Volume 21 Preprint 18
An Apparatus for the Study of Corrosion and Tribocorrosion of Metals Induced by the Flow of Highly Turbulent Aqueous Solutions
R.W. Hendricks, P.K.Todoroff, R.C. Taylor, W.C. Wenger and E.A. Cothron
Keywords: corrosion, corrosion loop, erosion, erosion-corrosion, tribocorrosion, flow accelerated corrosion, flow-related materials degradation, high turbulence flow, corrosion instrumentation, loss factors, fl
An apparatus to study tribocorrosion of materials in both laminar and highly turbulent flow of aqueous solutions under reliably and accurately controlled environmental conditions for long durations is reported. The apparatus simulates single direction, circulating turbulent fluid flow, as opposed to common laboratory â€œbathâ€ style configurations through the use of a custom-built corrosion resistant loop that comprises 11 m of DN 50 CPVC piping through which aqueous solutions are pumped via either a primary 3.7 kW centrifugal impeller pump or a secondary 54 W diaphragm pump. There is a 3.5 kW resistance heating element and a heat exchanger to provide a wide operating range of conditions to conduct environmental simulations to test both material performance, growth of passivation layers, and mechanisms of degradation. Flow rates ranging from 0-10 L/s and temperatures from 11 Â°C to 80 Â°C may be readily investigated. These flow rates correspond to Reynolds numbers from 200 to 300,000 at room temperature (900,000 at the maximum operating temperature of 80 Â°C). Environmental, hydrodynamic, and material performance properties are measured through a wide variety of in-situ sensors that are monitored and recorded via a data acquisition routine written as part of the apparatus monitoring system. Instrumentation includes three digital pressure gages for monitoring the line pressure at critical locations, a turbine flow sensor for monitoring the flow rate, as well as a dissolved oxygen sensor, two pH sensors, two conductivity sensors, and several thermocouples mounted at various locations to monitor solution chemistry. Two identical experimental bays allow for a wide variety of testing sections including a custom-designed universal sample chamber to study inserted orifice plates simulating valves and flow conditioners, commercial corrosion probes interfaced with commercial pipe components, or custom sensors to study the wall shear to remove protective coatings. The apparatus approaches tribocorrosion from an interdisciplinary approach by using computational simulations and measured conditions of hydrodynamics and electrochemistry to perform a comprehensive study with in-depth characterization of the environment. Performance of the apparatus is shown through an expansive study of its fluid dynamics behaviour including calibration of the relationship between the pump motor frequency and the volumetric flow rate and the effect of pressure thereon, the effect of pumping speed on the temperature of the fluid and the characteristics of the heat exchanger required for its control, the loss factors of the various components of the loop, and the flow factors of various experiments and the relationship between the loss and flow factors.
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An Apparatus for the Study of Corrosion and Tribocorrosion of
Metals Induced by the Flow of Highly Turbulent Aqueous
R. W. Hendricks, P. K. Todoroff, R. C. Taylor, W. G. Wenger and E. A. Cothron
Materials Science and Engineering Department, Virginia Polytechnic Institute and State
University, Blacksburg, VA 24061, USA,
An apparatus to study tribocorrosion of materials in both laminar and highly turbulent flow
of aqueous solutions under reliably and accurately controlled environmental conditions for
long durations is reported. The apparatus simulates single direction, circulating turbulent
fluid flow, as opposed to common laboratory “bath” style configurations through the use of
a custom-built corrosion resistant loop that comprises 11 m of DN 50 CPVC piping through
which aqueous solutions are pumped via either a primary 3.7 kW centrifugal impeller pump
or a secondary 54 W diaphragm pump. There is a 3.5 kW resistance heating element and a
heat exchanger to provide a wide operating range of conditions to conduct environmental
simulations to test both material performance, growth of passivation layers, and
mechanisms of degradation. Flow rates ranging from 0-10 L/s and temperatures from
11 °C to 80 °C may be readily investigated. These flow rates correspond to Reynolds
numbers from 200 to 300,000 at room temperature (900,000 at the maximum operating
temperature of 80 °C). Environmental, hydrodynamic, and material performance properties
are measured through a wide variety of in-situ sensors that are monitored and recorded via
a data acquisition routine written as part of the apparatus monitoring system.
Instrumentation includes three digital pressure gages for monitoring the line pressure at
critical locations, a turbine flow sensor for monitoring the flow rate, as well as a dissolved
oxygen sensor, two pH sensors, two conductivity sensors, and several thermocouples
mounted at various locations to monitor solution chemistry. Two identical experimental
bays allow for a wide variety of testing sections including a custom-designed universal
sample chamber to study inserted orifice plates simulating valves and flow conditioners,
commercial corrosion probes interfaced with commercial pipe components, or custom
sensors to study the wall shear to remove protective coatings. The apparatus approaches
tribocorrosion from an interdisciplinary approach by using computational simulations and
measured conditions of hydrodynamics and electrochemistry to perform a comprehensive
study with in-depth characterization of the environment. Performance of the apparatus is
shown through an expansive study of its fluid dynamics behaviour including calibration of
the relationship between the pump motor frequency and the volumetric flow rate and the
effect of pressure thereon, the effect of pumping speed on the temperature of the fluid and
the characteristics of the heat exchanger required for its control, the loss factors of the
various components of the loop, and the flow factors of various experiments and the
relationship between the loss and flow factors.
Keywords: corrosion, corrosion loop, erosion, erosion-corrosion, tribocorrosion, flow
accelerated corrosion, flow-related materials degradation, high turbulence flow, corrosion
instrumentation, loss factors, flow factors
Table of Contents
Note to Reader
Performance and Calibration
Turbulent Fluid Flow
Effect of Backpressure
Design and Construction
Laminar Fluid Flow
Loop Design Considerations
System Equilibrium Temperature
Electronic Control of Fluid Flow
Heat Exchanger Performance
Fluid Temperature at Sample
Thermal Properties of the Loop
Materials Test Equipment
Loop Component Loss Factors
Commercial Corrosion Test Samples
Experiment Loss Factors
Corrosion/Erosion with Orifices
Effect of Backpressure (reprise)
Pipe Wall Thickness Measurement
Device Loss Factor Measurement
Apparatus Monitoring System
System Block Diagram
Compact DAQ-9188 Modules
SMTP Communication Updates
III. Least squares fits to observed data
II. AMS Subroutine Descriptions
Emergency Safety Features
IV. Flow factors and loss factors
V. List of acronyms, alloys, pipe sizes
VI. List of Variables
Note to Reader
This paper makes extensive use of hyperlinks. All hyperlinks are blue, by convention, with
the exception that the underlines have been removed to improve clarity of the text. At the
highest level, each section in the table of contents (TOC) is linked to the referenced section.
Additionally, within the manuscript each section title is linked back to the TOC. Many of the
graphs of experimental data require least squares fits to the observed data. To eliminate
clutter on the graph, the least squares equations have been collected in a single table in
Appendix 3. Each graph is linked to the table and each equation has a return link to its
respective graph, thus allowing the reader to examine the equations at will and return to
the graph without thumbing through numerous pages of text. In several locations,
references (and accompanying links) are made to related sections of the paper but there are
no return links to the original location. Finally, we have provided external links to all of the
vendor-specific information and to numerous of the References. We hope that these links
make it easier for the reader to navigate within the details of the paper.
Materials dynamically interact with their environment with serious consequences for those
that are employed in harsh environments. Industries that involve materials for applications
such as gas and oil pipelines, nuclear power plant coolant loops, and naval heat exchangers
and propellers present highly aggressive conditions that degrade materials significantly.
Due to the nature of these applications, degradation poses both hazardous conditions to
human health and huge financial stresses for systems monitoring and upkeep. While some
materials may be nearly inert to a specific environment, many form a protective passivation
layer as an environmental response to retard further degradation. There is a complex
synergistic response that involves both electrochemical interactions at the molecular level
and mechanical interactions between a material and its environment at a macroscopic level.
In aggressively oxidizing or abrasive environments, this passivation film may be completely
removed; the adhesion between the passivation film and the substrate is often the point of
failure. Many such mechanisms may be considered to be yet another form of tribocorrosion
that can be added to the list provided by Ponthiaux, Wenger, and Celius (2012).
Worldwide, there are a few systems that have been developed to study materials
degradation from either the electrochemical or the mechanical abrasion perspective (e.g.,
Det Norske Veritas; Idaho National Laboratory; Institute for Energy Technology Norway;
Massachusetts Institute of Technology; Oak Ridge National Laboratory; Ohio University;
University of Michigan; Universiti Teknologi PETRONAS (UTP); University of Tulsa). These
facilities are major, expensive installations, constructed to serve a variety of research and
industrial needs. In contrast to these facilities, the instrument described here is oriented
towards developing the experimental data necessary for creating an understanding of the
complex interactions of fluid dynamics, the reactant diffusion, and the growth, spalling,
and regrowth of protective oxides as occurs in tribocorrosion in highly turbulent fluids. The
experimental data required to study these complex interactions include characterization of
both the fluid environment and material degradation from both in-situ and invasively
conducted micro- and macroscopic perspectives.
It is the purpose of this paper to describe the design, construction, and performance
behaviour of an instrument that fills this gap and meets these requirements. This effort
builds on the preliminary experimental apparatus designed and built in our laboratory
during the past several years (Cothron, Taylor, and Todoroff 2016; Cothron et al. 2017;
Lones et al. 2017) and on further improvements made during the course of our current
Design and Construction
Loop Design Considerations
The fundamental design of the apparatus can be divided among several different
categorical groups; structural, environmental housing, environmental control, experimental,
and user interface. Figure 1 presents a schematic of the apparatus while Figure 2 shows
associated images that illustrate the complexity and relationship between each group. With
the intention of developing an environment for the understanding of the degradation of
materials in a variety of aqueous media, commercially available, easily replaceable,
inexpensive, and broadly corrosion resistant materials of construction were chosen. A
structural skeleton of welded steel U-channel and Unistrut™,1 allows for inexpensive, nondestructive future modifications to the apparatus to accommodate additional desired
components. The structure primarily provides a platform to support the environmental
housing and associated control devices, experiments, and the user interface panels and
associated electronics. The environmental housing was composed of DN 50 (51 mm
Unistrut, a part of Atkore International: P1000 Channel, (http://www.unistrut.us, accessed 15-Feb2018).
Figure 1: Schematic of the loop.
Figure 2: The Virginia Tech high turbulence corrosion loop: (a) overview of
the system; and (b) view looking down the loop. Fluid enters from the lower
left and returns on the right. Heat exchanger is silver tube in outbound line
on left. (images courtesy LeeAnn Ellis)
diameter), chlorinated polyvinyl chloride (CPVC) pipe because of its low cost, high
temperature capability, strength, ease of “welding” and high resistance to chemical attack
(Boedecker Plastics; Knight).2,3 A total linear length of 10.8 m of pipe was utilized with
commercial fittings to attach sensors and join sections. Two sections of fabric-reinforced,
flexible PVC tubing reduced stress at several junctions, thus allowing for a decreased
pressure drop and therefore a higher performance of the apparatus. Environmental control
was composed of coupled actuators and sensors for various variables with instituted safety
feedback systems (whether automated or manual). The apparatus includes two identical
experimental bays installed in series that allow for a wide variety of customized
experimental sections. The experimental bays are connected to the apparatus via a ball
valve and union configuration, allowing for the sections to be removed easily without
having to completely drain the apparatus of fluid. The user interface has both manual and
computer-controlled sections. Fluid flow parameters (flow rate; temperature) are controlled
via a manual control panel while an apparatus monitoring routine (AMS) allows for
computer-controlled data acquisition of all environmental and material sensors remotely
with emergency shutdown features instituted.
