implemented using air, water, and direct immersion are in line to accommodate the increased heat flux trend in heat removal. They include conduction cooling, air cooling, micro channel heat sink cooling, micro heat pipes, pool boiling, Jet impingement, multi-phase flow etc. Although forced convection and boiling using a liquid medium offer the highest heat transfer rates, air cooling technique has been widely used for heat removal for a long period. The main advantages of cooling using air as fluid are ease of application, abundance and availability by Yoji Kitamura et al.(2005). The simplest method of cooling is by circulating air, facilitating natural convection. Natural-convection method of cooling of electronic equipment continues to be the effective way in their thermal management; because it provides the advantage of low noise and high system reliability and require the least maintenance. The forced convection air-cooled 130


systems, which are widely used, produces the acoustic disturbances in their operating environment and the reliability of the blower, which forms its main part present serious concerns as the air velocity has to be increased to increase the cooling rate. Thus, the interest in natural-convection air cooling is growing to take advantage of the absolute absence of noise and energy savings inherent in that cooling mode by Yoji Kitamura et al.(2005). Buoyancy-induced heat transport phenomena inside the casings of electronic devices have been concentrated on broadly for a long time. To understand this concept, there are several additional effects associated with this process has to be analyzed, such as heat-conduction mechanisms in solids, location of power input into the system ,radiations, vents, three dimensional effects etc. by chen Linhui et al.(2006). More recently, a large number of researchers have incorporated some of these parameters in an attempt to gather reliable data, which are valuable for thermal analysis and design. Elenbass (1942) conducted experimental investigations in laminar natural convection heat transfer using smooth parallel plate vertical channel and reported a detailed study on the thermal characteristics of cooling by natural convection. Wirtz (1982) have considered geometry, with constant heat sources placed over the entire length of the wall. Since the geometry could not simulate discrete placement of chips, the method of placing a number of discrete heat sources over a wall. Bodia and Osterle (1962) conducted numerical analysis on free convection heat transfer for development of boundary layer between parallel isothermal vertical plates and recorded results for velocity, temperature and pressure variation throughout the flow field .The numerical method used is Hybrid Finite Difference Method. Krishnan and Balaji (2004) conducted a synergistic approach to parameter estimation in multimode heat transfer. This paper reports the efficacy of the Least Square Residual Method in parameter estimation when more than one mode of heat transfer is encountered. Oztop et al (2001) carried out numerical investigation of natural convection heat transfer and fluid flow of two heated partitions which was placed within a square enclosure. The left wall and top wall provided Isothermal condition, while the bottom and the right side wall were adiabatic in nature. The two heated partitions were placed at the bottom of the square enclosure at different aspect ratio and the studies were focused on the effect of heights and position of heated partitions. The results were monitored at different Rayleigh number in the range of 104 to 106.The energy and flow equations were solved with TDMA, using finite difference equation, based on the finite control volume approach with non- staggered grid arrangement and SIMPLEM algorithm. Oztop et al (2004) investigated the characteristics of natural convection heat transfer in a square cavity, with a heated plate placed in vertical and horizontal manner. The governing equations were solved with TDMA using finite difference equation based on the finite control volume approach with non131


staggered grid arrangement and SIMPLEM algorithm. Computation was done with Rayleigh number ranging from104 to 106 at different aspect ratios and position of heated plate. Air was used as a working fluid (Pr=0.71).The effect of the position and aspect ratio of heated (Vertical and horizontal) plate on heat transfer and flow in square cavity were analyzed. The result showed that Rayleigh number increases with increase in mean Nusselt number at both vertically and horizontally oriented positions. At higher Ra numbers, when the plate is placed horizontally, heat transfer was found decreased as about 80% less than that of the plate placed in vertical position. Atchonouglo et al. (2008) used the Finite Element Method for an inverse analysis to identify simultaneously the constant thermal conductivity and heat capacity. Boukhattem et al (2013) conducted numerical investigation of natural convection heat transfer in a two-dimensional closed room, containing air, in the presence of a thin horizontal heater plate. The horizontal walls are kept isothermal, while the vertical walls are adiabatic. The flow and energy equations in the room are solved using a finite differential equation based on the finite control volume approach with non-staggered grid arrangement and the SIMPLEC algorithm. A thin horizontal plate is placed inside the square room with an aspect ratio equal to 0.5.The horizontal plate has higher temperature compared to the isothermal wall. Computation of Rayleigh number in the range of 104 to 106 has been performed. The result showed that Rayleigh number increases with increase in heat transfer coefficient which is marked above the heater plate region and this effect is attributed to Chimney Effect. From literature survey it can be inferred that free convection in vertical channel geometry with discrete heat source has drawn considerable attention. This geometrical configuration has physical relevance with respect to electronic chip placement inside an electronic device. This research aims at developing new concepts to find out the thermo physical parameters such as emissivity, thermal conductivity, thermal diffusivity etc. Such an attempt is made in this paper where, the unknown constant ‘C’ in the expression of Nusselt-number is determined from the known values of temperature from different experiments. The goal of the study is to compare the value of constant C in the Nusselt number in parallel Vertical plate with available correlation (Krishnan et al, 2004) using the Least Square Residual Method.

