Failure Analysis of a Diesel Engine Exhaust Manifold

Abstract

The diesel engine failed during a 700-h bench test due to the cracking of exhaust manifolds. To investigate the cause of exhaust manifold cracks, material chemical composition, hardness determination, metallographic examination as well as fracture observation were carried out. Besides, the corresponding stress states were calculated using the finite element method. The results show that the material chemical composition, metallurgical morphology and hardness of the exhaust manifolds A and B are in good accordance with the standard. The cracked failure of Part A is due to stress concentration caused by improper structural design. Increasing the fillet radius is a useful way to improve the quality of the exhaust manifold. When the fillet is R6, the stress concentration phenomenon is well resolved. For Part B, the crack is mainly caused by a porosity defects close to the crack initiation area, which has a diameter of 185 µm.

Full text

FAILURE ANALYSIS OF A DIESEL ENGINE EXHAUST MANIFOLD
Yuan Li, Jinxiang Liu, Weiqing Huang and Yuanhao Wu
School of Mechanical Engineering, Beijing Institute of Technology, Beijing 100081, China
Copyright Ó2022 American Foundry Society
https://doi.org/10.1007/s40962-022-00796-8
Abstract
The diesel engine failed during a 700-h bench test due to
the cracking of exhaust manifolds. To investigate the cause
of exhaust manifold cracks, material chemical composition,
hardness determination, metallographic examination as
well as fracture observation were carried out. Besides, the
corresponding stress states were calculated using the finite
element method. The results show that the material chem-
ical composition, metallurgical morphology and hardness
of the exhaust manifolds A and B are in good accordance
with the standard. The cracked failure of Part A is due to
stress concentration caused by improper structural design.
Increasing the fillet radius is a useful way to improve the
quality of the exhaust manifold. When the fillet is R6, the
stress concentration phenomenon is well resolved. For
Part B, the crack is mainly caused by a porosity defects
close to the crack initiation area, which has a diameter of
185 lm.
Keywords: diesel engine, exhaust manifold, failure
analysis, fracture examination, finite element method
Introduction
With the increasing demand for high performance and low
emission of diesel engines, the corresponding working
service conditions are becoming more and more severe,
which is undoubtedly a challenge for the key components
of engine.
1,2
As one of the most important parts of the
automobile engine, the exhaust manifold collects the
combustion gases from the cylinders and directs them to
the exhaust system of the vehicle.
3
The correct perfor-
mance of the exhaust manifold not only affects engine
power and fuel consumption, etc., but also directly influ-
ences environmental pollution.
4
However, the exhaust
manifold usually works under alternating thermal cycles of
fast cooling and fast heating as well as under vibration
conditions, particularly when starting or stopping.
5,6
Moreover, the heat flow and combustion pressure have
increased significantly because of the need to reduce pol-
lution emissions while controlling fuel consumption.
7
Working in this harsh environment with thermal shock and
cooling/heating cyclic loads, exhaust manifolds are highly
susceptible to fatigue failure.
8,9
Therefore, whether from an
economic or safety point of view, a failure analysis of the
exhaust manifold is very essential.
10
Due to the harsh operating conditions of exhaust manifolds,
high-temperature oxidation resistance and thermal fatigue
properties of the corresponding material are required.
11,12
Exhaust manifolds for mass-produced vehicles are usually
made of ductile cast iron, because compact and complex
geometries could be obtained from ductile cast iron by the
casting process. However, ferritic ductile irons grow slowly
at high temperatures.
13
Alloying with silicon and molyb-
denum greatly improves the high-temperature performance
of ferritic ductile iron and many advantages of conven-
tional ductile iron have been maintained.
14–16
In particular,
since 1980, the material requirements have become more
demanding as emission standards for pollutants have
become stricter, resulting in higher exhaust gas tempera-
tures.
17
Therefore, with the advantages of low cost, good
mechanical properties, high temperature oxidation resis-
tance as well as fair casting performance, SiMo ductile iron
has become the most promising material in the manufac-
ture of exhaust manifolds.
9,18–20
The exhaust temperature
of diesel engine is about 750–800 °C, and the exhaust
temperature of gasoline engine is close to or even more
than 900 °C. The diesel engine exhaust manifold was
analyzed in this paper, and the gasoline engines need to use
Received: 24 January 2022 / Accepted: 30 March 2022
International Journal of Metalcasting