For turbulent flow, the fluid is pumped from a 38 L stainless steel mixing tank by a six-
vane impeller that is driven by a 3.7 kW, 3-phase 247 V (5 HP) motor.4 The mixing tank, to
which the recirculating fluid is returned, is a custom-made stainless steel tank with four
ports for the connection or insertion of (i) the impeller pump, (ii) a 3 kW heating element,
(iii) a 54 W diaphragm pump for to provide low-velocity laminar flow, and (iv) a drain valve
from which fluid samples for chemical analysis may also be taken.5 Power for the main
pump motor is provided by a 15 kW autotransformer. The angular velocity of the motor is
controlled by a 7.5 kW AC variable frequency drive (VFD) that allows pumping speeds from
See, e.g., PVC Pipe Supplies, Olive Branch, MS; (https://pvcpipesupplies.com; accessed 01-Feb-
Spears Manufacturing Co, Sylmar, CA: CPVC-24 Gray/Orange Low VOC CPVC Solvent Cement,
Technical Specifications; (http://www.spearsmfg.com/solvent_cement_specs/CPVC24TS%20Technical%20Specifications_web.pdf; accessed 01-Feb-2018).
American Stainless Pumps, Los Angeles, CA: Model 2SSP/C C24660B5D3F SSP/C Centrifugal Pump,
(http://www.aspumps.com/products/sspc/; accessed 01-Feb-2018).
Indco, Inc., New Albany, IN: Model 22K1000 stainless steel mixing vat,
(https://www.indco.com/ten-gallon-stainless-steel-mixing-vat; accessed 01-Feb-2018).
approximately 0.32 to 10 L/s.6 For safety of operation, the power to the motor is switched
through a non-reversing definite purpose magnetic contactor that is controlled via a 24 V
For some materials such as Cu-Ni alloys used in marine heat exchanger applications, it is
necessary to pre-treat the inner diameter of the tubing under carefully controlled oxygen
and laminar flow conditions to develop a corrosion-resistant oxide layer (Tuthill 1987). To
accommodate such needs, a secondary 54 W diaphragm pump is installed auxiliary to the
primary 3.7 kW centrifugal impeller pump (see Figure 1).7 A series of ball and ball check
valves isolate the auxiliary pump from the primary pump, enforcing a unidirectional flow in
the loop during the operation of either pump. A 0-12 Vdc, ∆V=0.01 voltage regulator
allows for high-resolution control of the laminar flow.8
Passive heating of the system due to the turbulent pumping as found at low temperatures
and high flow rates is removed by a custom-designed 1.22 m long dual-pipe heat
exchanger that is installed immediately following the main pump to provide active cooling.9
Its fluid channel is a 51 mm diameter, copper-jacketed stainless steel tube, while the
exterior (coolant) channel is a 76 mm diameter concentric copper-jacketed insulated
stainless steel cylinder that is cooled with 70 psi, 10 °C water. The cooling water flow is
controlled from 0.0 to 0.6 L/s by an electronic valve and proportional-integral-differential
(PID) controller.10,11 For those cases where energy must be added to achieve a desired
Hitachi America Ltd., Industrial Components and Equipment Div., Tarrytown, NY; Model WJ200-75
10 HP AC variable frequency drive (VFD), (http://www.hitachi-america.us/ice/ac-drivesinverters/wj-200-series; accessed 01-Feb-2018).
Pentair SHURFLO, Costa Mesa, CA: Model 8009-541-236 12Vdc Diaphragm pump,
pump - current; accessed 08-Jan-2018).
TekPower US: Los Angeles, CA: Model TP3005T 0–30 V, 0–5 A DC regulated power supply,
(http://tekpower.us/tp3005t.html; accessed 01 Feb 2018).
Fluid Chillers, Inc. Lansing, MI: Pipe Heat Exchanger, Model VT021816,
Belimo America, Danbury, CT: Model B2050VS-02+LF24-MFT valve,
(https://www.belimo.us/shop/en_US/config?code=B2050VS-02%2BLF24MFT+US&siteName=Belimo+US+Official+Site; accessed 30-Jan-2018 ).
temperature (high temperatures at low flow rates), a 3 kW heating element mounted in the
recirculating tank provides active heating.12
To assure that the system pipe remains full at all times, a 0.6 m high vertical loop of CPVC
flexible tube was inserted between the downstream pressure gage (P4) and the recirculating
tank. This increased the load on the pumps by 5.9 kPa (see section on Performance and
The system is fully instrumented as indicated in Figure 1. The fluid flow is monitored by a
turbine flow meter that is mounted in-line,13 while dissolved oxygen, pH, and fluid
conductivity sensors are mounted in bushings that are inserted into the auxiliary ports of
inline tees or crosses.14 Each device has a built-in thermocouple for temperature
compensation. The fluid pressure at the four ends of the straight sections is measured with
three digital pressure gages and one (P2) analog Bourdon gage.15 The temperature of the
fluid in the recirculating tank is measured with a platinum resistance thermometer (PRT).16
This PRT is used as the control signal for a second, independent PID controller for the
Watlow Electric Manufacturing, Winona, MN: Model PM6C2FC-ARFCDAA EZ-Zone PM Temperature
Controller, (http://www.watlow.com/products/controllers/integrated-multi-functioncontrollers/ez-zone-pm-controller; accessed 30-Jan-2018)
Omega Engineering, Norwalk, CT: ARMTS-2305/208 water immersion heater,
(https://www.omega.com/pptst/ARMTS2_HEATER.html; accessed 30-Jan-2018).
Omega Engineering, Norwalk, CT: Model FTB-1441 liquid turbine flow meter with Model FTB-
1400-RD-A flow monitor digital display, (https://www.omega.com/pptst/FTB1400_SERIES.html;
Omega Engineering, Norwalk, CT: Model DOE-45PA dissolved oxygen sensor with Model DOTX-45
dissolved oxygen transmitter, (https://www.omega.com/pptst/DOTX45.html ); PHE-45P pH
electrode with Model PHTX-45 pH transmitter, (https://www.omega.com/pptst/PHTX45.html );
and CDE-45P conductivity sensor with Model CDTX-45 conductivity transmitter,
(https://www.omega.com/pptst/CDTX45.html; accessed 30-Jan-2018 ).
Omega Engineering, Norwalk, CT: Model DPG409 digital pressure gage with analog output,
(https://www.omega.com/pptst/DPG409.html; accessed 30-Jan-2018)
Omega Engineering, Norwalk, CT: Model PRCTL-2-100-A-1 platinum resistance thermometer
(https://www.omega.com/pptst/PRCTL.html; accessed 30-JAN-2018).
heating element.17 Four type-K thermocouples are surface-mounted to the exterior of the
inlets and outlets of the heat exchanger and are used to monitor its efficiency.
The loop contains two 1.52 m experimental bays as shown in Figure 1 where custom-
designed experiments may be inserted inline. These bays are isolated from each other
through valves in such a way that one may work on various sections of the loop or change
experimental sections without completely draining the system. Design of several such
modular sample chambers and experimental configurations will be described in the
The dissolved oxygen concentration is a critical variable that must be controlled, or at least
measured, if one is to develop models for the corrosion of materials (Jones 1992). An
appropriate gas handling system that includes the ability to bubble high-purity argon and
oxygen as well as compressed air has been included in the development of the loop.
Finally, to assure safety of personnel from scalding water and electrical hazards in the event
of a burst line, a series of aluminum and Plexiglas® shields reduce potential exposure to
Electronic Control of Fluid Flow
A sophisticated manual 24 V analog control system was designed and constructed for the
operation of the loop. It was our conscious decision that such a system would provide
maximum student safety for an instrument that is operated unattended for long periods of
time in a shared academic laboratory. No attempt has been made to provide computer
control of the loop safety relay system although all data acquisition is fully automated, as
will be described shortly.
The main control is separated on two panels, one of which handles all of the high voltage,
high power circuitry and the other of which handles the low-power logic control functions.
Cable bundles run between the two panels in cable trays built into the supporting frame.
The circuitry for both panels was designed using industry-standard 24 Vac analog logic
relays. The power panel delivers 110 Vac power directly to some of the instruments, to a
terminal strip that powers dedicated 110/24 Vdc step-down transformers required by other
Omega Engineering, Norwalk, CT: Model CN16PT-330 digital display controller,
(https://www.omega.com/pptst/CNPT_SERIES.html; accessed 30-Jan-2018).
instruments, and to a 110/24 Vac transformer that provides power for all of the control
logic. It also delivers 208 V single-phase power to the heating element and 247 V 3-phase
power to the VFD and thence to the pump motor. Each power source is isolated from the
building power supplies by appropriate contactors with 24 Vac coils that are controlled by
off/on switches on the operator’s (low voltage) panel. The heating element power is
controlled by two inline solid-state relays, one in each line, such that the PID controller can
switch the power off and on in fractions of a cycle.18
The operator’s control panel, shown in Figure 3(a), has a main power rocker switch that
arms the system by enabling the 110 V power, a switch that enables the 24 Vac power (all
of the data acquisition system and the analog logic circuits), a switch to enable the power
to the 3 kW heating element, and a switch to enable the power to the VFD and therefore to
the primary pump motor. The rocker switch is part of the safety control system and assures
that if the system loses power, it must be manually restarted.
For the convenience of operators monitoring research in the experimental bays, the digital
readouts for the various sensors (flow rate, dissolved oxygen, pH, and conductivity) are
mounted on a third panel located between the two bays as shown in Figure 3(b). Data
loggers for commercial electrical resistivity and linear polarization probes are mounted
below these panels.
As research progressed, it was found that new environmental sensors were required or that
some of the existing sensors failed and had to be replaced. With the data acquisition (DAQ)
chassis mounted on the back wall behind the flow loop, such changes were hard to reach
and difficult to make. A scheme was developed whereby all of the wiring for the 4-20 mA
current loops was transferred via a cable bundle from the DAQ chassis to three terminal
strips on the front of the operator’s panel (e.g., see Figure 3(b)). In this manner, a sensor
transmitter could be replaced in a matter of minutes by undoing three screws and
withdrawing two wires from the panel.
All corrosion mechanisms that involve the creation of a protective passivation film depend
on the concentration of dissolved species in the fluid (e.g., Jones 1992). A dissolved oxygen
Omega Engineering, Norwalk, CT: Model SSR330DC50 solid-state relay with Model FHS-6 heat
sink, (https://www.omega.com/pptst/SSR330_660.html; accessed 30-Jan-2018).
Figure 3: System control panels: (a) main control panel showing VFD (upper
right), power switches (lower left), heater PID control (center), safety system
(lower right), and argon, oxygen and air flow controllers (bottom panel); (b) fluid
sensors including, from the top, fluid flow turbine and ultrasonic thickness gage
and, clockwise from the middle left, dissolved oxygen, pH and conductivity
meters and (bottom) data loggers for electrical resistivity and linear polarization
probes; and (c) in the foreground, an alloy pipe with ultrasonic transducers, heat
exchanger with computer-controlled valve (orange), and top, auxiliary pump
voltage supply. (images courtesy LeeAnn Ellis)
sensor and display is included in the fluid chemistry monitoring system as discussed
above.19 A size 150A cylinder of ultra high purity oxygen with a 65 mm flow meter delivers
a controlled low flow rate of gas to the surge tank (see Figure 3a).20 The gas may be flowed
over the surface of the fluid in the tank or delivered by a small stainless steel tube to the
bottom of the tank where it is bubbled into the fluid. An identical system delivers a low,
controlled flow rate of high purity argon that is used to control the partial pressure of the
oxygen in the system. A controlled flow of compressed air is also provided to purge the
system of argon and return the dissolved oxygen to the concentration at atmospheric
Safety of personnel and of the system can be separated into physical and engineering
controls. As noted above, large safety shields protect personnel from scalding water in the
event of a major burst in the loop when operating at high temperatures. Engineering safety
controls have been implemented that trigger both software and hardware automated
emergency shutdowns of the system.