EXPERIMENTAL INVESTIGATIONS MATERIALS Experimental apparatus has been specially designed, fabricated and set-up to carry out investigations on different types of electronic devices. The layout of test- setup consist of an apparatus, Data Acquisition System, Computer, T-Type

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Thermocouple and AC power supply whose schematic diagram is as shown in Fig.1.

Fig.1. Schematic Diagram of Experimental setup The apparatus consists of a heat source placed inside a large rectangular box, which is open at top and bottom to simulate natural convection condition. Rectangular box has a dimension of 500 x 500x 1000 mm. The three sides of the rectangular box are made up of plywood and are supported by slotted LAngle. The fourth side is made up of acrylic sheet to have visibility. The central heat source plate, by itself, is an assembly of two Aluminium plates of dimensions 250 x 50 x 3 mm with a flat heater formed by winding a Nichrome wire over a Mica sheet sandwiched between two plain mica sheets. The surface of the central plates, those are exposed to the ambient air are given a suitable surface treatment like polishing by buffing. On the other side of the aluminium plate, i.e. on the side not exposed to the ambient air, two blind holes of 1.5 mm diameter and 1.5 mm deep are drilled at the points of temperature measurement, into which thermocouples are fixed. The bead of the thermocouples is attached to the slot in the aluminium plate by using thermobond. Thermobond is inorganic low expansion cold setting cement. Thermobond is an ideal adhesive cement for applications which require high resistance to electricity, chemicals and thermal shocks. It is suitable for a service temperature of 12500C. The plates that form the central plate, along with the thermocouples fixed at their respective position and taken out along the grooves for large and small heat source respectively. In these experiments temperatures were measured using T-type thermocouple which can withstand up to 4000C. Thermocouples thus fixed were laid into grooves, milled on the same side as that of the blind holes and taken out of the plate. As it could be seen there are 8 holes on each of the plate which house the screws that are used in fastening the plate together after the heater is sandwiched in between the plates. The whole assembly was highly polished on its outer surface to obtain an emissivity of 0.05. The final plate heater assembly is shown in Fig.4. At the two corners of the final heater assembly, two metal strips are attached, one on each side of the plate heater assembly. 133


The plate heater assembly is thus hung vertically using Teflon rods through thin metal strips. Teflon wire used has 10mm diameter and 5m length. The Teflon rods and the metal strips serve the purpose of minimizing the conduction losses. The Teflon rods in turn are fastened to slot L-Angle.The Teflon rods pass through holes in the slotted L-Angle, above which there is a hexagonal nut, which facilitate the adjustment of the position of the central plate. The lead wire from all the thermocouples is connected to a Data acquisition system. Data acquisition system consists of a temperature scanner having 40 channels; accuracy±3oC.The heater assembly is heated by using a regulated DC power supply, 0-600V, and 0-1.5A. METHODS Least Square Residual Method An inverse problem by definition is one in which the effect is known (or measured) and the cause(s) need(s) to be identified. Let us take a sample example of a person suffering from fever and going to a doctor for treatment.Invariably,the first parameter the doctor checks is the patients temperature(usually the oral temperature).if this quantity is more than 380C then the patient has fever. Fever is the effect and the doctor has to correctly identify the cause-or rather, has to identify the correct cause. The problem is “ill-posed”, as there could be several cause for the fever. The fever could be because of a viral infection, bacterial infection, inflammation or allergy or some very serious underlying disorder. In order to reduce the ill-posedness, the doctor either goes in for additional tests or simply starts treating the fever empirically, by guessing the cause based on his prior knowledge and applies mid-course corrections, if the patient does not feel better in; say 3-5 days’ time. In thermal sciences including radiation heat transfer as in other branches of sciences and engineering, there are several situations where the correct cause needs to be identified from the effect. The effect is usually a temperature or heat flux distribution. The cause we are seeking could be a thermo physical property like thermal conductivity, thermal diffusivity or emissivity or a transport coefficient like heat or mass transfer coefficient or could even be the heat flux(in this case, the effect is usually the measured temperature distribution). A familiar example is the surface heat flux on a vehicle that reenters the atmosphere. The velocities here are terrific and the kinetic energy of the fluid(because of the relative motion between stagnant air and the fast moving vehicle) is “braked” and converted to an enthalpy rise on the outer surface of the vehicle. This phenomenon is frequently referred as “aerodynamic heating” and results in a huge temperature rise on the surface.it is impossible to place a heat flux gauge on the outside surface of the re-entry vehicle as the temperature is of the order of a few thousand kelvin. In view of this, thermocouple measurements are made at convenient locations, where the temperature is “measurable”. From these measurements, one has to correctly estimate the heat flux at the outer surface. This flux is a critical design parameter that decides the cooling strategies. 134