materials with better heat resistance, such as high Ni aus-
tenitic ductile iron or steel. Several studies have reported
that the fatigue cracking behavior of SiMo ductile iron may
be influenced by factors such as the chemical composition
and microstructure of the matrix, as well as the properties
of the material, including hardness and tensile strength,
etc.
21–24
In addition, it was found that thermal stress has an
effect on the fatigue performance of the exhaust mani-
fold.
10,25–27
The steady-state temperature distribution of
the exhaust manifold was calculated using the finite ele-
ment method, which in turn determined the region most
prone to cause stress concentration and fracture.
28
In
summary, the harsh operating conditions of exhaust man-
ifold are likely to cause fracture failure, and there are many
factors contributing to fracture. Therefore, several dimen-
sions need to be discussed when determining the cause of
exhaust manifold failure.
In this work, the SiMo ductile iron exhaust manifold (the
material is SiMo51) was used. After a 700-h bench test, the
engine parts were disassembled for inspection, and two
exhaust manifolds (exhaust manifolds A and exhaust
manifolds B) failed the test exhibiting cracks. For the
convenience of description, exhaust manifold A is subse-
quently referred to as Part A and exhaust manifold B is
referred to as Part B. The whole engine after failure is
shown in Figure 1, and the location of the exhaust mani-
folds is also illustrated. In addition, it is interesting to note
that unexpected cracking failures of exhaust manifolds
were also observed frequently in previous durability eval-
uation tests. Therefore, it is particularly important to
investigate the cause of exhaust manifold failure. The
purpose of this paper is to determine the mechanism and
the cause of the exhaust manifolds failure. For this purpose,
the chemical composition and hardness of the failed
exhaust manifolds were measured. The detailed metallur-
gical observations as well as fracture morphology analysis
were also performed. Besides, to determine the failure
mechanism of exhaust manifold, the thermal–mechanical
coupling analysis was carried out using the finite element
method, and the stress state was calculated. Based on the
analysis results, improvement measures to reduce exhaust
manifold crack initiation were also discussed. This study
could provide guidance for the subsequent improvement of
exhaust manifold performance.
Investigation Methods
The exhaust manifolds were made in an industry foundry
(Xixia Intake & Exhaust Manifold Co., Ltd.) according to
the customer’s specifications for SiMo51 iron castings.
29
The durability test with engine performance as the moni-
toring index was conducted at Hebei Huabei Diesel Engine
CO., LTD. The durability test was conducted with the
engine performance as the monitoring index. There was no
abnormality in the engine performance during the whole
testing, but cracks were found in the exhaust manifold after
the testing. To conduct a macroscopic study of the exhaust
manifold cracks and thus make a preliminary determination
of its failure, the chemical composition of the failed
exhaust manifold was estimated by spectrographic chemi-
cal analysis method with a spectrum analyzer, and hardness
measurements were taken by an HBS-3000 type Brinell
hardness tester. The relationship between material quality
and fracture failure was studied by qualitative and quanti-
tative analysis of the microstructure around the exhaust
manifold cracks. A Leica optical microscope was used to
observe the metallographic distribution of the different
locations on the exhaust manifold. For quantitative analy-
sis, a microscopic image analyzer (SISC IAS8.0) was used
to determine some critical microstructural characteristics,
including graphite size, grade nodularization, pearlite as
well as ferrite content, etc. In addition, the fracture mor-
phology was analyzed by scanning electron microscopy
(SEM) in order to reveal the changes in macroscopic
properties. Besides, to determine the failure mechanism of
exhaust manifold, the thermal–mechanical coupling anal-
ysis was carried out using the finite element method with
commercial software ABAQUS, and the stress state was
calculated. The failure analysis and modeling were con-
ducted by Beijing Institute of Technology.
Failure Analysis Results
Failure Description of the Exhaust Manifold
Durability bench tests are often used in failure estimation
because it is a more effective and common method to test
the whole engine’s reliability and durability with critical
components under large loads.
30
The bench test of the
whole engine was carried out with a hydraulic
dynamometer (DT900), including a 500-h alternating load
test and 200-h performance test. The loading cycle of the
500-h alternating load test is shown in Figure 2, where
solid line indicates full throttle opening. Firstly, the engine
Figure 1. Macro-appearance of the failure exhaust
manifolds.
International Journal of Metalcasting