The fluid level in the tank is monitored by an ultrasonic level detect.21 This device is set to
assure that there is always at least 150 mm of fluid over the 3 kW heating element, thus
preventing the burnout of an uncovered heater or damaging the impeller by running it dry.
The recirculating tank temperature is also monitored by an externally mounted type-K
The specifications of the DOE-45P dissolved oxygen sensor assure us that this device operates
correctly in the temperature –flow rate space of the loop.
Airgas Inc., Radnor, PA: Model Y214651 65 mm flowmeter,
(https://www.omega.com/pptst/CN700_SERIES.html, accessed 30-Jan-2018).
Omega Engineering, Norwalk, CT: Model LVU-150-R ultrasonic level sensor with Model LVM-11
level track kit, (https://www.omega.com/pptst/LVU150.html; accessed 30-Jan-2018).
thermocouple in conjunction with a set point controller.22 This is usually set at 80 °C, thus
allowing for a 13 °C tolerance below the 93 °C maximum operating temperature for CPVC
(Boedecker Plastics). However, for lower temperature investigations, the set point may be
set to a temperature only 10 to 15 °C above the operating temperature, thus protecting the
experiment to some extent.
The outlet pressure for both pumps is monitored at location P1 by the data acquisition
system. A computer controlled, electromechanical relay is wired in series with other
temperature and level detect relay switches for the apparatus, allowing the data acquisition
system to trigger an emergency shutdown if an over-pressure limit is reached. Given the
primary pump’s operating curve, this over-pressure limit was flow-rate corrected for the
apparatus (see Figure 22).
Finally, there is a large, red emergency SCRAM button in the center of the panel as seen in
Figure 3(a).23 This switch has a key lock such that the system can only be started or reset by
an authorized operator. This switch may be activated in the event of cavitation or air
hammer, both of which have been known to cause violent vibrations of the system. These
three devices are connected in series and assure that in the event of any one of the alarms,
all power is disconnected from the system and it is locked out until an authorized restart. If
there is such an event, a very loud enunciator sounds and a red lamp is lit.24 Such a system
also assures that in the event of a building power outage, the system cannot restart
In addition to these user-designed safety features, the VFD has an internal safety feature to
prevent over-current to the pump motor through an immediate shutdown. This feature
protects the pump in the event that a significant back-pressure is developed resulting in
the primary pump stalling.
Omega Engineering, Norwalk, CT: CN702 thermocouple limit controller,
(https://www.omega.com/pptst/CN700_SERIES.html; accessed 30-Jan-2018).
According to Wikipedia, “a SCRAM is an emergency shutdown of a nuclear reactor.” We have
adopted a similar terminology.
Omega Engineering, Norwalk, CT: Model 70A-1 audible alarm enunciator,
(https://www.omega.com/pptst/70A_ALARM.html; accessed 30-Jan-2018).
Materials Test Equipment
To assure modularity and ease of entry of various corrosion experiments into the loop, two
identical sections, each 1.52 m long, were fitted with unions, thus providing two bays for
the insertion of different experiments into the line (see Figure 1). Several compatible
experimental sections have been developed that allow the tribocorrosion behaviour of a
wide range of metals and alloys to be studied across a wide variety of environmental
conditions. Among these are (i) an assembly that allows for the insertion of standard,
commercial corrosion samples, (ii) an assembly that allows sections of pipe or tube to be
exposed downstream from orifices, flow conditioners or valves, (iii) an assembly in which
the wall thickness of pipes and tubes downstream from various orifices can be measured,
and (iv) a device with which the differential pressure drop across various orifices of interest
can be measured and thus the loss factor for each can be experimentally determined. Other
devices, such as a wall shear stress monitor and a fluid flow hot film anemometer, are
under development, while an in-line electrochemical potentiostat is being planned.
Commercial Corrosion Test Samples
Two commercial corrosion probes were assembled into a single experimental section with
the probes spaced at a distance greater than 10 interior pipe diameters apart to ensure
hydrodynamic normalization into fully developed turbulent flow. An electrical resistance
(ER) probe was positioned upstream of the second probe, a linear polarization resistance
(LPR) probe (see Figure 4).25,26 The ER probe measures corrosion through measuring the
change in resistivity of a metal specimen that is mounted flush into the fluid flow.
Assuming the composition does not significantly change for a 250 µm thick disk, the
change in resistance is related to a change in geometry (material removal) for the specimen
(Metal Samples 2018a). The LPR probe measures corrosion through measuring the
polarization resistance across an anode and a cathode composed of the same material
(Metal Samples 2018c). The materials, shaped as cylinders and positioned perpendicularly
into the hydrodynamic flow field with one cylinder forced to become a cathode via an
Metal Samples Co., Munford AL: Model ER3322 Electrical Resistance Probe,
(http://www.alspi.com/erprobemenu.htm; accessed 15 Aug 2017).
Metal Samples Co., Munford, AL: Model LP3222 Linear Polarization Resistance Probe,
(http://www.alspi.com/lprprobemenu.htm; accessed 15 Aug 2017).
Figure 4: In-situ corrosion probes: (a) linear polarization resistance, (b)
electrical resistance, and (c) the devices installed in the loop. (images (a) and (b)
courtesy Metal Samples Inc; image (c) by LeeAnn Ellis.)
external power supply, forcing the counter cylinder to be an anode (Metal Samples 2018b).
Each probe is connected to a data logger that is read by the data acquisition computer via a
4-20 mA current loop interface.27 Images of these devices are shown in Figure 4.
A challenge with the probes was that they were mounted via commercial tees that alter the
cross section significantly and thus affected the turbulence significantly. Attempts were
Metal Samples Co., Munford, AL: Model MS3500x data logger, where X=E or L for the ER and LPR
probe, respectively. (http://www.alspi.com/erprobemenu.htm; accessed 15 Aug 2017).
made to reduce this effect by inserting 3D printed CPVC mounts that limited to the
clearance of the probe at the wall of the tube to a about 50 µm. However, flow disturbances
were still present. This experimental section is normally positioned in the upstream
experimental bay to normalize the incoming fluid flow.
Corrosion/Erosion in the Presence of Orifices
The current research in our group primarily focuses on the effects of various devices such
as valves and flow conditioners on the downstream tribocorrosion of alloy pipes and tubes.
A universal sample chamber was designed to allow the insertion of custom designed
orifices that simulate components of interest into the system without changes in cross
section associated with transitionary components. A drawing of the assembly is shown in
Figure 5. A 300 mm long section of pipe/tube directly downstream of the chamber allows
users to investigate the impact of different orifices on tribocorrosion phenomena. The
universal sample chamber was constructed of stainless steel for durability and allowed for
orifices up to 38 mm in width and 63.5 mm in diameter. The chamber is galvanically
isolated from the downstream material of interest with CPVC transfer unions. In the event
that a part must be longer than 38 mm, a 102 mm extension insert may be placed between
the two flanges, thus allowing components up to 140 mm long.
Pipe Wall Thickness Measurement
A key component of studying material degradation, or any dynamic process, is the
necessity for measurements to be conducted regularly and frequently over the duration of
the exposure. Such dynamic measures of material performance may be characterized by in-
situ, non-interfering ultrasonic wall thickness (UT) measurements utilizing an ultrasonic
transducer and an associated ultrasonic data acquisition and processing instrument.28 Here,
the time of flight of ultrasonic pulses is measured as the pulse is reflected from the inner
and out surfaces of the tube. With a well-calibrated acoustic velocity for the specimen, the
thickness is easily calculated. Accuracy, resolution, and penetration capability may be
altered depending upon the frequency of the signal and which time interval is measured.
For this application, 20 MHz transducers are used and the time interval between two
Olympus Corporation of the Americas, Scientific Solutions Group, Waltham, MA: Model 38DL Plus
thickness gage with optional Model DLP-HR high-resolution software, (https://www.olympus-
ims.com/en/38dl-plus/ - !cms[tab]=%2F38dl-plus%2Fcorrosion-measurement; accessed 08-Jan2018).
Figure 5: The testing module used to investigate 2024-T3 aluminum in highly turbulent
seawater: (a) assembly diagram of components; (b) isometric drawing of the assembly. In
the drawings, fluid flows from lower right to upper left.
Figure 6: UT transducer adapter, shown (a) as designed, (b) transparent with a
single attached UT transducer (blue) and associated delay line (yellow), and (c)
completely configured around a test specimen (orange).
successive backwall echoes in mode 3, which represents one round trip test in the material,
is measured, providing a resolution of 2.5 µm.29 Delay lines are spaced between the
transducer element and the specimen, further separating the transmission excitation pulse
recovery from backwall echoes to increase both accuracy and precision (Olympus Corp.
2017). Toughened PS delay lines (0.25 ms) were chosen for this application. If the
passivation films on the substrate are less than about 1 µm, as is often the case,
passivation film thickness measurements are outside the capability of the transducers.
Custom designed UT mounts fix the transducers against the experimental test sections and
thus eliminate systematic reapplication errors. The mounts were constructed from CPVC,
forming 19 mm wide rings that support up to four transducers spaced radially at
increments of 90º (Figure 6). Transducers thread into the mounts, pressing their required
time delay lines against the experimental wall. Petroleum jelly ensured a continuous, clean,
acoustic connection. With 14 constructed UT adapters, any configuration of azimuthal
positions (0º, 90º, 180º, 270º) could be measured with adapters incrementally positioned
axially (Figure 7). By convention, the 0° positions were positioned vertically with subsequent
positions following the right-hand rule. The axial position of a device is determined by
defining the origin, typically at the exit face of an orifice. These dimensions thus determine
Olympus Corporation of the Americas, Scientific Solutions Group, Waltham, MA: Model M208-RM
20 Mz single element transducers with Model DLH-1 delay lines, (https://www.olympusims.com/en/transducers-and-accessories/; accessed 06-Jan-2018).
Figure 7: UT transducer adapters (labelled 1 through 6 on top) with transducers (labelled A
through O) and clear plastic set screws mounted on an experimental section of nickel
aluminium bronze pipe assembly. Fluid flow direction is from right to left. An orifice that
simulates a valve is mounted in the stainless steel chamber on the right. (image courtesy
θ and z in a cylindrical coordinate system. The transducers are labelled alphabetically A-O
and adapters chronologically 1-14 with the smaller character configurations upstream.
Nylon fixing bolts held delay lines in place when a full adapter configuration was not
desired. The quantity and temporal spacing of UT measurements were limited by the
maximum capability to switch the transducers to a common instrument and physical
proximity interference between transducers during mounting and setup. Each transducer is
switched and connected to a common UT instrument via a 15-input multiplexer, thus
allowing for sequentially conducted thickness measurements to be performed in a 300 s
cycle.30 Details of the software routine are covered in an ensuing section of the paper.