For instance, the outer surface of a re-entry vehicle can be coated with an ablating material with a thickness that ensures the heat flux at the outer surface, is “melted” or more correctly “sublimated” away, thereby protecting the inside of the vehicle. A decision on the thickness of the coating critically hinges on our ability to estimate the heat fluxes. Experimental Procedure The experimentation consists of plate heater assembly which was given definite amount of power supply that was switched off after steady state was reached. Temperature measurements were done at regular interval of five minute. The experiment was carried out under controlled condition; simulating natural convection .The experiment was done for several power supplies. In this method, typically all the properties of the system will be known beforehand or a priori and the temperature response of the system will be sought. However in this case the experimentally measured temperature response is available and to estimate or retrieve the value of C in the Nusselt number

ANALYSIS AND DISCUSSIONS (1) Investigation of Natural Convection over a Vertical Isolated Plate The problem of free convection in a vertical plate was studied experimentally. The analysis was based on the method of calorimetry i.e.



Qrad  A Th4  T4

h

Qconv At T



Nu avg 

hl kf

These results are compared with the standard correlation for vertical channels in laminar region developed by Churchill and Chu as shown in Eq. (1). Nu  0.68 

0.67 Ra 0.25 4

9 9   16 1   0.492     Pr    

…………………. (1) Percentage of error can be calculated by the relation

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% Error 

Nu exp erimental  Nu correlatio

 100 Nu exp erimental All measurements were done after the system reached steady state conditions, which typically took four hours. The experiment was repeated for different power inputs and corresponding temperature were noted. n

The Rayleigh number is varied from 105 to 108 by changing the heat input and the Nusselt number obtained from experiment is compared in Table 1 with those obtained from correlation. A maximum percentage variation of 18.69 at a Rayleigh number of 105 is found. Table1:Comparision of Nusslet Number by Experimental and Correlation at Different Rayleigh Number ln Ra

ln Nu

% variation

Experiment

Correlation

5

1.23

1

11.5

6

1.45

1.21

12

7

1.62

1.46

8

8

1.84

1.57

13.5

Figure 6 shows the Comparison of Nusslet number in an isolated vertical plate. When input power of heater increases convective heat transfer coefficient by experiment increases. When input power increases, it will increase the temperature of the horizontal plate, which increases the heat transfer coefficient and there by increases the Nusselt number. When power input increases, the convective heat transfer coefficient by correlation shows a downward trend. This is because of two reason (1)the horizontal dimension of the heat source is small the viscous forces try to predominate over the buoyant forces, causing a Rayleigh number decreases and therefore decrease in Nusselt number (2) the temperature increases beyond a particular value.

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Fig.2. Comparison of Nusselt number in an isolated vertical plate

Fig.3. Variation of power with surface temperature Fig.7 shows the radiation heat transfer and convective heat transfer is depending strongly on surface temperature of the plate. The energy emitted depends on (1) the temperature of the body and (2) nature of radiating surface of the body. Both Convective and Radiative heat transfer increases non linearly with increase in plate temperature. At low temperature, radiation may be significant. So we expect that radiation heat transfer is directly proportional to temperature (i.e, Q α Tn for free convection, where 1.2 0.5 mm) , which promotes the rapid generation of reaction products [32]. By using sliding average of peak voltage as gap sensing parameter, the constancy and durability of controller are higher [33]. During machining of epoxy glass fibre it was clear that the overcut on hole radius is minimum when the gap between tool and anode is maximum at lower voltage of 55 V. The mrr throughout machining found to be lower at higher gap between the tool and workpiece [34]. Faraday’s law states that the removal of material is directly proportional to the current during machining but there is some contradiction on

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Machining of electrically non-conductive …..

removal of material at narrow IEG up to 15 µm because the polarization voltage is not constant which results in non-uniform entity elimination.

3.3 INFLUENCE OF NATURE OF ELECTROLYTE ON SURFACE ROUGHNESS The perfect control of the spark initiations in the electrochemical discharge machining process has been an exciting problem. The inspection of the processed job with a microscope shows the state of the each discharge striking the work piece layer, mass and dispersal over the surface. The range affected by the ejection is round and shiny [35]. As far as thickness of the work piece is concerned, it has fewer effects on surface roughness related to other process parameter [36]. The improper and insufficient flow rate of electrolyte during machining produces many fins and scattered along the work piece surface causing irregularity on surface [37]. The viscosity of the electrolyte can influence the fabrication of the micro sized channels. The channels with dissimilar forms and sharp ends can be attained by modification in the electrolyte viscosity and depth of the micro channels can be influenced by the tool-work piece gap [38]. As far as machining speed is concerned, the blending of capacitance and electrolyte with saline water gives twice fast machining speed as compare with the normal electrolyte [39]. The rate of electrolysis affects the machined surface which improves with upgrading the electrolytic concentration [40]. For the period of machining, the solution mixed with Sodium Dodecyl Sulfate provides increase in current density and bubbles around the electrode which enables more stable pulse current results in a less taper and a better quality but a little over size hole with a higher engraving speed [41]. The titrated electrolyte flow provides better machining performance and good surface quality with machining parameters of 36v, the electrolyte flow of 4.5ml/min, and 350 μm/min feed rate [42].The solutions mixed with conductive powders has enhanced the machining quality in electrical discharge machining techniques. The key reason behind use of powder mixed electrolyte is the ability of stabilizing discharge current. The models of hydrogen film and charge of single particle are able to explain the role of the elements in the electrolyte which diminutions the spark ignition voltage [43-46]. The spark energy for each discharge pulse decreases due to critical breakdown strength which is reduced by the existence of conductive elements in the hydrogen film. This breakdown is due to local electric field intensification due to attachment on the tool electrode and dynamic particle movements. The dielectric strength of the film is reduced by the presence of conductive particle. In the accepted ECDM process, the nearby focused spark energy causes unevenness of machined