speed was uniformly increased from the maximum torque
speed (n
t
) to the maximum power speed (n
p
), which lasted
1.5 min and runs steadily for 3.5 min. Subsequently, it
descends uniformly to n
t
after 1.5 min and runs steadily for
3.5 min at n
t
. Repeat the above alternating conditions until
the runtime is 25 min. Then close the throttle, reduce the
engine speed to idle speed (n
i
), and run to 29.5 min.
Increase the throttle and make the speed rise evenly to the
maximum speed n
M
(2300 r/min) within 0.25 min under
no-load condition. After that, turn down the throttle evenly
to make the speed drop to n
t
, lasting 0.25 min. The above
process lasted for 30 min, and repeated 20 times was
recorded as a cycle. The total duration of alternating load
test is 500 h for 50 cycles. The engine speed of the per-
formance test was 1350 r/min, and the engine continued to
work for 200 h under the condition that the power was
greater than 200 kW.
During the test, the exhaust temperature was measured with
a thermocouple, and the exhaust manifold was inspected
periodically to determine the appearance of cracks. An
unexpected crack failure occurred in the exhaust manifolds
after bench test. Locating the position of the crack is the
first step in failure analysis. At the end of the 700-h bench
test, the tested exhaust manifold is removed from the
engine so that the flanges between the exhaust manifold
and cylinder head could be visually inspected to determine
cracks within the exhaust manifold. The visual inspection
of the damaged Part A shows that the crack appeared in a
geometrically abrupt area near the pipe that connects to
Part B, as illustrated in Figure 3a. It can be seen in Fig-
ure 3a that the crack initiates from area 1 and propagates to
area 2. For Part B, the cracking area is located at the
reinforcement, as shown in Figure 3b. The crack exhibits a
‘herringbone’ shape, initiating from area 3, which was
directly below the reinforcement, and then propagating to
areas 4 and 5.
Chemical Composition of Material
and Hardness Measurement
The chemical composition of the failed exhaust manifolds
was determined by optical emission spectroscopy (OES)
through a spectrum analyzer. The determination of the
chemical composition is in accordance with the procedure
specified by the board of chemists of the VDEh (German
Iron and Steel Institute).
29
The test specimens were taken
from the regions near the crack sources of intake manifolds
A and B, respectively. The chemical composition results
and standard values are given in Table 1, where EM A and
EM B represent exhaust manifolds A and B, respectively.
Compared with the standard specifications for SiMo51 iron
castings, the chemical composition of the Part A and Part B
all meet the standard.
The hardness of the exhaust manifolds was determined
using Brinell hardness measurements according to Brinell
hardness tests criterion.
31
Brinell hardness tests were per-
formed at 15 s intervals with a 250-Kg force and a steel
indentor with 5 mm diameter. Three measurement points
were selected on each specimen, and the average value is
taken as the hardness of the measured specimen. The
details of hardness for the different exhaust manifolds are
also given in Table 1. The hardness test results show that
Part A and Part B have the same hardness, with a value of
229 HBW; both meet the requirements of the standard.
Microscopic Observation of Microstructure
Metallographic analysis was performed on the cracked
area, flange and normal pipe wall of the tested Part A. The
sampling locations are displayed in Figure 4a, b, where
ellipse A is the normal pipe wall away from the crack,
ellipse B is the flange, and ellipse C is the cracked area.
Similarly, Figure 4c, d shows the sampling locations of
Part B for metallographic analysis, where ellipse D is the
normal pipe wall away from the crack, ellipse E is the
flange, and ellipse F is the cracked area. All the sets were
etched with an alcoholic solution containing nitric acid (4%
nitric acid) before metallographic observation. The typical
metallographic structures of Part A and Part B are illus-
trated in Figures 5and 6, respectively, in which (a), (c) and
(e) are magnified by 1009, (b), (d) and (f) are magnified by
5009.
The microstructure of SiMo51 consists mainly of ferritic
matrix, graphite nodules and Mo-rich carbides. In addition
to the Mo-rich carbides, there are also magnesium (Mg),
chromium (Cr) and manganese (Mn) carbides. For Part A
and Part B, the pearlite at the normal pipe wall diffuses into
the matrix, as shown in (a) and (b) of Figures 5and 6. The
pearlite at the flange shows no tendency to decompose and
is distributed in a block shape, as illustrated in (c) and
(d) of Figures 5and 6. The metallographic morphology at
Figure 2. The loading cycle of the alternating load test.
International Journal of Metalcasting