To achieve maximum accuracy in our wall thickness measurements, a calibration tube with
a step in the inner wall diameter has been designed. Typically, such tubes are 250 mm long
and 50 to 64 mm OD with wall thicknesses of 5 to 6.5 mm and are constructed of the same
raw material as the experimental test sections. Using sections of the same tube as that
under investigation assures that the calibration of the velocity of sound in the material is
affected identically by the microstructure and by the parallax effects of the interface
between the transducer delay lines and the outer diameter of the specimens. One half of
each of these tubes is precision machined to have an ID that leaves a wall thickness of
approximately 1 mm. The inner diameters of the thick and thin wall sections are measured
with a precision dial bore gage to ±0.3 µm while the OD is similarly measured with a
precision outside micrometer to the same tolerance.31,32
Device Loss Factor Measurement
The performance of various orifices, valves and flow straighteners inserted into the
universal sample chamber are predicted via computations with a strong literature basis.
However, it is still necessary to measure the actual performance of a device in-situ. A
differential pressure gage was installed following proper ASME standards at locations of D
upstream and D/2 downstream of the orifice of interest utilizing a separate, custom
designed differential pressure chamber (ASME 2004, 2007). The chamber is configured with
spacers to position an orifice at the correct location with respect to a defined origin and
accommodates different device thicknesses. A schematic of this assembly is shown in
Figure 8. For any orifice or flow conditioner of interest, there are three components: (i) a
main housing that carries the fittings for the differential pressure gage and holds the device
Linkbone, Wieruszow, PL: Model 1-to-15 Dual BNC Switch, (http://linkbone.com/product/bnc-
switch-remote-1-15/; accessed 08-Jan-2018).
Mitutoyo America, Aurora, IL: Model 511-752 dial bore gage, (http://www.mitutoyo.com/wp-
content/uploads/2012/11/2005_BoreGage.pdf; accessed 08-Jan 2018).
Mitutoyo America, Aurora, IL: Model 101-120 outside micrometer,
(http://ecatalog.mitutoyo.com/Outside-Micrometers-Series-101-C1094.aspx; accessed 12-Jan2018)
Figure 8: Schematic of differential pressure device for determining
loss factor for various orifices and flow straighteners.
under test (DUT), (ii) a series of various sized spacers that are machined to assure that the
face of the DUT is at the location required by the ASME Standard and (iii) a capping tube
(ASME 2001). The two pressure ports are monitored by a digital differential pressure gage.33
The analysis of data from this device is discussed in the section on loss factors.
To fully characterize the fluid flow in the system described here, we are developing (a) a
device for directly measuring the wall shear stress using the techniques described by Schetz
(2010)and Meritt and Schetz (2016), and (b) an interface that makes it possible to traverse
Omega Engineering, Norwalk, CT: Model DPG409-030DWU differential pressure gage,
(https://www.omega.com/pptst/DPG409_DIFF.html; accessed 30-Jan-2018).
miniature thin film anemometer probes across the diameter of the tube and to thus provide
a direct measure of the orthogonal velocity components of the fluid flow and thus to
provide the degree of fluctuation of the Reynolds number at a given point in the flow
stream (Jorgensen 2002). The details of these devices are presented by (Todoroff 2018). An
inline electrochemical potentiostat is also planned.
Apparatus Monitoring System
An apparatus monitoring system (AMS) has been developed for remote and frequent data
acquisition (DAQ) of environmental conditions and ultrasonic thickness measurements, to
conduct emergency shutdown procedures (ESP), and to periodically communicate status and
recorded data updates with the operating user. The AMS was primarily designed around the
National Instrument CompactDAQ chassis platform which allows for several simultaneous
analog I/O, digital I/O, and counter/timing measurements utilizing the associated LabVIEW
virtual environment and visual programming language.34 The routine was written LabVIEW,
version 17.0f2 (64-bit) operating on Windows 7 Enterprise.35
System Block Diagram
A block diagram of the data acquisition system is shown in Figure 9. The outputs of the
DOTX-45, PHTX-45 and CDTX-45 series of data transmitters used to display and record
the output of the dissolved oxygen sensor, the pH sensors, and the electrical conductivity
sensors, respectively, as well as the output from the fluid flow turbine, the strain gage
outputs from the wall shear stress apparatus, the data loggers for commercial corrosion
probes, the absolute pressure gages P1 through P4, and the differential pressure gage used
to determine loss factors are all recorded via industry standard 4-20 mA current loops. Up
to 24 channels of data are acquired from these instruments via three 24-bit I/O modules
mounted in a compact chassis.36
National Instruments Inc., Austin, TX: The NI Platform, (http://www.ni.com/en-
s/innovations/platform.html; accessed 09-Jan-2018).
National Instruments Inc., Austin, TX: Model cDAQ-9188 compact DAQ chassis
(http://www.ni.com/en-us/shop/select/compactdaq-chassis; accessed 09 Jan 2018).
National Instruments Inc., Austin, TX: Model NI_9203 C-series current input module,
(http://www.ni.com/pdf/manuals/374070a_02.pdf; accessed 30-Jan-2018).
All four of the thermocouples mounted on the heat exchanger, as well as up to twelve more
thermocouples located strategically throughout the system are recorded with a 16-bit 16-
input thermocouple module with built-in cold junctions.37 This allows the efficiency of the
heat exchanger to be measured and thermal losses of the system to be determined.
The control voltage of the Belimo heat exchanger valve as well as the control voltage for the
auxiliary laminar flow pump are recorded with a four input voltage module.38 A signal to the
SCRAM shutdown system described above is provided by including a software-controlled
relay that is in series with the level detect and the surge tank over-temperature relays.39
This relay is opened under software control in the event that the pressure at P1 becomes
too large and thus provides an additional level of safety for the main motor and impeller.
This pressure is currently set at 200 kPa (absolute).
Finally, the multiplexor by which the various ultrasonic transducers are selected for reading
and processing and the ultrasonic thickness gage both have RS232C interfaces to the
computer via a USB to dual RS232C interface.40 This interface allows us to select any one of
up to 15 transducers in the array described above and read the tube wall thickness at the
location of the selected transducer, all under program control and without user
The “mini chassis” communicates with the computer via a CAT-5 Ethernet cable and all data
acquisition software is written in the LabVIEW© programming language which automatically
records the data to a comma separated variable (CSV) spreadsheet file.41 The program uses
National Instruments Inc., Austin, TX: Model NI-9213 C-series temperature input module
(http://www.ni.com/pdf/manuals/374916a_02.pdf; accessed 30-Jan-2018).
National Instruments Inc., Austin, TX: Model NI-9201 C-series voltage input module,
(http://www.ni.com/pdf/manuals/373783a_02.pdf; accessed 30-Jan-2018).
National Instruments Inc., Austin, TX: Model NI-9482 C-series relay output module,
(http://www.ni.com/pdf/manuals/373948b.pdf; accessed 09-Jan-2018).
National Instruments Inc., Austin TX: Model USB-232/2 Serial Interface Device,
(http://www.ni.com/en-us/support/model.usb-232-2.html, accessed 10-Jan-2018).
National Instruments Inc., Austin, TX: What is LabVIEW? (http://www.ni.com/enus/shop/labview.html, accessed 5-Feb-2018).
Figure 9: Block diagram of data acquisition system.
built in DAQ virtual instruments to start recording data and set the time interval between
readings. Further details of the DAQ modules are given in Appendix 1.
The AMS utilizes an evenly stacked, hierarchal, routine structure to improve user navigation
and ease of use by executing several nested and concurrent subroutines as shown in Figure
10. The AMS routine structure executes three sequential sections: Section A initializes user
defined parameters from a virtual user interface (Figure 11) and stores them to local
variables; Section B performs all necessary procedures for (i) DAQ, (ii) ESP, and (iii)
communication updates in a timed loop structure. Sequence Section C sends an SMTP
(simple mail transfer protocol) email to notify the operating user that the AMS routine has
ended and attaches the DAQ spreadsheet in xlsx format. LabVIEW protocol dictates that all
procedures run from left to right, requiring all signals to be present before conducting a
Data are taken from two primary sources; the NI CompactDAQ chassis platform and the UT
system as shown in Figure 9, executing sequentially in a timed loop at a user defined
frequency entered in Section A, referred to as the DAQ cycle. Environmental sensors are
recorded simultaneously via the NI CompactDAQ chassis that features eight module slots,
allowing for a wide variety of measured fluid and environmental conditions (see Figure 9).
The UT system switch connects each of the 15 UT transducers to a common UT instrument
via an RS232 serial commanded multiplexer which allows for sequentially conducted
thickness measurements. The arithmetic mean, standard deviation, skewness, and kurtosis
values for both the environmental conditions and the UT measurements were calculated
and recorded from 200 and 5 measurements taken at ∆t=0.1 s and ∆t=3 s apart,
respectively (e.g., Miller and Freund 1977, 72-74). At the end of the DAQ cycle,
measurements were time-stamped, counted, and appended to a common data file. The
minimum cycle period was 300 seconds. A typical e-mail message to the user on normal
completion of a data acquisition cycle is shown in Figure 12(a) while a section of the
attached data file is shown in Figure 13 as it would appear in MS Excel.
Figure 10: AMS routine and subroutine hierarchy (NI LabVIEW). The various
subroutines that are run at each level (I,..IV) are described in Appendix 2,
Figure 11: AMS routine User Interface (NI LabVIEW)
Figure 12: Text of e-mail messages sent to user after each data acquisition cycle: (a) on
normal completion of cycle; (b) on environmental change of 20%; and (c) DAQ abortion.
Figure 13: Screen shot (partial view) of data CSV file sent to user each data
cycle as it appears in MS Excel.
Emergency Safety Features
The AMS routine features dynamic responses to measured environmental conditions by
triggering the apparatus’ safety system via an NI electromechanical relay output module.
This is triggered if either (i) the fluid temperature rises above 50°C or the pressure at the
location of P1 rises above the maximum pressure capability or (ii) the pH, conductivity,
and/or flow meters depreciate by 75% of their range capability within an iteration of the
timed loop in Section B. The triggers are installed to prevent damage of system components
in the case of an overpressure/over temperature or loss of fluid. Once the safety system is
triggered, the main power systems are disabled; shutting down the Variable Frequency
Drive (VFD) and its associated 3.7 kW impeller pump, the 3 kW resistive heating element,
the 110 Vac lines, and all other lower power systems wired directly into the apparatus.
These updates to the emergency system are wired in series with the hard-wired safety
system (see Safety Considerations) and supplement the ability of the apparatus to protect
itself and the operating users through emergency shutdown procedures that must be reset
SMTP Communication Updates
A necessary feature for a routine that acquires data remotely, frequently, and for long
durations is periodic status and emergency updates. The AMS routine executes SMTP loops
by emailing the operating user an attached DAQ xlsx spreadsheet (i) during data updates at
environmental condition(s) have changed by 20% from the previous DAQ iteration
(Figure 12(b)), (iii) when software error codes are generated, and (iv) once the AMS routine
is aborted (Figure 12(c)). In each case, the attached file is the current xlsx file without
modification. The error code suggests where in the last few data records the user might
look for the source of the potential problem.
As user requirements and expectations changed during system development and testing,
the structure and complexity of the AMS routine evolved similarly. Several challenges were
presented during the process of producing the best possible routine. One consistent
challenge that could not be eliminated or significantly reduced via design changes was
signal noise. Data lines were shielded via grounded conduit where appropriate and were
rerouted to avoid AC power lines. Furthermore, a precise calibration signal was utilized in
order to determine the appropriate signal processing filter to stabilize each signal. The
continued existence of noise motivated us to include online computation of statistical
information of the signals; arithmetic mean, standard deviation, skewedness, and kurtosis
over a large population of 200/5 data points (environmental/UT) at a frequency of 10/0.33
While far from elegant, the current AMS routine allows for reliable data acquisition and
automated apparatus operation. The routine and subroutine structures involved are
described in detail in elsewhere (Todoroff 2018).