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Proceedings of ICETiME’15 December 16-17, 2015, CUSAT, Kochi, India. surface due to the micro cracks, rupture, and cracking of the work piece. The surface roughness is upgraded because of the abrasives which help in refining the microcracks and melting region which is formed by liberation of thermal erosion [47]. The nature of electrolyte affects the precision of machined hole significantly [48]. By using the conductive particles (graphite powder), there is reduction in micro cracks and breakages on the processed surface of the work piece. There is smoothness in the surface by using 0.5 wt. % and 1.0 wt. % of powder concentration. But as the concentration is taken above 2.0 wt. %, the micro cracks were produced [49]. The use of silicon carbide particle with electrolyte also improves the surface roughness and machining quality [50]. The generation of the hydrogen bubbles which affects the machining in ECDM process can be controlled by the concentration of the electrolyte [51]. 3.4 INFLUENCE OF SHAPE AND DIAMETER OF TOOL ON

PROCESSING OF MATERIAL The shape of the electrode plays a crucial role in machining of the material with better surface finish. The irregular surface and burs that are produced in the micro sliced surface can be effectively minimized by the proper setting of the process parameters [52]. With the use of tapered tool, the consistency of spark generations can be improved and the minor discharges can be suppressed [53]. The rotational tool with controlled feed improves the performance of the machining technique [54]. The small diameter of tool increases the wear rate and the clearance. For a precise hole with trivial quantity of heating damage, it is compulsory to keep frequency higher and duty ratio minor but MRR slightly decreases [55]. By using rotating micro drill as cathode electrode, the vibration does not put any significant effects on process and when a square wave form is applied to actuator to cylindrical rod, the metal removal rate increases [56]. Electrochemical discharge machining process is found suitable to process electrically non-conductive materials but still there are lots of difficulties to be removed. Initially time in machining and then diameter of machined hole increases with increase in depth of machining. The spherical shaped electrode gives better performance than the cylindrically shaped tool electrode. The spherical shaped electrode reaches to a superior machining depth and a smaller hole diameter as matched to cylindrical electrode. The spherical tool electrode has advantages of higher machining efficiency which continues with machining time because of having minimum contact area with the work piece. The inefficient material removal mechanism of the cylindrical tool electrode further impedes machining efficiency which worsens with the time consume in processing. The spherical shape allows the uniform growth of the bubbles which results in increase

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of the discharge frequency. The comparison between machining depth achieved by orthodox cylindrical tool electrode and the proposed spherical tool electrode shows the reduction of 83% in machining time and 65 % in hole diameter [57]. The diameter of the cutting tool also affects the cutting of the work piece. During machining of Al2O3 ceramics on TW-ECSM set up, the width of the micro slice which is initially equal to diameter of tool, increases along the depth of cut [58]. The production or different types of shapes or structures is deeply concerned with the diameter and shape of the tool used to machine the work piece. A number of different structures can be produced by the movement of tool electrode along a given direction [59-60]. These possibilities provides a lot of facilities regarding machining with desired output but with all these advantages there is limitation of decrease in machining rate with machining zone. There is decrease in cell impedance with increase in machining area or tool electrode diameter because of increase in double layer capacitance and decrease in electrolyte resistance. As the inductance effect increases, the rising time of the double layer potential will be longer. It was proved that electrode diameter influences the potential charging of the double layer which follows the equation Vc =1/C ∫i dt. The actual rising time of the potential increases with larger tool electrode but the potential does not rise sufficiently within the short pulse duration. As a result, the electrode diameter is responsible for the decrease in machining rate at similar pulse on time. When 50 ns pulse on-time was given to tool electrodes more than 160 μm in diameter, the machining was practically impossible. Therefore,the successful machining can be obtained by the larger tool electrode with longer pulse on-time [61] 3.5 INFLUENCE OF TREATED ELECTRODE ON MACHINING