Figure 3. Cracked areas on the exhaust manifolds.
Table 1. Chemical Composition (%) and Brinell Hardness (HBW)
Inspection item C Si Mn P S Mo Mg Cr Hardness
Trial value (Part A) 3.49 4.30 0.182 0.0282 0.005 0.821 0.042 0.016 229
Trial value (Part B) 3.44 4.54 0.148 0.0191 0.004 0.77 0.036 0.15 229
EN 10124 Standard 3.20–3.80 4.00–5.00 B0.70 B0.10 B0.015 0.75–1.20 0.03–0.07 B0.25 225 ±25
Figure 4. Sampling locations for metallographic analysis.
International Journal of Metalcasting

the cracked areas of exhaust manifolds A and B is given in
(e) and (f) of Figures 5and 6, respectively. The graphite
and gray Mo-rich carbides in the cracked area are either
embedded in the white matrix or gathered at the grain
boundaries, where the convergence of multiple grains is
mostly dominated by the gray Mo-rich carbides. In general,
the SiMo51 exhaust manifold is well spheroidized, and a
large amount of graphite is uniformly distributed on the
matrix and grain boundaries. The shape of the graphite is
almost spherical, with occasional polygonal shapes, and the
roundness of the graphite is high. The analysis of the dis-
tribution and size of SiMo51 exhaust manifold was carried
out, and the corresponding results are shown in Table 2.As
can be seen from Table 2, the graphite size and grade
nodularization as well as the composition of the matrix
microstructure are well in the rated range of standard.
32
SEM Investigation
In order to obtain detailed information about the failure
mechanism, the fracture surfaces of Part A and Part B
specimens were observed using a ZEISS SEM. The cut
fracture specimens were ultrasonically cleaned in anhy-
drous ethanol and then vacuum dried for inspection; the
corresponding results are given in Figures 7,8and 9.As
demonstrated in Figure 7a, b, the crack of Part A initiated
at the fillet of the pipe wall. It can be clearly seen that
although the fracture specimen of Part A is thoroughly
cleaned using dilute acid, it is still covered with oxide film.
However, multiple small steps at the crack initiation could
be seen in Figure 7a, and such steps could also be clearly
observed in the enlarged view of Figure 7b, i.e., Figure 7a.
Figure 8shows the SEM fracture morphology of Part A,
where (a) is the crack initiation and propagation zones and
(b) is the final fracture zone. The crack initiated in the
internal (combustion gas) wall. The initiated small cracks
keep propagating to the external (air) wall, and secondary
Figure 5. Typical metallographic morphology in different locations of Part A with
different magnification.
International Journal of Metalcasting