Performance And Calibration
Before experimental research on the corrosion and/or erosion of materials could be
performed in the instrument described above, a significant effort was required to calibrate
its fluid dynamics performance. Among the measurements necessary were: (i) calibration of
the volumetric flow rate of the fluid as a function of VFD frequency and experimentally
caused backpressures; (ii) determination of the passive heating curve due to generation of
energy in the fluid by the impeller and which delineates the boundary between the
conditions at which heat needs to be added or removed to reach a desired temperature at
each flow rate; (iii) the thermodynamic performance of the heat exchanger and of the CPVC
piping system; and (iv) the effects of backpressure on the fluid flow caused by orifices, flow
straighteners and valves used in the sample chamber. Each of these is discussed in this
Turbulent Fluid Flow42
A 1.52 m section of 102 mm diameter CPVC pipe was used in place of a sample chamber in
both experimental bays and the loop was filled with distilled water. The system was
operated at ambient temperature and the volumetric flow rate was recorded for various
motor frequencies as set on the VFD. Measurements were made sufficiently rapidly that the
fluid heating effect discussed in the next section could be ignored. This configuration
assured that, with the exception of a small contribution from the smooth CPVC pipe, all
flow losses were contributed by the permanent components of the system. The results are
shown in Figure 14. A least squares analysis of the data shows that they fit the linear
is the volumetric flow rate (L/s) and f is the VFD frequency (Hz). The R2 = 0.9999
for this fit is exceptional. Eq. (2) can be inverted to give
which is more convenient if one wishes to acquire data at specific volumetric flow rates. The
estimates of the standard error of the coefficients were determined following the analysis of
Miller and Freund (1977, 295 ff).
Both the density and the viscosity of water decrease with increasing temperature
(e.g.,Carney and Hendricks 2017). At this time, it is not known if either has an operationally
This section is based on the work of B.M. Greenblatt (2017).
Figure 14: Turbulent flow rate (L/s) versus pump motor frequency. The least
squares fit to the data is given in Appendix 3.
significant effect on the relationship between the pump motor frequency and the volumetric
Effect of Pressure on Fluid Flow
As various experiments are developed for study in the loop, the effect of restrictions and
constrictions in the tubes under examination (e.g., tube diameter, the presence of orifices,
flow straighteners, valves, etc.) will change the work required of the pump and will thus
affect the maximum flow rate that can be achieved. Details of how these effects can be
estimated quantitatively will be presented in a later section. However, to first determine the
significance of such effects, a straightforward experiment was performed in which the
pressure at P1 (Figure 1) was measured as a function of flow rate for various closure
positions of the last ball valve in the loop just in front of P4. This ball valve was used as a
Figure 15: Effect of closing a full-bore ball valve on the flow rate of water as a
function of pump frequency. P1: α=0.0 deg. (valve open); P2: α=22.5 deg; P3:
α=45.0 deg; P4: α=67.5 deg. The least squares fit to the data are given in
proxy for various experiments in the bays to provide data for the effect of introducing a
constriction into the loop. The results, shown in Figure 15, indicate a significant decrease in
achievable flow rate as the valve is closed. The data in Figure 15 can be related to the flow
, that are used to determine the maximum flow through a piping system and
thence to the loss factor,
as will be demonstrated shortly. It is important to note that
flow factors are dimensioned quantities that depend on the units of flow and pressure,
while the loss factors are dimensionless. The two quantities are related inversely. Details of
computations with both are given in Appendix 4. The slopes of the flow rate versus opening
angle shown above are related to the valve loss factor and, as will be shown, a quantitative
expression for the volumetric flow rate as a function of both the pump motor frequency and
the experiment loss factor can be developed. This relationship is invaluable in designing
the operating characteristics of various experiments in such a way as to not exceed the
operating point of the pump (see Figure 22).
Laminar Fluid Flow
To calibrate the laminar flow pump, the loop was set up in the same manner as described
above for the calibration of the turbulent flow pump except that the valves were adjusted to
allow all flow through the system to be controlled by the auxiliary pump. The flow rate was
determined by measuring the volume of fluid captured in a large graduated cylinder per
unit time as the flow rates were below the limit of sensitivity of the turbine flow meter. The
results are shown in Figure 16. A least squares analysis of the data show that they fit the
is the volumetric flow rate (L/s) and V is the applied voltage to the pump (volts).
As in the previous section, Eq. (4) can be inverted to give
an equation that that can be more convenient if flow at a well-defined rate is desired.
System Equilibrium Temperature
Although it was anticipated that the fluid temperature would rise somewhat during long-
term experiments as a result of energy deposited by the impeller, when the first
experiments were run, the magnitude of the increase was much larger than expected. To
quantify this effect, the system was filled with distilled water and the temperature of the
water in the recirculating tank was measured as a function of time for each of four pumping
speeds (flow rates): 10, 20 30, and 40 Hz (1.39, 2.88, 4.38, 5.94 L/s). The coolant flow to
the heat exchanger was turned off during these measurements.
The data at the three lowest flow rates were empirically found to follow the equation
is the initial temperature of water,
is the final (equilibrium) temperature of
is the time constant for the process.
Figure 16: Laminar flow rate (L/s) versus pump applied voltage (V). The least
squares fit to the data is given in Appendix 3.
Eq.(6) can be rearranged as
A plot of the left-hand side of Eq. (8) versus t should be a straight line with slope of
plot of how well this relationship is obeyed is shown in Figure 17.
Figure 17: Plot of temperature versus time for 20 Hz run. Data are in format
required by Eq. (8). The least square fit to the data is given in Appendix 3.
If data cannot be recorded when the water reaches
, then the final temperature may be
found as follows. A graph of the left-hand side of Eq. (8) versus should be a straight line
with an intercept of zero and a slope of
. Trial values of
are inserted into Eq. (8) and
the values of the statistical
fit (the largest
of the fits are recorded. The value of
that gives the best
) is taken as the correct value. Following this procedure, a value for
for the 40 Hz run was obtained. The high value of R2 indicates that Eq. (5) is a good model
for describing the experimental data. The equilibrium final temperature,
, versus pump
motor frequency is shown in Figure 18.
The slope of the curve in Figure 17at time t is
and at time
Table 1: Equilibrium properties of fluid flows.
Figure 18: Equilibrium temperature rise as a function of
motor frequency (flow rate). The line is the least squares fit
to the equation given in Appendix 3.
The time constant
may be found from the initial slope of the heating curve, provided the
final temperature of the water is known. A comparison of the values of
Eq. (8) and from the slope of the best fit to Eq. (10) gives an indication of how well Eq. (6)
fits the experimental data.
The measurement time was long enough to determine
rates. However, it was not possible to determine
directly for the three lowest flow
directly at the highest flow rates because
the equilibrium temperature was above the safe operating temperature of the system. The
results of these measurements are shown in Figure 18 and are tabulated in Table 1. In the
is the value obtained from Eq.(10). The
is the value obtained from Eq. (8) and
agreement is considered to be excellent, except for the data take at 20 Hz where they differ
by an acceptable 8%.
The power dissipated by the pump may be determined as follows. At time t, the heat
deposited in the water is given by
is the energy deposited in the time required to heat the water from the initial
is the mass of fluid in the system, and
is the heat
capacity of the fluid. It has been assumed that the heat capacity is constant over the small
. The power transferred to the water is the time derivative of Eq. (11) and is
The properties of water required to evaluate Eq. (12) are
kg/m3 (Eisenberg and Kauzmann 1969). Using these values, the power dissipated
in the water as a function of frequency (flow rate) may be computed from Eq. (12). The
results are given in Table 1.
The increase in fluid temperature as a function of motor frequency (flow rate) shown in
Figure 18 is well-explained by a power law of the form
This implies that if 80 °C is the maximum safe operating temperature of the loop and if the
ambient temperature is 22 °C, the maximum frequency (flow rate) of loop operation in the
absence of a heat exchanger capable of removing this excess heat is approximately 36 Hz
or 5.30 L/s, values well below the design specification for the instrument. These results
also show that in a configuration without a heat exchanger, it is not possible to operate the
system at any temperature—flow rate combination that falls below the curve given in
Heat Exchanger Performance43
To allow operation of the loop in the entire temperature-flow rate space, a double-pipe
heat exchanger was designed and installed (Holman 2002, 511). The inner tube through
which the fluid to be cooled flows is Schedule 40, type 304 stainless steel with an outer
diameter of 60.3 mm and a wall thickness of 3.91 mm, while the outer tube that defines the
annulus through which the coolant flows, is Schedule 10, type 304 stainless steel with an
outer diameter of 88.9 mm and a wall thickness of 3.05 mm. There are six baffles in the
annulus to help maximize the cooling surface area (Lawe 2017). Five type-K thermocouples
were mounted on the surface of the device, one each at the inlet to and outlet from the
main pipe, and one each to the inlet and outlet of the coolant annulus. An additional
thermocouple was mounted on the outlet to the main pipe and was used for electronic
control of the outlet fluid temperature.
For the proposed corrosion studies to be properly controlled, it is essential to know the
temperature of the fluid at the sample position in the loop. This can be accomplished by
interpolating between the temperatures determined by the thermocouples in the two
conductivity meters. However, it is also necessary to control the fluid temperature at the
outlet of the heat exchanger. Hence, it is essential to determine the difference between the
fluid temperature and the temperature of the externally mounted thermocouple at the heat
exchanger outlet. This calibration has been accomplished in two ways.
First, with the heat exchanger coolant valve fully opened, the flow rate of chilled (11 °C)
water was measured to be 0.60 L/s at a pressure of approximately 482 kP. Under these
conditions, the fluid in the system remained at 11 °C for fluid flow rates up to
approximately 6.31 L/s. The temperature at the sample position and the temperature of the
This section is based on the work of H. A. Vanhout (H. H. Vanhout 2017).
surface of the heat exchanger were the same within experimental error. These results
indicate the success of the heat exchanger design.
In a second measurement, the coolant flow to the heat exchanger was turned off and the
system was allowed to come to ambient temperature. The system was then operated at a
flow rate of 6.14 L/s and was allowed to come to thermal equilibrium as shown in
The heat exchanger outer wall temperature was 65.1 °C and the ambient
temperature was 23 °C. The temperature of the water flowing in the pipe may be
determined from these data.
The calculation proceeds as follows: (i) determine the
convective heat loss from the outer surface of the pipe to the ambient air; and (ii) determine
the heat loss from the flowing water to the outer diameter of the pipe by equating this loss
to the convective loss computed in the first step.
The heat flux per unit length of a horizontal pipe cooled by natural convection can be
determined in the following manner—first, the Grashof–Prandtl product must be
determined, from which the Nusselt number is obtained, from which the heat transfer
coefficient for convective heat transfer to air is found (Holman 2002, 328). The following is
an example calculation for the flow conditions quoted. Other results will be reported
without detail later.