PRECISION The machining precision is very important aspect in terms of reaching a good quality product. The tool electrode puts a significant effect on the precise machining of work piece material. By the time, the necessity of gaining superior machining accuracy and machining rate is increasing. The machining accuracy is improved during operation with micro-tool vibration as compared with operation without tool vibration. The trivial pressure waves which are produced in the machining zone terminate the inactive layer of work piece surface which advances the flow of current results in enhanced precision [62]. The stable cutting of glass is possible with the utilization of surface textured tool which decreases the minimum voltage by 10 VDC with enhanced frequency of spark [63]. The heat transmission from the machining zone to the surrounding electrolyte is affected by the thermal conductivity of electrodes which concludes that the less energy discharge will be

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Proceedings of ICETiME’15 December 16-17, 2015, CUSAT, Kochi, India. available for machining of materials for electrode of higher thermal conductivity [64]. The wear rate of electrode is deeply affected by the nature of the electrolyte. The tool wear can be minimized by dropping the electrolyte concentration because heat transmitted to the work piece rises with the raise in concentration of electrolyte which is responsible for the eroding of the tool [65]. For the finest machining performance in terms of the lowest hole diameter and minimum tool wear, tungsten carbide is paramount tool material for quick and accurate micro hoe drilling [66]. In comparison of machining ability of copper electrode and abrasive electrode, the abrasive electrodes enhances the cutting ability because of the presence of abrasive grains which results in better machining ability [67]. The coated electrode can be used in the machining of materials because electrodes with insulating coating layer can produce supreme machining rate as the increasing time of the double layer potential is diminished. The coating material used should not be thick as the thickness of the insulation ranges from angstrom to 50 µm. Enamel can be used as a coating material. The machining rate of the insulated electrode is far higher than that of the bare electrode. The machining depth can be improved because there is no size effect according to the machining depth. Therefore, the rate and gap of machining are uniform irrespective of the machining depth. During the machining of micro hole in Al2O3, the shape of the hole was exactly rounded initially and has same diameter as of electrode but later reduced along the depth of hole [68]. As there was increase in machining gap with machining time in ECM, a cylindrical tool electrode was used and a taper shape was generated [69]. Using the coated tool electrode, the local machining time is constant regardless of the machining depth which results in fabrication of high aspect ratio structures with uniform machining gap.

4. CONCLUDING REMARKS AND FUTURE SCOPE

 



The successive inferences can be drawn with concern to the effects of several parameters on the machining of electrically non-conductive materials: This machining technique is accomplished to machine hard and brittle materials which are non-conductive in nature and very difficult to machine with the help of orthodox machining processes. Significant efforts have been made to understand the phenomenon of the metal removal in the electrochemical discharge machining process. The core aspects which are conferred in the paper are insulating coating over electrode, diameter of the tool, shape of the tool, electrolyte mixed with conductive particles, Inter electrode gap, voltage and inductance. The controlled usage of inductance will increase the machining rate and increased tool diameter will require longer pulse on time. The rate of processing of materials by insulated electrode is far higher than that of the bare tool electrode and taper shape of structures can be effectively prevented.

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   

The applied voltage puts a significant effect on metal removal rate of the work piece. The metal removal rate varies with variation in the applied voltage range and decreases with increase in the inter electrode gap. The spherical electrode achieves a superior machining depth with smaller hole diameter as compared to cylindrical tool electrode with higher machining efficiency. The electrolyte mixed with the conductive powder reduces the micro cracks and fractures on the machined surface of the work piece which results a better surface finish than using traditional electrolyte. In the prior investigations, independent effects of these parameters on machining of materials were discussed. There is lack of analysis of combine effect of the multiple process parameters. The opinions regarding best voltage range are conflicting. The poor surface quality in slicing and cutting of workpieces is a point of further research. The machining ability of the electrolyte can be increased with electrolyte treated with different conductive materials. Also, there is limited research on the variation of shape of tool tip on MRR. These points could be the work for research to attain enhanced machining time, superior surface conditions and amplified metal removal rate in electrochemical discharge machining process.

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COMBINING FTA AND LOPA FOR IMPROVING RISK ANALYSIS STUDIES : A NOVEL APPROACH V. R. Renjith 1, Amal S.George 2 1

Associate Professor,2P.G. Student Cochin University of Science and Technology Kochi- 682 022, Kerala, India.

ABSTRACT Along with the rapid progress of industrialization, the risk of incidents (such as fire, explosion, and chemical release) is increasing as well. The release of the chemical methyl isocyanate in Bhopal in 1984 resulted in a catastrophe leading to thousands of deaths and more than ten thousand people were affected. The results of major industrial disasters can be vulnerable to human, as in the case Flixborough, England, which cost the lives of 28 people. LPG explosion in Mexico city resulted in hundreds of deaths and several thousands of injuries Therefore, it is essential to estimate the potential risk of hazardous process using various tools such as HAZOP, FMEA, FMECA, FTA, ETA, LOPA and QRA. Some of these techniques are qualitative in nature, whereas others are quantitative techniques. All the techniques have its own advantages and disadvantages. LOPA is the new methodology used for hazard evaluation and risk assessment. The LOPA methodology lies between the qualitative and quantitative end of the scale. LOPA can provide quantified results with less time and effort than other methods like QRA. For LOPA application, failure data are essential to compute the risk frequencies. But the application of LOPA becomes constrained when failures are compound and safety systems are integrated. In this work Fault Tree Analysis (FTA) is combined with LOPA to eliminate this drawback. This combined novel approach can be applied to common cause failures in all the LOPA analysis.