cracks are observed in the crack propagate area, as given in
Figure 8a. Besides, the equiaxed dimple could be observed
in the final fracture zone of Part A, as shown in Figure 8b.
These above fracture characteristics indicate that the Part A
is fatigue cracked and the failure mechanism is fatigue at
high temperature under vibration load.
The details of the Part B fracture are given in Figures 7c
and 9. Part B is also cleaned with anhydrous ethanol sev-
eral times, but the original appearance could not be
restored, and the fracture is still covered with oxide film.
From the fracture specimens, the crack initiation zone,
crack propagation and final fracture zone could be clearly
distinguished. In Figure 9, (a) is the crack initiation, (b) is
Figure 6. Typical metallographic morphology in different locations of Part B with
different magnification.
Table 2. Microstructure Characteristics of Part A and Part B
Inspection item Grade of
graphite size
Graphite
size (mm)
Grade
nodularization
Nodularity Composition of the matrix microstructure
Trial value (Part A) Grade 6 0.03–\0.06 Grade 2 90–\95% Ferrite = 89%; Pearlite
?Mo-rich carbides B11%
Trial value (Part B) Grade 6 0.03–\0.06 Grade 2 90–\95% Ferrite = 88%; Pearlite
?Mo-rich carbides B12%
EN 10124 Standard Grade 6–8 \0.06 Grade 1–2 [90% Ferrite matrix; Pearlite
?Mo-rich carbides B15%
International Journal of Metalcasting

the propagation zones, (c) is the final fracture zone, and
(d) is the higher magnifications of Ellipses A. The cracks in
Part B propagation from internal (combustion gas) to
external (air) wall. There are two obvious cracks in the
fracture surface expanding to the inner wall of pipe and
converging on the crack source, as illustrated in Figure 7c.
The crack initiation zone and crack propagation area of
Part B are magnified, and the corresponding results are
shown in Figure 9. As shown in Figure 9a, many defects
could be obviously observed in the crack initiation area.
The porosity defects are located near the internal wall with
a maximum diameter of 185lm. In addition, multiple small
cracks could also be seen in Figure 9a, and these small
cracks converge to form secondary cracks. This proves that
the Part B cracking has a multi-source character. Secondary
cracks can also be observed in the crack propagation area,
which is pointed out with arrows in Figure 9b. The specific
morphology of final fracture zone for Part B is given in
Figure 9c, d. The graphite particles are diffusely distributed
in the matrix, as shown in Figure 9d. In addition, river-like
ridges and striations are observed, indicating that the
fracture has a brittle quasi-cleavage fatigue fracture
morphology.
Discussion of Cause and Prevention
Effect of Microstructure on Crack Initiation
According to the above results, it can be found that the
chemical composition, hardness and metallographic struc-
ture of Part A and Part B are in accordance with the
standard requirements. The cracked area of the Part A
occurs in the reinforcement, which has abrupt geometrical
changes. Such a sharp structural change is prone to causing
stress concentration. In addition, microscopic observation
revealed that there is a main crack and several secondary
cracks in this region. Based on the above analysis, it is
inferred that the cracks in Part A may be caused by thermal
fatigue. The SEM results show that the cracking of Part B
has multi-source characteristics and the fracture exhibits a
brittle quasi-cleavage fatigue fracture morphology. Since
larger size defects exist in the crack initiation area, and
smaller defects are observed in the direction of crack
propagation. It is inferred that cracks initiated from a large
porosity defects close to the crack initiation area and then
propagate in the direction of the small defects.
Finite Element Analysis of Exhaust Manifold
Finite Element Model and Loading Conditions
In order to get information about the high stress region of
the exhaust manifold under working conditions and thus to
determine the possible failure, a numerical simulation of
the thermal–mechanical coupling process for the exhaust
manifold was carried out using the finite element method.
The finite element models are exhibited in Figure 10,
including the exhaust manifolds, cylinder heads, and
components. The meshing of these components is per-
formed with the commercial software Hypermesh, using a
four-node tetrahedral solid mesh. The mesh type is
assigned to C3D8T, which is a temperature-displacement
coupled mesh, meaning that it can perform coupled cal-
culation of temperature field and stress field.
33
The basic
mesh size of the whole assembly model is 3 mm. In order
to provide more accurate simulation results, the mesh near
the fracture position is refined, and the mesh size is set
about 0.5 mm. The whole assembly model contains
148,461 elements and 42,653 nodes.
The loads of exhaust manifold during operation conditions
consist of static load and dynamic load. The static load
includes the pre-tightened load force of the bolt connecting
the exhaust manifold to the cylinder head, while the
dynamic load is the vibration load generated during the
working process of the whole engine. Therefore, in order to
simulate the actual conditions of the exhaust manifold
more accurately, the temperature distribution, the assembly
load and the vibration load of the exhaust manifold were
considered in the coupled thermal–mechanical analysis. In
practice, the manifold is working in an unstable state. But
unsteady simulation calculation is complicated and time-
consuming. Another reason is that the engine was started
and stopped only once during the engine durability test,
and the loading frequency changes quickly throughout the
test. Therefore, the temperature field distribution of the
exhaust manifold is determined by the steady-state thermal
analysis.
26,34,35
During steady-state thermal analysis, the
internal pipe temperature of the exhaust manifold was set at
780 °C and the outer surface was set at room temperature
(25 °C). The heat exchange process of heat transfer occurs
due to the temperature difference between the internal and
external walls of the exhaust manifold. The external wall of
Figure 7. Macroscopic fracture morphology of Part A
and Part B.
International Journal of Metalcasting