The Grashof-Prandtl product is computed from
is the acceleration of gravity (9.81 m/s2),
temperature of the outer wall,
diameter of the pipe,
is the ambient air temperature,
is the kinematic viscosity of air, and
In the present case,
is the external
is the Prandtl number of air.
mm. For air evaluated at the
m2/s and Pr=0.704. The viscosity and the Prandtl
number for air are given in Holman (2002, Table A-5, 602). For the present case,
. The Nusselt number is then given by
where C and m are constants determined by the Grashof-Prandtl product. These constants
are given by Holman as either C=0.480 and m=1/4 or C=0.53 and m=1/4 based on two
different sources (Holman 2002, Table 7-1, 322). Using the average of these numbers, the
. By the definition of the Nusselt number, the wall to air
Nusselt number is
heat transfer coefficient is
is the thermal conductivity of air, which, for the present case, is by interpolation
in in Table A-5 of Holman,
W/m·K (Holman 2002, 328). For the present
W/m2•K. The thermal flux per unit length of the stainless steel pipe
of the heat exchanger is thus
The heat loss through the wall from the flowing fluid is given by
is the temperature of the internal water of the system and is the unknown to be
are the inner and outer radii of the heat exchanger tube, and
is the thermal conductivity of the 304 stainless steel wall of the heat
exchanger at 100 °C (AK Steel 2013). Since the system is at thermal equilibrium,
To evaluate Eq. (18), it is necessary to know the heat transfer coefficient
. Expressions for
for turbulent water flowing through a smooth pipe have been given as a function of the
Reynolds and Prandtl numbers of the fluid by Dittus and Boelter and by Gnielinski, as
summarized by Holman (2002, 268). The Dittus-Boelter equation for fully developed
turbulent flow in smooth tubes is
for cooling of the fluid. On the other hand, the Gnielinski equation, which is
For water at 65 °C flowing at 6.136 L/s) in a 52.5 mm ID smooth, Schedule 40 pipe, the
. The Prandtl number for water at 65 °C is
Reynolds number is
(Holman 2002, Appendix A-5, 602). With these values, the Nusselt number for the
conditions of the present example is estimated to be
Eq. (20). The thermal conductivity of water at 65 °C is
by Eq. (19) and
values, and using the inside diameter of the pipe, by Eq. (16) the heat transfer coefficient at
the inner wall is
W/m2·K according to Dittus-Boelter and
Eq. (18) can now be solved for the difference in temperature between the fluid and the
outer stainless steel wall temperature. It is found that, because the heat transfer coefficient
at the inner wall is so large, the first term in the denominator of Eq. (18) is negligible
°C. This difference is, for all practical purposes,
compared to the second and
negligible and the temperature reading on the externally mounted thermocouple may be
taken as the internal fluid temperature.
Similar computations at other, smaller flow rates indicate similarly small temperature
differences. Thus, these results show that the temperature measured by the surface
mounted thermocouple at the exit of the heat exchanger is a valid control signal for the
input to the PID that controls the chilled water flow for all operating conditions of the loop.
Fluid Temperature at the Sample
To be able to develop accurate models for the corrosion of metals studied in the loop, it is
essential to have a good measure of the temperature of the fluid at the sample. In the
absence of a thermocouple at the sample, this may be estimated as follows. The heat loss
per unit length of CPVC pipe may be calculated from the relationship
Is the thermal conductivity of CPVC and all other variables have been defined.
The values of
were found in the previous section and it was determined that, as
an excellent approximation,
is given by the exterior wall temperature of the heat
exchanger. For a fluid flow of 6.14 L/s of water at 65.1 °C, the heat loss per unit length of
Schedule 80 CPVC pipe is 40.8 W/m. This heat must come from the fluid. The heat loss per
unit length of fluid is
where ρ is the density of water at the fluid temperature, Cp is the heat capacity, and
is the temperature drop of the fluid per unit length. For a flow of water at 6.14 L/s and
65 °C, the temperature drop per unit length for Schedule 80 CPVC pipe is 0.0016 °C/m.
The two conductivity sensors (see Figure 1) have built-in thermocouples that measure the
fluid temperature and are used for internal temperature compensation. These sensors are
separated by 4.59 m. The predicted temperature drop between the two sensors is
0.0073 °C. The measured change is 1.8 °C. This discrepancy is outside the manufacturer’s
statement of the precision of the two devices and is, as yet, unexplained.
Thermal Properties of the Loop
The seemingly small values of the temperature difference between that of the flowing fluid
and the outer surface of the heat exchanger and the almost negligible temperature drop
between the two conductivity meters must be justified. This can be accomplished by
performing a system heat balance.
When operated at equilibrium with no heat added by the heating element and no heat
removed by the heat exchanger, the sum of all of the losses around the loop must equal the
heat input from the impeller. From Table 1, at 40 Hz (6.14 L/s) this heat input is 1034 W.
The heat loss per unit length of CPVC and stainless steel pipe (heat exchanger) were
determined in the previous section. All that remains is to calculate the heat loss from the
38 L recirculating tank that is 0.36 m in diameter and 0.41 m high. The computation
follows that for the heat loss from the exterior of the heat exchanger tube. For the vertical
wall of the tank, the Grashof-Prandtl number is 1.358 x 108. Also, for a vertical tank, the
Nusselt number constants are C=0.58 and m=1/4 as given by Holman (2002, Table A-9,
606) and the Nusselt number is
tank surface is 4.93
. The resulting heat transfer coefficient for the
from which it may be determined that the heat loss from the
tank is 94 W. For simplicity, we have ignored the heat loss from both ends of the tank. The
tank has a polished surface for which the emissivity
that results in a radiant heat
loss of 10 W (Holman 2002, 401 and Table A-410, 607). Thus, the total heat loss from the
tank is estimated to be about 104 W.
The total length of the piping in the system is 11.73 m, of which 1.22 m is the heat
exchanger that has only 0.23 m that is uncovered and thus looses heat to the air. The
balance is CPVC. Knowing the heat transfer coefficients for each of the surfaces, we find the
total heat loss for the system is 561 W for the conditions of this measurement.
compares in magnitude with the estimated heat input of 1034 W from the impeller and
validates that the computed temperature changes are realistic, but quite small.
Loop Component Loss Factors
The pressures in the piping system during pumping are critical parameters that control the
maximum flow rate that can be achieved at the site where corrosion samples reside during
an experiment. We have measured the static pressure as a function of flow rate at each of
the four locations identified in Figure 1 with Bourdon-type pressure gages as shown in
The head loss for any component, be they be pipe, fittings such as elbows and valves, or
in-line devices such as orifices, flow meters, or any such devices, the pressure loss can be
computed from the relationship
where KL is the dimensionless loss factor (Munson et al. 2013, 492). In the following, we
estimate the loss factor for each of the permanent components of the loop.
First, the pressure drop in CPVC pipe is calculated. Gages P3 and P4 are separated by
4.60 m of CPVC pipe. Thus, the pressure loss per unit length of CPVC is
where the pressure drop is given in kPa/m. The loss factor per meter for CPVC is then
found from Eq. (23) to be
. The large standard deviation results from
the imprecise pressure values recorded from the Bourdon gages at low pressures. This
value compares well with the value of
computed from the data from Spears
Another contribution to the pressure drop in the system is that caused by the four 90
degree elbows. The loss factors for the elbows may be calculated as follows. The distance
between gages P2 and P3 is 0.46 m with an elbow at each end (see Figure 1). The pressure
drop across each elbow is thus
Figure 19: Gage pressure at points P1, P2, P3 and P4 shown in
Figure 1. The lines are the least squares fits to the data, the
equations for which are given in Appendix 3.
is given by Eq. (24). Substitution of these data into Eq. (23) and solving for
the loss factor yields the result that
. Again, the relatively large standard
deviation results from the resolution of the Bourdon gages. This value compares well with
computed from the data of Spears (undated, 27).
The pressure drop across the flow meter as a function of flow rate is given in the technical
specifications for the device (Omega Engineering 2009). These data have been found to fit
an equation of the form
Figure 20: Pressure drop across the heat exchanger. The line is the least squares
fit to the data, the equation for which is given in Appendix 3.
where the pressure drop is given in kPa and the flow rate is in L/s. R2=1.000 for this fit. At
the flow rates under consideration here, the linear term is negligible. Solving Eq. (23) for
the loss factor, we find that
, and is independent of the flow rate.
Finally, the loss factor for the heat exchanger may be determined. Since there are 3.48 m
of CPVC pipe in the section between gages P1 and P2, the pressure drop across the CPVC
pipe between P1 and P2 is
and the pressure drop across the heat exchanger is
and is shown in Figure 20. As above, the loss factor is computed by inserting the data of
Figure 20 into Eq. (23). In this case, the loss factor for the heat exchanger is strongly
Figure 21: The loss factor of the heat exchanger, KL, as a function of flow rate.
The line is the lease squares fit to the data, the equation for which is given in
dependent on the flow rate as illustrated in Figure 21 and is given by
where all variables have been previously defined. For these data, R2= 0.953.
As expected, the heat exchanger optimizes the amount of heat transfer in a fixed length
through a geometrical design that increases both the heat sink surface area, as well as the
overall surface dwell time through redirection of the flow in a helical pattern. The nature of
this complex geometry affects the flow non-linearly, as illustrated by the large loss factor
and its flow dependency as shown in Figure 21.
Operation of the system requires that the head pressure at the pump be less than the value
specified by the pump specifications. The head at the pump outlet must include the
0.305 m difference between the height of the loop and the height of the fluid in the
recirculating tank (Badr and Ahmed 2015, 21). This is given by
where the pressure is given in kPa and is the equivalent to a pressure of 0.3 m of water at
4 ºC. This quantity is so small compared to other effects that correction for the change in
density of the fluid with temperature has been ignored.
Finally, the total length of CPVC pipe in the entire system is 11.43 m. This includes 1.22 m
of CPVC pipe from the pump to P1, 3.48 m between P1 and P2, 0.46 m of pipe between P2
and P3, 4.60 m between P3 and P4, and another 0.91 m between P4 and the recirculating
The total pressure drop in the system is thus
In deriving Eq.(31), frictional forces have been ignored in applying the Bernoulli equation.
These effects are estimated to be only a few percent (Panton 2013, 132). The total pressure
drop has been evaluated as a function of the fluid flow rate by using the least squares fits
to the various components as listed in Appendix 3 and is shown in Figure 22.
Self-consistency requires that the pressure drop in the horizontal part of the loop be
. A least squares plot of
calculated from Eq.(31) versus the values of
calculated from the least squares values fit to the observed data of Figure 19 (see Appendix
3 for equations) shows that
with R2 = 0.992, thus confirming expectation.
The pump curve, obtained from the ASP technical specifications, is also shown in Figure 22
(American Stainless Pumps 2012).44 The difference between the pump curve and the total
pressure in the system describes the flow rate space that is available to the experimentalist
in designing various sample configurations to be examined in the apparatus—the maximum
We are indebted to Luke Eck of ASP for providing the head versus capacity data from the technical
specifications in digital format.
Figure 22: Pump curve and total system pressure. Solid lines are the least
squares fit to the data and are given in Appendix 3. The intersection of the
two curves is the operating point of the system.
flow rate that can be accommodated is the operating point, the point where the two curves
These results show that the maximum flow rate through the system for any arbitrary
experimental configuration is easily determined by adding the loss factor for the
experimental apparatus to Eq.(31) and solving for the pressure and flow rate for which
Eq. (31) and the pump curve intersect. In our laboratory, where we are interested in the
effects of flow straighteners, valves, and orifices on the downstream corrosion and erosion
of various alloys and on the surface treatments of pipes, such calculations are always
performed in advance of building and mounting the experiment in order to assure its
feasibility. As noted in a previous section, we have developed a differential pressure system
for measuring the loss factors of our experimental devices.
Experiment Loss Factors
The decrease in volumetric flow rate as a function of increasing loss coefficient as
demonstrated in Figure 15 has significant implications for characterizing the fluid flow in
any tribocorrosion study undertaken in the VTHTCL. In this section we develop an empirical
relationship between flow rate and the loss coefficient.
Before beginning, it is essential to clarify terminology found in the literature. The loss
, is a dimensionless quantity defined by Eq. (23) in the previous section, and the
, is a dimensioned quantity defined by the Hazen-Williams equation (see
Eq. (35) below). The latter is used to define the maximum flow rate through a valve as well
as to define losses in other devices. In this section it will be necessary to use both
terminologies. Relationships between the two are developed in Appendix 4.