1


Combining FTA and LOPA for Improving Risk analysis studies …

INTRODUCTION Layer of protection analysis (LOPA) is a powerful analytical tool for assessing the adequacy of protection layers used to mitigate the risk involved in the process. It is a semi quantitative analysis of hazards that evaluates the frequency of the cause/s and the probability of failure of the protective layers. It was developed to determine the Safety Integrity Level (SIL) of Safety Instrumented Functions (SIF). LOPA is based on the concept of protective layers. In order to prevent the occurrence of an undesired consequence, a protective barrier is implemented. If this barrier works well, no more protection layers are required. However, there is no perfect protection barrier and several are needed to reduce the risk to tolerable levels. LOPA is useful to reduce the risk of a process to a tolerable level through the analysis of independent protection layers (IPLs). IPLs satisfy the criteria of specificity, independence, dependability and auditability. An independent protective layer (IPL) is defined as a device, system or action that is capable of preventing a scenario from proceddingto its undesired consequence, independent of the initiating event or the action of any other layer of protection associated with the scenarios [1]. CCPS [1] provides the required characteristics for IPLs as independence, Functionality, Integrity, Reliability, Auditability, Success surety, Management of change. Examples of IPLs are controls, alarms, Procedures, training and safety instrumented functions. An IPL has to be independent of other protection layers available against an undesired consequence, and also has to be independent of the initiating cause. The criterion of specificity indicates that an IPL detects, prevents or mitigates the consequences of specific hazardous events. Additionally, an IPL reduces the risk by a known and specific amount and it is designed to allow auditing the protective function [2]. Fault tree analysis (FTA) is a powerful diagnostic technique used widely for demonstrating the root causes of undesired events in a system using logical, functional relationship among components, manufacturing process, and subsystems [3,4,5]..

LOPA – An Overview LOPA is a simplified form of risk assessment. LOPA uses initiating event frequency, consequence severity, and the likelihood of failure of independent protection layers (IPLs) to approximate the risk of a scenario. LOPA is an analysis tool that typically builds on the. information developed during a qualitative hazard evaluation, such as a process hazard analysis (PHA). LOPA is implemented using a set of rules. Like many other hazard analysis methods, the primary purpose of LOPA is to determine if there are sufficient layers of protection against an accident scenario[6]. A scenario may require one or more protection layers depending on the process complexity and potential severity of a consequence. Note that for a given 334


scenario; only one layer must work successfully for the consequence to be prevented. However, since no layer is perfectly effective, sufficient protection layers must be provided to reduce the risk of the accident tolerable. LOPA provides a consistent basis for judging whether there are sufficient IPLs to control the risk of an accident for a given scenario. If the estimated risk of a scenario is not acceptable, additional IPLs may be added. LOPA does not suggest which IPLs to add or which design to choose, but it assists in judging between alternatives for risk mitigation. LOPA is not a fully quantitative risk assessment approach, but is rather a simplified method for assessing the value of protection layers for a well-defined accident scenario. Fig. 1 illustrates the different layers of protection. LOPA provides a risk analyst with a method to reproducibly evaluate the risk of selected accident scenarios. A scenario is typically identified during a qualitative hazard evaluation process such as a PHA, HAZOP or design review. LOPA is applied after an unacceptable consequence, and a credible cause for it, is selected. It then provides an order of magnitude approximation of the risk of a scenario. Another way to understand LOPA is to view it relative to quantitative risk assessment (CPQRA). In this context, a LOPA scenario represents one path (typically we choose the path to the worst consequence) through an event tree. Fig.2 shows an event tree for a given initiating event. An event tree shows all the possible outcomes (consequences) of an initiating event. LOPA is normally applied after a qualitative hazard evaluation (e.g., PHA/HAZOP) using the scenarios identified by the qualitative hazard review team. However, “typically” means just that LOPA can also be used to analyse scenarios that originate from any source, including design option analysis and incident investigations. LOPA can also be used as a screening tool prior to a more rigorous quantitative risk assessment (CPQRA) method. When used as a screening tool, each scenario above a specified consequence or risk level will first go through LOPA analysis, and then certain scenarios will be targeted for a higher level of risk assessment. The decision to proceed to CPQRA is typically based on the risk level determined by LOPA or based on the opinion of the LOPA analyst.