the exhaust manifold is exposed to the air and convective
heat transfer forms between it and the air. The high-tem-
perature exhaust gas flows directly into the internal wall of
the exhaust manifold, so convective heat transfer is formed
between the internal wall and the floating hot exhaust gas.
The corresponding heat transfer coefficients involved in the
above were set during the simulation. In addition, the
elastic-plastic behavior of the material was considered as a
function of temperature. Then, the calculated temperature
field was loaded into FEM software as a boundary condi-
tion, and the bolt pre-tightening force is considered in the
further stress analysis. In the finite element model, the
cylinder head and one side of Part A connected to the pipe
were fixed. Tie constraints were used between the exhaust
manifolds to simulate the actual threaded connection. The
exhaust manifold was connected to the cylinder head by
bolts, which were implemented with a connecter element.
The interface between the exhaust manifold and the
cylinder head is defined by the contact conditions.
36–38
The
pre-tightened load applied to the bolts is 65 Nm. Based on
this, the stress distribution of the exhaust manifold could be
obtained. After that, the calculated temperature field was
loaded into FEM software as the boundary condition, and
the thermal stress was used as the initial stress condition to
Figure 8. SEM fracture morphology of Part A.
Figure 9. SEM fracture morphology of Part B.
International Journal of Metalcasting