The fluid flow effects of an experiment with a given loss coefficient may be studied by
using the full-bore ball valve installed just before pressure gage P4 a proxy for the
experiment by adjusting its closing angle,
. Examination of the loop diagram (Figure 1)
shows that the following system of equations describes the pressure drop around the loop
is the density of the fluid (kg/m3), the various
are the loss factors for the heat
exchanger (he), the flow meter (fm), the CPVC pipe per unit length (m-1), and other
plumbing fixtures that make up the loop. The lengths l1…l4 are the lengths of CPVC pipe
(m) between the various pressure transducers,
is the mean velocity of the fluid (m/s),
is the pressure drop across the ball valve at an closing angle
is the height of
the exit tube (m) used to create sufficient backpressure to assure that at any flow rate the
tube is always full. The pressures are given in pascals. The exit tube drains by gravity from
into the recirculating tank, thus making the gage pressure at the exit zero. The
small head from the recirculating tank to the entrance of the loop at
has been neglected
as have the loss factors for the fully-open ball valves. Summing the four terms, the
pressure drop from P1 around the loop to the recirculating tank is
where the system loss factor is
The pressure drop across the valve is given by the Hazen-Williams equation (Williams and
is the specific gravity of the fluid (dimensionless) and
International System (SI) of units
is the flow factor in the
.45 Substitution of Eq.(35) in Eq.(33) and writing
all pressures in pascals, the final expression for the pressure at P1 is
If the value of the pressure for a fully open valve is taken as reference and subtracted from
the pressure for a valve with a closing angle α,
Eq.(37) is thus an expression for the pressure drop exclusively across the ball valve with a
given closing angle,
, and can be determined from the experimental data of Figure 15. By
We note that the units used in Eq.(35) are not the units commonly used in industry. The
relationship between the industrial SI units and those used here is given in Appendix 5.
Figure 23: Pressure at P1 versus flow velocity as a function of the closing
angle of a CPVC full-bore ball valve located just in front of pressure gage P4
in the VTHTCL. The least squares fits to the curves are given in Appendix 3.
Eqs.(36) and (37) we expect that if the head pressure of the return loop is small, a graph
versus the mean flow velocity should vary quadratically. That this is the case is
shown in Figure 23. The pressure drop across the valve as expressed in Eq. (37) is
presented in Figure 24. In this plot, the change in pressure was calculated at the various
flow velocities by subtracting pressures computed from the least squares fits to the data in
Figure 23. The predicted quadratic relationship between the pressure drop with flow
velocity is confirmed for both the closing angles of 45° and 67.5°. The flow through, and
hence the small pressure drop across, the valve at 22.5° is such that the relationship is not
well obeyed as seen in Figure 24.
Let us define the variables
From the data in Figure 24, the coefficients of the power law for the pressure given in
Appendix 3 yield
where the subscripts (1,2,3)
correspond to the valve closing angles 22.5, 45.0 and 67.5 degrees. Eqs.(38) are three
equations in four unknowns. In order to determine the flow factors from these data, it is
necessary to know the flow factor for a fully open valve. From Eq.(35), it is seen that if the
fully open valve is frictionless, the flow factor must be infinite in order that there be no
. With this
pressure drop across it. As an approximation, let us assume
assumption, and taking the specific gravity of 3.5 weight percent salt water to be 1.025, the
values of the flow factors at the other opening angles are readily found to be
From the least squares fit to the experimental data, it is seen that the extrapolated flow
factor for a fully open valve is expected to be
. Substitution of these values
into Eq. (38) allows a second iteration of the flow factor values. However, such an iteration
shows that, to the accuracy of the experimental results, the original hypothesis is valid. The
values of the flow factors are compared with those from the literature in Figure 25. The
results from the present experiment are approximately 47% of those given in MyDataBook
The loss coefficient,
, is defined by the equation (Munson et al. 2013, 415)
where, in SI units, the pressure drop across an assembly of pipes and fittings is given in
pascals, the density of the fluid,
in kg/m3, and the fluid velocity in m/s. The loss
Figure 24: Pressure difference (kPa) across a CPVC full-bore ball valve as a
function of mean flow velocity through the valve at three different opening
angles. The least squares fits to the data are given in Appendix 3.
coefficient is dimensionless. The quantitative relationship between the loss coefficient and
the flow factor is developed in Appendix 4. The loss coefficients for the ball valve under
consideration have been computed from the flow factor data shown in Figure 25 and are
summarized in Table 2. The conversion from the SI units for the flow factor used in this
paper and the industry standard (SU) units is also presented.
Figure 25: Flow factors, Kv for a full bore CPVC ball valve (SI units).
The least squares fits to the data are given in Appendix 3.
Table 2: Summary of experimental results for the loss
coefficient and flow factors for a full bore ball valve as
determined by pressure loss measurements in the
VTHTVL. The flow factors are given in both Standard
Units (SU) and in SI Units (SI) as discussed in the
0.1 5.01 344.4
Effect of Backpressure (reprise)
Examination of Figure 15 shows that the volumetric flow rate depends linearly on the pump
motor frequency and that the slopes of the curves are related to the loss factors of the
valve. Since the valve is in the location of typical experimental apparatuses, the loss factors
of the valve are proxies for the loss factors of experiments. The flow rate must be given by
is the motor frequency (Hz). The function
an expression of the form
may be found by plotting a graph of slopes of the curves in Figure 15 versus the
loss factors for the curves as is shown in Figure 26. From these results, the volumetric flow
rate may be written as
In Eq.(41) there is a small correction to the frequency to account for initial start-up of the
pump. This correction is the average of the values for the four different curves as found in
the least squares fits in Appendix 3. This equation fits the observed data to better than 3%
for P1, P2, and P4 and better than 8% for P3.
Figure 26: Semi-logarithmic plot of the slopes of Figure 15 versus
the valve loss factor,
. The equation of the least squares fit to
the data is given in Appendix 3.
The three gases of interest for work in this laboratory are oxygen, nitrogen and argon.
Oxygen (O2) is available both as an injected high purity gas and as a component of ambient
air, nitrogen (N2) is available as a component of both injected compressed air and of
ambient air, and argon (Ar) is available as an injected high purity gas. The delivery systems
for all three gasses allow for each to be either bubbled through a small stainless steel tube
at any level in the fluid or to be blown over the surface of the recirculating fluid in the surge
The solubilities of Ar, N2 and O2 in both pure water and in seawater have been summarized
in three IUPAC reports (Clever 1980; Battino 1981, 1982). Each is described by an equation
of the form:
is the mole fraction solubility at 1.0 atm (101.325 kPa) of the ith component and T
is the temperature in K. The constants A, B and C for the three gasses, as taken from these
references, are given in Table 3 and their solubilities as a function of temperature at 1 atm.
of pure gas are shown in Figure 27. We note that the calculated values for the solubility of
Ar in water using only the first three values given in Table 3 are significantly higher than
those presented in Table 1 of (Clever 1980). This inconsistency appears to be caused by the
omission of the linear term in Eq. (42) of the form D*(T/100). We have estimated that the
missing term is D=-0.2885. The data shown in Figure 27 have been corrected for this
The solubility of each gas is, by Henry’s Law, the value shown in Figure 27 multiplied by the
partial pressure of the gas. Since the recirculating tank is open to atmospheric pressure, the
sum of the partial pressures of all three gasses must be the local atmospheric pressure, or
are the partial pressures of argon, nitrogen, and oxygen, respectively.
Thus, for operation of the loop at any given temperature, the concentration of dissolved
oxygen can be controlled by controlling its partial pressure by adjusting the argon in the
space above the recirculating fluid in the recirculating tank.
Table 3: Water Solubility Constants for Eq. (42).
Figure 27: Solubility of argon, nitrogen and oxygen in water versus
temperature. Each curve is for 1 atm. of pure gas over the solvent. The
lines are given by Eq. (42) using the values of Table 3.
value) versus the equilibrium concentration as a
function of temperature.
As a test of this concept, the system was filled with distilled water and run at various flow
rates from 1.39 to 5.93 L/s with the heat exchanger operating its full capacity of 0.60 L/s.
Argon was flowed through the system at 40 ccm and bubbled from the bottom of the
recirculating tank. Under these conditions, the average temperature of the system was
of air is
at which the equilibrium concentration of dissolved oxygen under 1 atmosphere
. The measured concentration of oxygen under the Ar atmosphere was
, a reduction to 37% of the equilibrium concentration.
The identical experiment was run under the conditions of no chilled water flow in the heat
exchanger. Here, the system was allowed to run for several hours until the temperature in
the recirculating tank equilibrated (See Figure 18). The results of this experiment are shown
in Table 4. Here, it is seen that, except for the highest temperature, the dissolved oxygen is
also reduced below the equilibrium concentration, but that there is more uncertainty in the
data than at the lower temperatures. The excess oxygen at the highest temperature needs
While the aforementioned apparatus and its capabilities have been designed and developed
around commercially available two-inch pipe, the relevance and significance of this
apparatus relies upon the scalability of the data. In the case of corrosion and erosion
studies, the parameters of interest are the wall shear stress and the thickness of the surface
boundary layer. Thus, it is critical to understand how these parameters scale with the size
of the pipe or tubing under investigation. We have independently performed calculations
based upon generally accepted concepts of turbulent flow of incompressible fluids and
have verified our results with computational fluid dynamics (CFD) (Munson et al. 2013;
Panton 2013). This work suggests an expected power law relationship between the
thickness of the laminar sub-layer and the wall shear that is independent of pipe size. As
well, a linear relationship is expected between a modified Euler number, a ratio of the shear
, and the friction factor that is also independent of pipe
force over inertial force
size. Thus, the kinetics of growth of protective oxide layers should not be expected to scale
linearly with the Reynolds number of the fluid flow, but an appropriately more complex
scaling has been identified. Experimental research is underway to verify these anticipated
The Virginia Tech high turbulence corrosion loop (VTHTCL) is a unique new instrument that
has been developed for the study of corrosion, erosion and tribocorrosion of metal samples
under highly turbulent conditions during which the flow, the fluid chemistry and the
corrosion/ erosion of the material under investigation may be quantitatively monitored.
This system features:
Turbulent flow rates of aqueous solutions from approximately 0.32 L/s to
9.46 L/s through a 51 mm pipe over a temperature range from 11 °C to 80 °C. The
nominal Reynolds number of the maximum flow rate is 295,000 at RT and
approximately 800,000 at 80 °C.
using a small auxiliary pump.
Laminar flow rates in the range
A gas handling system that allows for the control of dissolved oxygen in the fluid
at levels below those of the equilibrium concentration of oxygen in the normal
Instrumentation to quantitatively monitor the fluid flow rate, temperature,
pressure, pH, dissolved oxygen, and conductivity via a computer data acquisition
Insertion bays for easy exchange of experimental configurations, thus allowing
several investigators to develop specimen chambers off line and mount them as
time becomes available in the loop.
A safety system that shuts the system down in case of a loss-of-fluid or a thermal
runaway accident and that has shields to protect personnel in the event of a burst
line filled with scalding water.
A suite of general-purpose sample chambers that allow the insertion of a wide
variety of devices for simulating valves and flow straighteners into the loop.
Experimental data and computations quantify and demonstrate several operating
The heating of the fluid due to the high turbulence of the fluid and the
performance of the heat exchanger in mitigating this heating.
The temperature of the fluid at any point in the system intermediate between the
locations of thermocouples.