335


Fig . 1 Layer of protection against a possible accident (Courtesy: CCPS /AICHE – Layer of protection analysis : A simplified risk

Fig.2 Event tree and IPLs (Courtesy: CCPS /AICHE – Layer of protection analysis : A simplified risk assessment)

336


Fig.3 Explain the spectrum of risk assessment tools: from purely qualitative to the rigorous application of quantitative methods. On the far left are qualitative tools; these are typically used to identify scenarios and qualitatively judge if the risk is tolerable. In the middle are semi-quantitative tools (or simplified quantitative tools); these include LOPA and are used to provide an order of magnitude estimate of risk. LOPA can be applied in a team setting, such as during or immediately following a HAZOP- or What-If based review used to identify accident scenarios. LOPA can also be applied by a single analyst; in this case, the scenarios have typically already been identified for the analyst (such as by a HAZOP team ). Fig. 4 shows the relation between HAZOP and LOPA.

Fig. 3Tools used for risk – based decision making. (Courtesy: CCPS /AICHE – Layer of protection analysis : A simplified risk assessment)

337


Fig. 6 HAZOP and LOPA (Courtesy: CCPS /AICHE – Layer of protection analysis : A simplified risk assessment)

Advantages of LOPA Some general benefits of LOPA include: LOPA requires less time than quantitative risk analysis. LOPA facilitates the determination of more precise cause–consequence pairs, and therefore improves scenario identification. LOPA helps identify operations and practices that were previously thought to have sufficient safeguards, but on more detailed analysis (facilitated by LOPA), the safeguards do not mitigate the risk to a tolerable level. 338


Information from LOPA helps an organization decide which safe-guards to focus on during operation, maintenance, and related training. Limitations of LOPA One limitation is that failure data required for a LOPA are generally available for component failure and human error failure [7,8], although many failures are so complex that there are multiple combinations of these basic failures. So it becomes extremely difficult to apply the conventional LOPA method. Secondly, in LOPA protection systems are taken as independent layer of protection (IPL) which satisfies conditions independent, dependable, auditable as per [9]. But in many cases the criteria of independence is not satisfied as protection layers are integrated or coincides with one another. It will be difficult to apply LOPA in the above scenarios [10]. FTA- LOPA combined approach approach

LOPA is a quick semi quantitative technique which provides results with less time and effort than other QRAs. For LOPA application failure data are essential to compute the consequence frequencies. One limitation is that failure data required for a LOPA are generally available for component failure and human error failure [7, 8], although many failure are so complex that there are multiple combinations of these basic failures. So it becomes extremely difficult to apply the conventional LOPA method. Secondly, in LOPA protection systems are taken as independent layer of protection (IPL) which satisfies conditions independent, dependable, auditable as per [9]. But in many cases the criteria of independence is not satisfied as protection layers coincides with one another. It will be difficult to apply LOPA in the above scenarios [10]. Fault tree analysis (FTA) can be integrated into LOPA to eliminate the above mentioned drawback. Fault tree analysis (FTA) is a widely used tool for system safety analysis [11,12]. It is a deductive (backward reasoning) logic technique that focuses on one particular hazardous event (e.g. Toxic gas release, explosion, fire, etc.) and provides a method for determining the causes of hazardous events. The basic process in the technique of FTA is to identify a particular effect or outcome from the system and trace backward into the system by the logical sequence to prime cause(s) of this effect [12,13]. This helps in analysing complex failures. But using FTA for analysing an entire process is Herculean task. So FTA application is limited to PFD calculations. Again, if pre-solved fault trees could be used by the analyst those can be inserted into LOPA even faster computing of consequence frequency. FTA when integrated into LOPA can be used in complex systems and integrated layer of protections. This 339


improves versatility of LOPA as a risk assessment tool without losing its swiftness and simplicity. Case study Result and discussions

Fig. 1 Flow diagram of tank filling

Fig. 2 FTA for tank –over fill 340


Fig. 1 Shows a flow diagram of a tank-filling operation. Once in a quarter a tank truck from a petrochemical industry arrives and refills the storage tank. Trained operators observes the level gauge and close the valve when the levels reaches the desired level. A Safety instrumented system consists of a sensor, logic solver and a shutdown system is also provided as a secondary protection. These two protective systems prevent the over fill of the storage tank. Otherwise there will be an overspill of toxic and flammable liquid. A fault tree is constructed (Fig.2) taking tank overfill as the top event. Intermediate events are connected with appropriate logic gates. All the intermediate events are connected with basic events with suitable logic gates. From Fig. 2 it is observed that there is an interdepence between two IPLs ( intermediate event 1 and 2). Thisis because they both IPLs relay on the level gauge.

CONCLUSIONS LOPA cannot be useful when the failures are interdependent.Many failures are so complex that there are multiple combinations of these failures. So it becomes extremely difficult to apply the conventional LOPA method. In LOPA protection systems are taken as independent which satisfies conditions independent, dependable, auditable etc. If FTA and LOPA are integrated, we could avoid the limitations of LOPA in dealing with compound failures and dependant failures.