perform the coupled thermal–mechanical analysis of the
exhaust manifold. To consider the effect of vibration loads
on the stress distribution in the exhaust manifold, the
amplitude measured experimentally with time is applied to
the exhaust manifold in the form of load. The properties of
the material are determined according to EN 16124, with a
yield strength of 480 MPa and ultimate tensile strength of
550 MPa. The stress relaxation during cycling is not
analyzed.
Temperature and Stress Distribution
The formation of cracks is influenced by thermal loads and
relatively large mechanical stresses.
39,40
The exhaust
manifold working in high temperature conditions is very
likely to result in transient temperature overheating, which
exceeds the rated capacity of the material, thus leading to
fatigue cracking. This reveals the importance of the tem-
perature distribution in the exhaust manifold. The simu-
lated temperature distribution of the exhaust manifolds is
given in Figure 11. The local temperature distributions of
Part A and Part B failure areas are given at the bottom of
the figure. From the overall view, the temperature distri-
bution in the exhaust manifolds is of a certain inhomo-
geneity. The regions with abrupt geometrical changes have
a higher temperature, reaching a value of 733 °C. This is
mainly due to the thicker walls in the areas, resulting in
poorer heat dissipation conditions. For the convenience of
analyzing the temperature distribution at the failure loca-
tion of the exhaust manifold, the failure areas of Part A and
B (indicated by arrows) are enlarged. From the enlarged
view, external and internal temperature distribution at the
failure area could be seen. The temperature at the failure
location of Part A is 652 °C and 723 °C internally. The Part
B failure region has an external surface temperature of 672
°C and an internal temperature of 712 °C. In addition,
during the actual durability test, the exhaust temperature
was measured using a thermocouple. Considering the dif-
ficulty of overall temperature measurement, only one point
was measured and compared with the simulation results.
The measurement location is at the flange, which is marked
with a box in Figure 11. The temperature at the flange is
about 462 °C, which is highly consistent with the tem-
perature (459 °C) measured in the durability test. This
illustrates the reliability of the finite element simulation
results.
Based on the above temperature field distribution, the stress
distribution of the exhaust manifolds is shown in Figure 12.
The stress in the exhaust manifold is time-dependent when
considering vibration loads. To facilitate the observation of
overall stress distribution in exhaust manifolds, Figure 12
illustrates the stress field when the maximum stress value is
observed during operation. The overall thermal stresses
have a large dispersion on the exhaust manifold. There is
an obvious stress concentration at the cracked location of
Part A. The maximum values are concentrated in the bolt
holes and geometrically abrupt regions. The stress ampli-
tude of Part A is 2.5 MPa and the average stress is 475
MPa. According to the GOODMAN equation, the asym-
metric equivalent stress was converted to the symmetric
equivalent stress, and the equivalent stress of Part A was
calculated to be 240 MPa. Based on engineering experience
and research, it can be approximated that the fatigue limit
of the material is one-third of the yield limit, which is about
160 MPa for the material studied in this paper.
29,41
It is
obvious that the equivalent stress of Part A has far
exceeded the fatigue limit of this material, which means
that it is easy to crack. However, no obvious stress con-
centration is observed at the fracture location of Part B.
The equivalent stress is only 50 MPa, which is far below
the fatigue limit of the corresponding material.
Analysis of Failure Causes and Prevention
According to the above analysis, it can be affirmed that the
fracture of Part A is mainly due to the stress concentration
at the failure region, which makes the stress at the fracture
location exceed the fatigue limit and cause fatigue fracture.
The stress concentration is mainly because the failure
region of Part A with geometric abrupt change, which is
easy to cause stress concentration phenomenon and thus to
Figure 10. Finite element mesh model of exhaust manifolds and constraint
components.
International Journal of Metalcasting

result in cracking. This is consistent with the described in
reference,
36
indicating that stress concentrations are likely
to occur in the structural conversion region, which in turn
lead to cracking. However, the calculation results of the
thermo-mechanical coupling show that the fracture region
of the Part B does not exhibit stress concentration and is
not susceptible to fatigue cracking. This is in good agree-
ment with the conclusion obtained from the fracture anal-
ysis, which demonstrates that the fracture of Part B is a
brittle fracture with quasi-dissociative properties and is not
a fatigue crack. In addition, based on above microstructure
results, the large size defects could be clearly found in the
Part B crack initiation area, as indicated in Figure 9.Itis
generally accepted that the presence of defects is prone to
causing fatigue cracks,
34
so it is inferred that the failure of
the Part B is attributed to brittle fracture due to cracks
caused by defects.
In order to eliminate the stress concentration in the failure
region of Part A as much as possible, the fillet radius of the
area is increased. In the original structure of the Part A, the
fillet radius (marked by double arrows in Figure 12)is0
mm, i.e., no fillet is used, which is noted as R0. In this
study, the fillet radius is changed to R2, R4, R6 and R8,
respectively. The thermal–mechanical analysis of exhaust
manifolds with different fillet radius designs was carried
out. This paper uses Mises criterion to describe stress state,
which is a simple method commonly used in stress state
description.
10,34
The maximum Mises stresses at failure
location of Part A with different fillet radius are given in
Table 3. As can be seen in Table 3, after increasing the
fillet radius, the stress in the failure region of Part A is
reduced from 483.2 MPa (R0) to 469.6 MPa (R8). When
the fillet radius changes from R0 to R4, the stress in the
failure region is reduced, while the maximum stress values
differ only slightly for fillet radius R4, R6 and R8.
Therefore, to determine the optimum fillet radius, the stress
field distribution of R4, R6 and R8 at the failure location
for Part A is illustrated in Figure 13. Compared with the
stress distribution of the original structure (shown in
Figure 11. The temperature distribution of exhaust manifolds under operating
conditions.
Figure 12. Typical thermal–mechanical coupling stress field in the exhaust manifold.
International Journal of Metalcasting