The effects of the pressure drop across various components, including the heat
exchanger, the flow meter, and the piping system on the operating capabilities of
the machine and on the constraints they impose on the design of experiments.
The measures of the flow factors and the loss coefficients of various components
that agree acceptably with the literature.
The development of a quantitative expression that gives the volumetric flow rate
of the fluid as a function of both the pump motor frequency and the loss
coefficient of the experimental apparatus.
The capability of using an argon gas cover that allows the dissolved oxygen to be
reduced by as much as a factor of three below the equilibrium concentration
under normal atmospheric conditions.
This research was supported in part by a grant from the US Department of Energy, Office of
Nuclear Energy, Scientific Infrastructure Award (Award#: DE-NE0008674; Project #: GSI-17-
13340), in part by the Naval Surface Warfare Center–Carderock Division through a
Cooperative Research and Development Agreement (NCRADA-NSWCCA-17-279) between
our organizations, and by the Materials Science and Engineering Department at Virginia
The instrument described here is the result of the efforts of three undergraduate Senior
Design teams spread over three consecutive academic years as well as independent
research by three other undergraduates and by the graduate research of three students.
Erik Cothron, Ryan Taylor and Peter Todoroff undertook the design and development of the
first iteration of the instrument as their Senior Project in AY 2015/2016. As a result of the
successes of the initial incarnation, each chose to continue his education as an MSc student
and continued their work on the instrument. John Lones, Trent Strickland, Trey Vanhout,
and Joel Zilke reconstructed the original 5 m long loop as the current 10 m loop, installed
the heat exchanger, and developed the first LabView control program during the
2016/2017 AY, while Justin Aird, Alison Carney, Dustin Rose and Will Wenger added more
instrumentation to improve the measurement and control of the fluid chemistry, made
significant improvements to the instrumentation wiring, added the capabilities to monitor
the wall thickness of the sample tubes via ultrasonic techniques, and made several
advancements in the LabView control code during the 2017/2018 AY. Ben Greenblatt, Trey
Vanhout and Will Wenger each made significant contributions to the calibration of the fluid
flow (Ben), characterizing the heat exchanger (Trey), and developing mechanical
components of the instrument (Will). As graduate students Peter, Ryan and Erik provided
their expertise and guidance to the incoming undergraduates in all aspects of the design,
improvement, and operation of the instrument.
An effort of this magnitude could not have been accomplished without the help of
numerous people and corporations. We wish to thank: Dr. Elissa Trueman of the Naval
Surface Warfare Center – Carderock Division for valuable discussions of corrosion under
high turbulence; Luke Eck of American Stainless Pumps, Inc. for assistance in selecting the
correct pump and motor system; Sam Mirza of Hitachi Industrial Equipment Systems, Co.
Ltd. for assistance with the selection and installation of the variable frequency drive; Tim
Ayres and Jason Lawe of Fluid Chillers for the design and implementation of the heat
exchanger, and Eden Klingenberg, Matt Pegram, and Kira Theuer of National Instruments
for assistance with the LabVIEW programming language. We would also like to give special
thanks to Dr. Tony Trueman for his valued guidance in the field of corrosion and
Various staff and faculty from the Virginia Tech university community provided invaluable
technical support: Jim McDaniel was the project manager for the installation of the power
systems and Joe Zokaites, the University Building Official, helped assure the safe design and
installation of the power control panels; James Lambert of the AOE machine shop
supervised the precision machining of the control panels, safety shields and the universal
sample chamber; Dr. Carlos Suchicital, Hesham Elmkharram, and Ibrahim Khalfallah of the
MSE facilities group were invaluable with their help in assembly of the system. Professor
Joseph Schetz provided valuable discussion concerning turbulent fluid flow in the system
and Profs. Thomas Staley, Sean Corcoran and Alan Druschitz assisted with a myriad of
issues associated with thermodynamics, fluid flow, and corrosion. Ms. LeeAnn Ellis provided
assistance with the photographs of the various components of the system. Finally, we thank
Prof. David Clark, Department Head and the MSE Department for financial support.
Appendix 1: Compact DAQ Chassis (cDAQ-9188) Modules
16-Channel, 75 S/s
1. Fluid Inlet
2. Fluid Outlet
3. Coolant Inlet
4. Coolant Outlet
electromechanical switch relay
1. Switch to analog relay
10 V, 500 kS/s
1. Heat Exchanger Valve Actuation
2. 54W Auxiliary Pump
3. 5V Safety Check
4-Channel, 50 kS/s strain gage 1. Wall Shear Sensor
20 mA, 200 kS/s
current meter module
1. Flow Meter
2. Dissolved Oxygen Meter
3. pH (upstream) Meter
4. pH (downstream) Meter
5. Conductivity (upstream) Meter
6. Conductivity (downstream) Meter
7. Metal Samples® LPR Data Logger
8. Metal Samples® ER Data Logger
20 mA, 200 kS/s
current meter module
1. Pressure Gage P1
2. Pressure Gage P3
3. Pressure Gage P4
4. Differential Pressure Gage
20 mA, 200 kS/s
current meter module
Reserved for parallel tube assembly
Appendix 2: AMS Subroutine Descriptions
Level, Purpose & Description
Level II: SMTP procedures to connect to a
specified email server with stored credentials,
compose message with attached files, and send
Level II: Concatenates entered strings into a
single string for the subject line of an email
User & Experiment
Level II: Concatenates entered experimental
information (as string form) into a single string
for the body of an email
Level II: Performs “Within Range” for select
environmental conditions to indicate a greater
than ±20% from the previous reading. Compiles
the names of each individual condition that
exceeds and concatenates into a single string.
Level II: Performs “UT Measurement – Single
Transducer” for a specified quantity of UT
transducers. Compiles individual channels into a
1D array, multiple channel dynamic data form
Lvel II: Generates a time stamp from the
computer’s system time in 1D array, multiple
channel dynamic data form
Level II: Returns the RMS (Statistics) of selected
module channels from the NI chassis (DAQ
Assistant) for 5 samples at 5Hz. Writes attributes
to each channel, and converts the data into
appropriate units from, takes the average of
similar data groups, and converts into dynamic
data form (to dynamic data).
Level II: Writes all multiple channel, dynamic data
Write to Measurement
to a single, user named spreadsheet file. Newly
acquired data is appended to the same
spreadsheet each iteration the subroutine is
Level III: Gives a Boolean response if a current
reading is within ±20% from the previous
Convert from Dynamic
Level III: Converts the dynamic data type to
numeric, Boolean, waveform, and array data
types for use with other VIs and functions.
UT Single Transducer
Level III: Selects the designated UT multiplexer
channel and transducer (A-O), transmits and
queries serial commands to the UT instrument,
writes the measurement command into 1D array,
single channel dynamic data form
Level III: Converts a Boolean, waveform, and or
Convert to Dynamic
array data to several different types of data;
single numeric value, single channel, multiple
channels for waveforms or indicators (numeric
Level III: Creates, edits, and runs tasks using NIDAQmx
Level III: Returns the selected parameter of the
first signal in a waveform. For RMS, it performs
for all signal values that enter.
Level IV: Establishes a connection to an
instrument via a physical communication port
and allows a user to send serial commands.
Appendix 3: Least squares fit parameters to observed data.
Least squares Fit to the Data‡
‡Units are defined in the figures.
Appendix 4: Flow factors and Loss factors
The pressure drop across a valve is given by the Hazen-Williams equation (Williams and
where, in the units commonly given in the literature,
specific gravity of the fluid (dimensionless), and
and is given in units of
is the flow rate (m3/hr),
is the flow factor for the valve at an
where the pressure drop is given in
bars. The subscript SU indicates that this value of the flow factor is given in “standard units”
found in the literature (e.g., MyDatabook.org 2018). For use in Eq.(33), the units of the flow
factor must be consistent with those of the remainder of the equation. Therefore, we define
a value of the flow factor by the equation
is the flow factor in more traditional System International (SI) units
We choose to use the mean velocity of the fluid (in m/s) to describe the fluid flow and to
express the pressure in kilopascals (kPa). The subscript SI is used to denote the flow factor
in these units. Performing the coordinate transformation between the SU and SI units, we
is the internal diameter of the pipe/tube. Again, the subscripts SU and SI define
the standard units of flow factors found in the literature (SU) and those that are used in this
paper (SI). For Schedule 80 tubes, the ID is 0.0492 m and
The loss factor across a ball valve is then found from the relationship
where the factor of 103 in the denominator of the middle term is to convert the pressure to
kPa and the right-hand term is Eq.(A.2). Solving Eq.(A.5) for the loss coefficient, we find
The standard measurements are made at 16 °C at which temperature the density of water is
999.1 kg/m3, and thus
If the loss coefficient is to be computed from the flow factor given in standard units (SU),
then the flow factor must be converted from SI to SU units as given by Eq. (A.4) and is
Appendix 5: List of acronyms, alloy designations, and pipe sizes
The Aluminum Association
American Iron and Steel Institute
Apparatus monitoring system
American National Standards Institute
Comma separated variable file
Device under test
Electrical resistivity (type of corrosion sensor)
Emergency shutdown procedure
International Standards Organization
Linear polarization resistance (type of corrosion sensor)
proportional-integral-differential (type of electronic controller)
Platinum resistance thermometer
Society of Automotive Engineers International
Shut down in an emergency (from nuclear reactor terminology)
Simple mail transfer protocol (standard protocol for sending
email across the internet)
Unified Numbering System (a system for identifying alloys)
Variable frequency drive
Virginia Tech high turbulence corrosion loop
Alloy Designations by Various Standards Organizations
(ANSI 1979; ISO 2007;
Aluminum Assoc. 2015;
(ANSI 1979; SAE 2017)
(ANSI 1979; SAE 2017)
Nominal Dimensions for NPS 2-in (DN 50 mm) Pipe
Appendix 6: List of Variables
List of Variables
A, B, C, D
Water solubility constants for gasses in water (Eq.(42))
Cubic centimetres per minute
Constant relating Grashof-Prandtl number to Nusselt number (Eq. (15))
Heat capacity at constant pressure
Inner diameter of pipe
Outer diameter of pipe
Pump motor frequency
Gravitational constant (=9.81)
Grashof number (Eq. (14))
Heat transfer coefficient of air
Thermal conductivity of air (Eq. (16))
Thermal conductivity of CPVC
Loss factor for CPVC (per meter)
Loss factor for elbow
Loss factor for flow meter
Loss factor for heat exchanger
Length of pipe
Mass of fluid
Constant relating Grashof-Prandtl number to Nusselt number (Eq. (15))
Nusselt number (Eq. (15))
Partial pressures of argon, oxygen and nitrogen (Eq. (43))
P1, P2, P3, and P4
Pressures at locations 1, 2, 3 and 4 in the line (Fig. 1)
Prandtl number (Eq. (14))
Heat loss per unit length (Eq. (17))
Heat loss per unit length (Eq. (18)
Inner radius of pipe
Outer radius of pipe
List of Variables (cont’d)
Equilibrium final temperature (Eq. (6))
Film temperature (Eq. (14))
Equilibrium ignition temperature (Eq. (6))
Wall temperature (Eq. (14))
Ambient temperature (Eq. (14))
Mean linear flow velocity in pipe
Volumetric flow rate of fluid in pipe
Mole fraction solubility of ith component of atmosphere (Eq. (42))
Inverse of absolute temperature (Eq.(14))
Pressure drop around the corrosion loop (Eq. (31))
Pressure drop of CPVC per meter
Pressure drop across the flow meter
Pressure drop across heat exchanger
System thermal time constant (Eq. (6))
Kinematic viscosity of fluid (air or pumped fluid, as appropriate)
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