REFERENCES Center for Chemical Process Safety, American Institute of Chemical Engineers, Layers of Protection Analysis : Simplified Risk Assessment, New York (2001). A.M. Dowell III, Is it really independent protection layer, Process safety progress Vol 3. No. 2 pp126-131.AIChE F.I. Khan, S.A. Abbasi, Analytical simulation and PROFAT II: a new methodology and a computer automated tool for fault tree analysis in chemical process industries, J. Hazard. Mater. 75 (2000) 1 – 27. W. Hu, A.G. Starr, A.Y.T. Leung, Operational fault diagnosis of manufacturing systems J. Mater. Process Technol. 133 (2003) 108 – 117. H.X. Li, M.J. Zuo, A hybrid approach for identification of root causes and reliability improvement of a die bonding process - a case study, Reliab. Eng. Syst. Saf. 6 (1999) 43 – 48. V. R. Renjith, Risk Assessment of LNG Storages using LOPA and FTA: An Integrated Approach,proceeding of the 18th international anaual sysmposium, Texas A & M university , Texas, USA, 2015. 341


Robert W. Johnson Beyond-compliance uses of HAZOP/LOPA studies. J.Loss Prev. Process Ind. 23, (2010) 727-733. ILO-NSC, Major hazard control- a practical manual, International Labor office, Geneva, 1996. Sam Mannan, Lees Loss prevention in ed.,Elsevier,2005.

the

process

industries,

Vol. 1,third

Geun Woong, Young, William J. Rogers, M. Sam Mannan.) Risk assessment of LNG importation terminals using the Bayesian–LOPA methodology. J. of Loss Prevention in the Process Ind. 22, (2009)91–96. M.H. Shu, C.H. Cheng, J.R. Chang, Using intuitionistic fuzzy sets for fault-tree analysis on printed circuit board assembly, Micro electron Reliab. 46 (2006) 2139 – 2148. V.R. Renjith et. Al. Two-dimensional fuzzy fault tree analysis for chlorine release from a chlor-alkali industry using expert elicitation, J. Hazard Mater. 83 (2010)103– 110. Dong Yuhua, Yu Datao, Estimation of failure probability of oil and gas transmission pipelines by fuzzy fault tree analysis. J. Loss Prev. Process Ind. 18 (2005) 83 –88.

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ELECTRON BEAM WELDING OF TITANIUM ALLOY- Ti6Al4V M. Gopalakrishna Pillai1, P.S Sreejith2, Bhanu Pant3, R.K Gupta4 1

Research Scholar,2Professor Cochin University of Science and Technology Kochi- 682 022, Kerala, India. 3,4 Vikram Sarabhai Space Centre Thiruvananthapuram, Kerala, India

ABSTRACT Ti6Al4V is an alpha beta alloy which is the workhorse material for the present day application in Aerospace area. This alloy has an excellent combination of strength, corrosion resistance, weld and fabricability. Many welding methods have already been developed for this alloy. Because of its high chemical activity, Electron beam welding (EBW) found to be the most suitable welding process. There are many factors which influences the weld quality and joint efficiency. The weld heat input which depends on welding speed, beam current, accelerating voltage, cleaning etc influences this. This paper aims at bringing out the heat input variation, tensile properties and microstructure of Ti6Al4V welds over a range of thickness.

INTRODUCTION Titanium alloys are the workhorse material in the area of Aerospace. Ti6Al4V is an alpha beta alloy which are metastable and include combination of both alpha and beta stabilizers - typically 6% Aluminium and 4% Vanadium respectively by weight. This grade has an excellent combination of strength, corrosion resistance, weld and fabricability (Irving Bob 1994). Many welding methods such as gas tungsten arc welding (GTAW), plasma welding, electron beam welding (EBW) and diffusion welding have already been developed (Irving 1994; Keshava Murthy and Sundaresan 1998; Qi et al 2000). Because of their high chemical activity, titanium 1


Electron Beam Welding of Titanium Alloy – Ti-6Al-4V

alloys readily absorb harmful gases (oxygen, hydrogen and nitrogen) and many problems such as low mechanical properties and unstable structures would appear. Electron beam welding, which offers a number of advantages and characteristics such as high energy density, narrow heat scope, rapid cooling rate over conventional joining process and also protect joints from gaseous contamination. In the present study, Ti–6Al–4V alloy plates of different thickness were welded by electron beam welding, its weld parameters were compared, tensile properties evaluated and the microstructure were analyzed and compared.

EXPERIMENTAL INVESTIGATIONS MATERIALS Ti6Al4V plates of size 100 x 125mm with different thickness were used for the study. 2 plates of this size is joined together to form a welded coupon of size 200x125mm. The chemical composition of the plates chosen is listed below: Table 1 Chemical Composition of Ti6Al4V samples C 0.0120.020

O 0.16 – 0.17

N H V Fe 0..005 – 0.00144.05 – 0.040.006 0.0044 4.14 0.05 Residual elements: each

data reconciliation and gross error detection in

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