Figure 12), the stress concentration areas in the crack
failure region of R4, R6 and R8 are significantly reduced.
Moreover, the stress distribution in the danger zone is best
when the fillet radius is R6, with only a small portion of the
region having higher stresses, as shown in Figure 13.
Combining Table 3and Figure 13, it can be found that the
best reduction of stress is achieved when the fillet radius is
R6. The overall typical stress distribution in the Part A at
R6 is also shown in Figure 13. A general decrease of the
stress values along the exhaust manifold is visible. The
suggested changes in geometry do not have much effect on
the simulated maximum stress. However, the change of
geometry effectively reduced the overall stress in Part A
and the stress concentration phenomenon in the sensitive
area of Part A, which contributes greatly to the reduction of
cracks, thus reducing the risk of failure. Therefore, the
transition fillet radius of the Part A could be appropriately
increased in subsequent production to reduce fatigue fail-
ure. In fact, the improved exhaust manifold was experi-
mentally verified, and no cracks were found. For Part B, the
cause of failure is the presence of defects. It is well known
that the exhaust manifold is generally made by casting
because of its complex structure, and the defects are usu-
ally caused by unreasonable process parameters during the
casting process.
9,42,43
Therefore, to reduce the failure of
Part B, the casting process parameters need to be improved
to reduce the defects.
Conclusion
To analyze the failure causes of exhaust manifold cracks in
a diesel engine, experimental analysis and finite element
simulations were performed. Based on metallographic
observations, fracture measurements and finite element
analysis results, the possible causes of cracks are assessed.
The main conclusions could be summarized as follows:
(1) The crack of Part A occurs in a geometrically
abrupt area near the pipe that connects to Part B.
The crack in Part B is located directly below the
reinforcement and has a multi-source character.
(2) The experimental measurement results show that
the material composition, metallographic mor-
phology and hardness of both exhaust manifolds
A and B are in good conformity with the
standard, which are not the cause of fracture.
(3) The crack failure of Part A is caused by thermal
fatigue. The main factor leading to the formation
of cracks is stress concentration, which is caused
by unreasonable structural design. The crack in
Part B is mainly caused by a large porosity
defects close to the crack initiation area, which
has a diameter of 185 lm.
(4) A useful improved design to reduce the stress of
Part A is proposed. Simulation analysis shows
that increasing the fillet radius of the failure area
in Part A can effectively reduce its stress. The
best improvement in reducing the stress and
stress concentration phenomenon is achieved
when the fillet radius is R6.
Acknowledgements
This work was supported by the Equipment Pre-
research Field Fund [Grant No. 61409220130].
Table 3. Maximum Mises Stress at Failure Location of
Part A with Different Fillet Radius
Fillet radius R0 R2 R4 R6 R8
Mises stress of EM A
(MPa)
483.2 476.1 470.3 470.0 469.6
Figure 13. The stress distribution of Part A with structure modification.
International Journal of Metalcasting

Conflict of interest The authors declare no conflicts of interest to
this work.
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